DESIGN OF PILES IN COHESIVE SOIL

Nguyen Truong rr

SGI, Linkoping, .3weaen, Sepi:( Tit•r-. .c 1-;, b 1

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DESIGN OF PILES IN COHESIVE SOIL

CONTENTS

SUMMARY 1 ACKNOWLEDGEMENTS 2 INTRODUCTION 3

1.

1. 1

1. 1. 1

1.1.1.1

1.1.1.2

1.1.1.2.1

1 . 1 . 1 • 2 • 2

1.1.1.2.3

1 • 1 • 1 • 2 • 4

1 • 1 • 1 • 2 . 5

1.1.1.2.6

1.1.1.2.7

1 • 1 • 1 • 2 • 8

1 • 1 • 1 • 2 • 9

Bearing capacity of single piles 5

Methods based on static formulas 5

Total stress analysis 5

End bearing 5

Shaft friction 8

Canadian Foundation engineering manual 9

Australian Code

Swedish Code

Danish Standard

Buildinq Code of

(SAA) 11

(SBN 75) 1 1

12

the Soviet Union 1 3

Experience of Thailand 1 6

Brom's recommendation 1 6

Vesic's recommendation

The CTH method 1 8

1 . 1 . 1 • 2. 1 0 The method of Caquot and Kerisel

1.1. 2

1.1.2.1

1.1.2.2

1.1.2.2.1

1 • 1 • 2 • 2 • 2

1.1.2.2.3

1 • 1 • 2 • 2 • 4

1.1.2.2.5

1 • 1 • 2 • 2 • 6

1 • 1 • 2 • 2 • 7

1.1.2.2.8

1 • 1 • 2 • 2 • 9

Effective stress analysis for bearing capacity of piles 24

End bearing capacity 27

Shaft friction resistance 29

Burland 30

Canadian Foundation enqineering manual 33

Meyerhof 34

Vesic's recommendation 37

Vijayvergiya and Focht 40

Flatte et al 43

Bozozuk et al 5 Blanchet et al

Esriq and Kirby 47

18

23

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1 • 1 • 2 . 2 . 1 0 J anbu 50

1.1.2.2.11 Parry and Swain 52

1 . 1 • 3

1.1.3.1

1.1.3.2

1.1.3.3

1.1.3.4

1.1.3.5

1 . 1 • 4

1 • 2

1 • 2. 1

1.2.1.1

1.2.1.2

1.2.1.3

1.2.1.4

1.2.1.5

1.2.1.6

1 • 2. 2

1.2.2.1

1.2-2.2

1.2.2.3

1. 2. 3

1 • 3

1. 3. 1

1.3.1.1

1.3.1.2

1.3.1.3

1.3.1.4

1.3.2

Discussion

End bearing capacity of a single pile

The a method

The B method

Relation between the B method and the A. method

Relation between the a method and the B method

Summary and recommendation for design

Determination of point and skin resistance from field test Static cone penetration test

Vesic

Nottingam ano Schmertmann

Broms

Tong et al

Sanglerat

Balasubramaniam et al

Standard penetration test

Meyerhof

David

Relationship between N and the undrained shear strength

Summary

Negative skin friction

Basic concept

Causes

Factor that affect the negative skin friction

Neutral point

Fellenius' observation

Design methods for negative skin friction

53

53

54

57

60

61

64

67

67

67

67

69

70

70

72

74

74 74

75

77

78

78

78

78

78 83

83

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http:1.1.2.2.11

3

1.3.2.1 Canadian Foundation Engineering Manual

1.3.2.2 Bozozuk

1.3.2.3 Broms

1.3.2.4 Fellenius

1.3.2.5 Kezdi

1.3.2.6 Auvinet

1.3.3 Reducing negative

1. 3. 3. 1 Bitumen coating

1. 3. 3. 2 Protection piles

1.3.3.3 Overlapping piles

skin friction

1.3.3.4 Change of the geometry of the pile group

1.3.3.5 Change of the shape of piles

1 .3.3.6 Reduction of point resistance

1.3.4 Summary

2. Settlement analysis of single piles

2.1 Vesic

2.2 Poulos

2.3 Summary

3. Pile grouns

3. 1 Ultimate bearing capacity of pile groups

3.1.1 Introduction

3.1.2 Design methods

3. 1 . 2. 1 Peck et al

3.1.2.2 Canadian Foundation Engineering Manual and Broms

3.1.2.3 Vesic

3. 1 . 2. 4 Morr .house and Sheehan

3.1.2.5 Brand et al

3.1.2.6 Meverhof 3.1.2.7 Australian Code(SAA)

3.2

3.2.1

3.2.2

3.2.3

Settlement of pile groups

Terzaghi and Peck

Tomlinson

Morgan and Poulos

83

85

8,

90

90

91

92

9'.?

92

93

93

93

94

95

97

97

100

101

102

102

102

102

102

104

105

106

106

108 109

11 4

1 1 4

1 1 5 11 7

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3. 2. 4

3.2.5

3.3

4.

Appendix

Apnendix

Anpendix

Appendix

Appendix

Appendix

4

Mattes and Poulos 1 21 1 24Vesic

125Summary of the design of pile group s

1 2 7 Conclusions

A Analysis of point resistance 129

B Typical values of soil pile adhesion 133

C Classification of the clay according to the consi_stency of the soil 13 5

D s values for piles in till (moraine clay ) 1 36

13 9E Values of Young's modulus 142F References

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SUMMARY

This report is made to review various design methods

in terms of total and effective stress analysis for

determination of the bearing capacity of piles in co

hesive soils. The relations between different methods

are commented and discussed. A summary of general re

commendations for calculation of the bearing capacity

of piles is presented. Also general methods for calcu

lation of negative skin friction and settlement of a

single pile are summarized.

Some papers and current methods for design of pile groups

are selected and reviewed. The report contains diagrams.

tables and typical values of parameters that can be used

for design purpose or as a guide in a preliminary design

of pile foundations in cohesive soil.

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I

ACKNOWLEDGEMENTS

This report was done at my visit at SGI during 1981

according to SAREC's (SIDA) programme to which appreci

ation is expressed.

Great thanks to Dr Jan Hartlen, Director of SGI for

his recommendation on the work's programme, and his

assistance and encouragement.

Grateful thanks to Dr Bo Berggren at SGI for critical

reading of the manuscript and invaluable discussions

and recommendations.

Gratitude is expressed to Mrs Eva Dyrenas for her expert

typing of the manuscript and Mrs Rutgerd Abrink for draw

ing the figures.

also express my thanks to other members of SGI for

their kindness and their assistance during my time at

SGI.

Linkoping September 1981

Nguyen Truong Tien

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INTRODUCTION

In the last decade, numerous studies have been performed

to determine the behaviour of single, axially loaded piles

in cohesive soil. There are many factors influencing the

behaviour of piles. The most important factors are: soil

conditions, pile dimension, installation methods, pile

material and stress-strain history of the soil. Therefore,

besides existence of a method which takes into account

the variety of conditions, the designer must possess a

good knowledge of engineering science.

The ultimate bearing capacity of a single pile in cohesive

soil is in general limited by the ultimate strength of

the surrounding soil. The ultimate bearing capacity of a

pile can be evaluated from

a) Calculation methods based on the measured or estimated

shear strength of the soil.

b) Static penetration tests where the resistance is measured

when a penetrometer is pushed down into the soil at a

constant rate.

c) Dynamic penetration tests where the ultimate bearing

capacity of a pile is calculated from the number of

blows required to drive a penetrometer a given distance

into the soil.

d) Pile load test.

The accuracy of different calculation methods depends to

a large extent on the measurements of the strength and

resistance. The most reliable method to determine the

ultimate bearing capacity of single piles is by pile load

tests.

Negative friction will be produced where the surrounding

soil exhibits a downward movement with respect to the pile

shaft, and this effect can cause excessive settlement of

the piles with severe damage of the structure. Consequently

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there is a great interest in practical methods of re-

ducing the negative skin friction.

The settlement analysis of pile foundations depends on

the position of the load transfer from the pile to the

soil; and this is a complicated problem. Therefore, only

approximate solutions of this problem are available.

The behaviour of a pile group differs from that of a

single pile. The ultimate bearing capacity of a pile

group depends on soil type, size of the group, spacing,

length of piles and the construction procedures. The

evaluation of the ultimate bearing capacity and the

settlement of a pile group is based on empirical methods.

A review of various design methods for determination of

the bearing capacity and the settlement of the piles is

presented in this report. In the report diagrams, tables,

empirical expressions for design purposes have been

collected. Relationships between different methods and

recommendations for design have been summarized.

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1 • BEARING CAPACITY OF SINGLE PILES

Vertical axial loads applied on single piles are trans-

mitted to the surrounding soil from the pile by skin

friction and end-bearing. Qf is defined as the ultimate

load where both the total shaft resistance Q and thes' point resistance Qp are mobilized simultaneously.

= A f + A q s s p p = Qs + Qp

where A, A are the shaft and point tip area of the piles p respectively and f, q are the unit skin friction and s p unit point resistance. According to the Canadian Foundation

Engineering Manual (1978)

if < 100 kPaCu Qf = Qs and C, > 1 00 kPa Qf = Qs + Qpu

where C is the cohesion in undrained condition. u

Two types of approach are currently in use to evaluate

Qf: the total stress analysis and the effective stress

analysis. Also two methods are generally used in pre-

diction of the ultimate bearing capacity of the pile:

the method based on static formulas and the method based

om the result of field tests.

1.1 Method based on static formulas

1.1. Total_stress_analysis_for_bearing_caEacity

1.1.1.1 End bearing

The end-bearing Q is related to the undisturbed undrained p cohesion c of the soil below the pile (Terzaghi, 1943) and

is given by the formula

Qp = c N A + 0 N A - WU C p V q p p

W is the weight of the pile, but in general W and p p

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a A are omitted because the weight of the pile is often V p

about equal to the displacement soil (N =1 in cohesive q

soil ( c/>= 0) )

A = cross sectional area of the pile tip, m2 p

CU = cohesion of the soil, kN/m2

N ,N = bearing capacity factors C q

If soft clays, N is often assigned a value of 9, but it C

can vary from 5 in very sensitive,normally consolidated

clays (Ladanyi, 1973) to over 10 in overconsolidated clays

(Skemton, 1951). It also varies with the internal angle

of friction. (Fig. 1 , Meyerhof, 1976)

Meyerhof (1976) limits the value of Q at the critical p depth D of the pile.

