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Appraisal of Reliable Skin Friction Variation in a Bored Pile
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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
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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,
16·25 m Anchored level A-
19·625 m Glostrext Sensor 4a,
23·0 m Anchored level A-
28·125 m Glostrext Sensor 5a,
33·25 m Anchored level A-
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)
Extensometer level 2 (RL 26·75 m)
Global strain gauge level C (RL 23·375 m)
Extensometer level 3 (RL 20·0 m)
Global strain gauge level D (RL 16·625 m)
Extensometer level 4 (RL 13·25 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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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
Total pile shortening: mm
PTP-2 – first cycle – TT1 – 5·56 mz �
PTP-2 – first cycle – TT3 – 17·816 mz �
PTP-2 – first cycle – TT5 – 32·068 mz �
PTP-2 – first cycle – TT2 – 7·816 mz �
PTP-2 – first cycle – TT4 – 22·816 mz �
(a)
0
2000
4000
6000
8000
10000
12000
14000
0 3 6 9 12 15
Applie
d load: kN
Total pile shortening: mm
PTP-2 – second cycle – TT1 – 5·56 mz �
PTP-2 – cycle – TT3 – 17·816 mz �second
PTP-2 – cycle – TT5 – 32·068 mz �second
PTP-2 – – TT2 – 7·816 mz �second cycle
PTP-2 – cycle – TT4 – 22·816 mz �second
(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
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz
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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12
Geotechnical Engineering Appraisal of reliable skin friction variation
in a bored pile
Nazir, Moayedi, Mosallanezhad and Tourtiz