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Statens Geotekniska Institut SGI Varia 95
DESIGN OF PILES IN NON-COHESIVE SOIL
Nguyen Truong Tien
SGI Linkoping Sweden
December 1981
1
CONTENTS
SUMMARY
ACKNOWLEDGEMENTS
INTRODUCTION
1 BEARING CAPACITY OF PILES
11 Methods based test (SPT)
on the standard penetration
1 11 The C~nadian FouManual method
ndation Engineering
112 The Schmertmann method
113 Experiences from United Kingdom
114 Swedish experiences
115 The Norwegian Pile Committe method
12 Methods based on the theory of plasticity and empirical formula
121 The Canadian Foundation Engineering Manual
122 The API method
1 2 3 The Vesic method
1 2 4 The Meyerhof method
1 bull 2 5 The Broms method
1 2 6 The Berggren method
1 2 7 The Australien Standard
1 bull 2 8 Swedish experiences
1 2 9 The Braatvedt method
1210 The Touma and Reese method
1211 Discussion
1 3 Method based on cone penetration test (CPT)
1 3 1 The Canadian Foundation Engineering Manual
1 3 2 The Meyerhof method
1 3 3 The Vesic method
1 3 4 The Thorburn method
1 3 5 The Tekam method
1 bull 3 6 The Begemann method
1 3 7 The Nottingham method
1 bull 3 8 The Norwegian Pile Committee method
2
1 3 9 Experiences from United Kingdom
1310 Discussion
14 Relation between the method based on static formula and static cone penetration
1 bull 4 1 Determination of Kstano from CPT-test
1 bull 4 2 Relation between Qs
and Qp
1 bull 4 3 Examples
15 Method based on pressuremeter test
1 bull 6 Method based on pile driving formula
1 7 Method based on stress wave measurement
1 7 1 The Case-method
1 bull 7 2 The CAPWAP method
1 7 3 Another application of stress wave measurement
18 Method based on the load test
1 bull 8 1 Method of load testing
1 8 2 Reporting of load testing results
1 bull 8 3 Failure criteria
1 8 4 Distribution of bearing capacity between the tip and the sides of the piles
1 bull 8 5 Safety factor
1 bull 8 6 Comparison between different methods of load tests
1 bull 8 7 Comparison between the creep load and the ultimate load of piles
2 BEARING CAPACITY OF PILE GROUP
3 SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 1 Settlement of a single pile
3 1 bull 1 Empirical method
3 1 bull 2 Settlement from load test
3 1 3 Method based on theory of elasticity
3 1 bull 4 The Vesic method
3 1 5 Method based on FEM
32 Settlement of a pile group
321 The method of Skemton
3
3 2 2
3 2 3
324
325
326
327
33
4
APPENDIX
APPENDIX
APPENDIX
APPENDIX
The Brezantsev method
Method based on theory of elasticity
Recommendations by Meyerhof
The Parry method
The De Beer method
The Schmertmann method
Summary of the methods for calculation of the settlement
CONCLUSION
A Relationship between of penetration test
the results
B Empirical methvalue of cp
ods to obtain the
C Safety factor
D References
4
SUMl-1ARY
The studies on the behaviour of piles in non-cohesive
soil have increased during the last ten years Different
methods to calculate the ultimate bearing capacity and
the settlement have been proposed This report makes
a review of the most common methods for design of
piles in non-cohesive soil The comparison between
different methods is discussed The relationship beshy
tween the CPT-method and the static formula can be used
for preliminary estimation of the bearing capacity of
the piles and the ratio between total skin friction and
total point resistance
The conclusions and recommendations for practical design
purposes are summarized
5
ACKNOWLEDGEMENTS
This report was made at my visit at SGI during 1981
according to SARECs (SIDA) program to which appreciation
is expressed
The writer greatly thanks Dr Jan Hartlen Director of
SGI for his recommendations on the program of the work
his assistance and encouragement
Specially thanks is expressed to Dr Bo Berggren at SGI
for giving recommendations invaluable discussions and
supervision to the work
The writer also gratefully thanks Mr Bergdahl Mr Hellman
of SGI and other members of SGI CTH and KTH for valuable
discussions
Gratitude is given to Mrs Eva Dyrenas for her expert
typing of the manuscript and Mrs Rutgerd Abrink for
drawing the figures
Thanks is also expressed to other members of SGI for
their kindness and their assistance during my time at
SGI
Linkoping December 1981
Nguyen Truong Tien
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
1
CONTENTS
SUMMARY
ACKNOWLEDGEMENTS
INTRODUCTION
1 BEARING CAPACITY OF PILES
11 Methods based test (SPT)
on the standard penetration
1 11 The C~nadian FouManual method
ndation Engineering
112 The Schmertmann method
113 Experiences from United Kingdom
114 Swedish experiences
115 The Norwegian Pile Committe method
12 Methods based on the theory of plasticity and empirical formula
121 The Canadian Foundation Engineering Manual
122 The API method
1 2 3 The Vesic method
1 2 4 The Meyerhof method
1 bull 2 5 The Broms method
1 2 6 The Berggren method
1 2 7 The Australien Standard
1 bull 2 8 Swedish experiences
1 2 9 The Braatvedt method
1210 The Touma and Reese method
1211 Discussion
1 3 Method based on cone penetration test (CPT)
1 3 1 The Canadian Foundation Engineering Manual
1 3 2 The Meyerhof method
1 3 3 The Vesic method
1 3 4 The Thorburn method
1 3 5 The Tekam method
1 bull 3 6 The Begemann method
1 3 7 The Nottingham method
1 bull 3 8 The Norwegian Pile Committee method
2
1 3 9 Experiences from United Kingdom
1310 Discussion
14 Relation between the method based on static formula and static cone penetration
1 bull 4 1 Determination of Kstano from CPT-test
1 bull 4 2 Relation between Qs
and Qp
1 bull 4 3 Examples
15 Method based on pressuremeter test
1 bull 6 Method based on pile driving formula
1 7 Method based on stress wave measurement
1 7 1 The Case-method
1 bull 7 2 The CAPWAP method
1 7 3 Another application of stress wave measurement
18 Method based on the load test
1 bull 8 1 Method of load testing
1 8 2 Reporting of load testing results
1 bull 8 3 Failure criteria
1 8 4 Distribution of bearing capacity between the tip and the sides of the piles
1 bull 8 5 Safety factor
1 bull 8 6 Comparison between different methods of load tests
1 bull 8 7 Comparison between the creep load and the ultimate load of piles
2 BEARING CAPACITY OF PILE GROUP
3 SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 1 Settlement of a single pile
3 1 bull 1 Empirical method
3 1 bull 2 Settlement from load test
3 1 3 Method based on theory of elasticity
3 1 bull 4 The Vesic method
3 1 5 Method based on FEM
32 Settlement of a pile group
321 The method of Skemton
3
3 2 2
3 2 3
324
325
326
327
33
4
APPENDIX
APPENDIX
APPENDIX
APPENDIX
The Brezantsev method
Method based on theory of elasticity
Recommendations by Meyerhof
The Parry method
The De Beer method
The Schmertmann method
Summary of the methods for calculation of the settlement
CONCLUSION
A Relationship between of penetration test
the results
B Empirical methvalue of cp
ods to obtain the
C Safety factor
D References
4
SUMl-1ARY
The studies on the behaviour of piles in non-cohesive
soil have increased during the last ten years Different
methods to calculate the ultimate bearing capacity and
the settlement have been proposed This report makes
a review of the most common methods for design of
piles in non-cohesive soil The comparison between
different methods is discussed The relationship beshy
tween the CPT-method and the static formula can be used
for preliminary estimation of the bearing capacity of
the piles and the ratio between total skin friction and
total point resistance
The conclusions and recommendations for practical design
purposes are summarized
5
ACKNOWLEDGEMENTS
This report was made at my visit at SGI during 1981
according to SARECs (SIDA) program to which appreciation
is expressed
The writer greatly thanks Dr Jan Hartlen Director of
SGI for his recommendations on the program of the work
his assistance and encouragement
Specially thanks is expressed to Dr Bo Berggren at SGI
for giving recommendations invaluable discussions and
supervision to the work
The writer also gratefully thanks Mr Bergdahl Mr Hellman
of SGI and other members of SGI CTH and KTH for valuable
discussions
Gratitude is given to Mrs Eva Dyrenas for her expert
typing of the manuscript and Mrs Rutgerd Abrink for
drawing the figures
Thanks is also expressed to other members of SGI for
their kindness and their assistance during my time at
SGI
Linkoping December 1981
Nguyen Truong Tien
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
2
1 3 9 Experiences from United Kingdom
1310 Discussion
14 Relation between the method based on static formula and static cone penetration
1 bull 4 1 Determination of Kstano from CPT-test
1 bull 4 2 Relation between Qs
and Qp
1 bull 4 3 Examples
15 Method based on pressuremeter test
1 bull 6 Method based on pile driving formula
1 7 Method based on stress wave measurement
1 7 1 The Case-method
1 bull 7 2 The CAPWAP method
1 7 3 Another application of stress wave measurement
18 Method based on the load test
1 bull 8 1 Method of load testing
1 8 2 Reporting of load testing results
1 bull 8 3 Failure criteria
1 8 4 Distribution of bearing capacity between the tip and the sides of the piles
1 bull 8 5 Safety factor
1 bull 8 6 Comparison between different methods of load tests
1 bull 8 7 Comparison between the creep load and the ultimate load of piles
2 BEARING CAPACITY OF PILE GROUP
3 SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 1 Settlement of a single pile
3 1 bull 1 Empirical method
3 1 bull 2 Settlement from load test
3 1 3 Method based on theory of elasticity
3 1 bull 4 The Vesic method
3 1 5 Method based on FEM
32 Settlement of a pile group
321 The method of Skemton
3
3 2 2
3 2 3
324
325
326
327
33
4
APPENDIX
APPENDIX
APPENDIX
APPENDIX
The Brezantsev method
Method based on theory of elasticity
Recommendations by Meyerhof
The Parry method
The De Beer method
The Schmertmann method
Summary of the methods for calculation of the settlement
CONCLUSION
A Relationship between of penetration test
the results
B Empirical methvalue of cp
ods to obtain the
C Safety factor
D References
4
SUMl-1ARY
The studies on the behaviour of piles in non-cohesive
soil have increased during the last ten years Different
methods to calculate the ultimate bearing capacity and
the settlement have been proposed This report makes
a review of the most common methods for design of
piles in non-cohesive soil The comparison between
different methods is discussed The relationship beshy
tween the CPT-method and the static formula can be used
for preliminary estimation of the bearing capacity of
the piles and the ratio between total skin friction and
total point resistance
The conclusions and recommendations for practical design
purposes are summarized
5
ACKNOWLEDGEMENTS
This report was made at my visit at SGI during 1981
according to SARECs (SIDA) program to which appreciation
is expressed
The writer greatly thanks Dr Jan Hartlen Director of
SGI for his recommendations on the program of the work
his assistance and encouragement
Specially thanks is expressed to Dr Bo Berggren at SGI
for giving recommendations invaluable discussions and
supervision to the work
The writer also gratefully thanks Mr Bergdahl Mr Hellman
of SGI and other members of SGI CTH and KTH for valuable
discussions
Gratitude is given to Mrs Eva Dyrenas for her expert
typing of the manuscript and Mrs Rutgerd Abrink for
drawing the figures
Thanks is also expressed to other members of SGI for
their kindness and their assistance during my time at
SGI
Linkoping December 1981
Nguyen Truong Tien
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
3
3 2 2
3 2 3
324
325
326
327
33
4
APPENDIX
APPENDIX
APPENDIX
APPENDIX
The Brezantsev method
Method based on theory of elasticity
Recommendations by Meyerhof
The Parry method
The De Beer method
The Schmertmann method
Summary of the methods for calculation of the settlement
CONCLUSION
A Relationship between of penetration test
the results
B Empirical methvalue of cp
ods to obtain the
C Safety factor
D References
4
SUMl-1ARY
The studies on the behaviour of piles in non-cohesive
soil have increased during the last ten years Different
methods to calculate the ultimate bearing capacity and
the settlement have been proposed This report makes
a review of the most common methods for design of
piles in non-cohesive soil The comparison between
different methods is discussed The relationship beshy
tween the CPT-method and the static formula can be used
for preliminary estimation of the bearing capacity of
the piles and the ratio between total skin friction and
total point resistance
The conclusions and recommendations for practical design
purposes are summarized
5
ACKNOWLEDGEMENTS
This report was made at my visit at SGI during 1981
according to SARECs (SIDA) program to which appreciation
is expressed
The writer greatly thanks Dr Jan Hartlen Director of
SGI for his recommendations on the program of the work
his assistance and encouragement
Specially thanks is expressed to Dr Bo Berggren at SGI
for giving recommendations invaluable discussions and
supervision to the work
The writer also gratefully thanks Mr Bergdahl Mr Hellman
of SGI and other members of SGI CTH and KTH for valuable
discussions
Gratitude is given to Mrs Eva Dyrenas for her expert
typing of the manuscript and Mrs Rutgerd Abrink for
drawing the figures
Thanks is also expressed to other members of SGI for
their kindness and their assistance during my time at
SGI
Linkoping December 1981
Nguyen Truong Tien
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
4
SUMl-1ARY
The studies on the behaviour of piles in non-cohesive
soil have increased during the last ten years Different
methods to calculate the ultimate bearing capacity and
the settlement have been proposed This report makes
a review of the most common methods for design of
piles in non-cohesive soil The comparison between
different methods is discussed The relationship beshy
tween the CPT-method and the static formula can be used
for preliminary estimation of the bearing capacity of
the piles and the ratio between total skin friction and
total point resistance
The conclusions and recommendations for practical design
purposes are summarized
5
ACKNOWLEDGEMENTS
This report was made at my visit at SGI during 1981
according to SARECs (SIDA) program to which appreciation
is expressed
The writer greatly thanks Dr Jan Hartlen Director of
SGI for his recommendations on the program of the work
his assistance and encouragement
Specially thanks is expressed to Dr Bo Berggren at SGI
for giving recommendations invaluable discussions and
supervision to the work
The writer also gratefully thanks Mr Bergdahl Mr Hellman
of SGI and other members of SGI CTH and KTH for valuable
discussions
Gratitude is given to Mrs Eva Dyrenas for her expert
typing of the manuscript and Mrs Rutgerd Abrink for
drawing the figures
Thanks is also expressed to other members of SGI for
their kindness and their assistance during my time at
SGI
Linkoping December 1981
Nguyen Truong Tien
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
5
ACKNOWLEDGEMENTS
This report was made at my visit at SGI during 1981
according to SARECs (SIDA) program to which appreciation
is expressed
The writer greatly thanks Dr Jan Hartlen Director of
SGI for his recommendations on the program of the work
his assistance and encouragement
Specially thanks is expressed to Dr Bo Berggren at SGI
for giving recommendations invaluable discussions and
supervision to the work
The writer also gratefully thanks Mr Bergdahl Mr Hellman
of SGI and other members of SGI CTH and KTH for valuable
discussions
Gratitude is given to Mrs Eva Dyrenas for her expert
typing of the manuscript and Mrs Rutgerd Abrink for
drawing the figures
Thanks is also expressed to other members of SGI for
their kindness and their assistance during my time at
SGI
Linkoping December 1981
Nguyen Truong Tien
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
6
INTRODUCTION
Pile foundations are frequently economical in many
countries The research on methods for prediction of
the bearing capacity and the settlement of piles has
increased considerably during the last 10 years Difshy
ferent methods to calculate the ultimate bearing cashy
pacity and settlement have been proposed Methods based
on static and dynamic penetration tests are commonly
used for non-cohesive soil Based on experiences and
theoretical studies empirical expressions and methods
based on the theory of plasticity or elasto-plastic
theories have also been recommended New methods based
on the results from pressuremeter tests and stress wave
measurements have been developed The reported results
show good agreement between calculated and measured
values However up to now a general solution for
computation of the ultimate load and settlement of
pile foundations is not yet available Because of many
uncertainties it is recommended to perform full scale
pile load tests
The methods used for prediction of the ultimate bearing
capacity and settlement of piles depend on soil conshy
ditions available equipments the characteristics of the
structure similar experiences and other factors
This report makes a review of all common methods for
design of pile foundations in non-cohesive soil In
the report diagrams tables typical values empirical
expressions for design purposes have been collected
A discussion and relationships between different methods
have been summarized The method based on static penshy
etration test and the method based on theoretical and
empirical expressions are suitable for Vietnam conditions
so they are concentrated in this study Further studies
in Vietnam to compare different prediction methods and
complementary load testing in the object to get suitable
methods for our conditions will be very valuable
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
7
1 BEARING CAPACITY
This chapter covers the calculation of the bearing
