Information Paper on Frictional Piles

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Information Paper - Shaft Friction Capacity of

Mini-Piles

CTW/MKL/CYK

Issue No./Revision No. : Draft 1 (September 2011) Issue/Revision Date :September 2011

Information Paper

Shaft Friction Capacity of Mini-Pile:

A Case-Study in Infrastructure Exhibition Gallery at City

Hall Annex

STRUCTURAL ENGINEERING BRANCH

ARCHITECTURAL SERVICES DEPARTMENT

September 2011



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1. Introduction

1.1 In Hong Kong, piled foundation is widely used to support building structures if

the structure is located on a site where surface sub-soil strata are not sufficiently

strong enough to support the loads from the building structures. Piles are either

end-bearing on hard stratum, or friction piles/floating piles relying on the shaft

friction between the soil and the pile shaft, or a combination of both. Engineers,

however, have uncertainties on the shaft friction between the soil and the pile

shaft, especially for cast-in-situ construction using augered or Odex methods

which may have caused disturbance to the soil along the pile shaft. Furthermore,

if the rock end bearing stratum is available at a reasonable depth, the proportion

of shaft frictional capacity will be small when compared with the end bearing

capacity. Hence, for end bearing piles (e.g. rock-socketted steel H-piles and

large diameter bored piles), the shaft friction component is usually neglected.

1.2 However, shaft friction is developed after small relative displacements between

the soil and the pile shaft, and hence shaft friction often contributes the bearing

capacity in practical situations; this contribution is important for floating piles

where the bedrock is at great depth. For percussive piles, engineers in Hong

Kong may rely on the Hiley Formula in predicting the load carrying capacity of

piles. For non-percussion cast in-situ concrete pile, the design shaft friction (in

kPa) varies from a maximum of 1.6×SPT-N values for continuous flight auger

piles to 0.7×SPT-N values for piles formed by boring with an auger and

temporary casing, and the design values have to be further verified by trial piles

before construction.

1.3 This paper presents the field measurements carried out to measure the shaft

friction in an instrumented frictional mini-pile in a project in a project of our

Department at City Hall Annex for the development of a permanent planning

and infrastructure exhibition gallery, and will present the findings for the shaft

friction correlated against the SPT-N values of the sub-soil. The ultimate shaft

friction has also been estimated based on the results of the proof load test and

with the help of the commercially available finite element program - Plaxis 3D

Foundation.

2. Theory and Literature Review

2.1 Figure 1 shows the forces acting upon an axially loaded pile. In theory, the unit

shaft friction, fmax, is a function of depth, and by integrating that resistance over

the surface of pile can give the total shaft resistance Rs.



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Figure 1 Stresses and forces on an axially loaded pile

(Source: O’Neill 2001)

Because relatively large displacements are required in floating/friction piles to

mobilize the end bearing capacity, the ultimate bearing capacity of a friction

pile may develop up to 80 – 90% of its capacity through shaft friction. The unit

shaft friction (fs) along the pile shaft is theoretically determined by the sum of

pile to soil cohesion and friction components in the following equation:

fs = ca + σ’h tans

where ca and s are respectively the adhesion and friction parameters between

the soil and the pile shaft, and σ’h is the effective horizontal stress due to

overburden. Numerous theoretical methods (e.g. Nordlund method (1963), α-

method (Tomlinson 1971), Burland -method (1973), Nottingham and

Schmertmann CPT method (1975, 1978), Meyerhof method (1976)) have then

been developed to compute the shaft resistance along a pile shaft.

2.2 The α (total stress) method suggests that the ultimate capacity of the pile is be

determined from the undrained shear strength (cu) of the cohesive soil

(Tomlinson 1971). This method further assumes that the shaft resistance is

independent of the effective overburden pressure, and the unit shaft resistance fs,

is therefore given by the following equation:

fs = ca = αcu

where α is an empirical adhesion factor to reduce the average undrained shear

strength along the pile length. The coefficient α depends on the nature and

strength of the cohesive soil, pile dimensions, method of installation, and time

effects. The α-method is, however, applicable to cohesive soil with s=0, and is

not applicable to the typical soil in Hong Kong, where both cu and s are present.

Burland (effective stress) method (Burland 1973), which is applicable to both

cohesive and cohesionless soil, models the long-term drained shear strength



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conditions of piles using the effective stress, and the unit shaft resistance fs is

calculated using the following equation:

fs = σ’v

where (Bjerrum-Burland beta coefficient) = Kstan, σ’v = average effective

overburden pressure along the pile shaft, Ks = earth pressure coefficient, and =

friction angle between the soil and the pile shaft.