C

For bored piles, the Canadian manual recommends

= N* c AC U p

where = ultimate point load, kN

= cross sectional of the pile point, m

= minimum undrained shear strength of the clay at the pile point level: kPa

N* = bearing capacity factor, which is a C function of the pile point diameter.

Point diameter N*_Q_

less than 0.5 9 0.5-1.0 7 greater than 1.0 6

In very stiff clays and till, cu can be measured by pressure-

meter. From experience, the Danish standard recommends that

Q = 18 Afp CU p

The Australian Code (SAA 1978) subjects according to

Whitaker & Cooke (1968) that the value of N varies with C

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L/B, where Lis the length of the pile and Bis the diameter.

If L/B > 4 N = 9 C

and L/B < 4 N = 5.6 C

Broms (1972) has pointed out that Qfp = 9 cuAp' where u

is obtained from fall cone tests, and the value of Qfp

generally corresponds to 10 = 20% of the total pile bearing

capacity.

1000

L ~r,,

,Jj/ ~;/,

J /

v/ 100

~.,,, ~ ,,✓ ~... /. // .// "'.,C: / //4' /

/,"7 r/ .. , ),"/ ,z ,,/ .. / ~ /.N .,,,, ~

/ u .,,,,,,, .,,,,,,,,I 1// 8. ,,. / 20 u Ne'"""., 1// // V00 ..C:

-::>-Nq ~

"'

i6 10 ....-

_,,,. .,,,., _,,,, ..[j....-,.,,,

/ /

V

;)/ /

V 1

/~.,... ..... // /

~

~ -

3

According to Caquot and Kerisel (1956) the point resistance

is given by the formulas

= C N + y D N . U C q

N and N are functions of the angle of internal friction C q

and are given in TABLE 1.

YD = overburden pressure at the level of the pile tip.

TABLE 1. Bearing capacity factors of point resistance (After Caquot and Kerisel, 1956)

1.1.1.2.1 Canadian Foundation Engineering Manual

Q = a c A s u s

The Canadian manual recommends a value of a according

to Tomlinson (1971) (see Fig. 2 and Table 2) . The value

of a is empirical, therefore the bearing capacity of

piles resulting from the above formula should be con-

firmed by load tests. For the case of bored piles in

clay, where c > 100 kPa, the shaft adhesion is calculated u by

Q = ac A s ua s

where c = ultimate adhesion kPa ua

Experience shows that

C =(0.3-0.4)Cua u

c is greatly affected by the excavation process. It is ua recommended that c is determined from the minimum un-ua drained shear strength c and that it is limited to a u maximum of 100 kPa. The ultimate load should be confirmed

by load tests.

TABLE 2. Design values of adhesion factors a for piles driven into stiff to very stiff cohesive soil (Tomlinson, 1971).

Case Soil condition Penetration Adhesion ratio factor

Sand or gravel, over- 20 1 . 25 1 lying stiff to very

stiff cohesive soil >20 see Fig.2

Soft clay or silt over- 8-20 0.4 2 lying stiff to very

stiff cohesive soil >20 1. 07

Stiff to very stiff 8-20 0.4 3 cohesive soil without

overlying strata >20 see Fig.2

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- -

1 0

Penetration ratio: Depth of penetration into stiff to very

stiff soil/Diameter of pile (relation between L/B in Fig.2).

Undrained Shear Strength (cul lb/ft 2

0 1000 3000 4000 5000 l.00

tj 0.75

0 0 ~ 0.50 C

-~ ., .c 0.25-0

1 1

1.1.1.2.2 Australian Code (SAA)

ct cA s

The values of a are presented in Fig. 3.

1.0

o.e \ l;j

~

0 \1-u 0. 6 ~

\.

'\ z ~ 0 ~ ~ u

0.4 -........ ----- -:::> 0 w ~ 0.2

0 0 100 200

AVERAGE UNDRAINED SHEAR STRENGTH, Cu,kPa

Fig. 3 Reduction factor a vs. undrained shear strength

for p iles in clay. (After Australian Code, 1978)

1 • 1 . 1 • 2. 3 SBN 7 5 ( 1 9 7 5)

The bearing capacity is calculated by

= a o P... ,...u ;::,

where A = shaft area of the pile s shear strenght determined by fall cone= CU test or vane test

a = 0.5, 0.8 and 1.0 for steel, concrete and timber piles respectively according to Broms' recommendation (1972).

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For a pile in tension the maximum load is 40 kN.

- SBN 75 does not allow the use of the upper 20% (or at

least 3 m when calculating the bearing capacity of a

floating pile).

- If a load test has been carried out at a constant rate

of penetration (CRP), the maximum allowable load corre-

sponds to 2/5 of the ultimate load with respect to soil

failure (FS=2.5).

- When the bearing capacity is calculated from the un-

drained shear strength the safety factor is equal to

3 • 0 .

A load test is carried out if the pile Class A is used.

(Q >600 kN). The procedure for pile driving tests is a 11ow-explained in Report No 59 of the Commission on Pile

Research.

1.1.1.2.4 Danish Standard

The value of a varies with the pile material as follows:

Q = ac A s u s

Timber pile a= 0.4

Concrete pile a= 0.32

Steel pile a= 0.28

The Danish standard requires that partial coefficients

in failure analysis is used. With a normal combination

of the loads (dead load+ live load+ snow or dead load

+ live load) the following partial safety factor should

be applied.

1tan

1 3

and the bearing ca~acity of the pile:

FS = 2 wi thou~ :toad testing__--

FS = 1 ,.6 with load testing

1.1.1.2.5 Building Code of the Soviet Union (Luga, 1965) see Kezdi (1975)

The maximum allowable load of the pile is calculated by:

P = nm (0Eaf. 1. + A q)max is i pp

where n = coefficient reflecting the scatter of the physical characteristics, usually n = 0.7

m = 1 for buildings, for bridges and hydraulic structures, see Table 3

0 = perimeter of the pile a = factor of safety (see Table 4) f.

lS = specific value of mantle

(see Table 5) friction (Mp/m2 )

1. l

= thickness of i:th l~yer A p = cross section of the pile tip

qp = ultimate value of (Table 6)

specific tip resistance

d

C=1rd A = 1rcl' /4

Fig. 4 Data for the pile formula published in the

Soviet Building Code.

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Table 3 Values of coefficient m

Structure

High piling Low piling

Table 4.

1-5

0.80 0.85

Number of Piles

6-10

0.85 0.90

0.90 1.00

31

1.00 1.00

Values of the coefficient a.

Vibrated Pile

Type of Pile Driven Pile Sand Coarse Silt Silt Clay

Sol id Pile 1 1.1 0.9 0. 7 0.6 Pipe pile 0.9 1.0 0.9 0.7 0.0

Table 5 Maximum unit values of mantle friction,

Sand, Fine Sand Silts and Clay Consistency Index le Average Depth Coarse

of Layer, to Rock m Medium Fine Flour 0.8 0.7 0.6 0.5 0.4 0.3

1 3.5 2.3 1.5 3.5 2.3 1.5 1.2 0.5 0.2 2 4.2 3.0 2.0 4.2 3.0 2.0 0.7 0.7 0.3 3 4.8 3.5 2.5 4.8 3.5 2.5 2.0 0.8 0.4 4 5.3 3.8 2.7 5.3 3.8 2.7 2.2 0.9 0.5 5 5.6 4.0 2.9 5.6 4.0 2.9 2.4 1.0 0.6 7 6.0 4.3 3.2 6.0 4.3 3.2 2.5 1.1 0.7

10 6.5 4.6 3.4 6.5 4.6 3.4 2.6 1.2 0.8 15 7.2 5.1 3.8 7.2 5.1 3.8 2.8 1.4 1.0 20 7.9 5.6 4.1 7.9 5.6 4.1 3.0 1.6 1.2 25 8.6 6.1 4.4 8.6 6.1 4.4 3.2 1.8 30 9.3 6.6 4.7 9.3 6.6 4.7 3.4 2.0 35 10.0 7.0 5.0 10.0 7.1 5.0 3.6 2.2

' 2 ton/m

Screw and Bored Piles

0.8 1.1 1.3 1.4 1.5 1.6 1.7 1.8 2.0 2.2 2.4 2.6

14

If the pile has an enlarged base, qp has to be multiplied

by the factor given in Table 7.

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Table 6 Ultimate value of specific tip resistance,ton/m2

Depth of Pile Tip, m

4 5 7

10 15 20 25 30 35

Granular Soils Gravel Coarse Sand Medium Sand Fine Sand

Cohesive Soils le 1.0 0.9 0.8 0.7 0.6

820 530 380 280 180 880 560 400 300 190 950 600 430 320 210

1050 680 490 350 :::40 1170 750 560 400 380 1250 820 620 450 310 1340 880 680 500 340 1420 940 740 550 370 1500 1000 800 600 400

Table 7 Reduction coefficients for

enlarged pile bases

Soil Type Beneath Base Ratio of

Base and Shaft Lean Clay Clay Diameters Sand Coarse Silt le~ 0.5 le~ 0.5

1.0 1.00 1.00 1.00 1.00 1.5 0.95 0.85 0.75 0.70 2.0 0.90 0.80 0.65 0.50 2.5 0.85 0.75 0.50 0.40 3.0 0.80 0.60 0.40 0.30

Coarse Silt

0.5

120 130 140 150 160 170 180 190 200

The following simple formula estimates the bearing capacity

of traditional piles under usual conditions:

in plastic clay Qf = 3 A s in mixed soil Qf = 6 A s in sand and gravel Qf = 1 0 A s

where A is s the mantel surface in m2 and Qf the failure

load in ton.

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1 • 1 • 1 • 2 • 6 Experiences from Thailand

Based on the test loading of piles in Bangkok clay:

Holmberg (1970) has obtained a relationship between

the adhesion factor a and the undrained shear strength

c as shown in Fig. 5. According to a new study of u Balasubramaniam et al(1981)this relationship is still

I

recommended for practical purpose.

l-1 0

.j.J u rd

4-l

Fig. 5

, 21----.----...----,,-----,----,--,------------,

o 0 25 - 30 cm. wooden piles.