capacity of piles founded in granular soils gravel
sand and silt where the granular soil extends to a
significant depth beneath the foundation level Several
methods based on different theories and experiences are
summarized here
1 bull 1 Method based on the standard penetration test (SPT)
111 The Canadian Foundation Engineering Manual (1978)
The ultimate bearing capacity of the pile is calculated
from SPT results according to the method suggested by
Meyerhof (1956)
where = ultimate pile load kN
N = average number of SPT at the tip level blow30 cm
-N = average number of SPT along the pile shaft blow30 cm
m2= cross sectional area of pile tip
= surface area of pile shaft m2
unit resistance at the point and along the pile shaft kPa
Factor of safety FS = 4
Meyerhof (1976) has suggested
= 40 N Db A + 2 NA B p s
-where f = 2 N (kPa)s qp = 40 N Qblt 400 N (kPa)
B -
where Db = embedment depth of the pile
B = diameter of the pile
and if the pile is driven in non-plastic silt q = 300 N p
(kNm 2) as an upper limit
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
8
An empirical relationship between N and q is shown in
Fig 1 Meyerhof (1956) has recommended
f = 01 MPa as an upper limit value for the s unit skin friction of driven piles
f = 0006 N (MPa) or maximum 004 MPa for bored s piles
1 bull 1 bull 2 The Schmertmann method
Schmertmann (1967) made suggestions for both end-bearing
capacity and side friction capacity estimated from SPT
data and the relationship between SPT and CPT data for
evaluation of the unit side friction on precast concrete
piles in Florida
Table 1 Usually conservative ultimate values for unit side friction and end-bearing correlated to SPT N-values (N-value range 560)
Type of soil Unified SCS description
q NC
Friction ratio Rf ()
Side friction (tsf) (kPa)
End-bearing
kN
Clean sand above and below GW all denshysities
GW SW
GP SP
GM SM
35 06 019
21
32
Clay silt sand mixed very silty sand silts and marls
GC ML
FC CL 20 20 04
44
16
Plastic CH OH 10 50 05 07
clays 056
Soft limeshystone lime rock very shelly sand
10 025 0 1
11
36
-For N less than -For N more than
5 use zero
60 use 60
Limitation As the SPT is subjected by many errors a high
factor of safety is recommended However many authors have
suggested not to use SPT as the only method for prediction
of the bearing capacity of the pile
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
9
1 bull 1 bull 3 Experiences from United Kingdom (Rodin et al 1974)
According to Gibbs and Holtz (1957) the penetration reshy
sistance increases with increasing relative density and
or overburden pressure Since the principal object of
standard penetration tests in non-cohesive soil is to
evaluate relative density the effect of overburden pressure
at the depth of the test must be taken into account The
correction chart (Fig2) was presented by Thorburn (1963)
and is widely adopted in the UK
With the value of N corrected the bearing capacity of
the pile can be evaluated by using the relation between
N and qc or alternatively the angle of internal friction
cp I bull
1 bull 1 bull 4 Swedish experiences (Broms 1981 Bergdahl 1980)
The bearing capacity and the length of driven reinforced
concrete piles are generally estimated in Sweden by ram
rounding The circular or square point is driven into
the soil using 32 mm diameter rods and a drop hammer
weighing 635 kg The number of blows N required to20 drive the penetrometer 020 miscounted The method has
an advantage that it is normally possible to reach much
deeper soil layers compared to weight sounding or cone
penetration test The relative density of non-cohesive
soils (sand and gravel) can be estimated from the following
table
Table 2 Ram penetration test Relationship between relative density and penetration resistance
Relative density Penetration resistance N20 blows02 m
Very loose lt5
Loose 5-12
Medium 12-35
Dense 35-60
Very dense gt60
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
1 O
The penetration resistance is thus about 11-20 larger
than that determined by Standard Penetration Test (SPT)
The relationship between N and N is also shown in20 30 Appendix A
The maximum depth to which a precast concrete pile can
be driven corresponds normally to a penetration resistance
of 200-300 blows02 mN20
1 bull 1 bull 5 The Norwegian Pile Commmitteemethod (1973)
According to Norwegian experiences the bearing capacity
of driven friction piles of wood and precast concrete
can be estimated by the Swedish ram sounding method A
(free falling weight=635 kg pressometer with square
point the penetrometer is rotated every 0 2 m to reduce
the skin friction along the rods The number of blows
required to drive the penetrometer 02 miscounted)
Fig 3 shows the relationship between the unit skin fricshy
tion resistance the surface area of the pile A skin and
the dynamic penetration resistance The indicated relationshy
ship has been evaluated from pile load tests The ultimate
bearing capacity was found to be higher for timber piles
than for precast concrete piles at the same penetration
resistance because of the conical shape of timber piles
The diagram is valid for timber and concrete piles with 2 a cross sectional area of 500-1000 cm and a pile length
between 12 and 15 m The equivalent skin friction is equal
to the ultimate bearing capacity of the pile divided by
the total shaft area
The bearing capacity of the pile can also be evaluated by
Q = A (00021+00022 Pd)s y
where
Q = ultimate bearing capacity of the pile MN 2A = total area of the shaft m s
and = average number of blows per 02 m of penetra-Pdy tion
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
1 1
It is also possible to calculate the ultimate bearing
capacity from the relationship between ram sounding
type A and SPT-test Test results indicate that the penshy
etration resistance from ram sounding type A (blows02 m)
corresponds to the N-value (blows03 m) from SPT The
same calculation methods can thus be used for ram sounding
type A as for Standard Penetration Tests
12 Method based on the theory of plasticity and empirical formula
General formula
= 0 +Q = f A + q A -s p s s p p
where f = unit shaft resistance kPa s
= unit point resistance kPaqp 2
A = pile shaft area m s 2
A = pile point area m p
= ultimate bearing capacity of the pile kNQf
1 2 1 The Canadian Manuals recommendation
a) Critical depth The bearing capacity of piles in granular
soils is not a continuous linear function of the overshy
burden pressure Vesic (1970) has demonstrated that the
skin friction and the point resistance become constant
below a critical depth D For practical pruposes D = C C
8-20 B (B=pile diameter)
b) Ultimate point resistance
The length of the pile L lt critical depth D p C
= L Nqfp p q
where
= ultimate unit point resistance kPaqfp 3
= effective unit weight of the soil kNm
L = the length of the pile in soil m p N = bearing capacity factor q
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
1 2
According to Brezantzev et al (1961) values of N are q
as listed in table 3
Table 3 Values of N according to the angle of friction~shyq
~ 25 30 35 40
N 15 30 75 150 q
The values of N~ can also be taken from Fig 4
For pile lengths exceeding D C
qfp = yDc N~
c) Skin friction
f = K 0 tano lt f1 S S V
K = average coefficient of earth pressure on the s pile shaft (see App B)
0 I = average effective overburden pressure alongV the shaft
= angle of friction of the sand-pile contact (see App B)
f 1 = maximum value off at and below the critical depths
Ks tano is influenced by the angle of friction~ the
method of installation the pile size and shape It inshy
creases with density For bored piles the values are equal
to 14 of the values of the displacement piles
d) Safety factor
Allowable load can be evaluated considering FS = 3
1 The length of the pile (L ) is less than the critical p depth (D )
C
1 TTD 2 fs = rD L )Qa 3(qfp -4- + 2 p
q f are calculated corresp_onding to thefp s pile length L
p
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)
1 3
2 The length of the pile (L) is longer than the p
critical depth (D)C
f 21 TTD
Qa = 3 (qfpL -
qfp and fs are calculated at the critical depth
D C
122 The American Petroleum Institute (ATP) method (1977)
a) Skin friction
In compression f = a Ktano K = 07 s V
In tension f = a Ktano K = 05 s V
a = vertical effective stress V
K = ratio of horizontal to vertical effective stress
0 = angle of wall friction between pile and soil o=cp-5 0
cp = angle of soil internal friction (limited to cp=35deg)
b) End bearing
Q p
= A p
crN V q
N q
= bearing capacity factor Caquot Kerisel (1956) or
Vesic (1967) (see Fig 4)
1 bull 2 3 The method of Vesic (1975 and 1977)
a) End-bearing
Q = A (1+2 Ko)a N p p 3 V q
where K is the ratio of at-rest horizontal to vertical0
effective stress N = bearing capacity factor (Fig 5)
When good information on the relative density is availshy
able Fig 6 or Table 4 can be used as a guid for evaluating
N q
b) Skin friction Fig 7 can be used as a guid for detershy
mination of the skin resistance
14
1 bull 2 4 The method of Meyerhof (1976)
qp = Po Nq 2- ql
= effective overburden pressure at the pile point
bearing capacity factor (Fig 8)
limit value of unit point resistance for D De B ~ B
De o DeAccording to Meyerhof B = 7 for~= 30 and B = 22 for
~ = 45deg Fig 8 shows that the bearing capacity factor
N varies with~ and the ratio between the bearing depthq
DB and the pile diameter B A relationship between the
critical depth De and the pile diameter is also presented
1 bull 2 5 The method of Broms (1978)
a) Skin friction
The skin friction resistance will increase linearly with
increasing depth down to 20B (5-10 m) when the relative
density of the sand is low Below this depth the skin
friction will be constant and the skin friction f down s
to a depth of about 20B can be calculated from the relation-
ship
f = K tan~ 0 1
S O a V
K = coefficient depending on the shape of the pile0 and the relative density (Table 5)
Table 5 Coefficient of lateral earth pressure K for friction piles in cohesionless soil~
Pile type Low relative High relative densitvDr=0 density Dr=1
Small displacement pile 05 1 0
Displacement piles 10 20
Conical piles 15 40
15
For small displacement piles (steel H piles) willK0
approach the coefficient of lateral pressure at rest
and increase with increasing relative density
For displacement piles the value of K0 is high because
of local arching around the pile
If a linear relationship between D and ~ is assumedr the values of K0 are given in Fig 14 a Recommended
values of~ are given in Table 6 a
Table 6 The friction angle~ as a function of the angle of internal friction of the soil (~)
Pile type ~a
Steel piles 20deg
Concrete piles 07 ~middot Timber piles 07 ~ I
For rough estimates K tan ~a= 030 can be used 0
b) Point resistance
The unit point resistance is calculated by
As L gtgtBand Ny ~ N the first term will generally be q
neglected and
qp = pag L Nq = a V
N q
where a~ is the effective overburden pressure at the level
of pile point The unit weight (p) of the soil correspondsa
to the apparent unit weight below water if the ground
water level is located at the ground surface
16
The net bearing capacity of the pile point will be equal
to net= p g L N - p g L
qp a q pile
If it is considered that the unit weight of the pile
corresponds to the unit weight of the soil
q = p g L(N -1)p a q
N can be obtained from Meyerhof (1953) (see Fig 4b)q
A safety factor equal to 3 is recommended
1 bull 2 6 The method of Berggren (1978)
The method is proposed for large diameter bored piles
in non-cohesive soils Based on test results Berggren
has stated
- In medium dense dense and very dense sand the ultimate
bearing capacity of the pile will be reached at a settleshy
ment of 5 of the diameter of the pile (5B) and the
settlement is equal to 10 Bin the case of loose and
very loose soil
- Maximum shaft resistance is reached at small displaceshy
ment of the pile (~5 mm) so a lower FS should be applied
to the shaft resistance than the point resistance Fig 9
is constructed with a safety factor of 5 on the point
bearing capacity psf thus giving the permissible point
pressure Psmiddot
- The permissible shaft resistance p is determined bym
Fig10 and the allowable skin friction resistance is
determined by
where A is the cross section of the pilep
- Based on the relationship between the angle of friction
~ and the bearing capacity factor N taking into account q
17
the critical depth according to Meyerhof (1976)
Fig11 has been drawn The unit point resistance
can be evaluated
qp = p sf = Nq CJ~
N = f(~ and L 1 B) where L1 = pile length in the q
bearing stratum and B = pile diameter
Note that for bored piles the bearing capacity factor
is 3 times lower than for driven piles
1 bull 2 7 The Australian Standard 2159-1978
a) Shaft resistance
The Australian Standard also considers the effect of
critical depth
for the depth z 0 lt z f = F 0lt Dc s V
for the depth z gt D f = F 0c s vc
0 1where is the effective vertical stress at the critical vc depth D The values of D and Fare given in Table 7
C C
b) Point resistance
The same consideration about the critical depth is taken
into account
q = N CJ p q vb
where CJ~b at the pile tip if L lt Dc
and 0 1 at the critical depth if L gt Dcvb and Lis the length of the pile
Values of N are given below q
Table 7 Values of D F and N C q
Soil condition Rela~ive rDcB F N q
density I I driven bored I driven bored I pile pile pile pile
25
Medium 04-075 8 10 05 100
iLoose 02-04 6 08 03 60
60
Dense 075-09 15 15 08 180 100
B = pile diameter
18
Comment The critical depth is between 6B and 15B for
loose and dense sand These values are smaller than the
values proposed by Meyerhof (1976)
1 bull 2 8 Swedish experiences (Hultsjo amp Svensson 1969)
Based on 32 load test results from different places in
Sweden good correlation was found between the measured
ultimate bearing capacity and the bearing capacity deshy
termined from a static bearing formula L
Q = No A + J Ktano o ~ du p p z~z z 0
where
N = point bearing capacity factor (N=8)
0 I = effective overburden pressure at the p pile point
A = cross sectional area of the pile pointp L = total length of the pile -a = effective overburden pressure at the z depth z
= circumference of the pileltp z Ktano = 035
The limit value of the unit skin friction is in general
equal to 5 MPa For the case of a steel pile in loose
sand the value of Ktano = 025 is recommended (Bergdahl
1980)
1 bull 2 9 The Braatvedt method (1976)
General bearing capacity formula
Qf = Qp + Qs
0 1Qp = (13 c N )+(04 ByNy)+ N Ab C V q
for sand Qp = (04 ByNy+ a N ) A q p
0 1= (ac + K tano) AQs s V s
N N bearing capacity factors depending on ltp I (Fig 1 2) Nq c
1 9
~ = average effective angle of friction of the soil over a length of 1B below and 4B above the tip of the pile obtained from laboratory tests or deep sounding May be reduced by s0 for bored piles and increased by 40 for driven piles
o = angle of friction between pile and soil
23 for timber piles
34 for concrete piles
20deg for concrete piles
K s = Coefficient depending or disturbance
on the degree of densification
05 for bored piles with no temporary support of the sides
1 bull 0 for bored piles with adequate support of the sides
20 for driven piles in sand
B = pile diameter at the pile toe
AbAs = area of the base and the pile shaft
0~ = effective stress at the pile toe
a = average effective overburden stress V
c = cohesive strength at the pile tip
c = average cohesive strength over the embedded length
Fig 12 also shows a correlation between N of SPT and~shy
Fig 13 shows the indicative ultimate pressure at the
base or the tip of piles in clean sand derived from the
above expression
Safety factor
Qs q = -- +
a 1 0
The settlement should be checked against allowable settleshy
ment
1210 The Tuoma and Reese method (1974)
A method to calculate the ultimate bearing capacity of
bored piles in sand is proposed based on the comparison
between calculated and measured load transfer in sand
Tuoma and Reese (1974) observed that the failure load
20
corresponds to a displacement equal to 5 of the pile
diameter The skin friction failure load Qf appears at
25 mm of downward movement of the shaft
1 Skin friction Qs
The ultimate resistance Q in sand is obtained as s
L a e f (5 tancp dl avg
0
where a = 07 for the length of the pile in sand avg not exceeding 8 m aavg may decrease with greater penetration of the pile
8 = perimeter of the pile
(5 = effective overburden pressure
cp = effective friction angle
dl = differential distance along the length of the pile
2 Point resistance Qp
The failure tip resistance in sand is obtained at 25 mm
downward movement of the pile
Ap QP = 50 (SI unit)13 qp
where B == diameter omiddotf the pile B ~ 0 6 m (expressed in m)
= cross sectional area of the pile
tip resistance at 5 B movement
0 for loose sand
15 kPa for medium sand
38 kPa for very dense sand
For sand of intermediate density linear intershy
polation can be used
The values of tip resistance is limited to
L gt 10B
21
1211 Discussion and conclusions about the static
method for calculation of ultimate bearing
capacity of piles in non-cohesive soil
(1) The critical depth should be considered in the evalushy
ation of the ultimate bearing capacity of the pile The
critical depth depends on the relative density of the soil
the internal friction angle and the diameter of the pile
The critical depth varies from 6 to 20 times the diameter
of the pile The critical depth values recommended by
Meyerhof (Fig 4) or the Australian Standard (Table 7) can
be used The values recommended by the Australian Standard
are more conservative The unit skin friction and unit point
resistance should be evaluated with the critical depth if
the length of the pile is greater than the critical depth
Below the critical depth the unit skin friction is conshy
stant The curves Figs 10 and 11 based on Meyerhofs
