Appraisal of Reliable Skin Friction Variation in a Bored Pile

Proceedings of the Institution of Civil Engineers

http://dx.doi.org/10.1680/geng.13.00140

Paper 1300140

Received 21/10/2013 Accepted 08/09/2014

Keywords: field testing & monitoring/geotechnical engineering/strength &

testing of materials

ICE Publishing: All rights reserved

Geotechnical Engineering

Appraisal of reliable skin friction variation

in a bored pile

Nazir, Moayedi, Mosallanezhad and Tourtiz

Appraisal of reliable skin frictionvariation in a bored pilej1 Ramli Nazir PhD

Associate Professor, Department of Geotechnics and Transportation,Faculty of Civil Engineering, Universiti Teknologi Malaysia, Skudai,Johor, Malaysia

j2 Hossein Moayedi PhDAssistant Professor, Department of Civil Engineering, KermanshahUniversity of Technology, Kermanshah, Iran

j3 Mansour Mosallanezhad PhDAssistant Professor, Department of Civil and EnvironmentalEngineering, Shiraz University, Shiraz, Iran

j4 Alireza Tourtiz MScLecturer, Department of Civil Engineering, Beyza Branch, IslamicAzad University, Beyza, Iran

j1 j2 j3 j4

The design of bored piles in Malaysia is usually based on the results of the standard penetration test. It is important

to predict the geotechnical capacity of a designed bored pile through the multilayer soil strata. The back-analysis of a

test pile is a reliable means of obtaining the range for the ultimate skin factor (Ksu) and the ultimate end-bearing

factor (Kbu). In this research, two case histories of maintained load tests on single bored piles (PTP-1 and 2) under

full-scale static load (up to twice designed load) are examined. Measurements are taken using various embedded

transducers, including both conventional instrumentation and a state-of-the-art global strain extensometer. The

results show the rates of the pile base and pile head load mobilisation with settlement, the variation of the skin

friction factors and stresses along the pile, and their proportion in relation to the total pile capacity. The Ksu and Kbu

factors for both tested piles are obtained and compared using a conventional vibrating-wire global strain gauge and

a global strain extensometer. It is also observed that for the stiff soil layers the skin friction is significant. However,

the increase in the applied load increases the proportion carried by the end-bearing.

1. IntroductionBored piles are commonly used as foundations to support heavily

loaded structures, such as high-rise buildings and bridges, in view

of their low noise, low vibration and flexibility of sizes to suite

different loading conditions and subsoil conditions. These piles

are sometimes referred to as ‘bored cast-in-place piles’, as

specified in BS EN 1997 (BSI, 1990). The bored piles are formed

by boring using a suitable type of machine. Subsequently, the

holes are filled with high-workability concrete and some

reinforcement. Their usual sizes are between 750 mm and

3000 mm diameter, with a capacity that can achieve a very high

working load depending on the pile size and geological profile

near the pile (Fang, 2002). A higher pile capacity will reduce the

pile cap size and the number of piles in the group.

It is well established that the ultimate bearing capacity of a pile

used in a design may be determined by one of three values:

(a) the maximum load, Qmax, at which further settlement (or

penetration) occurs without the load increasing; (b) a calculated

value which is required based on the sum of end-bearing and

skin friction (shaft resistances); or (c) the load at which a

settlement of 0.1 diameter (0.1D) occurs (when Qmax is not clear)

(Meyerhof and Yalcin, 1983; Poulos, 1989, 2007). For large-

diameter piles, settlement can be large; therefore, a safety factor

of 2–2.5 is usually used on the working load. Accordingly, a safe

load (or designed load) can be calculated from the working load

divided by the factor of safety specified for a particular project.

It should be mentioned that, in maintained load tests (MLTs), the

piles are loaded up to a point near the safety factor times the

maximum load transferred from the above structures. However,

the pile will not approach failure during the test. Prakash and

Sharma (1990) have stated that the design load may be

determined by consideration of either shear failure or settlement,

and that it is the lower of the following two values: (a) the

allowable load obtained by dividing the ultimate failure load with

a particular factor, or (b) the load corresponding to an allowable

settlement of the pile.