101----.---+----'----f----f----l-----l

)(!Concrete T' o 22 l 22 cm, prestressed concrete pile. 09 o.a l-----+-'.f----J-----.,1-----.,1-----.,----1 \ ~✓Wooden piles 01

o Octagonal ( 0 58 cm} conc,ele pile.

a 35 x 35 cm reinforced concrele pile

( ) lndicoles lime interval in weeks

between piling and load teslinQ, (4)1\~2

) 0 61------tt-

1 7

TABLE Sa. Adhesion factor

a) C-u < 50 kPa Adhesion factor (a)

Steel piles 0.5

Concrete piles 0.8

Timber piles 1. 0

b) C > 50 kPa C u

Steel piles 1 0 kPa

Concrete piles 30 kPa

Timber piles 50 kPa

ExperieBces in Sweden indicate that the undrained shear

strength of a clay is often overestimated by the standard

test methods (fall cone tests, vane tests, or unconfined

compression tests) when the liquid limit or the fineness

number exceeds 80 (LL ~ fineness number, Karlsson 1961).

The undrained shear strength for clay is generally re-

duced in Sweden as follows:

TABLE 8b. Reduction coefficient

Fineness number Reduction coefficient (Approx.equal to LL)

80-100 0.9

100-120 0.8

120-150 0.7

150-180 0.6

>180 0.5

The bearing capacity of a pile which has been driven into

a normally consolidated clay is approximately equal to

the calculated, when the undrained shear strength has been

evaluated by fall cone tests.

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1 e

If the undrained shear strength is determined by uncon-

fined compression tests the critical load is underestimated.

The table Ba shows that the upper limit of the unit skin

friction resistance is equal to 50 kPa. When c < 50 kPa, u

the unit skin friction resistance is approximately equal

to the undrained shear strength according to Broms (1972)

and Tomlinson (1957).

1.1.1.2.8 Vesic (1977)

For overconsolidated clay Vesic has recommended

a = 0.45

1.1.1.2.9 The CTH method (1979)

Bengtsson et al (1979) state that the shear strength should

be determined by vane tests because

- lower cost than fall cone test

- more reliable value of the shear strength than the fall cone test at greater depth.

No consideration should be taken to the pile material

because results of Torstensson (1973) show that piles

of different materials but with the same shaft area,

shape and dimensions have approximately the same bearing

capacity.

The surface of failure occurs at a small distance from

the pile in clay.

The shear strength of soft, highly plastic clay is de-

pendent on the rate of deformation. The shaft resistance

for a cylindrical pile one month after installation was

equal to 0.9 times the undrained shear strength determined

in a field test with the same time to failure.

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Torstensson (1973) showed that the shear strength deter-

mined with the field vane test varies with the time to

failure:

where

T /T = 1.21. (tto )-0.053 er o

T = critical shear strength er

1 9

t 0 = time to failure in a standard test (1 min)

t = any time to failure if t = 3 h, T = 0.9T er o

The displacement modulus can be calculated by Butterfield

et al (1971) and Torstensson (1973) for the case that the

pile base is negligible. The ratio of the stiffness of the

pile and soil is calculated and used in Fig. 6a.

0.7 ! l I! I I : I I I 1L/d l 0.6 I l ! I : I l ! I! I I 2oi ! r Ii

0.5 I I! 1 i I ' I ! i

14o,5d I 1 I! I I It l

0:4 i 60! ! , : I

• MOOI :i~ I l; l ! i '! i 0.3 I

I I/ 1 f I I I' I : . (!)(/) Q2 i I i: ' : ] l 1

I l I I I .....____, -a 0.1 I ' Ii! 1 (/) ' ~

I I! I 0 l ''

2 2 5 2 5 103 104 105 106

Ep/Gs

Fig. 6a Diagram for determination of the initial dis-

placement modulus.

Ks to be used for calculation of load/displacement curve.

d = equivalent pile diameter G = shear modulus of the clay Es= equivalent Young's modulus for the pile op= Tshaft/Ks (After Bengtsson et al, 1979)

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From results of tension tests on floating piles, Torstens-

son (1973) presented a normalized shaft shear stress dis-

placement curve (Fig. 6b).

1.0 Q85

0.5

---- -7---------------1------- A - I --j !

1--- I I I ~-----==-====.r------J B i r------1 C

I I

CU () '-..

. I I I I I

I- 0 O 025 C.6 1 2 3 4 5

6/8f

Fig. 6b Idealized relationship between shear stress ratio

.h/c ) along the pile shaft surface and relative a

displacement (o/of) of the pile with respect to the

surrounding soil.

= friction resistance Ca of= relative displacement at failure A = curve representing conditions at a low

rate of displacement B,C= curve representing conditions at a high

rate of displacement.

(After Bengtsson et al, 1979)

The procedure of calculation is:

1. Determine with help from Fig. 6athe initial displace-

ment modulus K from given data s

a) the length of the pile L

b) the diameter of the piled

c) Young's modulus of the pile material E If the pile is nonhomogeneous E = crois sectional axial stiffness divided by nd 2 /i

d) Shear modulus of the soil G s

For normally consolidated, soft, highly plastic clays in Sweden Gs = 15 0 c u

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2. The complete relation between the shaft shear stress

and the displacement can be calculated by Fig. 6b.

Note:

a) The failure load is calculated by

Q = f · 0L f s

f s

0L

0.9

vane = 0.9 Tfu tf

= shaft area of the pile

= the mean value of failure shear strength from vane test

= factor due to time to failure (=0.9 if time to failure= 3 h)

= due to the time of installation (1 month after installation)

b) The value of the displacement o can be evaluated from

K o s = Tshaft

where T is the shaft friction, assuming that the shaft

displacement oat a load equal to½ bearing capacity

is calculated by Ts/2 where Ts is the failure unit skin

friction of the pile. The initial modulus cannot be used

to directly calculate the value of ofailure.

c) To account the variation of if the shear modulus and

the undrained shear strength are not constant a finite

difference program can be used.

3. Simplified calculation method

As results from load tests show the displacement of

the tip is less than 25% of the displacement of the

pile head for a load in the permissible range. There-

fore, it can be assumed that the pile tip does not

move or the axial deformation of the pile is equal to

the pile displacement. The axial force can be expected

to fall between the type of stress distribution 7a

and 7b.

SGI Varia 65

Fig. 7

where

z ,r

. . . . . . . . . . . . . .

T

~ . . . . . . . . . . . . . . . .

T

.·· z

z ~,

Paxial

. . . ..·

. . . . . . . . . . . .

Paxial

. . .

. . . . . . . . . .

Distribution of shear stress and axial load for

piles where

22

a) the bearing capacity of the pile tip is neglected.

The dashed lines represent typical shear stress

distribution and distribution of axial force

obtained during load tests.

b) The bearing capacity of the shaft area is

neglected. (After Bengtsson et al, 1979)

Pl 0.5 c5 Pl EA < < EA. 1 • 0 elast

p = axial load on the pile head

1 = length of the pile

E = Young's modulus for the pile material

A = cross-sectional area of the pile

SGI Varia 65

Experience from behaviour of piles in soft, highly

plastic clays in Sweden shows that the point bearing

capacity of those piles is less than 10% of the total

bearing capacity.

Bozozuk (1979) has recommended that this method is use-

ful for primary design, for detail design it is necess-

ary to carry out load tests.

1.1.1.2.10 The method of Caquot and Kerisel (1956)

= A f s s

A = shaft area of the pile s f = unit skin friction s f = T in clay for cp = 0 s max

T = C + 100 c2

max 100+7c 2

and for cp > 0

f = T +T 1 s max max

where ,, , ( 1 . ,i., ) ( TI/ 2 + cp ) tan cp " = c +sin'!' e max

The relation of T /c is presented in Table 9. max

TABLE 9.

cp 0

1 0

1 5

20

25

30

35

40

Relationship between cp and T /c. max

Tmax/c

1. 06

2.06

2.70

3.62

5.01

7.27

10.30

23 SGI Varia 65

1 . 1 • 2 Effective stress analysis for bearing capacity of iles

It is recommended that the bearing capacity of piles is

calculated by effective stress analysis because:

.2 4

- skin resistance of piles is governed by the effective

stress conditions around the shaft, the increase in

bearing capacity of the friction piles in clay is essen-

tially a phenomenon'of radial consolidation of the clay.

The gain in resistance with time should be controlled

by the time factor Th defined by

in which eh is the coefficient of radial consolidation

and t is the elapsed time since pile driving and B the

diameter. Available field data on the subject are

assembled in Fig. 8 after Ve sic ( 19 7 7) .

The method of installing the pile and the sequence of

strata through which a pile penetrates has an important

effect of the relationship between available shear re-

sistance and undrained shear strength. A larger amount

of scatter about the average values is shown in Fig.10

after Platte et al (1977). Fig.9 after Vesic (1977)

also shows that no correlation exists between skin

resistance and undrained shear strength.

The variation of skin resistance of piles in clay could

be better understood if test results are interpreted in

terms of effective stress and the equation

f = K tano'a' S S V

The main difficulty in applying the effective stress approach

is to estimate the radial effective stress on the pile at

failure and the evaluation of Ks in the above formula, or

SGI Varia 65

u. 0

Fig. 8

Length Type Dia. ft.

Soil type ------

~} steel H 14 {191} 219

D. steel pipe 6

A steel pipe 12

@l precas t 14 @( concrete

:ts~el of Plpe 24

22 60

{m

1242} 316 300

silt

soft clay

soft clay

soft boulder clay

soft to stiff clay

TIME , SINCE DRIVING (days)

Location Source

Tappan Zee, N.Y. Yang 1956

San Francisco

Michigan

Horten Quay

Eugene Island

/

Seed & Rees~, 1957

Housel 1958

Bjerrum et al., 1958

}Mcclelland, 1969

Stevens, 1974 ----(theoretical prediction)

Field data on increase of bearing capacity with

time for friction piles in clay. (After Vesic,

1 977)

the state of stress around the pile and in the pile itself.

In the effective stress analysis, the end-bearing capacity

is related to the effective friction angle of the soil

and the vertical effective stress in the ground at tip of

the pile. The skin friction is related to the coefficient

of friction between the pile and the soil and to the normal

horizontal effective stress. The ultimate load is defined

by

Q = A f + A q = Q +Q f s s p"p s p

where f, q are unit shaft resistance and point resistance s p respectively, evaluated from effective stress criterion.