recommendation (Berggren 1978) are valuable for practical
purpose
(2) The unit point resistance is evaluated by
q = 0 1 N p V q
where 0 1 is the effective overburden pressure at the pointV
tip or the critical depth N is the bearing capacity factor q and can be evaluated by the plasticity theory and is
depending on the pattern of failure In general Berezantzevs
theory and Vesics theory have more support than other
theories in practical design The values of N can be ob-q
tained from Table 3 Fig 4 or Fig 5 It is useful to
take into account the values of Table 4 as a guide in
practical purpose
The expression for calculating the unit point resistance
is the same for cohesive soil but as the values of N q are greater the point resistance takes a more important
role in the evaluation of the ultimate bearing capacity
of the pile The values of N have a wide range of variation q and are very sensitive to the variation of the values of
22
internal friction angle On the other hand the point
resistance requires a greater displacement to mobilize
the ultimate bearing capacity in comparison with the
shaft resistance Therefore it is recommended to use a
safety factor equal to 3 for the point resistance
In the case of silty soil since high excess pore pressure
can develop when it is loaded it is recommended to use
~bull = 20-25deg to calculate the point resistance
(3) The unit skin friction is evaluated by
f = K tano aS S V
where a is the average of effective overburden pressureV
along the pile length or from the pile length to the
critical depth The factor Ks tano has the same meaning
as the factor Sin the effective analysis of piles in
cohesive soils It is interesting to note that the value
Kstano = 03-035 reported by Hultsjo et al (1969) and
used as a rough estimate of unit skin friction (Broms
1981) corresponds to the value of S reported by Burland
(1973) for normally consolidated clays
Table 8 contains a summary of values of Ks and Kstano
according to the recommendation of different references
The values of K and K tano have a wide range of~variations s and are depending on the relative density of the soil (or
angle of internal friction) material of the pile method
of installation and the length of the piles
Based on the recommendation of Broms (1978) the Australian
Standard (1978) and the relationship between the relative
density and the internal angle of friction~ (Terzaghi
and Peck 1948) a relationship between Ks and~ is presented
in Fig 14 together with data of Meyerhof (1976) Berggren
(1978) and the values measured by Christoulas (1981) and
Armishow (1980) The values of Ks of Armishow (1980) are
calculated based on the value of Ks tano measured and the
values of N from SPT
23
Table 8 Values of Kand Ktano for calculation of the unit skin friction
Reference K 0 K tano Remark s s
driven pile bored pilE
API ( 1977) 07 05 for tension
Australian 08 03 loose sand Standard (1978) 10 05 medium sand (SAA) 15 08 dense sand
Berggren (1978) (0 3) 0 1 cp = 30deg (0 6) 30 cp = 60deg
Braatvedt (19 76) 2 1 0 7cl timber pile Ks=05 for the cases 05 075clconcrete pilE of bored pile withoui
20deg steel pile support side
Brom ( 1978) 05-1 (Dr = 0-1) 20deg steel pile For rough estimation 1-2 (Dr = 0-1) 075 concrete pile Kstano = 03 15-4 (Dr = 0-1) 070cl timber pile
conical pile
Christoulos (9181) 15 06 cp gt 30deg K is measured
Armishow ( 1980) 09 NPS = 18 13 NPS = 54
Cooke (29 78) 07 L lt 75
05 120Lp gt p
Hultsjo et al 035 From load tests of
( 1969) driven piles
Meyerhof (19 76) 10 025 cl= 33deg 25 075 cl= 375deg
Tuoma amp Reese 07 Define as value (1974) decrease with Lgt8 m
Meyerhof (1951) 05 loose sand 10 dense sand
Kezdi (1957 Theory analysis~ Ireland ( 1957) 1 75-3 Pull out tests
Lambe-Whitman 2( 1969)
24
The values of Ks reported by Meyerhof have a big range
of variation with small variation of internal friction
angle The similar variation of Ks with angle of internal
friction can be observed from the recommendation of Broms
the Australian Standard and the values reported by Armishow
and Christoulos The values reported by Armishow are smaller
than values recommended by Broms and the Australian Standard
due to the different assumptions in the calculation and
that the values measured are carried out during the driving
of the pile
For practical purposes the values recommended by Broms
can be used
Values of friction between various construction materials
and sand have been reported by Potyondy (1961) and are
presented in Table 9 Ifmicro is defined as the ratio between
the friction between the pile material and sand and the
drained internal frictionmicro is in the range 06 i micro lt 10
and steel material has the lowest value of friction
API (1977) and Broms (1978) have recommended that o = s0
which is conservative in comparison with the values reported
in Table 9 except for the case of steel piles However
the criterium o = ~ 1 -5deg can be used in practical purpose
for concrete and timber piles because it differs very
little from the values given by Podyondy For steel piles
micro = 06 and o can be calculated from~bull given or
8 = 20deg the later criterium may be very conservative
Fig 15 shows the relationship between K tano and the s internal angle of friction~bull considering the values of
K from the recommendation of Broms and o = ~-5deg (API)s With given value of angle of internal friction the value
of Kstano can be obtained from the curve for concrete and
timber piles The values of Kstano measured during driving
of piles after Armishow (1981) and from test results after
Hultsjo (1969) are also plotted in Fig 15
25
The variation of Kstano for steel piles according to the
above criteria (o=20deg and M=06) and the values of Ks
according to Broms are also presented in Fig 15 In the
case of steel piles with o = 20deg K tano varies verys little with the values of angle of internal friction
Since o can be greater than 20 0 (Table 9) micro = 06 can
be chosen in the evaluation of o for steel piles
The rough estimation Kstano = 03 (Hultsjo 1969) reshy
commended in practice can be very conservative in practical
design This value can be used only in the case of loose
sand or in silty soil where high excess pore pressures
can develop when it is loaded
As the shaft resistance requires very small displacement
to mobilize the ultimate resistance therefore a smaller
safety factor can be applied in this case in comparison
with the case of point resistance It is recommended that
the safety factor is equal to 15-2 to obtain allowable
shaft resistance from the ultimate value calculated
(4) In general small variations of the friction angle
~bull of sand considerably influence the values of Ks and
N It is therefore generally preferable to use the q
results of subsurface sounding by means of penetration
tests for preliminary estimations of the point resistance
and the skin friction of piles in non-cohesive soil
The relation between two methods will be discussed later
26
Table 9 Coefficient of friction between various construction material and sand (Potyondy 1961)
qgt I 0 ITest material tano micro= tanlt)
Polished smooth steel 37deg 23deg30 057in saturated dense sand
Rough steel in dense sand 43deg30 33deg40 07
Pine wood in saturated dense sand test run parallell to 37deg 33deg 086 the grain
Ditto but test run in perpendicular to the grain 37deg 34deg20 1 091
Smooth concrete (made in wood form) in saturated 37deg 33deg20 1 087 dense sand
Rough concrete in dry 43deg30 42deg30 097dense sand
27
13 Methods based on cone penetration tests
Static sounding is used extensively in the Netherlands
France and Germany to estimate the bearing capacity of
friction piles in sand Examples of static penetrometers
are the Dutch cone penetrometer penetrometers type
Barros Nilcon and the Swedish weight sounding device
Different criteria on using results of static penetration
are summarized below
1 bull 3 1 The Canadian Manual (1978)
The ultimate load capacity is evaluated by
where Qf = ultimate pile load (kN)
q = unit point resistance from cone test kPa p
For a pile with a diameter Bgt050 m q is equal to the minim~m measured qp P
f = average skin friction measured by cone test kPa s 2A = cross-sectional area of the pile tip m
2Ap = surface area of the pile shaft m s
The safety factor is equal to 25 to 3 depending on the
number of cone tests performed and on the observed varishy
ability of the test results
1 3 2 Meyerhof (1976)
When the pile is driven into a stratified cohesionless
soil the ultimate point resistance and the skin friction
can only be calculated by semi-empirical methods from
static cone When the pile point is above the critical
depth in the bearing stratum the unit point resistance
has to be reduced from the limiting static cone resistance
q in proportion to the embedment ratio DBB
qc DB qp = 10B ~ ql
where qp = the unit point resistance of the pile
= limiting unit point resistanceq 1
The values of q 1 derived from the limiting static cone reshy
sistance q are shown in Fig 16 C
C
28
For bored piles roughly one third to one half of the
static cone point resistance and friction resistance
applicable to driven piles may be used for preliminary
estimations in cohesionless soil
For penetration of piles shorter than about 10B into
the bearing stratum the value q may roughly be estimated p by
in which q and q are the limiting unit point resistance 0 1
in upper weak and lower firm stratum respectively (Fig17)
When piles longer than about 15 to 20 pile diameter are
penetrated through a weak stratum into a thick firm deshy
posit the ultimate point resistance is increased with
DbB in this stratum but below the critical depth (10B)
the point resistance remains practically constant at the
limit value q 1 for this stratum If the pile point rests
in a relatively thin firm stratum underlain by a weak
deposit the ultimate unit point resistance in the bearing
stratum may be governed by the resistance to punching of
the piles into the underlying weak soil (Fig18) and
in which q and q are the limiting unit point resistance 0 1
in the lower weak stratum and the upper firm stratum reshy
spectively
1 bull 3 3 Vesic (1975)
The shaft friction can be determined by qc using the
following relationship
f S
= pqC
where P = 0 11 ( 1 0) -1 3tan
and ~=angle of internal friction expressed in terms of total stress
29
In case the penetration resistance is lower than 05 MPa
p can be determined by
3 p = --
Irr
where Irr is the reduced rigidity index (Vesic 1967)
and defined by
I rr =
~ = volume change
Ir = rigidity index (relative compressibility of the sand mass)
If no volume change (undrained condition) or little volume
change takes place
E ( 1 +v) (c+qtancp)
1 bull 3 4 The Thorburn method
Thorburn (1980) has recommended that the ultimate shaft
resistance can be determined by
Q = As qcs s 200
where = the embedment area of the pile shaft
the average static cone penetration resistance within the depth of embedment of the pile in the sand stratum
The ultimate base resistance is calculated for two cases
a) Deep embedment The pile penetration is at least 8B
into bearing stratum
Q = (025 q +025 q +05 qc )App Co C1 2
where = average static cone resistance over a distance of 2B below the pile base
= minimum cone resistance over the same distance below the pile base
= average of the minimum cone resistance over a distance of 8B above the pile base neglecting any value greater than q
C1
30
b) Shallow ernbedrnent The pile penetration is only 1-2B
into a fine grained non-cohesive soil
Q = (05 q b+05 q ) A p c ea p
= average q over a distance of 35B below the pile 5ase
qc +qc + qcnqcb = 2n +nqcn
= the cone resistance at regular intervals of 35B below the pile base and q is the lowest resistance en
= average cone resistance over 8B above the pile base neglecting any value greater than q bull en
1 bull 3 5 Te Karn WGB method (1977)
a) Skin friction In compression
q in MPa C
The resulting friction values are limited to 012 MPa
The method is based on the observed constant ratio beshy
tween q and the local sleeve friction measured in the C
cone penetration test L
and Q = TID 2 f f dx S O S
b) End-bearing
where I II and III are illustrated in Fig19
Generally in normally consolidated sands the CPT conshy
siders a limiting unit pile end-bearing capacity of 15
MPa However in overconsolidated sands a further limitation
might be applied (see Fig20) There are observations that
qcfloc ~ const floe is difficult to measure thus
according to Begernann f = q 1301OC C
31
In the case of the pile in tension L4 TID 3L4 L
Q = TID f fs dx + J fs dx + TID J f dx3s o o 3L4 s
The relationship between the point resistance q and the C
friction resistance f for a penetrometer measured with s
a separate friction sleeve is shown in Fig21 after
Begemann (1965) From the relationship shown in Fig21
it is possible to estimate the grain or particle size
when the point resistance and the skin friction resistance
are measured separately It is furthermore possible to
estimate the skin friction resistance from the point reshy
sistance if the average grain of the soil (gravel sand
silt or clay) is known Broms (1981) has reported the
ratio f q = 05-10 for sand and gravel and aboutS C
4-6 for clays
1 bull 3 7 The Nottingham and Schmertmann (19751977) method
a) Ultimate skin friction L
fsloc dx + J fslocnolO
82-8B
Qs =KSC 8B
or 8B L Q = K ( I f A + I f A )
S SC sloe s slos sl=0 8B
where 1 = depth to the f value considered s
A = pile soil contact area per f depth interval s s
Resulting skin friction values are limited to 012 MPa
f = unit local friction sleeve resistancesloe K = correction factor of the unit local friction s f in a sand layersloe K = correction factor of the unit local friction
C f in a clay layersloe K = function of the soil type pile material typesc
of cone tip etc see Fig 22
Beringen et al (1979) have suggested K-values as follows
In compression fsloc = 0007 qc (1143 qc)
In tension fsloc = 0005 qc (1200 qc)
32
Side friction calculation can be simplified if the
sleeve friction resistance does not vary significantly
with depth provided that L gt 8B
QS
= K frac12(f AS1
)0 8B +(fS
A )8B-
LS - S2
If two or more sand layers are invloved the above equation
can be used by considering each layer individually as shown
in Fig 23 The K-value should be the same for each layer
b) End-bearing
Fig 24 gives the basis for the calculation of the ultimate
pile end-bearing capacity using penetration data from a
Fugro-type penetrometer The procedure remains the same
for piles embedded in sand clay or mixed soil
138 The Norwegian Pile Committee (1973)
1381 Cone penetration test
The bearing capacity of friction piles in sand can be
calculated from CPT
fs = shaft friction factor depending on the soil density = 0005 q at q about 10 MPa = 001 q cat q cabout 25 MPa
C C
qcs= average static cone penetration resistance along the pile
point resistance factor= 05
static cone penetration resistance close to the pile point
1 382 Weight sounding
The bearing capacity of piles in sand can also be estimated
from Swedish weight sounding Fig 25 shows the relationshy
ship between the ultimate bearing capacity Qf the surface
area of the pile and the penetration resistance of th~ soil
expressed by the number of halfturns02 m of the Swedish
weight sounding The indicated relationship is based on
33
results from pile load tests In the evaluation of th~
test results the point resistance of the piles has been
neglected Normally the point resistance accounts for
almost half of the total bearing capacity The relationshy
ship shown in Fig 25 indicates that the ultimate bearing
capacity will be about 20 higher for timber piles than
for precast concrete piles of the same surface area of
the piles and the same initial relative density of the
soil As a guide in practice the following values can
be taken
When Dr is low (lt5 halfturns02 m) fs ~ 20 kPa for concrete piles
when Dr is medium (5-15 halfturns02 m) f ~ 35 kPa s and when D is high (gt15 halfturns02 m) fs ~ 50 kPa r
The diagram in Fig 25 is valid for ordinary timber and 2concrete piles with a cross sectional area of 500-1000 cm
and a length of 12-15 m
The ultimate bearing capacity of concrete piles in sand
can also be determined from the formula
= Af(00138 + 00016 P )Qu we = bearing capacity MNQu
2 = total shaft area mAf p = average number of halfturns per 02 m penetrationwe of the Swedish weight sounding penetrometer
It is also possible to calculate the ultimate bearing
capacity from the relationship between penetration reshy
sistance of the Swedish weight sounding and penetration
resistance of the static cone penetration test (Nht=3qc)
139 The United Kingdom experiences (Rodin et al 1974)
The ultimate bearing capacity of piles is calculated
34
Point resistance
1 Gravel for preliminary calculations piles in gravel
are designed as end-bearing and the following formula
is used as a guideline
Q = A n N (kN)p p
A = area of the pile base p
n = ratio of q to N C
N = standard penetration resistance at the pile base corrected using the Thorburn chart (Fig2)
2 Sand and silty sands
Using the Meyerhof (1956) formula
Qp = Apqc
q = average static cone penetration resistance c near the base as follows
for loose sand 35 B above pile base (acc to Van der Ven)
and 1 B below pile base
for medium sand80 B above pile base (acc to
35 B below pile base Begemann)
The skin friction is calculated
1 Total friction measured use the same formula as has
been recommended in Thorburns method (see 134)
and for fine grained non-cohesive soils (acc to
Thorburn et al 1970)
Q = s
2 LocaJ friction measured use the recommendations by
Begemann (136)
1310 Discussion and conclusions about the methods based on results from CPT
According to the review on the design of piles using reshy
sults from static cone penetration tests the following
points can be summarized
35