Numerous studies exist regarding the prediction of the geotechni-

cal capacity of bored piles through soft soil and weak rock

(Hooley and Brooks, 1993; Ng et al., 2001; Xu et al., 2009; Zou,

2013). There are also various studies on the long-term measure-

ment of strain in instrumented piles. Kister et al. (2007) used

Bragg grating sensors for the strain and temperature monitoring

1

of reinforced concrete foundation piling. They were able to

successfully measure the change in the strain distribution along

the whole depth of the foundation piles. Fellenius et al. (2009)

explored the long-term monitoring (200 days record) of strain in

two 31 m and 56 m long instrumented post-driving grouted

cylinder piles at a site west of Busan, South Korea. They

monitored the unexpected elongation of the pile, probably due to

swelling from the absorption of water; however, as the soil

reconsolidated, the elongation shortened, probably because of the

imposed residual load in the pile. Brown et al. (2006) have stated

that the existing methods for test analysis generally overpredict

pile capacities by up to 50% for clays. They studied the load-

transfer mechanisms of rapid axial loading on a full-scale

instrumented pile in a glacial lodgement till near Grimsby, UK.

In order to gain insight into the load-transfer mechanisms of a

rapidly loaded pile in clay, they compared the shaft frictions

derived from the strain-gauged reinforcement in the pile with

shear strains and stresses derived from accelerations in the

surrounding soil. It can be seen that the design and the construc-

tion of a bored pile are highly empirical and that they are,

perhaps, more an art than a science (Tomlinson and Woodward,

2003). In tropical soils, which generally have complex soil

characteristics, the construction of a bored pile is a preferred

option in comparison to other types of pile.

In Malaysia, the design of bored piles is usually based on the

results of the standard penetration test (SPT). The empirical

approach to ultimate unit skin resistance ( fs) relates to

Ksu 3 SPT, while the same approach to ultimate base resistance

( fb) relates to Kbu 3 SPT. Both relationships are widely used in

common design works (Hanifah and Lee, 2006). To evaluate Ksu

and Kbu, the values of the local soil conditions are required, and

vibrating-wire strain gauges (VWSGs) and mechanical tell-tale

rods are installed along the piles. The installed strain gauges

within the pile allowed the monitoring of axial loads and

movement at various depths down to the pile shaft and the pile

toe (Badrun, 2011). Recently, to address the challenges and

difficulties posed by conventional measurement methods, a

retrieval sensor – a global strain extensometer (GSE) – has been

used for instrumentation of bored piles (Aziz et al., 2005; Liew

et al., 2011). This technology consists of a deformation monitor-

ing system that uses advanced pneumatically anchored extens-

ometers coupled with high-precision spring-loaded transducers; it

is a novel analytical technique to monitor loads and displace-

ments down the shaft and at the toe of bored piles through sonic

logging tubes (Hanifah and Lee, 2006).

Generally, the main objectives of loading tests are: (a) to

determine the load–settlement characteristics of the pile at the

expected designed load using both conventional and GSE meth-

ods; (b) to check the ultimate capacity of the pile and to calibrate

the empirical design methods employed for the more accurate

assessment of the bearing capacity of the pile at a given site. The

main aim of the full-scale tests and analysis presented here is to

investigate: (a) the rates of the pile base and pile head load

mobilisation with settlement; (b) the variation of the stresses

along the pile.

2. Experimental set-upThe prediction of the load capacity for a pile foundation is most

quickly done through a field test accompanied by the semi-

empirical method (Abu Kiefa, 1998; Anoyatis and Mylonakis,

2012; Coop and Wroth, 1989; Robertson et al., 1985). Thus, any

prediction or calculation should be justified through a full-scale

test (Fellenius et al., 2009). Axial pile load tests are among the

design procedures of most major construction projects that

include pile foundations, and the aim is to determine both the

pile stiffness and the ultimate bearing capacity at the designed

load depth (Comodromos et al., 2003). An MLT is one of the

best tests for predicting the actual behaviour of the axial pile

capacity (Brown et al., 2006; Consoli et al., 2003; Dai et al.,

2012; Holscher et al., 2012; Salgado, 2013; Zhang et al., 2008).

In an MLT, the load is applied in increments (in the vertical

direction), each being held until the rate of movement at both the

top and base of the pile has reduced to an acceptably low value

before the next load increment is applied (Tomlinson and

Woodward, 2003). It is, however, important to mention that the

reliability of the result will depend upon the instrumentation used

to acquire the relevant data.

The MLT test presented in this research is based on the reaction

pile system. The test follows the method described in the ASTM

standard D1143/D1143M-07 (ASTM, 2013). The clear distance

between the edges of the reaction pile to the edge of the test pile

should not be less than five times the diameter of the largest pile.