2.3 In Hong Kong, Foundation Design and Construction (GEO 2006) published by

the Geotechnical Engineering Office recommends the use of either Burland -

method or Meyerhof method, with latter being more commonly adopted to

calculate the unit shaft friction and base resistance of friction piles. Meyerhof

method (Meyerhof 1976) estimates pile capacity based on semi-empirical

correlation between standard penetration test blowcount (SPT-N values) results

and static pile load tests. The advantages of this method are that it is very easy

to use and that SPT-N values data is typically available for a project. Meyerhof

(1976) provides the correlation factor of the average mobilized shaft friction

fmax (in kPa) with the SPT-N values to be 2 for driven pile and 1 for bored piles.

GEO (1996, 2006) summarizes the work of various researchers and in-situ tests,

and provides the correlation factor of the average mobilized shaft friction fmax in

(kPa) with the SPT-N values for different types of piles and methods of

construction. GEO (1996, 2006) further recommends the base resistance to be

ignored in calculating the load carrying capacity of the pile. Based on the in-

situ measurements, GEO (2006) suggests that the average mobilized shaft

friction fmax (in kPa) in Table 1 for different types of pile can be used.

Table 1 Suggested correlation between mobilized (maximum) shaft friction

and SPT-N values for saprolites of Hong Kong

Types of pile Suggested fmax /N Limiting N-values

Bored pile 0.8 – 1.4 200

Driven steel H-pile 1.5 – 2.0 80

(Source: GEO 2006)

2.4 GEO (2006) cautions that “[t]he design method involving correlations with SPT

results is empirical in nature, and the level of confidence is not particularly

where the scatter in SPT N values is large”. For shaft-grouted mini-piles, in-situ

measurements have been carried out locally to correlate the shaft friction

between the soil and the pile shaft. Chan et al (2004) used a correlation factor

of 2.85 for a frictional mini-pile with post-grouting construction method in a

project for the former Kowloon-Canton Railway Corporation (the “KCRC”) in

Tuen Mun, Hong Kong, and found satisfactory performance in the subsequent

loading test to twice the design working load. Littlechild et al (2000), based

from a number of loading test results on the foundations in a KCRC project,

reported that the correlation factors range from 1.3 to 3.6. Similarly, GEO

(2006) summarized that the correlation factor can range from 1.4 to 5.5.

However, in-situ measurements on the shaft friction have not frequently been

carried out for frictional mini-piles without shaft grouting. Because of limited

data in Hong Kong, this paper recommends that the average mobilized shaft

friction fmax (in kPa) for mini-piles without shaft-grouting can be taken as 0.8 to

1.4N for design, and that the limiting value for SPT-N values can be taken to be



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80. An instrumented pile had been installed with strain gauges in a project in

Central, Hong Kong in order to validate the recommended correlation factor

between the shaft friction and the SPT-N values.

3. Project Details

3.1 The project is to provide a new exhibition gallery at the existing City Hall

Annex for the Planning Department. In order to facilitate the public to access to

the new gallery, a new 5-storey block comprising a lift and staircase is added

adjacent to the main entrance of the City Hall Annex. The location of the

proposed new block is shown in Figure 2. Photo 1 shows the architectural

impression of the completed project.

Figure 2 Location of the proposed 5-storey block

Photo 1 Architectural impression of the completed project

Proposed 5-storey Lift and Stairs Block



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3.2 Ground investigation shows that the area is underlain by 15 metres deep of

fill/marine deposit/alluvium overlaying 33 metres of completely decomposed

granite. Bedrock of Grade III granite can only be found at about 48 metres

below ground. The ground water level is governed by the tidal effect of the

nearby sea and the highest is at about +2.5 mPD in accordance with the data

obtained from the Hong Kong Observatory. Table 2 summarizes the soil profile

from the ground investigation together with the range of SPT-N values. As

Grade III granitic bedrock can only be found at about 48 metres below ground,

end-bearing piles (in the form of large diameter bored, or pre-bored rock-

socketted steel H-piles, or mini-piles) with pile length of approximately 53

metres will be required, and this was considered not to be an economical option,

especially that a lightweight structural steel superstructure scheme will be

adopted.