(1) The static cone penetration test is a useful tool
for prediction of bearing capacity of piles in nonshy
cohesive soils The CPT compared to the SPT has the
following advantages it is more reliable more reshy
peatable giving valuable information on soil stratishy
graphy The results from cone penetration tests can be
related to many properties of the soil (density friction
angle Youngs modulus bull )
(2) The result of cone resistance is uncertain for silty
soils especially in the case that the soil is below the
ground water It is recommended to take samples in those
cases
(3) The result from static cone penetration resistance can
be used directly in the calculation of the point resistance
of the pile if qc lt 10 MPa and it is equal to 05 qc if
qc gt 10 MPa (see table 10)
(4) The value of q is calculated from the average value p
of q over a specific depth (see Table 11) For practicalC
purposes the criteria of Begemann and Van der Ven can
be taken into account for medium sand and loose sand
(Table 10) the same criteria is recommended by Broms (1978)
(5) For piles with a diameter greater than 05 m the value
q is less than q It is recommended to make a reduction p C
q in those cases For instance the minimum value of C
for calculation of q can be used p
(6) Experiences show that the values of qp is different
from q when q is greater than 15 MPa Thus an upperC C
limit value of q 15 MPa should be considered for p
practical purposes (see Table 12)
(7) It is recommended to use the static cone penetrometer
with local measurement of the skin friction Based on
Begemanns diagram the grain particle size can be deshy
termined and serve as a double check of the result from
36
the cone penetration test if the average grain size of
the soil is known
(8) The value of the unit skin friction can be evaluated
from the point cone resistance The value of f is 03-s 20 of the q value (Table 13) In general
C
f = 05 qc for dense sand (qc 1 0 MPa)s f = 1 qc for loose sand (qc ~25 MPa)s
where q is the average value along the pile lengthC
(9) The value of the unit skin friction is 025-05 of
q for piles in tension C
(10) If the local friction is measured by a sleeve the
unit skin friction is 05-10 of the value of the sleeve
resistance For practical purposes f = 07 f as s s 1oc
recommended by Begemann can be taken
(11) For bored piles the unit skin friction is 13 of the
values considered for driven piles
(12) The limit value of the unit skin friction is equal
to 012 MPa for silty soils For fine sand this value is
010 MPa
(13) The ultimate bearing capacity of piles can be evalushy
ated from results from the Swedish weight sounding penetro~
meter tests The methods based on cone penetrometer can be
used if the relationship between penetration resistance
from CPT and weight sounding tests is known In general
Nht = 3 qc where Nht is the number of halfturns and qc
is the point cone resistance MPa
(14) The ultimate bearing capacity of piles can be evalushy
ated from SPT test results and the relationship between
SPT and CPT (see Appendix A)
37
(15) It is important to carry out further comparisons
between CPT tests and loading tests The following
conditions have been fulfilled
- The CPT test has to be carried out before the pile
test and in the same position
- The type of CPT test and the tip geometry must be clearly
defined
- The geometry of the pile and its way of production must
be clearly described
- The failure criterion adopted for interpreting the test
load result shall be given
(16) The disadvantage of CPT is limited capacity to penshy
etrate dense layers For estimating the bearing capacity
of the pile it is necessary to carry out the penetration
test at least to a depth 4B below the pile tip It is
recommended to make a combination between CPT and SPT in
those cases A new type of static-dynamic penetration
equipment developed at SGI will be a good tool for this
purpose
(17) The CPT is not widely used in Sweden for calculation
of the bearing capacity of piles This is due to the fact
that 80 of the piles in Sweden are driven to a very hard
till or to the rock surface
(18) The factor of safety to apply to Qf should be 25
to 30 depending on the number of cone tests performed and
on the observed variability of the test results The minishy
mum factor of safety corresponds to a larger number of
results with a variability of less than plusmn10 from the
average
38
Table 10 Summary of the criteria for calculation of the unit point resistance q from the average value of cone penetration r~sistance
t
Reference qp Note
Norwegian Pile Committee 05 qc(1973)
Begemann (1965 05 qc Vesic (1977) qp=qc if qc lt 1 0 MPa
Broms (1978 1981) 05 q if qc gt 1 0 MPa C
Table 11 The depth to consider in the calculation of qp from the average value of static cone penetration resistance qc
Reference Depth above Depth below Note pile tip pile tip
Van der Ven 35B 1B Loose sand
Begemann 8 B 35B Dense sand
Te kam (1977) 8 B 07-4B 07B if qc in-creases with depth 4B if qcvaries
Broms (1978) 35B 1B lt 10 kPaqp Thorburn (1979) 8 B 2B Considers other
values and the minimum of qc
B = diameter of the pile
39
Table 12 Limit values of unit point resistance and unit skin friction
Author Limit value Limit value Note of qp of fs
Broms (1978) 1 0 MPa
Te Kam (1977) 15 MPa 0 12 MPa
Begemann ( 1965) Q f 12 MPa
Beringen ( 1979) 15 MPa 025 MPa
Ton et al 0 1 0 Silty and fine ( 1 981) sand
Table 13 Values of unit skin friction
Reference Values of fs Note
Meyerhof (1956) f = 05 qcs
Meyerhof(1976) f = f (with sleeve)s sloe
Norwegian Committee
Pile (1973)
f fs
s
= 05 = 1
qC
qC
for for
dense loose
sand sand
qc qc
~
10 25
MPa MPa
Te kam (1977) f fs
= 033 = 025
in compression in tension
s
Begemann (1977) f fs
s
= 08 qC
= 07 f l s oc (with sleeve)
After Beringen (1979)
Nottingham (1975) f fs
s
= 07 = 05
qc qc
for for
compression tension
After Beringen
Thorburn ( 1975) f s = 05 qc
Vesic (1977) f s = 0 11 (19)-13 tan~
qc
Tong (1981) f = 2 q If q lt 5 MPa fs = 01 MPa If qc c gt 5 MPa
s
Broms (1978 1981) f = 05 q Dense sand the s C
relative densitJ is high
f = 0 1 MPa Loose sand s
40
1 4 Relation between methods based on static formula and static cone penetration test
1 bull 4 1 Determination of Kstano by CPT-test
The relation between two methods can be derived based on
the following assumptions
The point resistance of th~ pile qp is equal to the
point cone resistance q q = qC p C
- The local shaft resistance f from the cone penetrations
test can be used directly for calculation of the unit
shaft resistance of the pile fsl = fs
or in other words the CPT represents a pile model test
the same failure pattern in the two cases
The unit skin friction can be evaluated by
f = K tano a s s V
or af K tano JI= s s 2
where a~ is the effective vertical overburden pressure
at the pile tip considering that the length of the pile
is equal or less than the critical depth The unit point
resistance can be calculated by
q = N a p q V
The ratio between f and isqps
f Kstano s fs = or K tano = 2 N
2 N s qqp qcq
The value of Kstano can be derived if N and f q are q S C determined
The values of N can be defined if the relative densityq or the angle of internal friction are knownmiddot The value
off q can be de~ived as followsS C
41
(1) Empirical ratio off qS C
In case of dense sand f q = 005 Thus K tano = 001 N S C S q
and in case of loose sand f q = 0001 thus K tano = S C S
002 N bullq
( 2) Determination of K tano from CPT test with measurings
of friction
The ratio off q can be defined from the CPT test S C
(3) Determination off q by CPT test with measuring of S C
q and using Begemanns curve (Fig22) or Figs 31-33 C
1 bull 4 2 Relation between Qs and Qp
From the relation between CPT and static formula the
ratio of Qs and Qp can be written as
9s = 4 Js Ngs ( 2L-Dcr)Qp B qp Nqr
where B = diameter or width of the pile
N N = bearing capacity factor of the shaft qs qrand the pile point
L = total length of the pile
= critical depth
= fs = R = ratio of skin friction andzrc- f point resistance
If the soil is homogeneous N = N and qs qp
If the length of the pile is less than the critical depth
The accuracy of the ratio~ depends on the accuracy of Rf
Those can be obtained from CPT test or estimated from
empirical expressions or empirical diagrams
42
The ratio Q Q can be used for rough extimation of the s p length of the pile and the load displacement curve
As a demonstration of the relation between the two
methods some examples and case records are presented
in 1 bull 4 3
1 4 3 Examples
Example No 1 Using the same data presented by Beringen
et al (1979) to evaluate the skin friction of two piles
The diameter of the first pile is cp = 356 mm and the length
of the pile is 7 m Fig 28 shows the result of the soil
investigation From the CPT values the ratios between fsl
and q are summarized in Table 14 for different depthsC
Table 1 4 bull Ratio between local friction and point resistance
Depth (MPa) f (MPa)qc fslqcsl
2 2 007 004
4 24 020 001
6 40 032 001
7 20 0 15 001
The ratio f q = 001 is used for calculation of the S C O
skin friction With cjl = 38 (from soil data) N = 127 q
(Berezantsev 1961) the value of K tano is calculated bys
K tano = 2bullN bull s_ s q qc
or
K tano = 2bull127middot001 = 254 s
The effective overburden pressures are calculated and
presented in Fig29
The skin frictionmiddotis evaluated by
-= K tano CJ I AQs s V s 0 3 5254middot1T 0356-35(~ +70+ ~- )Qs =
1235 kNQs =
43
The value of Qs
is approximately equal to 1310 kN whieh
is the value observed in the load test by Beringen et al
(1979) The value of Qp
is calculated by
= 0 1 N A V q p
where
0 1
V = 1085 kPa at the pile tip N
q = 127 A
p =01 2m
Qp = 1085deg127middot01 = 1371 kN
The measured value in the load test is somewhat smaller
1130 kN
Pile No 2 has the same diameter as pile No 1 356 mm but
the length of the pile is 65 m The pile is close-ended
A comparison of the measured and calculated pile capacities
is shown in Table 15
~xample No 2 (After Bergdahl and Wennerstrand 1976)
Two series of pile load tests have been performed at Albyshy
sjon The soil consists of cohesionless material mainly
sand the static cone penetrometer resistance is 2 to 3 MPa
down to 5 m and increases below that almost linearly with
depth and reaches 10 MPa at about 27 m below ground surface
The piles were steel pipe segments with~= 89 mm The
piles were pushed into the soil to different lengths and
after that they were pulled out the soil The results of
the test are presented in Fig 30 The ratio between the
specific skin friction resistance for the pile and the
point resistance of the static cone penetrometer gives
an average value of 00055
The shaft resistance of the pile is evaluated by two
methods
Pile Observed test Predicted value (kN) and predicted valueobserved test results test results
No Fugro-CPT method APT-method Begemann-method Nottingham-methoc Vesic method New method
kN kN kN kN kN kN kN
Skin frictior 1 1310 600 46 1 70 13 605 46 600 46 1235 94
2 1530 500 33 170 11 610 40 580 38 1084 71
Point 1 1130 1045 92 380 34 1000 89 13 71 120
resistance 2 1470 1500 102 370 25 990 67 1302 89
1 2440 1645 67 550 23 1650 68 1645 67 2606 107Total
2 3000 2000 67 540 18 2110 70 2080 69 2386 80
Table 15 Comparison of measured and calculated pile capacities
i
i
45
1 The Hultsjo and Svensson formula L
Q =AK tano fa dz S S S O Z
where K tano = 035 y = 10 kNm 3
s
A = TTBL s
The results are plotted in Fig 30
2 The relationship between the CPT and static formula
is
K tano = N fs s q qc
where N = 17(~~25deg) fs = 00055 q qc
K tano = 019 s
0 Vunit skin friction f = 019 2 and Q = A f
s 8 s s
The results are plotted in Fig 30
Note The results from the load tests (push) are similar
to the values calculated if K tano = 025 s
Example No 3 (After data of Bergdahl 1979)
The length of the pile= 152 m
Area of the point tip= 01middot01 = 001 m2
Shaft area of the pile= 4bull01middot152 = 608 m2
Submerged unit weight of the soil= 98 kNm 3
The ground water level is 05 m below ground surface
(1) Calculation according to the pile driving formula
(SBN 75) (see 15) the ultimate bearing capacity
of the pile is eavluated by
nh bull Q (1-01 Qp)e+c r Qrz
46
n = 10 (drop fall hammer)
h = height of fall of the hammer = 05 m
= weight of the hammer = 2 kNQr = weight of the pile = 1 9 kNQp
e = remaining penetration for the last blow = 3 mm
C = rebound of the soil = 4 mm
Qf = 08 10 middot~00 bull 2(1-01 ~) = 145 kN203 + 2
(2) Calculation according to Hultsjo et al (1969) L
Qf =NA bullO + K tano f ~ o dz p V O O Z V
A = area of the point tip = 001 m2 p
~2 = perimeter of the pile = 4 bull 0 1 = 04 m
N = K tano= 0358 0
o = effective overburden pressure at the pile tipV
o = yz = 152middot98 = 149 kPa V
221Qf = 8middot001middot149 + ~middot middot98middot04middot032
Qf = 12+145 = 167 kN
If the value of Ktano is reduced to 025 according to
Bergdahl et al (1979) the value of Qf is
Qf = 12+113 = 125 kN
(3) Calculation according to the relation between CPT and
static formula
Bergdahl and Wennerstrand (1976) have reported that the
ratio between the skin friction resistance of the pile
and the point resistance of the CPT is equal to 0005
The value is obtained from the two series of load test
piles at Albysjon Assuming this ratio for the relationshy
ship between the unit skin friction and the cone point
resistance or the relationship between the skin friction
of the pile and the point resistance of the pile the
value of Kstano can be derived In loose sand~ = 25deg
47
and according to Berezantzev et al (1961) N = 17 q
KstancS = 13 = 2middotN bull s_ = 2bull17bull00055 = 019 q qc
The ultimate bearing capacity is evaluated by
= N 0 1 A + frac12 0 1 bull K tancS A
q V p V S S
Qf = 17bull149middot001 + frac12 149middot017bull608
Qf = 2533 + 8606 = 1114 kN
(4) The load test gives Qf = 110 kN
Table 16 shows a comparison between measured and calculated
capacity of the pile
Table 16 Comparison between measured and calculated piles capacity
Results from Driving Hultsjo Bergdahl CPT and loat tests formula (1969) static
formula
Load (kN) 1 1 0 145 167 125 1 1 1 9c
QEred~cted 100 132 152 11 4 100 Qmesaured
1 4 4 Conclusion
Some examples above have shown the possibility to use
the relation between the CPT test and the static formula
to predict the pile capacity For this purpose some intershy
pretation charts after Searle (1979) are collected below
as a quide-line Relations between the two methods can
also be developed for cohesive soils
48
15 Method based on pressuremeter tests
When calculating the bearing capacity of friction piles
using results from pressuremeter tests the shaft reshy
sistance and the point resistance are added as usual
(Baguelin et al 1978)
1 bull 5 1 The point resistance
qp = qo + k(pl-po)
+ kqp = qo pl
AQp = kpl p
= measured limite pressureP1 = initial total horizontal pressurePo
p = net limit pressure determined by pressuremeter test1 k = bearing capacity factor
A = area of the point pointp
The bearing capacity factor k is dependent on a number
of different variables type of soil depth of embedment
shape of pile and method of installation The k factor
for driven piles is given by Baguelin et al (1978) see
Figs 34-37 or Table 17
The influence of method of installation on k is also shown
in Fig 38
152 Skin friction
The skin friction is also determined as a function of
the net limit pressure according to Baguelin et al (1978)
It is also dependent on the type of soil pile material
and method of installation The unit skin friction f s
for driven concrete and steel piles is given in Fig 39
Research in Sweden Sellgren (1981) shows a good agreeshy
ment between measured and predicted pile capacities
according to Baguelins method
49
Table 17 Bearing capacity factor (From Hansbo 1981)
Type of soil Net limit DB k max pressure Driven Bored(MPa)P pile pile
05 5 33 27Sand and gravel 1 6 48 42
2 8 68 57
4 83 70
6 1 0 90 73
Silt 01-03 2 22 18
05 3 27 24
1 4 30 27
2 45 34 3 1
3 5 36 33
Clay 01-02 2 2 18
1 35 29 26
2 4 34 30
4 5 38 33
16 Method based on pile driving formula
Pile driving formulas are often used for point bearing
piles or friction piles in sand During the driving work
it is important to check that all piles in a pile group
have been driven to the same resistance and that the piles
have not been damaged The ultimate bearing capacity of
a pile can be evaluated by the formula in SBN 75
Q = 08nh Q (1-01 Q Q )(e+cs)u r p r
= ultimate bearing capacity should be at least 3 times larger than Qa 11ow
n = correction coefficient for drop hammer
= 1 free falling hammer = 08 hammer of single line
= height of drop of hammer m
= weight of the hammer t
= remaining penetration for the last blow
C = rebound of the soil (see Fig40)
50
If it is not possible to measure the rebound it can
be assumed for concrete piles that pule
C = AE p p
1 e = Qrqp independent of the length of the pile
qp = the weight of pile per meter
lf = length of follower
ApAf = average cross sectional and the follower
area of the pile
EpEf = modulus of elasticity of follower
the pile and the
The formula can be used for concrete piles if Qall~450 kN
the penetration should exceed 2-3 mmblow and for timber
piles with Qall~150 kN and the penetration is about 4-5
mmblow
Limitation
The main limitation of these formulae is the calculation
of energy transmitted to the pile and the change of the
bearing capacity that takes place after the driving
- Indicates only the load distribution from the pile
and the resistance of the soil during the driving
- Can only be used for limited allowable load and limited
penetration
17 Method based on stress wave measurement
171 The Case method
The Case method developed at Case Western Reserve University
by Goble et al (1980)
The bearing capacity of a pile can be determined from
top measurements of force (from strain) and velocity
(from integration of acceleration) The soil resistance
R was first computed from the rigid body equation