In the set-up used, the piles were loaded using hydraulic jacks

acting against the main beam. The jacks were operated by an

electric pump. The applied load was calibrated using vibrating-

wire load cells (VWLCs). To ensure the stability of the test

assembly, careful consideration was given to the provision of a

suitable system. The geometry arrangement should also seek to

minimise the interaction between the test pile, the reaction system

and the reference beam support. The capacity of the reaction

against the maximum test load should be 10–20% higher. A

typical load application and measurement system consists of

hydraulic jacks, a load-measuring device, a spherical seating and

a load-bearing plate. The jack used for the test should preferably

have a large diameter with a travel of at least 15% of the pile

diameter. Pressure was applied using a motorised pumping unit.

Pressure gauges were fitted to permit checking of the load. In

addition to the independent load-measuring device, linear variable

differential transducers (LVDTs) and optical levelling systems

were also used during the load test. All the devices were

calibrated before each series of tests.

2.1 Site condition

In this research, two series of full-scale MLTs were performed on

a bored pile. The first full-scale test was conducted at Cadangan

Pembangunan 2, Lorong Stonor, Kuala Lumpur, Malaysia, and is

denoted by ‘PTP-1’. The test pile was a preliminary and was

2

Geotechnical Engineering Appraisal of reliable skin friction variation

in a bored pile


loaded up to twice the pile’s structural capacity. It should be

mentioned that for test pile PTP-1, the required structural

capacity was 22 200 kN. PTP-1 was designed for a nominal

diameter of 1800 mm and a penetration depth of 36.95 m from

the existing piling platform depth of 36.25 m. The pile was tested

up to 44 400 kN (twice the designed load) in two loading cycles

for the initial test programme. The location of the second full-

scale project was at Utama Lodge, Jalan Senangria, Kuala

Lumpur, Malaysia, and is referred to as ‘PTP-2’.

From the subsurface investigation, Table 1 presents a summary of

the soil type and the standard penetration test (SPT-N) values,

respectively. The depth of the borehole in the vicinity of PTP-1

was 31.66 m and the soil profile at the borehole comprised very

stiff, sandy silt (at a depth of between 0 m and 24 m) and

fractured limestone (at a depth of more than 24 m). When there

is a rock layer more than 3 m thick, it is assumed that the

mentioned layer can be considered as a bedrock. In the case of

PTP-1, from z ¼ 24 m, limestone rock appeared (where z is depth

below ground level). As it was fractured, the SPT could not give

a reliable value. Therefore, core samples were taken from rotary

drilling, which show that the fractures in the limestone continued.

Accordingly, the layer below that was assumed to be a sedimen-

tary rock (fractured limestone). However, in depths lower than

32 m, a softer layer was found and the SPT was applied once

again. The depth of the borehole in the vicinity of PTP-2 was

47.5 m and the monitored soil profile at the borehole comprised

sandy silt, sandy clay, hard silt (at a depth of between 0 m and

23 m), with completely weathered sandstone (at a depth of more

than 23 m).

Table 2 presents a summary of the instrumented bored pile

load test. The 1800 mm diameter bored pile (PTP-1) was

instrumented using a seven-level vibrating-wire (VW) global

strain gauge and an eight-level VW extensometer. However, the

1000 mm diameter bored pile (PTP-2) was instrumented using

a five-levels VW strain gauge and a mechanical extensometer.

For PTP-2, the designed load was 6750 kN. Static load was

applied by hydraulic jacks acting against the reaction pile

system. The piles were loaded up to two times the designed

load, which was near to the safety factor of 2.5 considered.

For each loading, calibrated load cells were used to measure

the actual applied load on the pile head. The general soil

profile and SPT value (SPT-N) at the project site of piles PTP-

1 and PTP-2 are shown in Figure 1(a) and Figure 1(b),

respectively.

2.2 Bored pile construction and instrumentation

In this study, the bored piles were installed and concreted directly

into the study area. To install a bored pile, a borehole of a

specified diameter and depth – based on the required depth and

diameter for PTP-1 and PTP-2 – was drilled. Next, the borehole

was reinforced with a metal frame of a required cut and filled

with fine-aggregate concrete. As stated, both of the bored piles

were tested using the MLT method through the reaction pile

system. All of the instruments were logged automatically using a

Micro-10 data logger and multilevel software. The conventional

method for instrumentation using a VWSG and mechanical tell-

tales was employed. The VWSGs were attached to the steel cage

of the bored pile (used for PTP-2). The VWSG and mechanical

tell-tales were embedded in the concrete permanently. The second

Test pile Soil stratum Depth: m SPT-N values Average SPT-N

PTP-1 L1 Stiff sandy silt with little gravel 0–8 3–16 15.50

L2 Very stiff sandy silt with little gravel 8–10 16–50 27.5

L3 Hard yellowish sandy silt with little gravel 10–17 50–111 110

L4 Hard yellowish sandy silt with little gravel 17–24 111–150 122

L5 Fractured limestone 24–36.95 143–150 150

PTP-2 L1 Sandy silt 0–12 4–30 30

L2 Sandy clay 12–17 19–39 39

L3 Silt 18–23 54–125 122

L4 Weathered sandstone 25–31.65 176–200 195

Table 1. Summary of soil profile for test pile locations

Pile No. Diameter: mm Working load: kN Pile length: m Test load: kN Type of instrument No. of instrument

