technical paper
Pile test at the Shard London Bridge
David Beadman, Byrne Looby Partners, Mark Pennington, Balfour Beatty
Ground Engineering, Matthew Sharratt, WSP Group
Introduction
The Shard London Bridge, designed
by Renzo Piano and developed
by Sellar, is adjacent to London
Bridge railway station. The tower
will become the tallest building
in Western Europe standing at
310m and will include 60,000m2
of office, hotel and residential
accommodation.
Sellar appointed WSP as the
structural consultant for the project
and MACE as principal contractor.
Stent Foundations (now Balfour
Beatty
Ground
Engineering)
undertook the design-and-build
contract for the foundation piles
and the secant pile retaining
wall, with Byrne Looby Partners
undertaking the detailed design of
the foundations and retaining wall
for Stent Foundations.
The foundations comprise up
to 60m deep 1.8m diameter bored
piles founded in the underlying
Thanet Sand. This paper reports on
a 30MN preliminary pile test carried
out on a 53.8m deep 1.2m diameter
pile. The pile was instrumented
in detail to record settlements and
strains at various levels, enabling the
mobilisation of shaft friction with
settlement to be determined through
the various strata.
Notation
mOD metres Ordinance Datum
mSD metres Shard Datum
(0mSD=4.3mOD)
FOS factor of safety
cu undrained shear strength
α adhesion factor for undrained
soil, where skin friction = α.cu
Nq* equivalent bearing capacity
factor for Thanet Sand
σav’ average effective stress
σv’ effective vertical stress
Ks effective lateral stress
coefficient at perimeter of pile,
where lateral stress = Ks.σv’
Ko ratio of horizontal effective
stress to vertical effective stress
under at rest conditions
Φ’ angle of friction of soil
δ angle of friction on perimeter
of pile
εy average strain reading of the
four strain gauges placed at a
distance y from ground level
AG cross sectional area of the pile
EG gross stiffness of the pile
Ec stiffness of the concrete (varies
with strain)
The presence of such geological
features within this area of London
is documented in the Ciria report
Building Response to Tunnelling
(Burland et al, 2001) on the Jubilee
Line construction where similar
features were noted during the
tunnelling works. The fault lies
below the footprint of the tower
and resulted in displacement of the
London Clay, Lambeth Group Beds,
Thanet Sands and the Upper Chalk.
The typical ground conditions on
either side of the fault are listed in
Table 1. Data from either side of the
fault was plotted and indicated that
a good match was achieved when
plotted as metres below the top of
the London Clay rather than metres
above the Thanet Sands or chalk,
Table 1: Strata levels
Strata
Elevation of top of strata mSD
Ground conditions
West of Fault
A site investigation for the
development was completed in
2007 and the ground conditions
encountered in the exploratory holes
comprised made ground overlying
alluvium, River Terrace Gravel,
London Clay, Lambeth Group Beds,
Thanet Sand and the Upper Chalk.
A fault is present below the site
trending in a general north-to-south
direction. This fault has a 5m throw
and a downthrow to the east of the
fault line.
Made Ground
Alluvium
Terrace Gravels
East of Fault
0
0
-4.3
-4.3
-4.4 to -5.5
-4.4 to -5.5
Preliminary pile
record
0
Not recorded
-4.0
London Clay
-9.3 to -10.3
-9.3 to -10.3
-10.5
Lambeth Group
-30.0 to -30.5
-34.1 to -35.5
-31.5
Thanet Sand
-45.05 to -45.5
-50.3 to -51.5
-46.0
Chalk
-58.5 to -58.9
-62.5 to -65.5
Not recorded
Table 2: Soil strength parameters (pile design)
Strata
Density
(kN/m3)
Angle of
friction
(degrees)
Effective
cohesion
(kN/m2)
Undrained
cohesion
(kN/m2)
Maximum
shaft friction
limit (kN/m2)
Maximum base
bearing capacity
limit (kN/m2)
Made Ground
18
n/a
n/a
n/a
n/a
n/a
Terrace Gravel
20
38
0
n/a
n/a
n/a
London Clay
20
n/a
0
90+9.5/m
140
n/a
Lambeth Group
20
n/a
0
400
140
n/a
Thanet Sand
20
36
0
n/a
200
20,000
24
suggesting that the faulting occurred
relatively early in the depositional
history of the London Clay.