Table 2 Summary of sub-soil strata

Soil Type Depth interval (m) SPT-N values

Fill/ Marine Deposit/ Alluvium 0 – 15m 6 – 18

Completely Decomposed Granite 15 – 48m 43 – 200

Grade III Bedrock > 48m -

Having considered various factors (including subsoil profile, limited site access,

limited working space and the concern on noise and vibration problems

affecting the nearby City Hall), the piling option with 14 nos. of frictional mini-

piles (whose loading capacity would be provided by the shaft friction between

the soil and the pile shaft) was adopted. If the present design method is not

followed, the mini-piles will be installed at a depth of 7 or 8m more. Moreover,

the congested site also permits the use of small drilling equipment for mini-piles

rather than the heavy machineries for large diameter bored or pre-bored rock-

socketted steel H-piles.

4. Design and Construction of Mini-Piles

4.1 The frictional mini-pile for this project consists of a 152×152×37 kg/m UC

installed in a prebored hole formed into soil with a temporary steel casing with

minimum internal diameter of 305 mm and then grouted with cement grout

followed by extraction of temporary steel casing before the setting of grout.

Shear bars are provided to the steel sections. The maximum theoretical safe

loading capacity of the frictional mini-pile was 580kN. In the design, the

average mobilized shaft friction fmax (in kPa) for the mini-piles was taken as

1.4N and the SPT-N value was limited to 80. With a factor of safety of 3, the

design shaft friction fs (in kPa) was therefore taken as 0.467N. The shaft

friction of the pile section in the soil strata of fill and marine deposit have not

been included in the design, and the end-bearing capacity of the pile has also

been ignored. Typical details of the frictional mini-pile are shown in Figure 3.

During the construction, the casing of the upper portion of the instrumented

mini-pile was accidentally left in to a depth of 27m below the ground level, due

to the early setting of the grout before extraction. The founding depth of the

instrumented pile was at 45m below the ground level. The construction of the

14 nos. of frictional mini-piles commenced in May 2010 and was completed in

July 2010.



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Figure 3 Typical details of pile

4.2 Instrumentation

One of the 14 nos. of frictional mini-piles was designated as the instrumented

pile in order to verify the actual loading behaviour and to validate the shaft

friction values adopted in the design. A total of 20 nos. of Geokon vibrating

wire strain gauges were installed at ten levels along the pile length with 2

numbers of strain gauges at each level. The set-up of the instrumented pile is

shown diagrammatically in Figure 4.

Figure 4 Diagrammatical representation of the instrumented pile



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4.3 Field Measurements and Analysis

The shaft friction along the instrumented pile was calculated by measuring the

pile head movement and strains along the pile during a static loading test to

twice the theoretical load carrying capacity (i.e. 1160 kN). The load applied to

the pile head through four loading cycles, and readings of the strain gauges and

pile head movements were taken at each load increment throughout the loading

test. The load settlement curve of the pile is shown in Figure 5. The maximum

pile head settlement at a load of 1160 kN was found to be 11.132mm and the

residual pile head settlement was 2.515mm which are far below the limiting

values of 28.065mm and 6.542mm respectively in the contract.

Figure 5 Load-Settlement Curve of the Instrumented Pile

Figure 6 plots the variation of the strains along the pile length, and shows that

the strain along the pile decreases with the depth of pile in general. This

observation is in line with the expectation as the pile load was transmitted from

the steel sections to the cement grout and then to the surrounding soil by shaft

friction along the pile shaft. From the distribution of strain distribution, the axial

load distribution along the pile length can be deduced from the elastic modulus

of the pile section.



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CTW/MKL/CYK


Figure 6 Distribution of Average Strain Distribution along Pile Length

Figure 7 shows the axial load distribution along pile length, and also shows the

SPT-N values along the pile shaft. The relationship between the shaft friction

and the SPT-N values can then be deduced accordingly. Table 3 shows the

calculation of the axial load distribution and the shaft friction to the SPT-N

values along the pile length. The ratio of the shaft friction to the SPT-N values

for the pile section in the CDG layer of the soil stratum is found to range from

0.44 to 0.50 for the load tested. The test results well proved the construction

method and performance of the frictional mini-piles in this project. However,

the static loading test revealed that there could be much more reserve capacity

for this pile type due to the relatively small maximum and residual pile head

settlement obtained, and either a shorter pile length or a higher capacity for the

same pile length is feasible.



Mini-Piles

CTW/MKL/CYK


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Mini-Piles

CTW/MKL/CYK


4.4 Finite Element Analysis

With the data of the instrumented pile obtained from the static loading test, the

stress distribution along the pile shaft was further analysed by the commercially

available finite element program Plaxis 3D Foundation. By adjusting the soil

parameters in the model in the program, the load-settlement curve from the

computer analysis matches with that of the loading tests. Figure 8 compares the

load settlement curve for the pile from Plaxis 3D Foundation and that from the

static loading test.