levels

PTP-1 1800 22 200 36.95 44 400 GSE 7

PTP-2 1000 6750 32.56 13 500 Conventional 5

Table 2. Summary of instrumented bored pile load test

3


in a bored pile


instrument used to measure the axial load and settlement

distribution along the bored pile was the GSE (used for PTP-1).

During static load testing, the deformation of the pile under

loading produces relative movement between each anchored

interval, causing a change in the strain gauge wire tension and a

corresponding change in its resonant frequency of vibration. The

resonant frequency is measured by plucking the GSE sensors/

transducers through a signal cable to a read-out box/data logger,

which also measures the frequency and displays the shortening

reading and the strain reading.

With the installation set-up as described above, this state-of-the-

art GSE system can measure shortening and strains over an entire

section of the test pile during each loading step of a typical static

pile load test; thus, it integrates the strains over a larger and more

representative sample. With the proper implementation of an

instrumentation scheme, the collected data from an instrumented

pile are more reliable, and a better and more meaningful

interpretation can be made. The obtained results from the GSE

method (PTP-1) were compared with the bored pile with conven-

tional instrumentation (PTP-2) results. For PTP-2, the Geokon

VWSG and tell-tale extensometers were installed internally in the

test pile to monitor the strain development and shortening behav-

iour of the pile during testing.

GSE instrumentation has been placed at seven levels for PTP-1.

The number of required GSEs depends on the length of the pile

and the vertical variation of the subsoil conditions, through sonic

logging tubes (Figure 2). A calibrated GSE sensor was installed

near the pile head (where no interaction from the soil friction to

the pile shaft is expected) for the calibration of the applied axial

load and the measured average strain. The GSE sensors measure

the strain and the axial load transferred through each section of

the pile shaft. In addition, the GSE sensor at the toe of the pile

measures the load contributed by the toe or else by end-bearing

resistance.

The VW extensometer was installed at eight depths at the

anchored intervals (Figure 3). Deformation of the pile under

loading produces relative movement between each anchored

interval. This causes a change in the strain gauge wire tension of

the VW transducers and a corresponding change in its resonant

frequency of vibration. The VWSG instruments for PTP-2 were

also installed at five levels (levels A through to level E), with

four per level (as shown in Figure 3). A schematic view of the

VWSG attached to the steel cage can be seen in Figure 4.

The gauges were checked before and after installation, after the

placement of the cage in the borehole and after concreting. For

the rod extensometer, galvanised iron (GI) pipes were tied to the

main reinforcement cage with steel wires at each terminating

depth, as shown in Figure 5. The 10 mm mild steel rod was

inserted until it touched the bottom of the pipe. A steel plate was

welded onto the end of the rod for the plunger to sit on during

the load test.

The pile head displacements were also measured by dial

gauges and LVDTs with readings to an accuracy of 0.01 mm.

These displacement measurement instruments were mounted

on stable reference beams, and the whole system was

protected from direct sunlight and disturbance by the person-

nel who were performing the pile testing and instrumentation

work. Settlement measurements using a precise levelling

technique were also taken as a useful backup, as well as to

check the movement of the reference beams. The VWLCs,

strain gauges, retrievable extensometers and LVDTs were

logged automatically using a Micro-103 data logger and the

MultiLogger software at 3 min intervals for close monitoring

0

12

17

23

Sandy silt

Sandy clay

Silt

Weatheredsandstone

(b)

34

32

30

28

26

24

22

20

18

16

14

12

10

8

6

4

2

00 50 100 150 200 250

SPT-N

Depth

: m

0

8

17

24

Stiff, sandy siltwith little gravel

Very stiff, sandy siltHard, yellowish,sandy silt withlittle gravel

Very stiff, sandy siltHard, yellowish,sandy silt withlittle gravel

Fracturedlimestone

(a)

34

32

30

28

26

24

22

20

18

16

14

12

10

8

6

4

2

00 50 100 150 200 250

SPT-N

Depth

: m

Figure 1. Variation of SPT with depth for soil in the vicinity of:

(a) PTP-1; (b) PTP-2

4


in a bored pile


during the loading and unloading steps. Only precise level

readings were taken manually.