The made ground comprises
hardstanding surfaces over soft
to firm slightly sandy gravelly
clay which in turn overlay soft
brown gravelly clay (alluvium) and
medium dense to very dense sandy
gravel (River Terrace Gravel). The
London Clay comprises firm to
very stiff brown grey to grey brown
or grey fissured or locally thinly
laminated clay.
The Lambeth Group beds
comprise the following typical
sequence of strata:
n Very stiff multicoloured fissured
clay (Upper Mottled Clay);
n Very stiff grey clay with lenses/
pockets of silt and sand (Laminated
Beds);
n Very stiff grey or black shelly clay
(Lower Shelly Beds);
n Very stiff multicoloured fissured
clay (Lower Mottled Beds);
n Clayey gravel, green gravelly sand,
green shelly gravelly clay (Upnor
Formation);
The underlying Thanet Sands
comprise a very dense silty fine to
medium sand. Pressuremeter testing
using a Menard pressuremeter was
undertaken within Thanet Sands and
the results from these tests plotted
against depth below the top of the
Thanet Sands are presented as Figure
1. The results showed the expected
decrease in limit pressure with depth
in the Thanet Sands, as described by
Nicholson et al (1992).
Groundwater
monitoring
indicated that hydrostatic conditions
were generally present through
the London Clay and underlying
Lambeth Group beds with under
drainage in the Thanet Sands and
Upper Chalk. Within the River
Terrace Gravels, groundwater was
present at -4.8mSD. The lower
aquifer level, in the Upper Chalk,
was recorded at approximately
-38mSD.
Pile construction
The Shard has a three-level basement
over the entire footprint of the site.
The piles to support the structure
were constructed from the existing
ground level with approximately
15m of empty bore above the cutoff level of the pile. Many piles
were cast with plunge columns to
facilitate top-down construction.
ground engineering january
2012
Elevation mOD
-40
BH7
BH8
BH9
-42
-44
-46
-48
-50
-52
-54
-56
-58
0
5
10
15
20
25
Limit pressure MPa
Figure 1: Pressuremeter limit pressure v elevation (mOD)
The test pile was constructed
west of the fault from the existing
ground level and debonded to the
proposed basement level to model
the actual loading applied to the
piles in the permanent condition.
The debonding was achieved
using a bitumen coated steel tube.
The photograph (Figure 2) shows
the bitumen coated liner prior to
installation.
All the piles including the test pile
on the site were constructed using
standard rotary boring techniques.
Temporary casing was installed
through the made ground, River
Figure 2: Slip liner prior to installation
ground engineering january
2012
Terrace Gravels and approximately
1m into the London Clay to provide
a water seal during pile construction.
The piles were excavated in an open
bore to a depth of approximately
35m below ground level. Bentonite
support fluid was introduced into
the pile bore to provide support to
the pile during construction to the
required toe depth. The majority of
piles were founded in the Thanet
Sand at a depth of approximately
55m below ground level.
For the test pile, a 1,650mm OD
temporary casing was installed to a
depth of approximately 12m below
ground level. The pile was then
excavated at 1,500mm diameter
through the casing to a depth of
19.5m. The bitumen-coated slip
liner was lowered into the bore. The
annulus around the liner was then
grouted in two stages to secure it
into position. The first-stage grout
was poured to a level just below
the temporary casing. This was
allowed to reach initial set prior to
the temporary casing being extracted
and the second stage of grout was
poured. The pile was then excavated
through the bitumen-coated liner to
the toe depth of 53.8m below ground
level. Bentonite support fluid was
introduced into the bore during the
excavation.