SGI Varia 65

N .µ 4-l ....____

i::

I.S I SOURCE OF DATA:

S.. S...l.h{l95n B • a; .. ,_ (19Sll E • E;.i. ... 1 (1961! G • Gol.., (1913)

Gol.., l L'°"°'d {1914) L .L,&5-,«(1~) w • w..,.1,o1 (19SJ) R • Rot.lift & TOMliRSOft (19SJ)

26

I SYMl!Cl.l:

0 CAST-IN SITU CONCRETE PILE 0 DRIVEN COHCRETE PLE 0 STEEL PLE 0 TIMBER PILE

.. .. 0 .µ 1.0 ~s . s.-- (1959) .. (I)

0 i:: m .µ Ul

·r-l Ul 0.1 (I) H

i:: ·r-l ~ Ul

Fig. 9

Fig. 1 0

T • TOfllllU\MWI (1953) u • U.S.,-, Watorwoy, E,... Ito. (1950) .. w • Woocfwcnf et al (1961) h .. .. .. .. .. .. .. .. .. .. .. ,: .. .. .. .. .. .. - .. .. .. .. . . ... .. . . . . ... .. .. .. ... ''al .. .. .. .~ a, •• r ... - .. ..

=..- ... : ... 1 . ., g: a:.. ••.• ..

II a: 81 -· -.. ,: .. ,,,, .. . .. .. .. .. :r:r.:: - ., .. , ... -a. .. .. .. . .

1 z "' z 0

5 ~ w 0 .;; w

"

1.1.2.1

where

Fig. 11

End-be aring capacity

a' V =

= N o'A q V p

effective vertical stress in the soil at the tip of the pile

2 7

N q

= bearing capacity factor (Berezantzev et a 1 1 9 6 1 , Ve s i c ( 1 9 6 3 ) ) ( s e e F i g.. 1 1 )

tJ' z H 0 +l u m ~

:>-i +l ·r-1

A p

= cross-sectional area of the tip of the pile

10.000,-------- --r----~-- - ~-~-~

1000

De Beer. J ;i k y

Me yerhof

Beresan tsev. v~siC

U 100 m P., m u b, ~

·r-1 H m Q)

P'.l

TL"rzaghi

10~---~ ----

In compressible silty clay, the bearing capacity factor

has a value of about 10 (Blanchet, 1979) when the pile

ends in a saturated clay the above equation gives a

reasonable estimate of the point resistance of the pile,

(Bozozuk, 1979).

28

Vesic (1975 1 1977) has been working on the expansion theory,

and has recommended the following formula for calculation

of the point resistance

Qp = (cN* + a N*)A C O q p

in which N* and N* are appropriate factors, related to C q

each other by

N~ = (Nq-1) cot

1.1.2.2 Shaft friction resistance

The effective unit shaft resistance fs on a pile in

homogeneous clay is given by

where

f = c' + K 0 ' tan o ' S a S V

c' = unit adhesion between the soil and the pile, a which is independent of the normal stress

K s = earth pressure coefficient on the pile shaft

() I = effective angle of friction between the pile and the soil

Clark and Meyerhof (1972) measured the friction between

the soil and a steel plate and showed that as the shear

rate was reduced, c' decreased, and in a drained test it a

29

became equal to cero (Fig.12). Bozozuk (1979) came to the

. 0.,

5...------------------,

4

3

z

(o} • Und

30

same conclusion in a soil pile friction test. Meyerhof

(1976) and Vesic (1977) suggest that c~ can be neglected and

f = K 0 1 tano' S S V

or f = S0' S V

Some criteria on the calculation off are summarized s

below.

1.1.2.2.1 Burland

Burland (1973) followed Chandler's (1968) approach

and suggested the equation:

Q = A f sf s s

where f = K 0 1 tan~d = S0' S S V V

K = earth pressure coefficient. s

For driven piles, K s

K = K (safe side), S 0

clay:

K = 1-sin~'d 0

> K, so it is assumed that 0

and for normally consolidated

the effective overburden pressure

the remoulded drained angle of friction of the soil

(According to Tomlinson (1971), it is assumed that

the failure takes place in the remoulded soil close

to the shaft surface so o~ ~d).The reduction factor

Scan be written:

S = (1-sin~d)tan~d

Sis not very sensitive to clay type.

For normally consolidated clays S = 0.24-0.29

(~=20-30°).

SGI Varia 65

Fig.13 shows the relationship between the average

shaft friction (fs) and the average depth below the

ground surface, and Fig.14 shows the observed side

friction versus the effective vertical st~ess. Most

Cl) Q.) ... ... Q.)

E

Q.) u

2

4

2 :5 6 Cl)

"'O C: :::i 0 ... Cl

~ 8 Q) ..0

.r::. ... a. Q.)

"'O Q.)

~ 10 ... Q.)

~

12

14

Average shaft Friction - KN/m2

10

0

20

• •i ~

\ X II

\ . DV 6

I II

0

* 6

D

'v

X

+

• II

" ...

30 40 50 60

Steel ! Concrete Tomlinson (1957) Timber H.R.B. (1961) Sharman (1961) Brand (1971) Fellenius (1971) Eide etal (1961)

Concrete I r b Hutchinson and 1m er

S I

Jensen (1968) tee

_

000

\13=0·40 /{3:0·25

+

Fig. 13 Relationship between average shaft friction T

31

s and average depth for driven piles in soft clay.

(After Burland, 1973)

SGI Varia 65

values are between B = 0.25 and B = 0.4, with an average of approximately B = 0.32. It is reasonable

to take B = 0.3 for design purpose.

However, the frequency curve, Fig.~b for the quotion

of calculated and observed side friction, as presented

by Flatte et al (1977) r shows that the method sometimes

overestimated the skin friction resistance, but it is

better than the method based on total stress analysis

(see Fig.23 for comparison).

40

N

~ 30 z "' ;;i 0 ;:: u a:

20 ... w a in

"' ·~

33

1 . 1 . 2. 2. 2 The Canadian Foundation Engineering Manual (CFE.M)

CFEM recommends that the skin friction resistance can

be calculated by Burland's method:

Q = A f sf s saverage

Qsf = ultimate load capacity, kN

As = surf ace area of pile shaft, m2

f = average effective shaft friction, kPa savg fsavg is computed from the shaft friction fs at various

depths along pile shaft.

f = 0 ' K tan of' S V 0

or = (30 V

0 1 = effective overburden pressure at the considered V

depth, kPa

= rest earth pressure coefficient

= effective angle of friction between the clay and the pile shaft

K0

and of are difficult to measure. However, available

test results indicate that the factor K0tan of (or S)

varies only from 0.25 to 0.40 for normally consolidated

to slightly overconsolidated clays with cu less than

100 kPa. A typical value of 0.3 can be used for design

purpose or

Qsf = 0.3 0 1 A V S

It is recommended that the calculated value is confirmed

by a load test and in this case FS = 2.5 is applied. In

cases where no load test is performed FS = 3.0 should be

applied. CFEM recommends that if c > 25 kPa, the u effective stress analysis appears more rational.

SGI Varia 65

1.1.2.2.3 Meyerhof (1976)

The skin friction resistance is also calculated by

where

or

= A f s s

f = K o'tan~•

35

d evelops between the soil and the upper part of the pile,

when long piles are driven and this would reduce the

v alue of S.

0

5

50

75

,;" ~100 11 8-0 ~

l 125

150

175

20 0

I 0 • Cylindrical

., Tapered

o Negat ive skin frict ion

Theory, • I I

15° 20• 30• • • •• • •• -

• •• •• • • . ., -. • ' • • . . · f. • • . ., =·· •• • • ' •o •1 f •• • •• - Je --l·o • 0 • .., 0

0 References

1.0 • Beeemann (1969) ~ - Blessey (1970) I • Bjerrum, et al. (1969) -0 0 I

Bozozuk (1972) • Bozozuk and Labrecque (1969) •

Burland (1973) •• •• Darraeh and Bell (1969) -Eide, et al. (1961) I 0

I 0

Endo, et al. (1969) 0

Fellenius (1955)

I Fel lenius (1972) - . - Garlaneer (1972) -

I • Garneau (19H) fo

Mansur and Focht (1953)

Mccammon and Golder (1970)

I McClellan

1311tl f21B1tf 15011+ +m tt 0.1 0.2 0,3 0.4 0,5 0.6

Skin friction factm, i

Fig. 1 5 Positive and negative skin friction factors of

driven piles in soft and medium clay.

(After Meyerhof, 1976)

For stiff saturated clays, Meyerhof (1976) estimates

K from: 0

where

K (1-sin ip' ) i,1R 0 0

R 0

the overconsolidation ratio of the clay .

SGI Varia 65

3 6

Analyses of load tests on piles in stiff clay show that

B increases with the average undrained shear strength

(Fig.16). In the case of driven piles B = 0.5 for long

piles in a lightly overconsolidated clay and B = 2.5 for

short piles in a heavily overconsolidated clay. For

bored piles B = 0.5-1,5. As preliminary values K = s f or driven piles and 0.75 K for bored piles can be

0

1,5 K 0

taken.

If the K0-value is not known, the following relation can

be used:

and

f = 1 • 5 c tanqi s u for driven piles

for bored piles f = c tanqi s u

3,0r----T"""",---,,,-----,1----,,---... 1-------Sho,t piles ( D- 10 11-50 II): • Cylindri cal • Tape1ed Lone piles !D- 50 11-IOO II): o Cylind1ical '1 Tapered H H-piles

Oto (D >IOO It): D Cylindrical

1.51----+---1--/ _ _jl------l----l.----1--.--_J _

~ 1.5

j k 1.01-------1----~'----- London x. -3 +------1-----1-----l 5 , J. - • __ _ ,,, ,... :!

I . . H • ~ 0 - 4 ~ I - [ • App1ox1mate K 0-2

/. ________ .___ ~

• • o - 3 0

1.0 f----+-,tf---l------JI--H _.:_•_-1-__ _j. ___ _j_ __ H 1 • /"o_. 9o •• '?,

0

0 __ A_p~r~im•~•-,:_ __ +!

V V 0 0 e O H o.51-----l.hv;_•_::____JiL--•----J,----,.K>----J-----0-----!.--_j

o o D -~ ..... •--k>---~---

......... K.-0.5 O'----L----L---..J..----l---..l...--....l.-----lo

0 0.5 1.0 1.5 1.0 1.5 3.0 3.5 Mean undrained shear strength, ..:u, in tonrp·er square fool

References

Ballisager (1959) Ostenleld, et al. (1968) Burland (1973) Peck (1958)

Clark and Meyerhol (1973) Schlitt (1951)

Fellenius and Safflson (1975) Sherman (1969)

Fox, et al. (1970) Stermac, et al. (1969)

Kerisel (1964) Tomlinson (1957) and (1971)

Meyerhof and Murdock (1953) Woodward, et al. (1961 )

Fig. 16a Skin friction factor of driven piles in stiff

clay. (After Meyerhof, 197 6)

SGI Varia 65

where~ and c are from undrained shear tests. u

37

These expressions are found to be in agreement with some

load tests on piles in stiff clay.