Figure 8 Validation of the Finite Element Model

With the soil model/parameters established, the finite element model can then

be used to simulate the load-settlement curve of the pile for an applied load

beyond the maximum test load of the loading test until its failure. The load-

settlement curve of the pile in the finite element model is shown in Figure 9. By

plotting the commonly adopted Davidson’s off-set criterion for loading test (i.e.

PL/AE + D/120 + 4mm), it can be found that the ultimate load carrying capacity

of the pile is about 2600 kN.

Figure 9 Load-Settlement Curve of the Pile by Finite Element Analysis

The relationship between the maximum mobilized shaft friction and the SPT-N

values can therefore be deduced, and the mobilized shaft friction fmax to the



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CTW/MKL/CYK


SPT-N values for the pile section in the CDG layer of the soil stratum ranges is

found to be about 1.1. The result is in line with the in-situ measurements

summarized in GEO (2006), where the average mobilized shaft friction fmax to

the SPT-N values ratio ranges from 0.8 to 1.4 for the bored piles without shaft-

grouting in Hong Kong.

5. Concluding Remark

Shaft friction along a pile is hard to be estimated accurately especially for cast-

in-situ construction using augered or Odex methods which may have caused

disturbance to soil along the pile shaft. This paper presents the use of the

frictional mini-piles founding on soil in a project of our Department at City Hall

Annex for the development of a permanent planning and infrastructure

exhibition gallery. The relationship between the design shaft friction fs to SPT-

N values is obtained from the analysis of the results from the instrumented pile.

The ratio of shaft friction fs to the SPT-N values is found to range from 0.44 to

0.50 for the load tested. Finite element analysis with the use of the data from the

pile load test is used to estimate the maximum mobilized shaft friction of the

pile, and the result shows that the correlation factor between the maximum

mobilized shaft friction fmax to the SPT-N values is about 1.1, which is in line

with GEO (2006) that the ratio ranges from 0.8 to 1.4 for the bored piles without

shaft-grouting in Hong Kong. The test results in this project further well proved

the construction method and performance of frictional mini-piles. It also

demonstrates the benefit and option of using this pile type against the

conventional pile types which have to be founded on or socketted into the rock

where the rockhead level may be very deep.

References

Burland, J B (1973), “Shaft Friction of Piles in Clay”, Ground Engineering, 6(3), pp. 30-42

Chan, C K, Tsang, A H K, Chow, R N, and Tam, J Y C (2004), “Prebored Friction Mini-pile

Foundation for Light Rail Grade Separation”, The Structural Engineer, 82(20), pp. 24-7.

GEO (1996), GEO Publication No. 1/96: Pile Design and Construction (Hong Kong: GEO).

GEO (2006), GEO Publication No. 1/2006: Foundation Design and Construction (Hong Kong:

GEO).

Littlechild, B D, Plumbridge, G D, Hill, S J, and Lee, S C (2000), “Shaft Grouting of Deep

Foundations in Hong Kong”, in N D Dennis, Jr et al (eds), New Technological and Design

Developments in Deep Foundations (Houston: University of Houston), pp.33-45.

Meyerhof, G G (1976), “Bearing Capacity and Settlement of Pile Foundations”, Journal of the

Geotechnical Engineering Division, ASCE, 102(3), pp. 195-228.

Nordlund, R L (1963), “Bearing Capacity of Piles in Cohesionless Soils”, Journal of the Soil

Mechanics and Foundations, ASCE, 89(SM3), pp. 1–35.

Nottingham, L C (1975), Use of Quasi-Static Penetrometer Data to Predict Load Capacity of Piles

(Gainesville: University of Florida).



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O’Neill, M W (2001), “Side Resistance in Piles and Drilled Shafts”, Journal of Geotechnical and

Geoenvironmental Engineering, 127(1), pp. 3–16.

Poulos, H G and Davis, E H (1974), Elastic Solutions for Soil and Rock Mechanics (New York:

John Wiley & Sons).

Schmertmann, J H (1978), FHWA-TS-78-209 Report: Guidelines for Cone Penetration Test,

Performance and Deign (Washington, DC: Departments of Transportation).

Tomlinson, M J (1971), “Some Effects of Pile Driving on Skin Friction”, Proceedings of

Conference on Behavior of Piles, ICE, London, pp 107–14.

Tomlinson, M J (1994), Pile Design and Construction Practice (London: E & FN Spon, 4th

ed).

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Information Paper on Frictional Piles