51
R = F(t) - m a(t)
F(t) a(t) = the pile tip force and the acceleration
m = the pile mass
Later studies of the elastic pile produced
R = frac12 (F(t1)+F(t2) + ~~ v(t1)- v(t2)
where
F(t 1 ) = the force at the pile top at time t1
v(t 1 )
t2
t1
=
=
=
the velocity at the pile top
t1 + 2L C
some selected times during that the first maximum of force
at time t1
e blow (usually and velocity)
L = the pile length below the poiment
nt of measure-
c = velocity of wave propagation measured at the record or calthe relation
c can be culated from
C =
E = Youngs modulus
p = the mass density of the pile
The resistance R is made up from a static portion R and s
a dynamic portion Rd It is assumed that Rd is propor-
tional to the pile bottom velocity
EAR = J Vd c c bmax
where J is the damping constant depending on the soil C
type at the bottom of the pile recommended values are
(see also Hermanson amp Gravare 1978)
for sand 0-0 1 5
for sandy silt 015-025
for silty clay 045-070
for clay 090-120
The exact value for a special site can be obtained from
a static load test
52
From wave theory
V = 2 V - __ Rbmax top me
V = pile top velocity at time t 1top
The static resistance can be derived to
R = R - J 2Fti) -R I s max c maxJ
The correlation between the static capacity as obtained
in a static load test and the Case method is given in
Fig 41 bull Fig 4 2 shows the schematdc of instrumentation
A typical record of pile top force and velocity is shown
in Fig 43
1 7 2 The CAPWAP method
(Case Pile Wave Analysis Program)
The method is used in the laboratory schanatic of the
processing system is shown in Fig 44 Either pile top
force or pile top velocity can be used in a dynamic analyshy
sis as a boundary value This method divides the pile
into a number of mass points and springs (Fig 45) bull In the
computation a reasonable assumption is made regarding
the soil parameters and then the motion of the pile is
input using the measured top acceleration as a boundary
value The agreement between the measured and the calshy
culated forces can be improved iteratively by changing
the assumed soil resistance parameters (See Fig46)
The method provides a direct estimate of bearing capacity
and of skin resistance distribution No soil constant
needs to be assumed in order to make the analysis Fig
47 shows the result of a static load test (with and
without creep) compared to the CAPWAP computation
53
1 7 3 Other applications of stress wave measurements
(1) Integrity control
The force F and the velocity v caused by a stress wave
are related by constant FAc Resistance effects cause
the force to increase relative to the velocity A cross
sectional reduction causes the opposite effect The
impedance is defined by
I = me L
= EA -
C
and F(t) = v(t)I
If the pile changes its impedance from I 1 to I 2 the
impact wave having the force F will generate F l U
The corresponding velocity is
Given F and V through measurements one may also de-u u termine the lower cross section or impedance from
I2 = I Fi+Fnlpi-Fu
Figs 48-50 give an example of a pile which is breaking
during redriving
(2) Stresses Tension stresses occuring at a distance
below the point of measurement is given by
T (x) =frac12I (t2)-F(t2)-I (t3)-F(t3)V V
54
(3) Driving system performance
The energy transferred to the point of measurement can
be determined by using t
E(t) = f F(T) V (T) dT 0
where v(t) is the integrated acceleration The maximum
of the E(t) function occurs just before the pile reshy
bound starts
18 Methods based on load tests
The design of piles on the basis of theoretical or
empirical methods are subjected with some uncertainties
- The soil proterties cannot be measured with great
accuracy and are always variable within a building
site
- The correlation between the soil parameters and the
bearing capacity of a pile includes a margin of error
- The actual driving or placing condition varies from
pile to pile
Therefore the best method of assessing the bearing
capacity of the piles is to load test typical units
181 Methods of load testing
1811 Load testing with a constant rate of penetration
The method is usually employed to determine the ultimate
load and mode of operation of a pile
End-bearing piles should be loaded until failure or until
the maximum available load has been applied
Friction piles should normally be loaded to achieve at
least 60 mm of settlement
55
The rate of penetration is normally chosen at approxishy
mately 05 mmmin It is advisable to record the force
required and the magnitude of the movement every second
minute
Unloading is interrupted for 1-2 min ie the time
required for the rebound to be stationary and to permit
reading of the dial gauges when the load has fallen to
approximately 75 50 25 10 and 5 of the maximum
applied load Typical working curves are shown in Fig51
1812 Load testing with stepped load increments (the ML method)
This method is designed to permit determination of the
ultimate load and the creep load of the pile
The load is increased every 15 minutes by a constant
amount approximately 5 of the estimated ultimate load
generally rounded of 10 20 50 or 100 kN Dial gauges
are read 3 6 9 12 and 15 min after application of a
new load
To be able to evaluate the creep load the load measureshy
ment and the maintenance of a constant load must be
carried out with great accuracy
1813 Cyclic load testing
This method is designed to permit the determination of
the creep load of the pile The load test is carried out
by alternating the load between a high load and a low
load Each load is maintained for 10 minutes One cycle
comprises a period with a high load directly followed
by a period of a low load and is thus of 20 minutes
duration The high load is generally twice the lower
one It is advisable to measure and record the pile settleshy
ment at the head of the pile very two minutes
56
1814 Long-term load testing
Long-term load testing is designed to check the creep
of the pile Load durations of the magnitude of one day
up to six months or a year may be required
1 8 15 Combination of methods of load testing
18151 Stepped load increment combined with constant
rate of penetration loading
This combination is often used in Sweden The constant
rate of penetration loading is carried out in the vicinity
of the ultimate load This provides a continuous and
detailed load settlement curve
The load testing is carried out in accordance with 1812
until the load under which the head of the pile sinks
approximately 01-02 mmmin during the later part of a
15-minutes period is reached The pile is then loaded to
achieve a constant rate of penetration of 05 mmmin
1 bull 8 2 Reporting of load testing results
The scale which is chosen for the presentation of the
results will also in many cases influence the calculated
ultimate strength as illustrated in Fig52 The recommenshy
dation of the Commission on Pile Research (Report 59
1980) is as follows
The load should be set out along the horizontal axis
and the settlement of the pile head (and pile tip)
along the vertical axis using a load scale of 10 kN =
1 mm and a linear settlement scale of 1 mm measured=
2 mm on the graph The scales may be varied but the
relationship between them must be maintained Example
of results (working curves) are shown in Fig53 and
Fig54
The creep load (using 1812) should be plotted under the
later part of the load steps (for example the settlement
taking place between the readings taken 9-15 or 12-15 min)
One example is shown in FigSS
57
The Canadian Manual (1978) recommends that the scale
for the load and the settlement should be selected so
that the line representing the elastic deformation 6
of the pile will be inclined at an angle of about 20deg
to the load axis The elastic deformation 6 is computed
from
1 bull 8 3 Failure criteria
1831 The Canadian Manual (1978)
The failure load Qf of a pile is the load which produces
a settlement of the pile head equal to
Bs=6+3()
s = settlement at failure m
B = diameter of a pile m
6 = elastic deformation m
Fig 56 presents the failure criterion
1832 Kezdi (1975)
Some rules to determine the design load from pile test
diagrams are summarized in Fig57 Kezdi (1975) reshy
commends the criterion (3)
1833 Vesic (19751977)
Some failure criteria summarized by Vesic (19751977)
are presented in Table 18 Vesic (1977) recommends the
criterion 16 to be used in the following corrected form
- Unless the load settlement curve of a pile does not
show a definite peak load the ultimate load is defined
as the load causing total pile settlement equal to 10
of the point diameter for driven piles and 25 of the
point diameter for bored piles
58
1834 90-criterion
The ultimate load is often defined in Sweden and Denmark
from a criterion where the deformation of the pile head
at 09 Qult is half the axial deformation at the ultimate
load
1835 The Commission on Pile Research (1980)
A new definition of the ultimate bearing capacity is
+ PL + 20 (mm)0ult = B
20 AE
where settlement of the pile head at the0ult = ultimate load (mm)
B = deiameter of the pile (mm) (For square cross section B = 113 x the side of the cross section of the pile)
PLAE= Compression ofthe pile when it is loaded as a column
The simplest method of determining the ultimate load is
shown in Fig54 where
B a= 20 + 20 (mm)
1836 Values form practical experiences
Tuoma and Reese (1974) Andreasson (1976) Berggren (1981)
have reported that the deformation at the bearing of the
failure is 5 of the pile diameter For loose non-cohesive
soils the settlement is about 10 according to Berggren
(1981)
1 bull 8 4 Distribution of the bearing capacity between
the tip and the sides of the piles
A reliable determination of the distribution between the
bearing capacity of the tip of a pile and its sides is
only possible if the pile is fitted with a tip-force
gauge and if a load testing with subsequent test pullingshy
out of the pile is carried out
59
Table 18
Rules for determination of ultimate load
1) Limiting total settlement
a) absol 0-1te 10 in (Holland New York Code) b) relative 10 of pile tip diameter (England)
2) Limiting plastic settlement
025 in (AASHO) 033 in ( Magne 1 19 4 8) 050 in (Boston Code)
3) Limiting ratio plastic settlementelastic settlement
15 (Christiani anJ Nielsen)
elastic settlement increment4) Maximum ratio plastic settlement increment
(Szechy 1961 Ref 15)
5) Limiting ratio settlementload
a) total 001 inton (California Chicago) b) increment al 003 inton - Incremental (Ohio)
005 inton - Incremental (Raymond Co)
6) Limiting ratio plastic settlementload
a) total 001 inton (New York Code) b) increment al 003 inton (Raymond Co)
settlement increment7) Maximum ratio load increment
( Ve s i c 1 9 6 3 Re f 16 )
8) Maximum curvature of le w log Q lire
(De Beer 1967 Ref 17)
9) Van der Veen postulate (1953) (
~ ( - _-__)w = 1bull ln 1 Qmax
(From Vesic 1977)
60
By measuring the compression of the pile during load
testing it is possible to assess at which load the
entire load increment is carried by the pile tip
(Fig 58)
When the ratio i~ (compression of the pile in relation
to applied load increment) becomes almost constant the
load increments are transferred directly to the pile tip
Weele (1957) stated based on observations that
- after a certain settlement of the pile the total skin
friction remains constant and the increase in load at
the pile point is the same as the pile top (see Fig54)
- If only the displacement of the pile top is measured
the elastic compression of the pile together with the
elastic compression of the soil can be obtained by means
of the recovery of the pile top during unloading (Fig59a)
As the elastic compressions have a linear relation to the
pile load after the skin friction reaches its ultimate
value The skin friction corresponding to every pile load
can be obtained drawing a straight line through the origin
parallel to the linear part of the load elastic compression
(Fig 59b)
1 bull 8 5 Safety factor
SBN 75 recommends that load tests are carried out at a
constant rate of penetration (CPR) and that the allowable
load corresponds to 25 of the ultimate load with respect
to soil failure
1 8 6 Comparison between different methods of load tests
Bergdahl and Hult (1981) made the following conclusions
- The ultimate bearing capacity determined by CRP tests is
about 10 higher than that by ML tests
- The cyclic load test reduces the ultimate bearing capacity
by about 15
61
1 bull 8 7 Comparison between the creep load and the
ultimate load of piles
According to the Commission on Pile Research (Report 59)
the bearing capacity of a pile in soil may be expressed
as an ultimate load or as a creep load
- The ultimate load of a pile is the load at which failure
is reached along the sides of the pile and under the tip
- The creep load of a pile is the maximum load that can
be applied without greatly increased deformation under
continuous loading or cyclic loading
The creep load is normally assumed to be 80 of the ultimate
load determined from constant rate of penetration test
A large difference between these values may indicate that
the pile is bent
- Berggren (1981) and Sellgren (1981) have recommended
the creep load (Fig 55) as the interpretation of a
failure because it is in general best defined
62
2 BEARING CAPACITY OF PILE GROUPS
Piles in a group in granular soil develop a larger load
capacity than isolated piles their group efficiency is
greater than 100
Influence of spacing and pile cap (the Canadian Manual)
Piles in group
- Act as individual piles if s gtgt B
- Act as a group at 25B lt s lt 7B
- Should not be installed at s lt 25B
SBN 75 Spacing center to center
L (m) S
lt10 3B
1025 4B
gt25 SB
Fig 60 shows the relation between the pile group efficiency
and the pile spacing according to different authors
Fig 61 after Kezdi (1975) also presents the relation
between the efficiency n and the spacing of the piles in
group The results of different authors confirm that the
efficiency factor for a pile group in sand is greater than 010 (except for ~=45) For practical purpose it is re-
commended to taken= 10
63
3 bull SETTLEMENT OF A SINGLE PILE AND A PILE GROUP
3 bull 1 Settlement of a single pile
3 1 bull 1 Empirical method the Canadian Manual
For normal load levels the settlement of a displacement
pile may be estimated from the empirical formula (Vesic
1970)
S = settlement of the pile head cm
B = pile diameter cm
o = elastic deformation of the pile shaft cm
= 100 QLp_0 AE
A= average cross sectional area of the pile m2
E = modulus of elasticity of the pile material in kPa
The ultimate or failure load produces a settlement approxshy
imately 3 times greater
For a bored pile S 1 = 35
3 1 bull 2 Settlement from load tests
Settlement during a load test (ML) can be considered as
representative of the long term behaviour of the pile
This method is recommended by the Canadian Manual
3 1 3 Method based on the theory of elasticity
3131 Mattes and Poulo s method
The settlements of a single pile in an elastic medium
can be estimated from the following relationship as
proposed by Mattes and Roulos (1969) and Poulos (1972)
s = Q I EsL s
where Q is the applied load Es is the elastic modulus of
the soilLis the pile length and Is is an influence factor
that depends on the length and diameter of the pile
64
Poissons ratio of the soil and the stress distribution
along the pile (see Fig62) For rough estimates I = s
18 can be used The most difficult of these factors to
evaluate is E and the stress distribution along the s
pile
For the case of an end-bearing pile on rigid stratum
PLs = EA MR p p
where MR is the movement ratio (see Fig63) depending
on the pile stiffness factor K determined by E
K = l RE a
s
E = Youngs modulus of the pilep E = Youngs modulus of the soil
s
Table 19 Typical values of E (driven piles) (From Poulos 1977 ~ustralian Standard 1978)
Soil E (MPa)s
Loose sand 42
Medium sand 70
Dense sand 80
Medium gravel 200
nd 2
Ra= Area of the pile section - 4-
The method based on the theory of elasticity is recomshy
mended by the Australian Standard (1978) Vesic (1975)
and Broms (1981)
65
3132 The method of Berggren
Berggren (1981) had proposed a relationship between the
pressure and the settlement for a circular rigid foundshy
ation on non-cohesive soils
S-1 s TT E (1-V) 2 1 1+)= -- -- a + a(p-0 1
)B 04 m (1-2v) 03 1-V O 0
s = settlement
B = diameter of the foundation
p = contact pressure MPa
m = compression modulus number
v = Poissons ratio (030-035 according to Barkan et al)
0 1 = initial effective overburden pressure MPa 0
a= pressure factor (varies with SB from model tests)
S = pressure exponent
The equation can by model scale tests be used to establish
the relationship between the pressure and the settlement
for foundations of different diameters and different foundshy
ation depths Values of Sand m can be evaluated by the
empirical expressions (Andreasson 1973)
m = 295 C -078 e -264 U 0
S = 029 log (i~ 1middot)-0065 log Cu
where C = uniformity coefficient dso u d10
e = void ratio 0
dso= grain size in mm (diameter corresponds to 50 of the soil)
The expressions are valid for dsolt5 mm
A good agreement between measured and computed contact
pressure and settlement has been obtained by Berggren
(1981)
66
3 1 bull 4 Vesics method (1977)
s = s + s + s e s p
s = elastic compression of the pilee
= settlement of the pile caused by loadSS transmitted by shaft friction to the soil
s = settlement of the pile due to the point load p
s = (Q + Qs) L
e p AEp Qs Qss
s =
B qo
and C C s = p p
p D qo
where = pile settlement coefficient dependingCscp on soil type and method of installing the pile
B = diameter of the pile
D = depth of pile embedment
= actual point and skin loads transmitted by the pile in the working stress range
= ultimate point resistance
The value of C is given in Table 6 for total long-termp settlement of the pile where the bearing stratum under
the pile tip extends at least 10 pile diameters below the