1.0

1.5

... 0 tl .!! I 1.0

C

a

0.5

See Fig. 14 for symbols References

Burland (1973)

Chandler (1%8)

t------+----...,___•_-+---.--;-- Meyerhol and Murdock 11953) O'Neill and Reese (1972)

Skempton ll919)

I Touma and Reese (1974) ..!. _! _• L~o~ 0 •3 Wall, et al. (1969) 1-----+--,f--+---+----j-Whrlaker and Cook 119661

I •

Woodward, el al. (19611 •

0

0 0 0

• 0 • • Approx1mat~ K 0 -2 --- 0

• ~K0 -0.5

~ ---•

---Approximate K 0 ==l

oL-----L---....L----'-----'-----..._ __ __, 0 0.5 1.0 1.5 1.0 1.5 3.0

Mean undrarned shear strength, cu. in tons per square foot

Fig 16b Skin friction factor of bored piles in stiff clay.

(After Meyerhof, 1976)

1.1.2.2.4 Vesic 1 s recomendation (1977)

a) Normally consolidated clay

According to Burland (1973) and Chandler (1968) Vesic

(1977) recommends that

f = So' S V

where f3 = bearing capacity factor.

For a normally concolidated clay it can be assumed as

Burland (1973) did

K = K (1-sin~') S 0

or that

f3 = (1-sin~)tan¥

SGI Varia 65

where cp' = angle of friction of remoulded clay in a

drained condition (for 15° < cp' ~ 30°, B =.2-0.29). Vesic has also observed that B varies very little with

3E

the soil and the pile typ~ (see Fig.17) and he recommends

B = 0.29 for preliminary design. It is assumed that the

vertical component of soil stress remains unchanged during

pile driving. although the pile skin friction becomes a

slip surface, B can be calculated by:

sin"'' cos"'' B = 'I' 'I'

The above expression has been proposed by Vesic and gives

values of B that are 20% higher than from Burland's formula.

Based on an analytical approach Parry (1977) has obtained

the same formulas as those of Vesic.

co.... 1-z w 0 U 0.5 ii: ~ 0.4 0 u

0.3

~ a: 0.2 et w a:, 0.1

z 5,2 0.0 Cl) 0.0

0Detroit

eHorganza

111 San Francisco

0

C.Cleveland, Burnside

•Dra11111en '70rayton

50 100 kN/m2

150 200 250 I . I

- -• 0

Oo~ ~- 0 c. c. ~ A ~ «> 'o Oe c. Ill x~ @ - 0

0.5 1.0 1.5 2.0 2.5 3.0 VERTICAL GROUND STRESS ( t/sf) House 1, 1950 o Lemoore Woodward , 1961 Mansur, 1956 @Khorramshahr Hutchinson & Jensen, 1968

Seed & Reese, 1957 8 Oonaldsonville Darragh, 1969

Peck, 1958 Cll British Colt.J11bia HcCanmon, 1970

Eide et al., 1961 X New Orleans Blessey, 1970

Peck, 1961 o Misc. Locations Mcclelland Engrs.

Fig. 17 Observed values of skin bearing capacity factor

Bin normally consolidated clay. (After Vesic.1977)

SGI Varia 65

39

b) Overconsolidated clay

According to Burland (1973), the B-values can be calculated:

B = tan-"' u

.!: 6 J:.

a. Q)

"C Q) Cl

"' di ~ 8

10

12

o Wembley I Whitaker and • Tension { Cooke (1966)

\ ' ' \

II \

11 Moorfields .. D

Barbican Hayes St.Giles

Burland. Butler

and Dunican (1966)

\ ' \\ '\ \ '\x :

X Kensal green l Finsbury Skempton (1959) Millbank

\ . ',

O

'f \\. O+ \ \ I() ·, \\

\ ''\ \ \ •\ X \\

"'o\ oo ' . \ \, '(l ' • .-.__ \ \ ,,,_ . \

D \ \

eo\ .\ \ o •o '

\ \\ ~ \ \ '\\. -

\

"v" \ \~:-1.20

\,o_m overburden \

0:08\

Relationship between average shaft friction and

average depth in clay for driven piles in London

clay. (After Burland, 1973)

SGI Varia 65

40

The values of S evaluated by the above expression also

were plotted in Fig.18. Use of the value of S = 0.8 for

preliminary design is conservative. Vesic (1977) has

recommended the use of Fig.19 as a guide for design.

0 50 100 5

en. 4

kN/m2 150 200

0 North Seo boulder cloy (Fox, 1970)

6 Columbia Lock cloy ( Sherman, 1969)

• London cloy, Stonmore ( Tomllnsa,, 1971) 0 8Q900let cloy ( K,rlsel, 1964) I

250

I- 4D Ogeechee River sand ( Veslc, 1970) ~ 31-------1,------t,-------- Bored plkls In London cloy ( Bunand, 1973 ) (Eq 16 u I LL LL w 0 (.)

(!) z a:: ~ 11----=~-"G?----.=±----~~---+------i----7 w CD

~ :x:: ~ Ot-----L-------~---...i------'-------------~

0.0 0.5 1.0 1.5 2.0 2.5 3 .0 VERTICAL GROUND STRESS ( t/sf)

Fig. 19 Observed values of S for driven piles in stiff,

overconsolidated clay. (After Vesic, 1977)

1.1.2.2.5 Vijayvergiya and Focht (1972)

(The A correlation method)

Vijayvergiya and Focht (1972) summarize data from 42

load tests and relate available shear resistance to ver-

tical effective stress and undrained shear strength using

an empirically determined correlation factor A.

f = A(0 1 + 2 c ) S V U

SGI Varia 65

The values of fs' a~ and cu are average values over the

embedded length of the pile.

The correlation coefficient A was plotted against the

depth of penetration (Fig.20a).

41

Fig.20b presents a frequency curve according to Flatte

(1977) and shows that the A method generally overpredicts

the capacity of the pile.

The A method suggests that the available shear strength

of a pile in a normally consolidated clay may decrease

with increasing pile penetration (see Fig.20a).

The ultimate load can be evaluated by

Q = Q = A f f s s s

neglecting the point resistance.

According to Jimenes Salas (1976) the value of A can be

evaluated by

A= 0.0897 L + 2.781 L + 5.563

where Lis the embedded length of the pile, and can be used

for practical purpose in normally consolidated clay.

The frequency curve (Fig.20b) shows poor corre-

lation between the measured values and the predicted

values, since the tests reported in Fig.20b were generally

performed on 9~15 m long timber piles in normally con-

solidated or slightly overconsolidated. clay. Accord~ng to Fig.

20a, the values of A also have a big range of variation

in this range of pile length.

SGI Varia 65

Fig. 2 0

0 0,1 0,2 0,3 0,4 0,5

15

7T 30

E w· I- o, 0 ~= ..J 15 (am+ 2cm)As a: 45 ..J Detroit D Housel w 0 0 Morganza • Mansur z 12 0 Cleveland 0 Peck u

43

1.1.2.2.6 Plate et al (1977)

After a study of the data from load tests Platte et al

haveevaluated the value of the shaft friction resistance

by:

where

= f A s s

f = µL ( ( 0 . 2-0 , 0 0 1 I ) "7.'f::" o ' + 0 . 0 0 8 I c ) S p O V p U

= reduction of mobilized side friction with increased pile length

µL = (L+20)/2L+20) (see Fig.21)

R = overconsolidation ratio 0

I = plasticity index p

a• = effective overburden pressure V

c u = undrained shear strength

1,50r----

1,00

0,75

o\$2 N + + ..J 0,50 ..JN

" ~

0,25

0 10 20 30 Pile length,m

40 50

Fig. 21 Reduction of mobilized side friction with in-

creased pile length. (After Flaate et al, 1977)

SGI Varia 65

The formula is applicable to both overconsolidated and

norma lly consolidated clays.

Fig.22 shows the relation between the observed shaft

fric tion and adjusted effective vertical stress.

{FLATTE ANO SELNES, ·• :

•o

N

! z " ~- JO ;:: u

~ w a 0 in 20 w

" < ::; > <

10 illfflQ

e NC- CLAY 0 DC-CLAY

0 :"o '------'----~.o-----;so~--;:eo~----:-!:,oo::----:,~20:-------,-,.1:-0--__J,eo'---_J,eo

'ADJUSTED EFFECTIVE VERTICAL STRESS,'t'OCA•d'"..,O' KN/m2

Fig. 22 Observed side friction vs. adjusted effective

vertical stress. (After Flaate et al, 1977)

4 4

Fig.23 gives frequency curves for the quotient of cal-

c ulated and observed shafr friction for various formulas.

A= total stress analysis

B = effective stress analysis (Burlands 1973)

C = effective stress analysis (Vijayvergiya et al 1973)

D = effective stress analysis

( a = 1 )

(S=0.32)

f = A(o '+2c) S V U

f = So ' S V

E = effective stress analysis equation of Platte et al

SGI Varia 65

Legend: c fs=J\(pv'+ 2sul D OC·c I a y 14 pi I es t----+--+----,e----l

NC-clay 30piles

30

"' A fs = Su .! 20 -a. 0 .; 10

..Q

E :, 0 z

B fs = 0.32 pv' 20

10

0 0 50 100 150 0 50 100 150 200

Computed I observed side friction, percent

Fig. 23 Frequency curves for the quotient of calculated

to observed side friction for various formulas.

(After Flaate et al, 1977)

From comparison with Flatte et al (1977) it can

be suggested that S = 0.32 ~ (Fig.22) gives good pre-o diction of available shear strength resistance and this

value of Scan be used for design purpose. The reduction

factorµ can be taken from Fig. 21.

1.1.2.2.7 Bozozuk et al (1972,1979)

Bozozuk et al define the factorµ asµ= tano I tancp'

Note that for steel concrete or timber piles in clay,

silts or sand the value ofµ is

0.6 < µ < 1.0

For cylindrical piles driven into normally consolidated

clays, and the ground water table at the ground surface:

5 SGI Varia 65

where

f = K y'tano' s s

K = K S 0

46

y' = effective unit weight of the soil at the depth z.