3. Results and discussion

3.1 Bored pile deformation

Figure 6 and Figure 7 show the variation of applied load plotted

against pile top and base settlement, respectively, for two

continuous cycles on PTP-1 (Figure 6(a)) and PTP-2 (Figure

6(b)). During the first cycle, the observed maximum pile top

settlement at a loading of 22 418 kN was 9.60 mm. Upon

unloading to zero, the pile rebounded to a residual settlement of

0.36 mm. However, during the second cycle the observed maxi-

mum pile top settlement at the peak load of 44 036 kN was

24.63 mm. Upon unloading to zero, the pile rebounded to a

residual settlement of 5.34 mm. The relationship between applied

load plotted against pile base settlement obtained from the pile

load test is presented in Figure 7. The irregular shape in the base

settlement at each step of loading – particularly the first cycle –

at the location of sandstone or limestone might be the result of

rock particle rupture. The continuance of such an irregularly

shaped settlement, however, might also be due to the high excess

pore-water pressure (because of applied stresses from the pile)

produced in the small fractures of the ruptured rock leading to

non-uniform sliding of small particles.

As can be seen, the maximum pile top settlement for the two

designed loads is about 24.6 mm, which is very small in compari-

son to the length of the installed bore pile. Faisal and Lee (2013)

have stated that the critical shaft displacement should be rel-

atively small (in order to fully mobilise the shaft resistance)

compared to the large movement that is needed to fully mobilise

end-bearing. Excessive settlement and differential movement can

cause distortion and cracking in structures (Salgado et al., 2007).

0·0 m Anchored level A-0

1·0 m2·0 m

5·75 m

Glostrext Sensor 1a,Anchored level A-

9·375

Glostrext Sensor 2a,

9·50 m Anchored level A-

12·875 m Glostrext Sensor 3a,






34·60 m Glostrext Sensor 6a,35·95 m Anchored level A-636·45 m Glostrext Sensor 6a,36·95 m Anchored level A-

Pile toe at 36·95 m depth (RL 0·7 m)�

RL 36·74 m (pile top)

RL 36·25 m platform level

Global strain gauge level A (RL 35·25 m)

Extensometer level 1 (RL 34·25 m)

1800 mm Bored pile

Col. RL 26·875

Global strain gauge level B (RL 30·50 m)


Global strain gauge level C (RL 23·375 m)


Global strain gauge level D (RL 16·625 m)


Global strain gauge level E (RL 8·125 m)

Global strain gauge level F (RL 1·65 m)Extensometer level 6 (RL 0·3 m)Global strain gauge level G (RL 0·2 m)�

Extensometer level 7 (RL 0·7 m)�

denotes Glostrext anchored level (two sets per level)

denotes VW Glostrext Sensors (two sets per level)

Figure 2. Arrangement of the instrument at different levels for

GSE in pile PTP-1

5


in a bored pile


RL 70·724 m (pile top)

RL 70·392 m (existing platform)

Bored pile

4·286 m

VWSG and extensometerlevel A

5·559 m

VWSG and extensometerlevel B

7·816 m

VWSG and extensometerlevel C

17·816 m

VWSG and extensometerlevel D

23·816 m

VWSG and extensometerlevel E

32·068 m

COL RL 66·106 m

TT1

TT2

TT3

TT4

TT5 Pile toe at 32·568 m depth(RL 37·824 m)

Galvanised item pipefor extensometer rod

Reinforcement bar

Attached

VWSG: (four sets per level)

Tell-tale extensometer(TT, one on each level)

Figure 3. Arrangement of the VWSGs and tell-tale extensometer

in pile PTP-2

Figure 5. Galvanised iron pipes for tell-tale extensometer were

pre-installed at VWSG levelFigure 4. Schematic view of the VWSG attached to steel cage

6


in a bored pile


3.2 Measuring the axial load-carrying capacity of the

bored piles

The load distribution curves for the test cycles are plotted in

Figure 8 and Figure 9. The load distribution curves – capable of

indicating the load distribution along the shaft and the base –

were derived from computations based on the measured changes

in the strain gauge readings and estimated pile properties (steel

content, cross-sectional areas and modulus of elasticity). The

computations made for PTP-2 were based on as-built details

(including concrete record) known from the construction record.

The difference between the loads at any two levels (levels are

given from the top and bottom of the pile) represents the shaft

load carried by the portion of the pile between those levels.