pile point
L = length of the pile shaft
A= cross sectional area
a= factor depending on the distribution of skin friction 06 in sand
E = modulus of elasticity of the pilep
The value of C can be evaluated from s
C =(093+016 Vr57B)cs p
The values of C are listed below p
67
Table 20 Typical values of the coefficent C p
Soil type Driven pile Bored pile
Sand (dense to loose) 002-004 009-018
Silt (dense to loose) 003-005 009-012
Clay (stiff to soft) 002-003 003-006
315 Method based on the Finite Element Method (FEM)
Stress distribution along a loaded pile can be determined
by FEM using the non-linear stress~strain relationship
The settlement of a single pile can be determined with a
given condition of the load the geometric of the pile
boundary and stress-strain relationship of the soil
The accuracy of the method depends on the degree of accuracy
of the parameters of the soil
32 Settlement of a pile group
321 The method of Skemton
The settlement of a pile group is often calculated by the
method proposed by Skemton et al (1953) This method is
based on settlement observations of actual structures
The observations indicate that the settlement will increase
with increasing size of pile group and the settlement of a
pile group S is always larger than that of the in-group dividual piles forming the group
s _ ags group
S = settlement of a single pile under its allowable load
a = group settlement ratio a function of the g dimension of the group and of the pile spacing or of the ratio BD of the width of the pile group to the diameter of the piles as follows
68
Table 21
BD 1 5 1 0 20 40 60
a 1 35 5 75 1 0 1 2 g
322 The method of Berezantsev
In the USSR and Poland the method proposed by Berezantsev
et al (1961) is used The settlement is assumed to inshy
crease linearly with the width of an equivalent area
located at the base of the piles The stress distribution
within the pile group is assumed to correspond to an angle
cp 4
323 Method based on theory of elasticity
The Australian Standard recommended Poulos method for
evaluation of the settlement
S = R S group s
S = settlement of a single pile
Rs = settlement ratio (see table 22)
For a number of piles different from 16 R can be s
extrapolated
R25 = settlement ratio for a group of 25 piles
R1 s = settlement ratio for a group of 15 piles
n = number of piles in a group
--
-- ----
--
69
TABLE 22 a THEORETICAL VALUES OF SETTLEMENT RATIO R FOR
FRICTION PILE GROUPS WITH RIGID CAP ON DEEP UNIFORM SOIL lIASS
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Lid sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 if
2 183 225 254 262 278 380 442 448 376 549 640 653 475 720 848 868 10 5 140 173 188 190 183 249 282 285 226 325 374 382 268 398 470 475
to 121 139 148 150 142 176 197 99 163 214 246 246 185 253 295 295
2 199 214 265 287 301 364 484 529 422 538 744 810 540 725 1028 l 125 25 5 147 174 209 219 198 261 348 374 246 354 496 534 295 448 650 703
10 125 146 174 l78 149 195 257 273 174 246 342 363 198 298 428 450
2 256 231 226 316 443 405 411 615 642 614 650 992 848 840 925 1435 100 5 188 188 201 264 280 294 338 487 374 405 498 754 468 518 675 1055
10 147 156 l76 228 195 217 273 393 245 280 381 582 295 348 500 788
TABLE 22 b THEORETICAL VALVES OF SETTLEMENT RATIO R FOR
END-BEARING PILE GROUPS WITH RIGID CAP BEARING ON A RIGID STRA TCM
Settlement ratio R Length Spacing
dia dia Number of piles in group n ratio ratio
4 9 16 25
Ld sd Pile stiffness factor K
10 100 1000 00 10 100 1000 00 10 100 1000 00 10 100 1000 00
2 152 l14 100 100 202 131 100 100 239 149 100 100 270 163 100 100 10 5 l15 108 100 100 123 112 102 100 130 114 102 100 133 115 103 100
10 102 101 100 100 104 102 100 100 104 102 100 100 103 102 100 100
2 188 162 105 100 284 257 116 100 370 328 133 100 448 413 150 ICO 25 5 136 136 108 100 167 170 116 100 194 200 123 100 215 223 128 100
lO 114 115 104 100 123 126 J06 100 130 133 107 100 133 138 J08 100
2 254 226 181 100 440 395 304 100 624 589 461 100 818 793 640 100 100 5 185 184 l67 l00 271 277 252 J00 354 374 347 l00 433 468 445 100
10 144 149 t46 100 l84 199 l98 100 221 248 253 100 253 298 310 100
70
324 Meyerhofs recommendation (1976)
Using the concept of an equivalent pier foundation the
settlement of a pile group in a homogeneous sand deposit
not underlain by more compressible soil can be determined
from the result of the standard penetration test
s = 2p VB I N
S = settlement in inches
p = net foundation pressure in tons per square foot
B = width of the pile group in feet
N = average corrected standard penetration resistance (blowsfeet)
I = influence factor of effective group embedment
DI= 1 - 8I3 gt 05
The settlement can also be evaluated from the result of
static cone penetration test
BI ps = 2q
C
q = average static cone resistance C
Both methods appear to give roughly reasonable estimates
for practical purposes Fig64 shows a comparison between
calculated settlements from these methods and observed
settlements of foundations supported by driven and bored
piles
325 Parrys method (1977)
S = 300 pBN
(mm) m
p = net pressure at the level of the footing (MPa)
B = width of the footing (m)
N = number of blows for 20 cm penetration with m the Swedish ram sounding or number of blows with SPT
71
326 The method of De Beer
The settlement can be evaluated by 0 1 0 1z
s = ~ 236z log( v ~ v V
where qc E
C = 15 (1 = CJ (E=15 q) kPa C
V V
point resistance from CPT test
effective overburden pressure at the level considered
60 1 = additional load V
327 The method of Schmertmann (19701978)
The settlement is evaluated from the expression 2B
s = C1C26P L (z_ 6z)o E
where E = 20 q (Schmertmann 1970)C
influence factor
1-05 ( t) = effective initial pressure at the level
of foundation
6p = net foundation pressure increase (p-p =6p)0
Note C1 takes account to the strain relief due to embedment
C2 = influence factor of the time rate in deshyvelopment of settlment in sand (creep effect)
The procedure of the calculation is the following
1 Obtain the static cone resistance (qc) profile over a
depth interval from the proposed foundation level to a
depth 2B below this B = pile diameter
2 Note the data of the pile foundation design (width
depth of embedment average foundation contact pressure)
3 Obtain the unit weight of surcharging soil
4 Divide the qc profile into a number of leayers each
with constant qc
72
5 Obtain E from qS C
6 Draw the assumed (sB06) triangular distribution for
the strain influence factor Iz (see Fig65) Calculate
(Ii 6z)
7 Calculate C1 C2 and the settlement
Experience shows that Schmertmanns method is better than
the method of Buisman and De Beer and is recommended by
many authors and codes
Recently (1978) Schmertmann et al based on results from
research by the Finite Element Method and model tests
(Fig66) recommended a new distribution of the strain
(Fig67) and the values of E calculated as follows
E = 25 qs(axisym) c
E = 35 qs(plane) c
This method has also been used to some extent in Sweden
Bergdahl (1980) and the experiences are so far rather
good but calculated settlements are on the safe sideIt
is interesting to note that Trofimenkov (1974) has reported
the value of E = 3 q according to experiences in the USSR C
33 Summary of the methods for calculation of settlement
1 Four common methods have been used for evaluation of the
settlement of single piles The empirical method (Vesic
1970) recommended by the Canadian Manual assumes that
the pile compresses elastically as a supported column
with no load transfer due to skin friction and the
settlement of the tip is constant for any normal pile
and is equal to 10 of the diameter of the pile The
method can be used as a guideline in practical purposes
The methods based on the theory of elasticity and FEM
requires good information about the properties of the
soil The accuracy of these methods for prediction of
73
the settlement will be very good because the variation
of different layers and boundary conditions can be taken
into account
The method recommended by Vesic (1977) and the method
based on load tests can be used for design purposes
Those methods will be in agreement with practical obshy
servations and dont need special information about the
data of the soil
2 Two approaches are generally used for evaluation of the
settlement of pile groups The method based on the settleshy
ment of the single pile or the equivalent area located
at the base of the pile group and the method based on
results from SPT or CPT The Schmertmann method is conshy
sidered as useful Practical experience shows good
agreement between the valvulated and measured settlements
74
4 CONSLUSIONS
(1) The SPT method can be subject to many errors during
the test It is recommended to use SPT in combination
with other methods In silty soil SPT will produce
high pore pressure it is therefore suggested to take
soil samples and carry out laboratory tests
(2) The critical depth could be taken into account for
evaluation of the bearing capacity of the soil The
point resistance is more important than the skin friction
in a non-cohesive soil than in a cohesive soil The
bearing capacity factor after Berezantsev (1961) Vesic
(1969) or Berggren (1981) can be used for evaluation of
the point resistance For calculation of the skin fricshy
tion the value of Kstano in Fig15 after Broms and the
recommendation of API can be used It is obvious that
a small variation of the values of angle of internal
friction will produce a big variation in the value of
bearing capacity of the pile Evaluating the skin
friction of bored piles K -values after Berggren (1978)s can be used (Fig 14)
(3) The use of CPT for the prediction of the bearing capacity
and the settlement of the pile is recommended As CPT
can be looked upon as a model test of a pile it will
be expected that the stress-strain behaviour of the
soil surrounding the pile is similar to that when the
cone is penetrating the soil Different methods have
been recommended for the evaluation of the bearing
capacity of the pile based on CPT test results
(4) The methods based on the results of pressuremeter tests
and the stress wave equation show very good agreement
with the values measured However those methods reshy
quire special equipments and specialists
(5) The load test is the best method for prediction of the
bearing capacity of piles The constant rate penetration
75
test (CRP) is recommended for this purpose To define
the ultimate load the following expression can be used
B PL0ult = 20 + 2p + AE
It is interesting to take into account that the settleshy
ment of the pile at failure is 005 Bin medium dense
dense and very dense sand and 01B in loose and very
loose sand
(6) It is important to separate the shaft friction and the
point resistance in the load test results and in the
calculations Different safety factors for shaft load
and end-bearing load are recommended It is common
practice to take FS = 3 for point resistance and FS = 15 for shaft friction
(7) Methods based on pile driving formulae can only be used
for piles with limited allowable load According to
SEN 75 Q lt 450 kN for concrete piles and timber pilesa -
Q lt 130 kN for timber pilesa -
(8) The relationship between the CPT method and static
formula for prediction of the bearing capacity of piles
may be interesting for practical purposes Some examples
with case histories show a good agreement between the
measured and calculated pile capacities
(9) The bearing capacity of a pile group can be evaluated
by the efficiency factor n = 1
(10)The calculation of the settlement of a single pile and
a pile group can be made by the use of Vesics method
(1977) Poulos method (19721977) and Schmertmanns
method (197 0 1978)
(11)For practical purposes it is useful to take into account
relationships between the results of different methods
of penetration tests
APPENDIX A
Relationship betwen results of penetration tests
1 Meyerhof (1956)
q = 04 N where q is the cone resistance in C CPT tesfs (MPa) and N is the number
of blows from SPT tests
2 Schmertmann (1970) gave the expression q N = n C
where n varies with type of soilp see Table A1
Table A1 Values of n
Soil type n
Silts sandy silt and slightly cohesive silt sand mixture 02
Clean fine to medium sand and slightly silty sand 03-04
Coarse sand and sand with little gravel 05-06
Sandy gravel and gravel 08-10
Sanglerat (1972) recommended this relationship enabling
the usage of theories from CPT tests to calculate bearing
capacity and settlement
3 Dahlberg (1975) Based on the results obtained from
penetration tests the relation between q (SGI mechanical)C
and N of SPT is established as follows
log(q) = 0999 log(N)-0131 plusmn0037 C
and
log(N ) = 0919 log(qc)+0212 plusmn0003630
The relationship is presented in Fig A1 n = q N is 06-07C
which is about the same (05-06) as Schmertmann (1970)
suggested for sand and sand with little gravel
4 Moe et al (1981) In accordance with results from soil
investigations the following relationship between qC
and N is suggested
A2
as
q = 05 N or n = 05 C
5 Berggren (1978)
Based on the results of Dahlberg (1975) recommended the
following simple relationships between different penetrashy
tion tests
= 070 or = 1 4 (5 1 ) qc N20 N20 qc M = 212 or = 047 M (52)N20 N20w w
09 or = 1 11 (5 3)N30= N20 N20 N30
whre number of blows per 020 m ram soundingN20= (Swedish method A)
= cone resistance in MPa
= number of halfturns per 020 m weight sounding penetration
N = number of blows for 03 m of penetration30
The above relationships will give
q N = 077 (MPa) ( 5 4)C
that is the same value recommended by Dahlberg (1975) for
coarse sand of low gravel content
However if taking the relationship between N and N30 20
according to load test results on bored piles on silty
sand by Tuoma and Reese (1974)
(5 5)
which will give
q N = 027 (5 6)C
The range of variation of q N is from 027-077 that is C
similar to the range of variation reported by Schmertmann
(1970) The relationships 51 52 and 55 are presented
in FigA2 For practical purposes it is recommended to
take the following relationships
A3
= 37 qcN30 M = 30 qcw
= 1 4 N20 qc
where is measured in MPaqc
6 Relations between N q and some parameters of the C
soil according to Terzaghi and Peck (1948) Sanglerat
(1972) and Bergdahl (1980)
Table A2
SPT Relative Description of Static cone Angle of density compactness resistance internal
N Rd MPa friction
lt2 02 Very loose lt2 (lt25) lt30
5-10 02-04 Loose 2-4 (25-5) 30-35deg
11-30 04-06 Medium dense 4-12(5-10) 35-40deg
31-50 06-08 Dense 12-20(10-20) 40-45deg
gt50 08-10 Very dense gt20 ( gt 20) gt45deg
These values are approximate the values inside the parenthesis are according to Bergdahl (1980)
The ration= q N is almost constant and is approximatelyC
equal to 04
7 Sanglerat (1972)
Sanglerat has been collecting the relationship between qC
and N according to various authors and presented the re-
sults as in Fig A3 The range of variation of the value
n = q N is about 025-08 which is about the same as C
suggested by Schmertmann (1970)
8 Thorburn (1971)
According to results of various authors n (q N) depends on C
the particle size of the soil The range of variation of n
is about 02 for silt to 06 for gravel The result of the
comparison is presented in Fig A4
A4
9 Penetration testing in Greece (Tassios ESOP 1974)
As a cross-check of site data the following values are
frequently used (q =MPa) Table A3 C
Table A3
Soil type n=q NC
Clay 02
silty clay and sand 03
Silty fine sand 04
Sand 05-08
Sandy gravel 08
The values of n are similar to those of Schmertmann (1970)
but it is interesting to note that the value of n = 02
is used for clay
10 Schmertmanns research (1977)
Fig AS presents a semi-theoretical correlation between N
and qc The use of Fig AS requires a knowledge of Rf as
obtained from a Begemann type of cone tip If only SPT
data available then the ratio NOGinN12 _ 18 in andor
N6_12 inN12 _ 18 in can provide a measure of Rf (Table AS)
Summary
There is no unique relation between qc from cone penetration
tests and the N-value from standard penetration test The
ratio between qc and N is dependent on the type of soil
The range of variation of n (q N) is from 02 to 10 the C
largest value for a coarse soil
For practical purposes the ration can be obtained from
Table A1 or Fig A4 The angle of friction can be obtained
from Fig A2
B1
APPENDIX B
Empirical methods to obtain the value of angle of internal
friction
B1 Brinch Hansen (1950)
~ = 26deg + 10 ID+ 04 Cu+ 16 lg (dm)
ID= relative density of the soil
C = uniformity coefficient (should not be chosen u higher than 15)
dm = average grain sice (mm)
The new code of practice for foundation engineering
(Danish institute 1978) recommended the similar expression
o 3 4 ~ = 30 - - + (14--) I
Cu Cu D
B2 The angle of internal friction can also be derived
from CPT test according to Meyerhof (1976) see Fig 16
B3 Estimating the angle of internal friction from the
relative density Schmertmann (1977) recommended a modified
chart from Burmister (1948) as in Fig B1
B4 Estimating the angle of internal friction from CPT test
Trofimenkov (1974) shows in Fig B2 the relation between the
angle of internal friction of sands and the results of
static sounding The diagram takes into account the value
of effective overburden pressure
B5 Estimating the angle of internal friction from pressureshy
meter test Baguelin et al (1978) recommended to use the
Centre dEtudes Menards equation
~-24Pi= 25middot2 4
and the results are shown in Fig B3
C1
APPENDIX C
Values of safety factor
C1 Broms (1981)
FS = 2 is frequently used ot calculate Q bull a 11ow However it is common to use
FS = 3 with respect to point resistance
FS = 15 with respect to shaft resistance
C2 Berggren (1978)
For bored piles FS = 5 for the pile point
and FS = 1-2 for the shaft
C3 De Beer (1979) recommended FS for very dense sand
where F1 = 14 n1 = 15 and n2 = 13
C4 The Canadian Foundation Engineering Manual (1978)