Q can be calculated: s

where

Q = 1 0K y' tano I (L 2 -D 2 ) S

2 S

0 = the perimeter of the pile

L = the length of the pile

D = any depth above the tip level of the pile

The value ofµ can be taken according to Blanchet et al

(1980).

1.1.2.2.8 Blanchet et al (1980)

Q = A f sf s s

f = K tano'0' =B0' S S V V

For rough timber or precast concrete

For steel piles in clay

piles in clay o'= ~ 3 tano' =4 tan~'

For straight walled piles K =K =1-sin~• S 0

K =2K S 0

For tapered piles

B=(1-sin~')tan~'

B=2(1-sin~)tan~'

Note that the values of Bare recommended for the case

of the piles in soft and normally consolidated clay. When

the length of the pile is more than 15 m (L>15), use the

reduction of factor B according to Meyerhof (1976), see

Fig.24. In Fig.24 values of B calculated by Vijayvergiya

and. Focht (1972) are presented.

Note that in this Fig. the relation of

0.25 to 0.30 and y' = 5.5 kN/m 3 •

C

"1 is assumed to a'

V

SGI Varia 65

Fig. 24

0.....------,,----,----.,--r----ir----r----r LIMIT'S o, 0.-SE"°"ATI I ANO AVf:lltAG( R'ELATIONSHlfll fll1tOf>OS[O IY lll£Y£RHOf' (r1'7'1)

I 0 I

I I J..lli!!Q

I ·• TIMHR ltlL[ NOi• I MIO. 0 TUIIKII ttn .. f(LOUtHVtUl 11ft) • flM:CAff CONCM"T't ~ IIIO.S

•o,+---r--- #-f.----+-0 NECAST COftC'lltCT( ,u 11110-!l G ITUL l'1fl't: JIILf lfO 40• l..OA0IPNI) 0 ITUl. ,,..- '1t..l lt0.4(2MI \..Ql.0ffel) I

I ao+----+-

j

/ A 11ETH00 .ut1Ut11NG C!flO'W •oa TO 0.10 AIC) y• • I I ICN/ .. tYIJAYYOtfMYA a POCHT •1z)

LJIIITI o, oaH•vATIONI

LL.U+t=::P:,,;;,..-i ANO AVlltAO( IU:LATK>flftHt,

~110 IY IIIIYIRHO' 0971)

Skin friction factor

Frictional capacity of piles in normally con-

solidated or lightly overconsolidated clays.

(After Blanchet et al, 1980)

1.1.2.2.9 Esrig and Kirby (1979) (critical state method)

Qf = Qs + Qp

Qp = 9 C A u p L

Qs = D Sf dz 0 S

L = penetration length of the piles D = pile diameter

The skin friction along a vertical failure surface

f = s

1 µ0 cos.+.' 2 vf 'V

coscp' = factor to convert from maximum shear stress

to shear stress of the vertical plane.

47 SGI Varia 65

Assumed that after reconsolidation, the vertical and

circumferential stresses are equal and less than the

radial stress, Mohr-Coulomb's failure criterion gives:

_ 6 sin' µ -(3-sin

Fig. 25

Fig. 26

.8 ,-----,------.,-----,1---~l---~I---~

.6 f-

0

.4 f-

/ 0 / /

/ //.

/ . / __./ o

0

8 FROM DIRECT

0

o/(:-/ RECOMMENDED

·2

CORRELATION SIMPLE SHEAR TESTS _

0 FROM TRIAXIAL TESTS AND DIRECT SHEAR TESTS

Q ,...._ ___ ,...._ __ __,lc__ __ ..--Jt ___ ..--Jt ___ __,_1 ___ ~

20 40 eD 80 lOQ 120 140

LIQUID LIMIT

Summary of p /p- from laboratory tests on - cs nc twelve clays. (After Esrig and Kirby, 1979)

6 .::

~c, • f P (max) l( P(;:x)) ~

LOG MEAN NORMAL EFFECTIVE STRESS, LOG (i

49

Basis for derivation of expression for Pcs for

driven pile in overconsolidated clay. VCL = virgin compression line CSL= critical stress line

(After Esrig and Kirby, 1979)

SGI Varia 65

If sbbstitute~.-the expression ofµ into the

expression off will give s

and then

f = s 3cos~'

(3-sin~')

s = 3cos ~' sin~' (3-sin~')

50

For common values of~' (20-30°), the range of variation

of Sis 0.36-0.53. It is interesting to note that those

values are the same as have been reported by Vesic (1977)

for both normally consolidated and overconsolidated clays.

The values of Sin this @ase is apparently independent of

K. Note that a'f is the mean normal effective stress 0 V

(with the shaft pile) or ah. This approach works promising

because it is not necessary to evaluate Ks and o.However,

it is necessary to carry out comparisons with field

observations.

1.1.2.2.10 Janbu (1976)

Based on the classical criterion of failure of Mohr-

Coulomb and the degree of mobilization of the skin friction:

Janbu (1976) Grande et al (1979) have recommended that:

f = s (o +a) S V V

o' = average vertical effective stress V

a = soil attraction= c cot~

Sv = skin friction number

S = IrIµK V

K = coefficient of lateral earth pressure

r = a roughness number

51

z (m) 0 5 10 20 50 100

r 1.0 0.9 0.8 0.7 0.6 0.5

Q

• Q

/ _,,_, ': //_ -. I

z: 15 l p ' z

..Lill Jl !o ~ 0.. V

I I

I

I

C: I

p' Q) u 5

Jill __i._ LI... >---+I

• Qp

0,50

lv = Sv (p; +a)

i i 0.40

> I V') r-: cj ..0 ... E Q) :, ..0

C:

0,301 0,9 E

C: :, C:

.Q ~ u 4'

,.;: r---- C: 0.8 ..c CJ)

C:

-~ :,

.,:;. 0.20 0 V') - c:: 0.7

0.6

0,5

clay - silt - sand

0 0.2 0.4 0.6 0.8 1.0

fv'.obilized soil friction µ = f· tan \f)

Fig. 27 Static long term skin friction along piles.

(After Janbu, 1976)

The evaluation of the above formula is more

complicated than other methods. Numerical examples show

that the value obtained by this method is similar to

the S method, using the values of S derived according to Vesic (1977) and Parry (1977a) for normally consolidated

clays.

SGI Varia 65

52

1 . 1 . 2 . 2 . 1 1 Parry and Swain (1977b)

Parry and Swain (1977b) made the following observations:

1. The stress condition on the soil pile interface is

between (0'=0 1 T ) and (0'=0 1 T=O) where 0 1 > 0 1 h v' max h 1 ' h v (Mohr-Coulomb criterion of failure).

2. The failure occurs on the soil element and not on the

soil pile interface.

3. With the soil element at failure, the friction angle

development on the soil pile interface is o < ~•.

and

f = !30 1 S V

where 0 1 = vertical effective stress after installation V

of the pile

0' = m 0' (0 I is V VO VO

the stress before installation

of the pile)

m = 3-4 for stiff clays, m ~ 1 for soft clays

Parry (1980) has recommended m = 2.5 for tentative values

of i3 and the value of i3 can be defined as:

i3 =

where

m sin(o+y) cos ~•-cos(y+o) ec

siny = sino'/sin~•

and the value of o can be calculated from:

tano = ½ +l sin 2 ~'-n I

where the value of n = ~~ V

1 n

SGI Varia 65

53

The value of Sis also a function of OCR (R), Parry 0

assumed that S varies lineary with R0

in the range of

R0

= 1-12 and is constant after R0

= 12. As R0

varies

with depth, the value of S must be evaluated at different

depths·.

This method is very complicated and it is necessary

to make assumptions about some parameters. Numerical examples

show that the value of fs calculated by this method is

similar as calculated by the S method, using the correction

of Sin the case of overconsolidated clays according to

the recommendation of Meyerhof (1976).

1 • 1 • 3 Discussion

1.1.3.1 End-bearing capacity of a single pile

As has been reported by different authors, good agreement

between measured and calculated point resistance is ob-

tained by two formulas:

(total stress analysis)

(effective stress analysis)

The values of Q are depending on the values of the un-p drained shear strength c and the vertical effective u stress o~. The two formulas give the same result only

in the case that c ~a•, or if c increases with depth, U V U

and has the same value as a~ at the tip pile level.

However, it has been shown that this result may be otained

only in the case of overconsolidated clays (Parry, 1980).

In the case of normally consolidated clays the ratio

6u_!a~ generally is between 0.25-0.30 as reported by Bjerrum (1973), Esrig (1977). In normally consolidated

clays Blanchet (1980) has shown that the values of Q p

obtained by effective stress analysis are always greater

than the values of Q obtained by total stress analysis. p

SGI Varia 65

But as the value of Q is p 6-15% of Qf) and does not

small compared to Q (Q is s p have any great effect on the

54

total bearing capacity of the pile, the two above formulas

still can be recommended for design practice.

The method based on the expansion theory by Vesic (1972,

1977) takes into consideration factors that makes the

failure pattern similar to real conditions. Methods for

evaluation of end bearing are summarized in Table 10

1.1.3.2 The a-method

Table 11 shows that the values of a have a big range of

variation (0.2-1.5). It can also be seen that a is a

function of the pile types, soil conditions, material

of the pile, time to failure and the method of installation.

Available data from load tests also show that the corre-

lation between the skin friction and undrained shear

strength of the clay has only a limited meaning. However,

Fig.9 shows that for soft clays (c~ < 50 kPa) the

coefficient a should be equal to 1 or f ~c • This con-s u

clusion is confirmed by the work of different authors

(Tomlinson, 1957, Broms 1972) and the a-method can there-

fore still be recommended for design purpose. It is

interesting to note that the Canadian Manual makes the

comment that the effective stress analysis will give a

rational value only if c > 25 kPa, or in other words u it is recommended that the a-method is only used in soft

--:lays.

~his case of very stiff clays or overconsolidated

(c > 100 kPa) it is doubtful to use the effective u

analysis, because the values of Shave a very big

- variation, S reaches the value of 5 in some

is difficult to evaluate the value of K and s

~ese cases, because it is very complicated

d~ne the real stress state of the soil around

~~le at the moment of failure. On the other hand

SGI Varia 65

Table 10. Summary of the methods for calculation of end-bearing capacity.