For instance, for PTP-2 when the 6735 kN test load during the

first cycle was applied, almost 99.78% of the test load was

carried by skin friction (the portion of the load carried

between depths of z ¼ 0 m and z ¼ 32.06 m in comparison to

the applied load); the remaining 0.22% test load was carried

by end-bearing, as shown in Figure 9(a). For the second cycle,

the maximum applied load was 12 904 kN while approximately

95.58% of the test load was carried by skin friction; the

remaining 4.42% test load was carried by end-bearing, as

shown in Figure 9(b).

It can be concluded that the measured skin friction resistance

between 0 , z , 15 m and 0, z ,5 m in the soil at the vicinity

of PTP-1 and PTP-2, respectively, was much lower in comparison

to the designed load. The variations of the portions of skin

friction resistance were calculated at the end of each test (i.e. for

PTP-1 at depth z ¼ 36.45 m and for PTP-2 at depth z ¼ 32.06 m),

as shown in Figure 10. It can be seen that the higher load reduces

the effect of skin friction while increasing the influence of the

end-bearing portion. For example, the skin friction effect in PTP-

1 varied from the portion between 98.1% and 87.1% when the

applied load increased from 3313 kN to 22 418 kN (Figure 10(a)).

This, based on the evidence presented by Chin (1970) and

Fleming (1992), is true of piles that carry most of their load by

shaft friction. Owing to the observed low values for the base

resistance, it is suggested that the end-bearing resistance of the

bored pile should be eliminated in the design, especially when the

wet drilling method should be used.

0

5000

10000

15000

20000

25000

30000

35000

40000

45000

50000

0 5 10 15 20 25 30

Applie

d load: kN

Total pile top settlement: mmPTP 1 – first cycle PTP 1 – second cycle

(a)

0

2000

4000

6000

8000

10000

12000

14000

0 5 10 15 20 25 30

Applie

d load: kN

Total pile top settlement: mmPTP 2 – first cycle PTP 2 – second cycle

(b)

Figure 6. Variation of applied load plotted against pile top

settlement for two continues cycles: (a) PTP-1; (b) PTP-2

0

5000

10000

15000

20000

25000

30000

35000

40000

45000

50000

0 2 3 5 7 9 10

Applie

d load: kN

Total pile base settlement: mm1 4 6 8

PTP 1 – first cycle PTP 1 – second cycle

(a)

0

2000

4000

6000

8000

10000

12000

14000

0 1 2 3 4 5 6 7 8 9 10

Applie

d load: kN

Total pile base settlement: mm

PTP 2 – first cycle PTP 2 – second cycle

(b)

Figure 7. Variation of applied load plotted against pile base

settlement for two continues cycles: (a) PTP-1; (b) PTP-2

7


in a bored pile


Based on the soil properties in the vicinity of PTP-1 and PTP-2,

the pile may not be able to provide significant load capacity or

stiffness at 8 m depth below the platform. However, at depths

below 8 m, the long-term settlement of incompressible underlying

layers (e.g. the very stiff, sandy silts in PTP-1 and the weathered

sandstone in PTP-2) will increase the contribution of the pile in

relation to the long-term stiffness of the foundation.

For a bored pile installed through soft soil, the focus is

mainly on the skin friction. An increasing proportion is taken

up by the end-bearing, as the shaft has been fully mobilised.

Since the pile base was located on limestone (PTP-1) or

sandstone (PTP-2), the effect of the end-bearing capacity

should be carefully considered. In the present study, the piles

were loaded up to twice their designed loads. Under such

conditions, the skin friction may not be fully mobilised and

the point where the end-bearing capacity becomes significant

may not be reached. For example, the end-bearing capacity

portion after the application of twice the designed load was

18% and 12% for PTP-1 and PTP-2, respectively (Figure 7).

As such, considering displacement at the head of the bored

pile, a much higher applied load is needed if the end-bearing

tends to be significant.

3.3 Pile’s vertical shortening

The results of this research show the importance of considering

both elastic and plastic deformation behaviours during the axial

loading of a pile test. The variation of applied load plotted

against the measured total pile shortening by the GSE for PTP-1,

and the use of tell-tale sensors for PTP-2, are presented in Figure

11 and Figure 12, respectively. The pile shortened significantly,

up to 8.68 mm and 18.89 mm, when the applied vertical load

reached a maximum of 22 390 kN and 44 000 kN, respectively.

The plastic deformation behaviour of the test pile for high static

loads was 1.51 mm which, in comparison with the total length of

the pile, is insignificant. However, when the plastic deformation

results from the first cycle and the second cycle are compared,

there is a much higher incidence of permanent deflection in the

vertical axis of the test pile (Figure 11).