recommended FS = 3
APPENDIX D
References
American Petroleum Institute (1977) Recommended practice
for planning designing and constructing fixed off
shore platforms APT Also see Beringen et al 1979
Andreasson L Kompressibilitet hos friktionsjord Laborashy
toriestudier Chalmers 1973
Andreasson L Hansbo s and Hartlen J (1976) Results
of loading tests on bored piles Proc of 6th Europ
Conf on Soil Mechanics and Found Engn Wien 1976
Vol 1 2
Australian Standard or SAA Piling Code (1978) Rules for
design and installation of piling
Armishow JW and Cox DW The effect of change in pore
water pressure on the carrying capacities and settleshy
ment of driven piles end-bearing in sand and gravel
stratum Recent Development in the Design and Conshy
struction of Piles London 1979
Baguelin F Jezequel JF and Shields DH (1978)
The pressuremeter and Foundation Engineering Trans
Tech Pub
Barkan DD (1962) Dynamic of bases and Foundation
Mc Graw Hill Book Comp p 13
Begemann The friction jacket cone as an aid in determining
the soil profile 6th ICSMFE Vol
Begemann (1977) Foundation Building Research Soil Mechanic
Aspect on Pile Foundation Samson Alphel (Also see
Beringen 1979)
Beringen FL Windle D and Vanhooydonk WR Results
of loading tests on driven piles in sand Recent deshy
velopment in the Design and Construction of piles Proc
of the Conf in London
D2
Berggren B Large diameter bored piles Commission on
Pile Reserach Report No 58
Berggren B Gravpalar pa friktionsjord Sattningar och
barformaga Department of Geotechnical Engineering
Chalmers University of Technology
Bergdahl U (1979) General discussion Proc of Recent
Development in Design and Construction of Pile London
Bergdahl u and Wennerstrand J Bearing capacity of driven
friction piles in loose sand Proc of VIEVSMEF 1976
Wien
Bergdahl u and Moller B The static and dynamic penetroshy
meter Proc of X ICSMFE 1981
Bergdahl u (1980) The use of penetrometer in Sweden
Bergdahl u Hult G Load test on friction piles in sand
Field tests with different test procedures Report No
56 Commission on Pile Research 1979
Bozozuk M 1979 Bridge foundation move Trans Research
Record G78 Washington
Braatvedt I 1976 Franki pile
Berezantzev VG Khristoferov VS and Golukov VN (1961)
Load bearing capacity and deformation of pile foundation
Proc 5 ICSMFE Vol 2 Paris
Broms B (1966) Methods of calculating the ultimate bearing
capacity of pile Summary Soil and soil No 18 1966
Broms B (1978) Precast piling practice
Broms B (1981a Pile foundation General report X ICSMFE
Broms B (1981b Field and laboratory method in Sweden
Bustamante MG 1981 Reajustement des Parametres de
Calcul des Pieux Proc of X ICSMFE
D3
Canadian Foundation Engineering Manual Part 3 Deep
foundation CGS March 1978
Caquot A and Kerisel J (1956) Traite de Mecanique
des sol 2nd edition Paris
Christoulos A Tassement de Fondations sur Pieux Proc
of X ICSMFE Stockholm 1981
Commission on Pile Research Recommendation for pile drivshy
ing test with subsequent load testing Report 59
Stockholm 1980
Danish Geotechnical Institute Code of practice for foundshy
ation engineering Bulletin No 32 1978
Dahlberg R and Bergdahl U (1974) Investigation on
Swedish ram sounding method Proc ESOPT Stockholm
Dahlberg R (1975) Settlement characteristic of preshy
consolidated natural sands D11975
De Beer EE (1967) Bearing capacity and settlement of
shallow foundation in sand Proc of Symposium on
bearing capacity and settlement of foundations Duke
University pp 15-33
De Beer (1979) New development on design and construction
of piles London (1979)
Franke E Point pressure vs length and diameter of pile
Proc X ICSMFE Stockholm 1981
Gibbs HH and Holtz WG (1957) Research on determining
density of sand by spoon penetration testing Proc
4 ICSMFE Vol1 London
Gravare CJ and Hermansson I Practical experiences of
stress wave measurement Application of stress wave
theory on piles Edited by Bredenberg 1981
D4
Goble GG Rauche F and Linkins GE 1980 The
Analysis of pile driving A state of art Int
Seminar on the application of stress wave theory
on piles Stockholm June 4-5 1980 pp 131-162
Hermansson I Static bearing capacity of piles from
dynamic measurements Vag- och vattenbyggaren 8-9
1978
Hultsjo S and Svensson The bearing capacity of friction
piles in sand Swedish Commission on Pile Research
Report No 16 1969
Hansbo s 1981 Grundlaggning av byggnader och maskinshy
fundament
Ireland HO Pulling tests on piles in sand Proc 4
ICSMFE Vol 2 London
Jamiolskowski M and Pasqualini E Discussion Proc
Conf on Recent Development in Design and Const Pile
London 1979
Kezdi A Pile Foundation Foundation Engineering Handbook
1975
Mattes NS and Poulos HG (1969) Settlement of a single
compressible pile Proc ASCE Journal of SMFE Vol95
SM 1
Meyerhof GG (1956) Penetration test and bearing capacity
of cohesionless soils Proc ASCE Vol 82 SM 1
Meyerhof GG (1959) Compaction of sand and bearing
capacity of piles ASCE 85 SM 6
Meyerhof GG (1960) The design of Franki Pile with
special reference to group in sand Proc of Sym
on design of pile found Stockholm
Meyerhof GG (1963) Some recent research on the bearing
capacity foundation Canadian Geotechnical Journal
Vol 1
D5
Meyerhof GG (1976) Bearing capacity and settlement of
foundation Proc of ASCE Vol 102 GT3
Michel JK and Burgunoglu HT (1973) In situ strength
by static cone penetration test Proc 8 ICSMFE Moscow
1973
Moretto o (1971) Cimientos profundos Revista Latinoshy
americana de Geotecnia 1 No 2
Moe D et al (1981) Friction bearing pipe piles at
Calabar Port X ICSMFE Stockholm
Norwegian Pile Committee (1973) Recommendation for pile
foundation See also Broms 1978 Bergdahl 1980
Parry RHG (1977) Estimating bearing capacity in sand
from SPT values Journal of SMFD ASCE GT9
Potyody JG (1961) Skin friction between cohesive granular
soils and construction material Geotechnique Vol11
No 4 pp 339-353
Poulos HG (1972) Load settlement prediction for piles
and piers Proc ASCE Journal of SMFD Vol98
Poulos HG (1977) Estimation of pile group settlements
Ground Engineering March
Rodin et al 1974 Penetration testing in UK ESOPT 1974
Stockholm Vol 1
Schmertmann J 1967 Guidelines for use in soil inshy
vestigation and design of foundation for bridge structures
in the state of Florida Also see Schmertmann 1975
Schmertmann J 1977 Guidelines for cone penetration test
US department of transportation
Schmertmann J 1975 The measurement of in situ shear
strength State of the art report to ASCE conference
on in situ measurement of soil properties Raleigh
North Carolina Vol 2
D6
Sanglerat G 1972 The penetrometer and soil exploration
Amsterdam Elsevier
Skemton AW et al (1953) Theorie de la force portante
des pieux dans le sable Annales de Institute Tech du
Batiment des Travaux Public March-April
Senneset K Penetration testing in Norway ESOPT Stockholm
1974
Sellgren E The pressuremeter and pile foundation Proc
of X ICSMFE Stockholm 1981
Sellgren E Friction piles in non-cohesive soil evaluation
by pressurmeter test Thesis CTH 1981
Searle IW The interpretation of Begemann friction jacket
cone result Proc 7 ECSMF Brighton 1979
Swedish Building Code 1975 Chapter 236 Pile Foundation
Tassios TP and Anagnostopoulos AG Penetration testing
in Greece Proc of ESOPT Stockholm 1974
Te kam WGB 1977 Static cone penetration testing
and foundations on piles in sand Also see Beringen
1979
Terzaghi K and Peck RB (1967) Soil mechanics in
engineering practice Wiley New York
Thorburn S Tentative correction chart for standard
penetration in non-cohesive soil Civile Engineering
and Public Work Review Vol SB No 683 See also
Rodin et al
Thorburn S and Buchanan N-W Pile embedment in fine
grained non-cohesive soil Proc of Recent development
in the design and construction of piles London 1979
Tong YX et al (1981) Pile foundation in soft soil Proc
of X ICSMFE Stockholm
D7
Tuoma FT and Reese LC (1974) Behaviour of bored
piles in sand Journal of GE ASCE Vol100 GT7
Trofimenkov JG Penetration testing in the USSR ESOPT
1974 Stockholm
Van Weele AF A method of separating the bearing capacity
of a test pile into skin friction and point resistance
IV ICSMFE London 1957
Vesic AS (1967) Ultimate load and settlement of deep
foundation in sand Proc Syrop on Bearing Capacity
and settlement of foundation Duke University
Vesic AS (1970a) Test on Instrumented piles Ogeechee
river site J SMFD ASCE 96 SM2
Vesic AS (1970b) Load transfer in pile soil system
Design and installation of pile foundation and cellular
structure Pennsylvania
Vesic AS (1975) Principles of pile foundation design
Soil Mechanics Serie No 3B Duke University
Vesic AS (1977) Design of pile foundation Transportation
Research Board
JOOr-----r----~--------~---~---~ 0nw11 p1ts uI
A Giavt o
bull150 t----+----+----+-- X Sill ~10o Sofedp1tes1nsari
References ~~lOunltss-to~
Cie~sen~al11373) llalch Jnd Ste1nac fl971) McCWgt$Qn arid Codel 11970) ______________
lle1(tfl97l) ~~t2 v 1153 lt-~lt~(_ Slitutaetalfl913l deg1ri
~1a~~~mn bull ~t toodwardttJt11)
Fu1the1 teferttices 1n F12 J
bull6
bullJ 6
2--10 5021) JO deg
Slmtd ptnet11hoo 1es1stmce N in blows per loot
Fig 1 Empirical relation between ultimate point reshy
sistance of piles and standard penetration
resistance in cohesionless soil (1 tsf = 958
kNm 2 1 blowft= 1 blow03 m) (After
Meyerhof 1976)
I
~ z r 0
I gtO
~ r ~
0ii
0 l
~ 101 z
0 0 0 20 gtO 00
Fig 2 Relationship between N-value overburden pressure
and relative density for non-cohesive soils
(After Rodin et al 1974)
RAM SOUNDING Rods diam 32mm Rom weight 635kg
0 Wood
10
J 8
0 I
C o o 00
C
c ltI)
0 gt oshy
w
0 __ __________________
0 7 3 4 5~H =- (tvpmm)s Average
0 v
D
va Iue over pile Iength Wood pile Wood pile spliced Concrete pile
Fig 3 Equivalent skin friction derived from penetration
tests (After Senneset 1974)
1000
bull C z
0_ u lt lL
gtshy_ 0 lt a lt u ~ z a lt UJ co
100
10 25deg 30deg 35deg 40deg 45deg 50deg
ANGLE OF INTERNAL FRICTION -
Fig 4 Bearing capacity factors for circular deep founshy
dations (After Vesic 1975)
600 ~ +qNq400
200
100 80 60
z O 40 a 0 t-u 20 ~ gt-t-u 10 ~ 8 lt( u 6 (9
a z
4 Ir lt( w l+IrllCD
2 GI
= c +q tan qgt
1 --------------------____1-_1-_ ___ _1
0 5 10 15 20 25 30 35 40 45
ANGLE OF SHEARING RESISTANCE qgt
Fig 5 Variation of bearing capacity factor N q with
Irr and cp (After Vesic 1975)
1000 loctt1on 0 Sourlaquo
0 18bull stttl pfpe Ogeechce 861 bullo Yuff (1967)800 A 12 stttl H I
9 6 11bull p~~~ con-I St CMllts 591 25 Ttnnu (1971)
600 II abull precut con- RoorkH Ag9lt1~1 ( 1966) crtte J bullekelt bull 24bull st1 pfplt Arro Llke 631 45 McC Goldltr (1970) 13 30-36 concrete Kirr1sbortd
lt) B 30 concrtte y-__9y Kirr1s
bored Rffs (1974)
a z430bull concrete Live Ok300 bored i32 0 SS concrete Abidjn so rfs1 (1957) Jbullcked
driven200 i) 40 concrete HarlCiibo 100 Qrfsel (1961)bored
ltBgt 59 concr-ett bored
o 8S penetrometer Chevreu~ n- ~ laquots1 (196lt)~er 0 emiddot steel pfpe Tokyo 781 SCP C(lflllllitte-e (1971)z
bull 8 stee 1 pipe Tokyo 651 t a 0 100 I-()
80it -----bull Chattahoochee - - - -
gt- 14 Dr =80 60 15 145 ()
14~ 52
Ef 2 lt( 10 132 bull 213 () i
40 35 0 125 52 ] () 12 ] 125z iE 46 ~ lt( 30 78 w (Il 42~6 bull
B3220
NUMBERS INDICATE DB RATIO
10________-__-~2~0~181-------------~-=---- 0 05 10 15 20 25 30 35 40
18131
VERTICAL GROUND STRESS ( TSF)
Fig 6 Experimental values of N in sand from different q
investigations (After Vesic 1977)
Table 4 Experiemental values of N in sand q
Fig 7
Ns
SAND RELATIVE BORED COMPACTNESS DENSITY() DRIVEN PILES PILES
Very dense gt80 60-200 40-80 Dense 60-80 40-80 20-40 Medium 40-60 25-60 10-30 Lpose lt40 20-30 5-15
2 I
100 - E Iamp 5 05z ~
t lampI () lampI
()z 20 02 z ~ tJ) ~
tJ)iii 10 01 iii ~ lampI
0 z
5 lii tJ)
tJ)
2 0 02 04 06 os
RELATIVE DENSITY DR
y-Floridbull (lrlaquoI~ ls-7) A- alaquoeh Aivr (Selby 170)
- Pbr9yen1U (~llt1r __ 1SI) 6- St Charles kiver (hvetin 111)
reg-~ru (Yull 1611) -Piftlretih (hrry 171) 0- 1oorlcN (~bullnql 14) e- Slbullyerry (teo WilllllgtCY ltd)
e- Oguchu kiver Vuic 167) - Wuhi9ton DC SdviaHI Anoca) -middot-ChemiddotHWU (~rlul 1amp-l - Corgibull Tech (Vupound 14i7)
S-Toko ~P CO-lttM 171) bull -uiwoer Arrow Lbullkbull lcC-gttl degbull 1570)
Variation of skin resistance of piles
with relative density (After Vesic
in sand
1977)
1000 16 e__
j 12 c
~~
~t Y
[f- 1-1 o r7
I 100 J 7_ 7~ w 7
J
j1 __ 11
Nc---s_r -- t=-Nq_
~ V
20
1 - V
_ V v V~ c-lt~
ye--- ~ D---~p 4
B
1
L V 1 1 0 10 15 20 25 30 35 40 45
Angle of internal friction bull In degrees
Fig 8 Bearing capacity factor and critical depth ratio
for driven piles (After Meyerhof 1976)
so
0 0 01 O -
Ps Pst
Fig 9 Relationship between pressure settlement and
pile diameter ps is the permissible pressure at
the pile point (After Berggren 1978)
200--------------------------+---i 60
---+----55
-1--- 40 LU __ ~
lt( a r1
P m -middot -----+--7 35 o (J)
v z er LU z 0
5---+---r---+---~----+-----+--~ (J)
30 f5---+---7 25 2
----+--7 2C o
2 5 10 20 50 100
L
Fig 10 Relation between pressure pile diameter pile
embedment length and angle of friction~ (~Ns)
p is the permissible shaft stress expressedm
as load per aren of pile cross section
(After Berggren 1978)
45
~ v -
--C z
Ps1v~ a UJ or f-
( No UJ
- 2c lt c f if c z a UJ z 0 (j
z UJ 2c ci
0 5 10 15
Fig 11 Relationship between pressure pile diameter
pile embedment length and angle of friction
psf is the failure at the pile point (After
Berggren 1978)
V V
I
I ~
soJ--4----+---+--+---t--+----1rl~-1----+--- o - middotv LO+---+---+---+-- -- - -ltlt--V--+---i----- -lt~ ~-
JJ +--4----+---+------t ~
I v P d 20 _______--+-gt----middot - ~--1t---+---+---i---
10 i===l=~~==l==-middot-=-tiC-t-- - - - - - - - - o middot o 1
t-=_middot-t----t---_1-zttj K-- - - - _ -- obull2s ~ ~ 0 _bull _L_ I i------txi-u s ___ ___ ___-+-v_-+- 1 H middot JO 60
L ___ ___ _____~middot___ _____ I ~ - - -1- - -_ ___v__ bullbull
1 l ____1 I I -r- L0-7)
I I I-- - - - SOmiddotll
i I I 1 - bullMo I I i I -j--10V 11I 1 I Ii
10 40 50 I
I I 0 I V 0 I
CLEAN $AHO OENSHY
f = Avlaquorag ttttet1v d soil rtsulting from pHe instaHation owr Ht19ht B below and LB above toe of ptc~middot
Wtt Cast Bore Pile with Standard FRANKi PILE ftmporary sluvmg but with with cnlorgcd Base ckgru of BOILING-during bOftng
22bull_
Fig 12 Values of bearing capacity factors (After
Braatvedt 1976)
~bull = average effectice ~ soil resulting from
pile installation over height B below and 4 B
above toe of pile
INDICATIVE ULTIMATE PRESSURE AT BASE CR TIP OF PILE
IN CLEAN SAND DERIVED FROM EXPRESSION
10L D~lbullIPd Nq)
2 --r711-~--r-l -Ii t
21 - xf middot1-1---f- -- e-- -- _ t
I ~ I I I _ I II I I I 1 I V J
I ltiI I 12 I ---I I I A
w 0 0
0
f Jw It V CD II I
I
I I I w II I ~ I I I AI i
~ 1 I i 10 000 2C 000 JO 000 c 00)
100 200 JOO LOO
ULTIMATE POINT PRESSURE
~ t gmiddot Average d1et1vc (J sod resulting from pde mstallal10( over h1ght B
olow and LB JbOIC toe of pile 2 Figures are based on clean sand and Natural SpWt Any reductoo in
1 0 th ~~~~ f1oe~hi~e ~~~~tbull~~ cJ~~erto~ ~
0~an~~~~H r~0u~h in ~uJ~tonal
rrcJuchon of intlicated figures (Su Consistency Chart tor Appoiumote Specific Wughts)
3 If embedded Length 1s less than 100 ttwn the abollt point pressuru must have the following tact0lt applied
Em~dded length x 1
P1 le dam at Bast fa
Fig 13 Ultimate pressure at base or tip of the pile
(After Braatvedt 1976)
Meyerhof (1976)
SAA(1978)- 2 i--------
~ - -------+~ - -------+- -- -- __ + ~ ---- +--
~ ---~~+--- Broms(19661~-o ___+ I bull
~ 11----------- _________ j-------~en(1978 ) OJ t bull Bored piles 0 4-
I -~ +-c ~ QJ
() I Meyerhof ( 1976 ) Bored piles
35 40 Angle of friction (fJ
Fig 14a Values of coefficient of earth pressure K for s
driven and bored piles
I I e Driven cylindrical piles
H Driven H-piles I - + Jacked piles
o Bored pries ISee Fig 3 for references
I - 2i + v+II
middotv I
+
H
+
V
0 30 35 40
Angle of internal friction bull rn degrees
Fig 14b Coefficient of earth pressure on shaft of piles
above critical depth in sand (After Meyerhof
19 76)
2~----------------r-----~
Broms+API
Steel piles(6=20deg) tO C a
U1
~ 030 35 40 45 Angle of friction (Ji
Fig 15 Variation of K tano with internal friction angle5
400
I I I bull Kltlrl ll181 J o Kmselll11 7 ~ 6 lluhsmdirnsil91l -I-JIO X lt1wllal~ 0
-ii V -
~ -bull----i JOO
I I
I I
I X xllo
I x ~
il i 100
7 cl X6
IO bulldVMyloo~ i j gt-- 06--middot I 0
JO JS 40
1[
Fig 16 Approximate relation between limiting static cone
resistance and friction angle of sand (After
Meyerhof 1976)
Wrik cn1
Fig 17 Relation between ultimate point resistance of
pile and depth in sand stratum beneath weak soil
layers (After Meyerhof 1976)
Weak soil
Dense ~ind
Weak soil
Fig 18 Relation between ultimate point resistance of
pile and depth in thin sand layer overlying
weak soil (After Meyerhof 1976)
I
Clp bull f f + 11 ) 2 + Ill 2
K 0 O~ofthopi I Awwago cone rustanco ~ tho tlp of the pi- ovw a CMpth
whkh may wry blaquowffn 070 and 40 II Minimum ~ ianot rtcOrded ~ tM pUbull dp o tho
_ dopth of 07D to 40 UI A~ of the ~ of minNffum coN rocistancoc rtltOfdocl
MgtOH the piM Op owr bull hotifht which ITWY VMY b4ttwMn 60 and 80 In dttlaquommint thia n lopo ~ abo tho minimum fflu-e sctad uncSr II are to b4 diiw~
tip Utdrnate u~_PltgtMt rUnoe of thct pUo
Fig 1 9 Application of CPT-method (After Beringen 1979)
20 Limit value It 15 MNm2
for all cohnionltH 1 z E 15
() z lt ~ ij 10 a ~ z ~ ~ 5 ltE i= OCR bull OyenlaquoCOltUltgtlidatlon ratio
0 10 15 20 25 30 35
THEORETICAL POINT RESISTANCE It ( MNm2 I 0 5
Fig 20 Limit values of point resistance for driven piles
(After Beringen 1979)
15 bull
25 bullbull
a 3S bull
200 i-------+------+--~ir---h-C----c+---c---Jcc--c--~-l 45 u
ltT
g SSbull 65
2 75 bull middot ISbull _ 100 1------11---- -------middot----~-=--middot- --c=---1 95 bull p__~~-= 100 bull 0 C
u-_-- lt16 tU
tocQl friction f1 in bar
Fig 21 Graph showing the relationship between qc f s
and soil type (After Begemann 1965)
K for Steel Pipe Piles K for Square Concrete Piles
0 0
10 i I
20 30 00 10 20