Qp Apqp A (CN + iDN )A p C q p

4U)alysing Reference Based on the criterion Typical value of Nc,Nq methods

etmethocl Blanchet, 1980 Available data of load N =l, N =9

I test q C

I

Qp cN A Caquot Kerisel N =1.84 ,j, = 5o C p 1956 Nq=9.66 + YDN

C

q Ladany 1973 N = 5.0

C

Brems and Experiences in N = 9.0 SBN -75 Sweden C

Skemton 1951 N = 10.0 C

Danish Standard Experiences N = 18.0 1978 C

Australian Increasing Ne with N = 9 D/B > 4 Code 1978 depth Ne= 5.6 D/B-< 4

C

Canadian Diameter of the pile Ne= 9 B < 0. 5 Manual 1978 7 B = 0.5-1

= 6 B > 1.0

13-method Berezantzev 196 Theoretical analysis N = f (,P•)

q Blanchet 1980 Results of load N = 10 Q =N I A

tests q p q V p

Vesic 1977 Expansion theory Nq,Nc= f(,j,,Irr) qp =f(Oa, Oy_ )

Formulas of Qp Recommendations

(9c +YD)A Use in the case the weight u p of the pile is carried by

skin friction

(9.66 cu+l.84 YD)Ap

5.0 Sensitive clay

9 c A Soft clay u p

10 CA Use in overconsolidated clay up

18 c A Very stiff clay up

9 Cu Ap 5-6 CUAP

Bored pile cu > 100 kPa 9 CA for B

56

the driving of piles in stiff clays can be similar to

the driving of piles in dense sand. It may be for this

reason some authors have recommended the use of

the "a-method" for heavily overconsoidated clays. For

this purpose it is useful to make a statement about the

variation of a with penetration depth of the piles.

For practical purpose a~values recommended by Meyerhof

(Table 11) can be used. In the case of overconsolidated

clays a= 0.45 is recommended by Skempton (195). Vesic

(1977) gives a value of 0.55 for driven piles and 0.36

for bored piles. Meyerhof recommends~= 20° for bored

Table 11. Summary of the methods for calculation of skin friction resistance. Total stress analysis - a-method.

Q"'fA•OCA s s s u s

Reference Criterion on the Typical value of a Range of variation Comments and value of a for dcsian notes

Tomlinson, 1971 The values of a vary with For piles driven through soft Applicable for Canadian Manual depth of penetration in clay: driven pile. Fa, [978 stiff clay and type of L/B < 20, a = o .4 0.4-l .25 bored pile, " .

soil overlying and the L/0 > 20, " . l.07 0 .J-0 .4 values of c0 •

Brems, [972 a varies with pile ma- for LL < OD: 0.25-l .0 u • SO kPa a:: upper SBN, 1975 terial and LL > 80 steel pile: Cl = 0.5 o.=O. 25 {LL> 180, stcelpile) llmit and ocu •SO,

concrete pile: a .. o.a Ct""1.0(LL

57

piles and since a= tan~ this gives a= 0.36. On the

other hand the value a= 0.4 (L>20B) is recommended by

Tomlinson (1970) and a= 0.5 is recommended by API (1976).

All recommendations indicate that it is useful to take

a= 0.45 as a guideline for design purpose in overcon-

solidated clays. For "boulder clay" or glacial tills

where cu can reach the value of 80-350 kPa, a-values

after Weltman et al (1978) can be used for design purpose,

but they should be verified by load tests. The a/c u relationship is presented in Appendix B.

The values of a for long piles in soft clays shall be

reduced.

1.1.2.2 The S-method

The values of Sare evaluated based on empirical formulas

obtained from the results of load tests or on simple

theoretical assumptions. For normally consolidated clays,

the range of variation of Sis small (0.25-0.40) (see

Table 12) in comparison with the range of a in the total

stress analysis (0.2-1.5). The correlation between 0 1 V

and f is better than between c and f. s u s

Sis a function of the internal angle of friction, soil

pressure coefficient, length of the pile, type of piles

ratio of overconsolidation and other factors. Different

values of Sare recommended by different authors. Some

values of Sare evaluated from different criteria and

presented in Table 13. The values of S calculated by

Burland' s method are 20% lower than those of Vesic' s

method. However, Vesic's method gives values in better

agreement than Burland's with observed values by different

authors (see Fig.13 and Fig.14a). It is recommended to

use the formula of Vesic (1977) to evaluate the coefficient

s.

SGI Varia 65

Table 12. Summary of the methods for calculation of skin friction by effective stress analysis. 6-method.

Type or soil Reference

:

Normally Burland, 1973 consolidated

Canadian clay Manual NC 1978

Meyerhof 1976

Vesic, 1977 Parry, 1977

Burland, 1972

Ovcrcon- Vesic, 1977

solidated

clay Meyerhof, 1976

oc

Flaate, 1977

Esrig, 1977

Note: BP• bored pile DP• driven pile

Q9 • f 5 A 9 • K 8tan Q•o~ •$a~

Based on the method, Typical value of Formulas B for design

6 is based on the re-sults of the load test and assumption Ks = Ko o = q,d, B= (1-sinq,')

0.32

tan

D B= tan• f 9v K d 0.8 for bored

D qa 0 0 Z piles

Available data from Fig. 15' load tests

Available data from 2,5 for 060 (DP) B •(1-sinq,')tan

Table 13 •. Values of Sin function of$'.

Author Expression Angle of internal friction, drained remoulded of

50 10° 15° 200 25° 30°

Chandler, 1968 (1-sin$') tan$' 0.08 o. 15 0.20 0.24 0.27 0.29 Burland, 1973

Vesic, 1977 sin$' cos$' 0.09 0 .17 0.23 0.29 0.32 0.35 Parry, 1977

!+sin'$

Esrig, 1977 3sin$' cos$' 0.09 0.18 0.27 0.36 0.44 0.53 3-sin$'

Comments

S=0,24 ($=20°) for preliminary calculation of NF or pile in tension

S=o.29 ($=20°) for preliminary calculation of skin friction

Use of 0 1 VO

U7 '-.0

SGI Varia 65

60

It is very complicated to evaluate the value of S for

heavily overconsolidated clays because the value of Ks ~ K0

in these cases. According to the recommendation of Meyerhof

(1976) it can be expressed.

where S is used for calculation of skin friction for oc overconsolidated clays where S can be calculated by the nc method of Vesic (1977) or Burland (1973).

The values of Scan be calculated by Flaate's (1977)

formula. If R =1, S according to Flaate will be reduced 0

to the upper limit of Burland's observation for the value

of Sin normally consolidated clay. As the values of S

in heavily overconsolidated clay are very scattered, the

total stress analysis is recommended.

1.1.3.4 Relation between the S-method and A-method

The unit skin friction is calculated by

and

or

f =S0' S V

f = A(0'+2c) S V U

S0 1 = V

A(0 1 +2.c) V U

s X = 1 + ~ Cu 0'

V

or s =

The realtion between Sand A values also depends on the

relation of c and 0' u v·

If the values of S of Meyerhof (1976) and A of Vijayvergiya

et al (1972) are used the relation of 2c /0 1 can be . U V established. This relation varies with penetration depth

of the pile.

SGI Varia 65

61

Fig.28 shows the relation of c /0 1 and the length of U V

the pile derived from the A-method and the S-method.

Numerical examples show that A and Shave the similar

effect on the reduction of the skin friction when the

depth of penetration of the pile increases (see Fig.24).

0,6 r-----r---,----.-----.------.-----

0,2 ~ > 0 '-::, u 00 10 20 30 40 50 60

pile length, m

Fig. 28 Variation of c /0' with depth calculated from U V

data from Meyerhof (1976) and Vijayvergija et

al ( 1972) .

Below the depth L = 22.5 m the ratio of~ /0 1 will de-~u V crease. In general the values of c and 0 1 increase

U V

with depth, so the values of a will be reduced for long

piles as in the case of the S-method.

1.1.3.5 Relation between the a-method and the S-method

The unit skin friction is calculated by:

f = ac in the a-method s u and f = Sa' in the S-method

S V

In real conditions, independent of the method of cal-

culation, the value off is unique. Assume that the s

value off is equal in both methods or: s

SGI Varia 65

f = ac = So' s u V

or Q..J.l .@_ = 0 I V

Cl

For normally consolidated clay s, the relation of c /a' U V

can be defined according to Esrig (1978)

~ = (0.11+0.0037 I) a' P

V

The variation of cu/o' with I is plotted in Fig.29. V p

From the two expressions above:

Cl = s (0.11+0.0037 I )

p

6 2

In this formula it can be seen clearly that the values

of a will vary with S and with the consistency of the

soil. Fig. 29 also shows the variation of the ratio c / a ' U V

0,6

,,,,.,.✓

/ /

/ /

0,4 / /

-0,2

~,

Bjerrum 11~;~-=>--_,-

// V C 0 u =0,11 + 0,0037 IP ( Skempton l V

t:::,,..

20 40 60 80 100

Fig. 2 9 Relation between c /a' and I according to Skemton U V p

(1954) and Bjerrum (1973).

SGI Varia 65

(or S/a) with I according to Bjerrum (1973). p

63

As the a-method still continues to be used in practice

it is useful to take into account the variation of a

according to the length of the pile and the consistency

of the soil. Taking into account the consistency of the

soil (LL) it is recommended by Broms (1972) that the

values of a will be reduced at increasing LL values

(indicates generally an increase in I). p

For normally consolidated clays and silt where the value

of c /0' increases with depth and with I the two methods U V p

will give the same result. The ratio of c /0 1 = 0.25-0.30 U V

is reported by Blanchet (1980). For soft clays it is

common in Sweden that thts ratio generally is about 0.35.

If the average value of S = 0.32 as recommended by Burland,

and the range of c /0 1 as mentioned above are taken the U V

values of a should be:

C /0' = 0.25 a = 1 . 28 U V

C /0' = 0.30 U V

a = 1 . 0 7 C /0' = 0.35 a = 0.91

U V

This shows clearly that the value of a is approximately

1.0 which is in agreement with the value of a reported

by Tomlinson (1971) and Broms (1972).

Fig.29 shows the relationship between the ratio c /0 1 U V

and Ip according to Bjerrum (1973) and Skempton (1954).

The figure gives I = 40-65 for c /0 1 = 0.25-0.35. p U V

Hansbo's formula (1957) is widely used in Sweden:

,., ~ = 0.45 WL

C

where WL = liquid limit.