As stated earlier, the tell-tale sensors were installed in five

different positions along the vertical axis of the test pile. Figure

12 shows the influence of the applied load on total pile shortening

in PTP-2 for the two continuous cycles. As shown in Figure

12(a), the higher depth of the test pile (z , 7.816 m) resulted in

less shortening in the piles. The total elastic shortening deforma-

tion of test pile PTP-2 (for z ¼ 32.068 m) was 5.04 mm and

40

35

30

25

20

15

10

5

00 5000 10000 15000 20000 25000

Load registered: kN

Depth

belo

w p

latf

orm

leve

l: m

PTP-1-C1-3313 kN

PTP-1-C1-6883 kN

PTP-1-C1-11061 kN

PTP-1-C1-16056 kN

PTP-1-C1-20370 kN

PTP-1-C1-4698 kN

PTP-1-C1-8897 kN

PTP-1-C1-14016 kN

PTP-1-C1-17973 kN

PTP-1-C1-22418 kN

(a)

40

35

30

25

20

15

10

5

00 10000 20000 30000 40000 50000

Load registered: kN

Depth

belo

w p

latf

orm

leve

l: m

PTP-1-C2-5627 kN

PTP-1-C2-16664 kN

PTP-1-C2-24479 kN

PTP-1-C2-29181 kN

PTP-1-C2-33186 kN

PTP-1-C2-11351 kN

PTP-1-C2-22391 kN

PTP-1-C2-27138 kN

PTP-1-C2-31072 kN

PTP-1-C2-35465 kN

PTP-1-C2-37716 kN

PTP-1-C2-42184 kN

PTP-1-C2-40475 kN

PTP-1-C2-44036 kN

(b)

Figure 8. Load distribution curve for PTP-1 in: (a) first cycle and

(b) second cycle, computed from VWSG

8


in a bored pile


11.61 mm for the applied loads 6735 kN and 12 904 kN, respec-

tively. However, the corresponding plastic deformations of test

pile PTP-2 were less than 1 mm and 2 mm for the same loading

conditions, respectively.

3.4 Back-analysis of the full-scale pile-load test

As stated earlier, for bored piles the axial load capacity can be

evaluated empirically from the correlation of SPT-N values using

the modified Meyerhof method, where the ultimate bearing

capacity of a pile in compression is given by Equation 1

Qu ¼ KsN sAs þ Kb(40N b)Ab1:

where Qu is the ultimate bearing capacity of the pile (in kN); Ks

is the empirical design factor relating the ultimate shaft load to

SPT values (kN/m2 per SPT blow); Ns is the SPT value for the

pile shaft (blows/300 mm); As is the perimeter area of the shaft

(m); Kb is the empirical design factor relating the ultimate end-

bearing load to SPT values (kN/m2 per SPT blow); Nb is the SPT

value for the pile base (blows/300 mm); and Ab is the cross-

sectional area of the pile base (m2).

Generally, the results of the load-transfer parameters for each of

the soil layers are summarised in the corresponding correlation of

SPT-N values plotted against maximum mobilised unit shaft

resistance. The skin friction factor will be calculated as the

changes in the mobilised unit friction resistance over the changes

in the SPT-N for a 0.3 m penetration. A summary of the results of

the back-analysis of the ultimate skin friction factor (Ksu) for

PTP-1 and PTP-2 is given in Table 3.

The results of back-analysis of the ultimate end-bearing factor

Kbu for PTP-1 and PTP-2 are summarised in Table 4. The Kbu

values corresponding to the allowable settlement of 40 mm for

PTP-1 and PTP-2 were 7.8 kPa and 2.23 kPa, respectively. The

expected ultimate end-bearing capacities from the SPT-N results

for both PTP-1 and PTP-2 were 2977.3 kN and 309.2 kN, respec-

tively. Compared to the obtained values for skin friction from the

SPT-N results for PTP-1 and PTP-2, the end-bearing values are

considered quite small. It is important to note that the base

35

30

25

20

15

10

5

00 2000 4000 6000 8000

Load registered: kN

Depth

belo

w p

latf

orm

leve

l: m

PTP-2-C1-738 kN

PTP-2-C1-2013 kN

PTP-2-C1-3320 kN

PTP-2-C1-4741 kN

PTP-2-C1-6051 kN

PTP-2-C1-1316 kN

PTP-2-C1-2759 kN

PTP-2-C1-3986 kN

PTP-2-C1-5596 kN

PTP-2-C1-6735 kN

(a)