I
10 11 (
1i
i
i I
10
Q)
I
c)
- 20
I l
I I I I
I - c) - 20
()
5
tJ C I
C
30
40
II I -- pound1cctricil II I l enetronetcr+L- middot--middot1middot middot-middotmiddotmiddot - middotmiddot-----middot middotmiddot--r - tbullechinicalPenetrometer
I middot1- middot-middotmiddot-middotmiddot ---~+ I I I I I1--t- I
I I I
C)
30
40
Q) c
u
C
c u CJ = I L-
CJ C (lJ c
umiddot-i u CJ
w
I r-I I r-- i I
Ktimber = 125 Kpipe
Fig 22 Penetrometer design curves for pile side friction
in sand (After Schmertmann 1977)
qc
lt r 86
ro r Ti2Jill
~ F = Kf 1A(8B) +f AT - 8B) +middot5 s s 2 s1 s 1 B
~ Kff A (T1 - 4B) + ~ f AT ~ 51 5 sr s n
where r J c n = layer number c
lt 0 gt
gt Te layer thickness f = average unit sleeve friction
5
lt A= pile soil contact area per unit depth
~i
-
lt For T lt 8B --1--
Use Eq to a depth of 8B and theL__~ average sleeve friction concept below 8~
Fig 23 Side friction computation method for layered
soils (After Schmertmann 1977)
qcl + qc2 q =---p 2
qcl = Average qc over a distance of yd below the pile tip (path abullbbullc Sum qc values in both the downward (path amiddotb) and upward (path b-c)
directions Use actual qc values along path abullb and the minimum path
rule along path bbullc Compute qc for y-values from 07 to 40 and use 1
the minimum Qct value obtained
qc2 = Average qc over a distance of 8d above the pile tip (path c-e)
Use the minimum path rule as for path b-c in the qc1 computations
Ignore any minor x peak depressions if in sand but include in minimum path if in clay
Fig 24 Dutch procedure for predicting pile tip capacity
(After Schmertmann 1977)
WEIGHT SOUNDING 1 JO 1-------------------+---t-----1
-
f ~ 8 Q- Q Wood ---
1 0 61------------+---+----~--~
C 0 u 41--~1~-11-c---c+---c~I
I PQ
c 2 -~~t bull__-+-i---gt--+-----11----+ v I
I Jr o --middot- -middotmiddotmiddotmiddotmiddot-~------_____L
0 20 t) ffJ 80 100 120 Number of half turns per meter average over pi I e Iength
Fig 25 Skin friction (After Senneset 1974)
- -
g
t bull
I
10
11
NATURAL UNIT
WEIGHT (kNm3)
15 20 ~
~
WATER VOIDS CONTENT RATIO (1
04 0608 1 02 0 20
I
Ilt
bull
Fig 26 Combined boring log
UPPER LAYERS 10 bull 25ml PARTICLE SIZE O lmml
sect 0
0 -lgtlt gt ~i
~Kf ltIgtI
rq
i
lgt)0
~ 1~ -t ~
IO P jgtlt~ gt ~
Lgtlt ~
INE MEOH U NAASpoundCLAY SILT
LOWER LAYERS (25-95m) PARTICLE flZE O fmml
fi
l--+--1-4-l-+-l-l-+--l-+--I-__
0 degf--+--++H-H++-++--+-1-
~ f---+-+--+++t+H--l-1--1--1-
i =l---1---l--l-H-H+4--l--l---l--l-~ l---+-+--I---H-++H--1--1--1-l-
i~ 1--+--1-4-1-+-1-1-+--l-+--I--+-~
f--l--l--+-14+-+-l--l--1--1-1-+---+--1--l-l-+-+--+-___
CLAY SILT
Fig 27 Particle size distribution
CONE RESISTANCE IMNm2)
02 0 06 08 10 12 1 1 1tS
UNITSLEEVE FRICTION (MNm21
Fig 28 CPT-profiles at test locations
bullbull 2I_
~ lS =20 kNm3 a= tZ =20 middot 35 =70 kPa
70 kPa
35m 1s11kNm3 I
a= a +11middot 35 =1085 kPa
1085 kPa
Fig 2 9 Distribution of effective overburden pressure
Shaft load kN
00 20 40 60 80 100
o Push down test o Pull out test 0 CPT bull Hulsjos formula
E
0 tJJ
0
0 0 151----+----+----+-----f---4
lt-0
c Ol C
-5 20~--~--~--~--~~-~
Fig 30 Comparison between measured and calculated pile
capacity (After Bergdahl amp Wennerstrand 1976)
-- 10- u 0
~
bull Hard
~ -~ Very= ~
i ltI)
C 3 stiff I H- - 0 ~ u ~~ Stiff in u
_ in -- -5 0 u Firm
-Soft
-Very wit
o 2 10 100
()Friction Ratio Rf
Fig 31 Interpretation chart based on the work of Begemann
(After Searle 1979)
(Oenbullbull or
c~tedJ
10
i vty aheHy
bullIi -middot I
Sood limofOcka
c
E (lOOH I
I
Soft
Voy
soft
01
01 1 ) 4 10010
Friction Ratio Rt ( )
Fig 32 Interpretation chart based on the work of
Schmertmann (After Searle 1979)
10
u O
i u
e a
C e () = C
lti () C E
() e
I iii
Very Sensitive Soils
01 ------~---------------------J--l-------~---~1-
01 10 100
Friction Ratio Rf ( )
Fig 33 Proposed interpretation chart for normal soils
(After Searle 1979)
4 4
300
ppound-kPaltI 1000
1000 3
k IClay SiltpkPa~ I i I
I i30021 I I I200 I 2
I i 3 4 5 G 8 3 3 4 5 6 7 8 3
Fig 34 Fig 35
k-values for driven piles (After Baguelin et al
1978)
21-+---+-+---t--+--t-i--1
3 4 5 6 7 8 9 10
Fig 36 k-values for driven piles
(After Baguelin et al 1978)
k
z 3 4 6 7 8 9
Fig 37 k-values for driven piles
D = depth of ernbedrnent B = width of the pile
(After Baguelin et al 1978)
10 I
I -- I Sand i I -e
8 ~ ]
-- pf bull6000 kPav
k v
v
V k
4
Driven lay --i----~-- -- --- C
[=1000 kPa1PV Cast
0 0 2 6 8 10 12
Fig 38 Influence of method of installation on k
(After Baguelin et al 1978)
~(kPa)
150 i I Ir
i i I I I I
I i soi bull~a(--~I i i ~~
l i 1- I te100 lI coomi -11
I I ~i tion procedures slt-a a I I i ~r degI I i olt- ~ o I (I ir ltcohesive soil _e I elle
0V gt ~svfJ-n l50 I i iv j--j j J t piles in bullnv soil
I oisl3~~ - I (onmiddot middoti- 1 - IIYV1
Y I IP( I I ~ 0 500 IOOO 1500 Pt (kPa)
Fig 39 Adhesion or friction on the sides of a pile
(After Baguelin et al 1978)
DETAIL
HAMMER
PILE
I SUITA3LE TAPE
-==-J [ ~~---HGLOSSY PAPER
- ~~~- ---~- -----=- ----=- - - r--~ _-
lJ
Fig 40 Principle for measuring pile rebound
(After Swedish Commission on Pile Research Report 59 1980)
CASE METHOD CAPACITY PREDICTION IN MN 0 CJI o CJI
0 o I
X)tX Y
X
X
0 X
X
X X
(JI gth ~x
xx )xlten XX gt-X X
--1 X X~ Igt o lgt--1 X0 x X
) X Xr X
0 gt
X X
X
--middot-0 CJI X
X i X--1
X tlt Xm )(() = I X--1 X
X
I) YX Xn 0m X gt X() l
C r gt
--1 () l
I) X
z CJI X X
s z
cgt o
cgt (JI
Fig 41 Case method and static load test capacity
correlation (After Goble et al 1981)
r 2 ST PDA
J ACC CJ 51i_ I
bull
TR
07[0
l ltCJ MIC
16a I SCOPE
L-lt
ST STRAIN TRANSDUCER ACC ACCELEROMETER PDA PILE DRIVING ANALYZER TR TAPE RECORDER (ANALOG 4-CHANNEU
Fig 42 Schrnatic of instrumentation (After Goble et al
1981)
F I MpJ vt -middoti
I 100
eo 2Lc bull 26 ms
60
15
10
20 os
lmx
-~-----------------~middot
Fig 43 Typical records of pile top force and velocity
Pile 235 x 235 mm length 46 m
Hammer Hydraulic free falling
(After Hermansson 1978)
TR
OSOOO
MP PR
AD
CPU 1-----------1 l-I s
TR TAPE RECORDER SYSTEM middot1 middotiUL T PLEXER tJD rlLOG-TO-DGTAL COINERTER CiU bull 3K lmiddotIORD MINICOMPUTER DS DISC STORAGE R PRirJfER L ~)LOTTER S SCOPE
Fig 44 Schematic of laboratory processing system
(After Goble et al 1981)
(A) ACTUAL SYSTEM (B) MODEL
Fig 45 Hammer - pile soil model of wave equation
MN 3 ----------------------------
2-+-----
3
10 40 MN
I 2----
10 20 3G 40
TIME IN MS
Fig 46 Pile top force matches for four different sets
of soil resistance parameters
1 Measured force curve 2 Low damping 3 High static resistance 4 High skin friction low end-bearing 5 Final solution
(After Goble et al 1981)
F Mp)
200
no
creep
100
0 L------------------------ 0 10 20 30 40
DISPLACEMENT mm Fig 47 Result of static load test (with and without
creep) compared to CAPWAP-computation
(After Goble et al 1981)
eT F
l~~-------0
10
il1Lc
Fig 48 Force and velocity records from a bearing pile
The proportionality is broken before the time
2 Lc Pile length 195 m c = 3000 ms
Blow No1 of redriving (After Gravare et al
19 81 )
F (l2pl
120
BO
40
Fig 49 Same pile as above Blow No 42 (After Gravare
et al 1981)
F (Mp)
n t I I f -
~ 120 ~
I bull _
BO
I 40
F __
10 20 30
Fig 50 Same pile as above Blow No 45 The pile is
broken (After Gravare et al 1981)
LOAD
a COLUMN IN AIR b END-BEARING PILE
c FRICTION PILE IN SANO d FRICTION PILE IN SAND
e FRICTION PILE IN CLAY
Fig 51 Typical working curves for piles with different
modes of operation (tested as a constant rate of
penetration) (After Swedish Commission on Pile
Research Report 59 1980)
LGAO ( TOi~ 5 i
351--+-+---+---t-middot-+---+---i---l--l---t--l
EEFJ I I ttfLl I J
Fig 52 Load-displacement diagram of a test pile drawn
in two different scales (After Vesic 1977)
LOAD kN
EALOAD TESTED 17-10-1978
Q
25 -1---
1 --~-
50 LOAD TESTED 1J-J-1979__j
75
PILE Top segment of precast concrete 200 m C450 Bottom segment of timber 181 m Osp = 134 mm Total length 381 m
PILE DRIVEN 15-06-1978 PILE LOAD TESTED 17-10-1978 and 13-03-1979
SOIL PROFILE 0 - 2 m Dry crust 2 - 55 m Firm clay E6 HIGHWAY MALMO - GOTHENSURG
55 - 60 m Friction soil SECTION ASKLOSTER - FRILLESAS
ULTIMATE 4 months after driving 1230 kN Load testing of piles at LOAD 9 months after driving 1380 UI bridge N 582 Test pile 2
Fig 53 Example of results (working curves) reported from
load testing with a constant rate of penetration
(After Swedish Commission on Pile Research Report
59 1980)
LOAD kN
Q
~ 25 a lt w c w _
-w
_
50
75
PILE TIP SIDES TOTAL 100
PILE Precast concrete 270 x 270 rrm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand
BRIDGE OYER NYA KARLVIKSVXGENEVALUATED ULTIMATE LOAD 1440 kN LULE4 MUNICIPALITY
Fig 54 Example of results (working curves)reported from
load testing with stepped load increments
(After Swedish Commission on Pile Research Report
59 1980)
----
Fig
15 gt-z w w 0 u z
Cl c 0 - 10
0 w c w a E u Llt 0 I
N w I-c z a w
w a gt-w 5 co
i s rmin
0 0 500 1500 2000
QKRLOAD kN
PILE Precast concrete 270 x 270 nm L = 110 m PILE DRIVEN 10-02-1978 PILE LOAD TESTED 28-03-1978
SOIL PROFILE 0 - 4 m Organic silt 4 - 21 m Sand BRIDGE OVER NYA KARLVIKSVAGEN
LULEK MUNICIPALITY EVALUATED CREEP LOAD 1250 kN
55 Example of reported 11 creep load curve from
testing with stepped load increments
(After Swedish Commission on Pile Research
59 1980)
load
Report
0 omiddot ELASTIC COMPRESSION
------_jOF PILE QLP 02 ----- _ 8 = ----- M 05
-----04 10
z D30
E u
I- 06 FAILURE 15 I-z uJ
f
CRITERION z w w
uJ J I-I-uJ VI
08 J20 I-I-w ()
LU 1 0 25 w J
J Cl a
l 2 30
l 4 35 0 25 50 75 l 00 125 15 0
APPLIED LOAD TON
0 200 400 600 800 1000 1200 1400
APPLIED LOAD kNI
Fig 56 Example of recommended failure criterion
(After Canadian Foundation Engineering Manual
1978)
--0
[k~n ltVJd
Ollt---P-_bullP_f_1_-i-
012 111h m Z-t hr
I JinH4111dl
Im -t hr
Lou1S-1a11i mn001 PJ HI -t hr
bull y Y y
01PbullP1~ p
y y
Fig 5 7 Rules to determine design load from the pile
test diagrams (After Kezdi 1975)
E E
t 20 ____
amp 2SP
mmMN
~middot LEA__a________ -middot--~____
0 1 f-j Lu
_J bull-ltl 0
I 1JJ c f-
Lt
10 5)0
i()~
SETTLrnENT OF THE 30~ lJJ
20-1f-
10 ~ V)
0 05 10 5 20
P LOAD AT THE PILE HEAD MN
Fig 58 The elastic compression of a pile in relation
to the applied load increment and the settlement
of the pile tip as a function of the applied load
(After Swedish Commission on Pile Research
Report 59 1980)
0 - - - - - -0 c===s--------------------------------
I 10 J
ComWIOgttgtrlwtft1nbull-d ca01om1 1 tlgt110
Jemmlt1bullbull~ubullbulllaquol-d I
motllttlt(lCohtltOPlttUlaquoj Lbulltxbullbulltc-va1mrgttdtJdbullbullbullbulld c(SltnbullmiddotWtt bull JE-S~t
c11110d111dtltgtt middot JOOO H~ 2400 bull N 11 l0ltlt11
I I
ArMI OOmThGL
I oo FmMQttllaquol(lbullv I
ffom10luuo1111gtbulltt
61 middot Sltfv(l1vvd9bullhbullf I
Sm Blm rm11vc1v
110 101m 1-ldibullVbulllblto U
Fig 59a A elastic recovery (After Weltman 1978)
6__-~--J---JL----I--L-------_
Fig 59b Separating the skin friction and point resistance
(After Van Weele 1957)
20 0 z lt
f (I)
gt- (J)u 0z 0w u J 15Li
lL 0 w z a ~ ) w0 (J)a z l w a w 10 a _J w 0 I
2 3 4 5 6 7
PILE SPACING IN DIAMETERS
NO OF PENETRATION PILES IN PILE IN BEARING
SYHBOL GROUP DIAMETER STRATUH(OIA) SOIL TYPE SOURCE
c 4 4 20 LOOSE SANO KEZDI (1957)
bull 4 HED DENSE VESIC (1968)4 15Ill 9 SAND
DENSE SAND bull 4 4 3 OVERLAIN BY VESIC (1968)
LOOSE SANO
0 4 I 4 15 LOOSE SANO TEJCHHAN (1973)9
0 4 HEO DENSE1411 15 TEJCHHAN (1973)(] 9 SAND
0 4 12 11 20 LOOSE SANO KISHIDA (1967)IZJ 9
Fig 60 Observed efficiencies of square pile group in
sand (After Vesic 1975)
3
0 2 4 6 8
ad
3
d~a=h
2 +middotgt
~
~
0 2 4 6 8
aid
Fig 61 Group efficiency in sand (After Kezdi 1975)
p 200 lf
4 d
100 er 0 f--u ltl u w50 u z w -- u ~ f--z w
~ 20 -- f--f--w VJ
5
10
2
05
10 100 1000 10000 ex
PILE STIFFNESS FACTOR K
Fig 62 Displacement factors for top of single comshy
pressible floating pile (After Mattes amp Poulos
1969)
0 ~ 0 6 1---+lt---+~lt+--+--lt--+--+--l
f-shyz w ~ o 4 f----t----L----1-------1--+---+--+--l gt 0 E
1000 10000 PILE STIFFNESS FACTORK
Fig 63 Movement ratio for end-bearing pile on rigid base
(After Mattes amp Poulos 1969)
o Standard penetration test
bull Static cone penetration test
i Jl-----1-----+1----------- References Blessey (1970)= bull OeBeer (1975)
~ Eide (1956) la Koerner and Partos (1974) ~ 21-----1--------P------II---- Leonards (1972)
I bull a Malich and Stermac (1971) g 0 Meyerhof (1953)
0 bullbull Meyerhof and Fraikin 11957)J11--~-~---f----1f--- Meyerhof and Sebaslyan (1970)
bullo ur Shibata et al (1973)degImiddot 0 Vargas (1948)
o_______________________
0 Observed maximum settlement S in inches
Fig 64 Comparison of estimated and observed settlements
of pile foundations in cohesionless soil
(After Meyerhof 1976)
C 0
C 0-2 ~
0-0
L -Q
IO
I I M - I L 0 J
L -- I M
~-------------------~----~ 0 02 OJ 05 08
vertical strain influence factor f z
Fig 65 Some theoretical and experimental distribution
of vertical strain below center of loaded area
a = I 2 elastic theory (v=04) b = Iz elastic theory (v=05) C = DAppolonia stress-strain path test on overshy
consolidated dune sand d = Eggestads tests (Dr=44) e = Eggestads tests (Dr=85) f = assumed simplified Iz distribution recommended
by Schmertmann (1970)
C(Hf[ill LIN[ V(1TftA snu- II 06 o
i ~ bullt--__~lY--t---1 g ~ ~ J i ~ ~ -bullbull--bullbull6
llbull161bull200f11((h
lfl ~~-=~b~~bull~~f t
Fig 66 Comparisons of vertical strain distribution
from FEM studied and from rigid model tests
(After Schmertmann 1978)
Rigid footing vertical strain influence factor= 12
B2
B
gt ~ 0
s ] ~ 2B 0
1 c a O middot gt
3Ba
4B
gt]
4 plane strain
LIBgt10
(sae (b) below)
B = least width foundation L = length foundation
(a) Simplified strain influence
factor distributions
depth to Izp
(b) Explanation of pressure
terms in Eq
Fig 67 Modified strain influence factor diagrams for
use in Schmertmann method for estimating settleshy
ment over sand (After Schmertmann 1978)
15 ----
0
0 log qd0999 tog (N3ol-0131 0037 00 0 ~
() ~10
w u LEGEND z ltI r= 0959 0 Tr5middot6-SPT2 (f) n=12 o T r S 11 - SPT 3 (f)
bull T r S 4 - SPT 4 ~ 5
z w log N30) =0919 log qc5+0212 0036 0 u
0 _____J_______________
0 5 10 15 20
POINT RESISTANCE N30 SLOWS30 cm
Fig A Relationship between cone resistance (SGI penetroshy1
meter) q and standard penetration resistance CS
(N ) (After Dahlberg 1975)30
I
N30 N20 Mw Qci )111 I l
-I 1
g )1 ifI I I
I 10~ I verf4~nse __
Q) ()
C - = Ill i41se- I 50 I
() ~middot ~middot-middot -middotmiddot-middot- l ampl
Q) ~ ~ ~- I JQ ~i
1amp-1J mediumI dense 22 Cl Cl C CC Cl deg C o o
I I
I 0 9 o sect - C
-JQ
1C ---111 - - - o 0
-Ill_ lt)0 - loose
Q) lt)
I ~ C I- E ~ t2 Q) Q ~Ill--middot-
~Q () 5 1 very loose I ~i I
30deg 40deg 50deg 60deg
angle of friction q
middot middot Mw Qcvery densa
dense Q) () C co () -middot Q)
C 0
--co Q)
IshyC QQ) ()
I Q 30deg 40deg 50 60
I angle of friction q
Relationship between penetration resistance and
angle of friction ~- (After Berggren 1978)
3 50
JOO
C
u (T
abull 250 u C
0
~ 200
C 0 C u
a 0 150
100
50
50 60 70
N SPT ( blows per ft)
Comparision between static penetrometer resistance
q and SPT by various authors (After SangleratC
19 7 2)
o ~~~r-r-rr-rrrr-r-r------------r-r--r-r--i-rr-rr-r-r--rrrrrrr---i-r-1--rrrrrrr--III
i I 1
I II il 1
Ji II II iYj micro+U-+-I-IH--++t-HiHH--l-+I-H-H1-H-l-+H-f+tt-++++++ti+t-+t-r-H-t+tt+--+-i
h I++++-11
-++-+-H++++I---Hf--+-++H-i-1--i-+--i1-H+++H-IH---jftt-f+t-+t--i1r-H+++tt-1___ 1
~ l 1 I I ~ 1 i = ii I I 1- )I I sH-i-H-++-f-H+f+H-l-+--4----+41f+li-1f I ibullC-f-+tl-Hlri-1-++-H+ttt-++-1i--H+t+-t-li-jll-1I I -+I+-t--1 1 i ~ ~ bull i I I Ult I I
I -H1i - u1 Ii i I i--1-1-+-+++++++ltI~i +---+H-HI++1-f-+-l----i+tttt--+-+-+-++-++t-H--H-+-++tttt1t-H1 I I Ii 11
I I 1111 I I I Il I I i Ii 111 i bulll-++-+-1-1-++t+tti--i-+--+-ri-t++t+--+--t--+--+-tttt+--t-t-H+ttttt-tt-r-t-H-ti+t-
i i I 1111 I j I I iiI
O OL4-__0_middot002L--l-LJOmiddotLdegJUL-LOLmiddot02_JO-middotO~---OJmiddotl------OW Jc--l-l-ltgt---bull-Lltgt---c__~l0-------0~~--l00
Values of n and type of soil (After Thorburn
197 0)
100 For case of 50 bull of maximum hammer energy (4200 in-lb) entering SPT sampler rods
80
~ c 0 C
co
co 60
~ ~
C
a zmiddot c 0
40 u 0
0 f--a
20
0------~----~----~----~ 0 50 100 150 200
qc from Begemann mech friction-cone (kgfcm2)
Fig AS Experiemental theoretical relationship between
q and N using liner SPT sampler without liners C
and Delft mechanical cone (After Schmertmann
1977)
Table A bull3 Method for estimating Rf from 6 incremental
SPT data using the same equipment as in Fig
Asmiddot (After Schmertmann 1977)
Rf 6_ N0-6 6N12-18 fN6-12 fN12-18
frac12 085 093 1 076 088 2 065 083 4 053 077 6 046 073 8 0425 071
BEARING CAPACITY ON COHESION LESS SOILS
Chart for the approximate evaluation of the peak angle of internal friction after the relative density has been evaluated Modified from Burmister Donald M The Importance and Practical Use of Relative Density in Soil Mechanics ASTM Proc Vol 48 1948
X
E e
1 0
middotx Q E ~ l 36 s
0 Q)
0 34 C
laquoI ltgt 32cc
30
Relative Density in
Note of caution
In problems where the sand may strain past the peak strength value before a general failure occurs then a reduced value of P must be used (particularly in the denser cohesion less soils)
peak ifgt
odeg I
I strain
Estimating sand~ from estimate of
density (After Schmertmann 1977)
-shy
relative
()01_ 4o 6o flO f60 200 21o 280 uo 360 4co Jcmc
-middot __ __ -- middot-
0 ( [ t
-c--
- ~
N r-r- r- i-k k4 f 1 -oz r-- r--~ - - r-y- --r- r-- ~-03 r- ~ h
~ ~~ -middot-r1-1 t -~ middot-middotr--K - -middot )04 _ fs ~ -- - -N- ~-
[le- fD I k05
L-~ middot I i-I --~ 1-fi- 1
06 middot-
~
~__ 0 I J
07 middoto-
iLc k I
~ - f~ J 0
rlt (gt~ 1 ~ -lo ~o
08 a l I
09
H-Hmiddot- 7-~~ i-middot r ) ~
cm 2
Relation between the angle of internal friction
of sands and the results of static sounding
(After Trofirnenkov 1974)
j cp(O)
401------------------------( from Centre ciEtude5 Menard D 3863) 7 7
V
I
250 500 1000 2000 4000 Pi ( k Pa)
Fig Menards graph to determine~bull from PibullB3 (After Baguelin et al J978)