SGI Varia 65

The relation between a/S can thus be written

a = s

1 • 1 • 4 Summary of recommendations for design

1. The point resistance of piles in clays can be evaluated

by the a- or the S-method. N = 9 and N = f(~) according C q

to Berezantzev (1961) and Vesic (1977) can be used.

2. In soft clays (cu < 25 kPa) the a-method is recommended.

The value of a in those cases can be taken approximately

equal unity. In general the value of a decreases with

increasing I (plasticity index) and shear strength of p the soil. The influence of the length of the pile on

the a-value needs future development.

3. In the case 25 kPa < cu < 100 kPa it is recommended to calculate the skin friction by both the a and the

S-method. Note that for c < 50 kPa the value of a u is still approximately equal unity (see Fig.8). The

factor of Scan be calculated according to Vesic (1977)

or Parry (1977). The value of a can be calculated

according to Tomlinson (1971). It is interesting to

check the relation of c /o' and S/a. The same value U V

of ratio means the same value of skin friction.obtained by

both methods. The two methods give similar results in

the case of soft clays and silt where cu/o~ increases

with depth and with I . p

4. In the case c > 100 kPa the a-method is recommended, u where a can be taken according to Tomlinson (1970).

The value of a= 0.45 is useful as quideline in the

design of bored piles.

SGI Varia 65

65

If the soil pressure is known from previous experience or

from load tests the value of Scan be evaluated and the

S-method can be used. If the ratio of overconsolidation

of the clay is determined, the value of S can be oc

calculated according to Meyerhof (1976)

6oc = 6nc ~

where S can be evaluated by Vesic's method (1977). nc A typical value of K and Scan be found in Appendix D.

0

5. For long piles (L>15 m) it is important to account

for the reduction factor µL applicated on the skin

friction factors. The value of Scan be derived from

Fig.14 according to Meyerhof (1976) or from Fig.20a

according to Vijayvergiya (1972), but it is more use-

ful to use the expression of Flaate (1976) or from

Fig.21 where the value of Scan be expressed in a

general formula.

where S can be evaluated by common procedure (S=K tano') nc o or Vesic (1977). R

0 is the ratio of overconsolidation as

before and µL the reduction factor of S.

6. For long piles (L>15 m) the value of a should be re-

duced. The reduction factor can be taken from the S

nethod.

7. It is useful to take into account that when

C < 100 kPa Qu = Qs u C > 100 kPa Qu = Qs+QP u

and in general Qp =(6-15)% of Q . u

8. The safety factor should be more than 2.5. In the case

SGI Varia 65

that the allowable load Q 11 will be confirmed by a ow load test FS = 2.0 can be chosen. It is recommended

that partial FS will be applied to Q and Q p s Q = _Q_p + _Q_s_

allow 3.0 1.5

66

9. The shear strength c. can be determined in the field u

by vane test or in the laboratory by fall cone test.

10. For piles in glacial till of moraine clays the a-values

can be taken from Wettman (1978) (see Appendix B).

SGI Varia 65

1.2 Determination of point and skin resistances

from field tests

The calculation of point and skin resistance of piles

67

by static formulas requires not only a detailed knowledge

of the strength and deformation characteristics of the

soil strata, but also knowledge of variation of density

and water content within these strata. However, according

to Vesic (1972) the number of samples is often prohibitive

and cannot be justified except in the case of very important

structures. In all other circumstances, it may be prefer-

able to estimate q and f directly from field penetration p s or expansion test.

1 . 2. 1 Static cone penetration test

1.2.1.1 Vesic's experience wiht CPT shows that for driven

piles:

where q is the cone resistance. C

Except for highly sensitive clays the skin friction is

equal to the shaft resistance of the friction sleeve or

electrical penetrometer

or f = l.f s 2 cd

where fed is the average shaft friction of classical

Dutch cone.

1.2.1.2 Nottingham and Schmertmann (1975,1978)

The total ultimate skin friction resistance for piles in

clay layers can be determined from CPT results according

to Nottinghamas method (1975)

SGI Varia 65

8B z::

l=0

l fA +~fA) 8J3 SS t.. SS 8B

Q = total ultimate skin friction resistance s

K ,c = correction factor (see Fig. 30) s

B = the diameter of the pile

f = unit local sleeve friction resistance s

L = total embedded length of the pile

68

As= the pile/soil contact area per fs depth interval

1.4

1.2

.s LO +l

69

Qs = I:L',Qs = I: q . L',A s C s

where qc = the cone resistance at the increment depth

L',A = the perimeter area of the increment s s = a correction factor, see Table 14

Table 14. Correction factor S

Type of Type of soil Value of s friction cone

-

Fugro Sand 1. 6

Clay 1.0

Begemann Sand 1 • 6

Clay 0.6

1 • 2 • 1 • 3 Brorns ( 1 9 81 )

The cone penetration test can be used for cohesive soils

to estimate the undrained shear strength by the relation-

ship:

where qc is the average cone resistance.

Both the mechanical and electrical cone penetrometer with

separate friction sleeve (A=150 crn2 ) can be used to

estimate the shaft resistance of piles. The same criterion

in determination of c from q was given earlier by U C

Sanglerat (1972).

SGI Varia 65

70

1 • 2. 1 • 4 Tong et al ( 1981)

For normally consolidated or slightly overconsolidated

clays down to a depth of 6 to 8 m, when q < 1 MPa, the C -

shaft resistance can be determined by:

f = qc s 20

This value corresponds to a= 0.8 to 1.0. For stiff

clays, when q > 3 MPa f = const = 100 kPa. C S

For clays with 1 MPa < qc < 3 MPa

1 fs = 25 + 4() qc

For soft to medium cohesive soils (or cohesionless soils)

down to a depth of 6-8 m below the ground table

f = 20-25 kPa. s

1.2.1.5 Sanglerat (1972)

A statical analysis of tests performed in France, Belgium

and Holland was made at the Ecole Centrale Lyonnaise and

the results of the statical analysis are summarized in

Table 15.

Table 15. Local side friction f as a function of static point resistan~e q.

C

Clay and peats

where q < 10 bars q /30 < f < q /10 C C s C clays q /25 < f < q /12.5

C s C clays, silt, sand q /100< f < q /25

C s C sands q /100< f < q /50

C s C coarse sand and gravel f < q /150 s C

For clays the result above coincides with that of Begemann

(1965) where he recommended

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C u

71

The result is in very good agreement with local sleeve

friction (f) measurements and vane tests. s

According to Kerisel (1969) the maximum value of qc is

as follows

qc < 10 bar for soft clay or peat ( coincides with table above)

qc < 30 bar for medium clay

qc < 60 bar for silt

qc = (50-300) bar for sands depending on their

compactness.

2.5 N

8 0

...._____

tn ~

2.0 ~

·r-1

;::J 0

1.5

3 !

10 20 30 40

q in kc:r/cm 2 -c

Fig. 31 Relation between static cone penetration qc and

undrained shear strength cu in cohesive soils of

Colombia. (After Liem, 1970)

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According to Liems (1970) with the Delft cone

qc/cu = 10

qc/cu = 18

for

for

qc < 5 bar

qc > 27 bar

The result is presented in Fig.31.

1.2.1.6 Balasubramaniam et al (1981)

The ultimate load can be calculated by:

where

A

A

0 = the perimeter of the pile

= cumulative local friction

a = friction factor ( see Table

A = bearing factor (see Table

= cross sectional area p = shaft area of the pile s

qc = average cone resistance w = weight of the pile

1 6)

1 6)

Table 16. Friction factor and bearing factor for driven piles in the Bangkok area.

a A Investigator Medium Stiff Soft clay stiff clay Clay

clay

Pham (1972) 1.4 1. 4 0.7 0.33

Juta (1972) 1.0 1 • 0 1.0 0.33

Chotivitayatamin (1977) 1.1 0.7 0.5 0.33

Phota-Yanuvat (1979) 1.0 0.7 0.5 0.33

I

Sand

1 . 0

1. 0

0.5

0.5

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...,..., I ..)

Thsomputed bearing capacity and the measured_ultimate

102 from load tests have a good agreement(see Fig.32

Therefore the method is recommended for prediction of

the bearing capacity of piles in Bangkok clay.

Ul Ul .::: .::: 0 0

..j.J ..j.J LONG PILES ..

~15 SHORT PILES 'D m 0 0

r-1 r-1

Q) Q) ..j.J ..j.J m m 8 8

·r-1 ·r-1 ..j.J ..j.J r-1 r-1 :::J :::J

'D 5 A Stoel Pile 'D Q) o Concrete Pile Q) 100 H H AStffl Pile :::J o Wooden Pile :::J Ul Ul o Concrete Pile m m Q) 0 Q) :s 0 5 10 15 ~ 0

0 JOO 200 ~ 400 Predicted ultimate load,tons Predicted ultimate load,tons

Fig. 32 Measured vs. predicted ultimate loads.

(After Balasubramaniam et al, 1981)

David (1975) has found that

f = 1. 24 c or s u = 0.81 f s

where f is the sleeve friction and the undrained shear s u strength. The value of 0.81 is similar to the empirical

values of a of Bangkok clay.

Thomas (1965) suggests the following empirical relation-

ship for London clay:

C = q /18 U C

That is the same expression obtained by Liem (1970).

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1 • 2 • 2 Standard Penetration test

1.2.2.1 Meyerhof (1956)

Experience shows that the point resistance of driven

piles q (ton/ft2 ) can be related to N by:

p

N

-in which N = N when N < 15

and N = 15 + ½(N-15) when N > 15

74

S varies with the soil type and stress level. S = 2 for saturated clays and S = 4 for sands have been suggested

by Meyerhof (1956).

1.2.2.2 David (1979)

The ultimate capacity is defined by:

where

= Qs + Qp

= f A + q A s s p p

f = shaft friction s

f = ac s u a = the empirical adhesion coefficient, can be

estimated by Fig.33

qp = N =

C

C = u

c N U C

bearing capacity factor, N = 9 for deep foundations c

undrained shear stren~th estimated from SPT-N value

N 2 cu = 1.4 (t/m)

or c = 7.14 N (kPa) u

N = average value above the tip and below the tip for calculation of the skin friction and the tip point resistance respectively.

The author has obtained a good agreement between results

from load tests and prediction of the bearing capacity

SGI Varia 65

" c -~ 1.0 :i: 8 C

-~ .. .&:. -0

< o. 75

0.50

0.25

0 0

SPT N-Value, blows/ft 3

0.6