PTP-2-C2-1750 kN

PTP-2-C2-5025 kN

PTP-2-C2-7321 kN

PTP-2-C2-8661 kN

PTP-2-C2-10072 kN

PTP-2-C2-3381 kN

PTP-2-C2-6714 kN

PTP-2-C2-7976 kN

PTP-2-C2-9370 kN

PTP-2-C2-10584 kN

PTP-2-C2-11429 kN

PTP-2-C2-12714 kN

PTP-2-C2-12123 kN

PTP-2-C2-12904 kN

(b)

35

30

25

20

15

10

5

00 5000 10000 15000

Load registered: kN

Depth

belo

w p

latf

orm

leve

l: m

Figure 9. Load distribution curve for PTP-2 in: (a) first cycle and

(b) second cycle, computed from VWSG

9


in a bored pile


resistance of bored piles is usually ignored, since in comparison

to the magnitude of skin friction (particularly at the top of the

bored pile) the amount of end-bearing resistance in the soft soil

is negligible. In addition, it is difficult to obtain a clean base

during construction to ensure suitable end-bearing capacity.

Generally, the back-calculated Kb values represent a conservative

approach to end-bearing resistance factors, as the majority of the

piles were not tested to full failure. However, it can still serve as

a useful initial design guide for shaft resistance factors. The

back-analysis of a pile load test allows the evaluation of the soil

modulus and, consequently, the more accurate prediction of the

pile response. However, the obtained Ksu and Kbu values –

particularly for these case studies – could be different for other

80

84

88

92

96

100

0 10000 20000 30000 40000 50000

Skin

frict

ion p

ort

ion: %

Applied load: kN

PTP-1-C1-skin friction PTP-1-C2-skin friction

(a)

88

92

96

100

0 2000 4000 6000 8000 10000 12000 14000

Skin

frict

ion p

ort

ion: %

Applied load: kN

PTP-2-C1-skin friction PTP-2-C2-skin friction

(b)

Figure 10. The portions of the skin friction and end bearing varied

with the applied load for: (a) PTP-1 depth z ¼ 36.45 m; (b) in

PTP-2 depth z ¼ 32.06 m

0

10000

20000

30000

40000

50000

0 5 10 15 20

Applie

d load: kN

Total pile shortening: mm

PTP-1 – first cycle PTP-1 – second cycle

Figure 11. Effect of applied load on measured total pile

shortening by GSE for PTP-1

0

2000

4000

6000

8000

0 1 2 3 4 5 6

Applie

d load: kN


PTP-2 – first cycle – TT1 – 5·56 mz �





(a)

0

2000

4000

6000

8000

10000

12000

14000

0 3 6 9 12 15

Applie

d load: kN


PTP-2 – second cycle – TT1 – 5·56 mz �

PTP-2 – cycle – TT3 – 17·816 mz �second


PTP-2 – – TT2 – 7·816 mz �second cycle


(b)

Figure 12. Effect of applied load on measured total pile

shortening by TT system for PTP-2: (a) first cycle; (b) second cycle

10


in a bored pile


projects, depending upon the soil layer characteristics, loading

conditions and site effects. The available data are limited, and

thus more instrumentation data need to be combined to obtain

closer range values for the skin resistance factors and base

resistance factors. The use of the suggested values in this project

should be applied with caution, and establishing an MLT as a

prove test is recommended.

4. ConclusionThis study can help engineers evaluate a pile under designed load

conditions. From the preceding analysis and discussion, the

following conclusions can be derived.

(a) The GSE method significantly simplifies the effort involved

in instrumentation by enabling the sensors to be post-installed

after casting the piles. This method also minimises the risk of

the instruments being damaged during concreting work,

compared with the conventional method.

(b) For the stiff soil layers where skin friction is significant,

increasing the applied load reduces the effect of skin friction

while increasing the effect of the end-bearing portion. It can be

concluded that an increasing proportion is taken up by end-

bearing as the shaft is fully mobilised. The obtained unit skin

friction resistance in soft soil layers between 0, z , 15 m (for

PTP-1) and 0 , z , 5 m (for PTP-2) was insignificant in

relation to the total resistance capacity of the pile.

(c) The results imply that when the pile is loaded higher, the

influence of shaft friction is lower.

AcknowledgementsThe authors would like to thank the Research Management

Centre of Universiti Teknologi Malaysia (UTM) and Ministry of

Higher Education (MOHE) for providing financial support

through research vote R.J130000.7822.4L130, thereby bringing

the idea into fruition.

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in a bored pile


Appraisal of Reliable Skin Friction Variation in a Bored Pile

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