It was anticipated that each
plunge column pile would take two
days to construct, being partially
excavated during the first day
and then left overnight prior to
completing the excavation, placing
the
reinforcement,
concreting
and installing the plunge column
on the second day. To assess the
impact of leaving the pile partially
excavated overnight, the test pile was
constructed over a similar two-day
period. In the event, a problem with
the piling rig during the construction
of the test pile meant that the pile
was left excavated to approximately
1m above the final toe level for two
nights. The bentonite support fluid
was maintained in the bore to a level
inside the permanent casing during
this time.
A full-length reinforcement cage
with four vibrating wire strain gauges
at six different levels was installed
into the pile bore. Strain gauges
were positioned at the top and the
bottom of the slip liner to establish
the shaft friction along the slip liner
and at other levels to identify the
mobilised shaft friction in the various
strata. Two extensometers were
installed into the pile to measure
the movement of the pile toe during
loading.
Following
the
cage
being
positioned, the concrete was
tremied into the pile displacing the
bentonite support fluid. The test
pile was constructed between 25
and 27 March 2009 and tested over
a four-day period from 23-27 April
2009. Concrete cubes were crushed
to ensure the concrete had reached
adequate strength to resist the
applied stresses.
The proposed pile test load meant
that the concrete stresses under
the jacks applying the load and the
concrete stress in the top section
of the pile were high. A heavily
reinforced pile cap was designed and
constructed to spread the loads from
the jacks into the pile (Figure 3).
Pile design
The soil strength parameters adopted
for the design are tabulated in Table
2. In the London Clay, the undrained
shear strength was assessed based
on the Stroud correlation between
SPT and undrained shear strength
of cu = 4.5N. The skin friction in
London Clay and in the cohesive
Lambeth Group was based on α.cu
where cu is the undrained shear
strength. α (alpha) was taken as
0.5 assuming pile construction
within a single shift. Where pile
construction was predicted to extend
beyond 24 hours, the value of α
was reduced to 0.45. The maximum
shaft friction was limited to 140kN/
m2 as recommended by the London
District
Surveyors
Association
(1999).
The angle of friction for the
Thanet Sand was based on a typical
value from published data (see
references, specifically Chapman et
al, 1999).
The ultimate end bearing
was limited to 20,000kN/m2
(Suckling and Eager, 2001) and the
ultimate shaft friction to 200kN/
m2 (Chapman et al, 1999). The
calculated base capacity was based
on Nq*.σav’ where Nq* = 47 and σav’
= average effective stress at the pile
25
technical paper
Figure 3: Pile cap and jacks
toe (Troughton and Platis, 1989),
which was based on a Ko value (ratio
of horizontal to vertical effective
stress under at rest conditions) of
1.15 by Troughton et al.
The base capacity exceeded the
limiting 20,000kN/m2, particularly
when the existing lower ground
water pressure was applied (as
noted below). The base capacity
did not, however, exceed that
indicated from the pressuremeter
testing, which would be of the
order of 32,000kN/m2. The shaft
capacity was based on Ks.tanδ.σv’
where lateral stress coefficient, Ks
= 0.7, pile interface friction angle,
δ = angle of friction of the soil, Φ’
and σv’ = effective vertical stress.
The calculated skin friction at the
depth of the Thanet Sand exceeded
the limiting 200kN/m2.
The Thanet Sand on this site was
deeper than for the majority of the
case histories reported to date and
therefore the vertical and average
effective stresses were much higher
in the Thanet Sand. Pile design
theory relates the shaft and base
capacities directly to the effective
stress as noted above and therefore
it is important to be cautious when
extrapolating pile capacity from
existing case history data.
This was the reason for applying
the conservative limits on the shaft
friction and end bearing capacities.
It is important to recognise that
the theoretical design pile capacity
is based on the worst case conditions
anticipated during the working
life of the pile, specifically ground
water conditions. Conditions during
the pile test may be less onerous.
The ground water level in the
Terrace Gravels was measured as
4m below ground level during the
pile construction. The water level
in the underlying Thanet Sand was
recorded in the pile drilling log
as 46.5m below existing ground
level (-46.5mSD), although the
site investigation and the recent
deep aquifer monitoring data both
indicate the water level to be higher,
at approximately -38mSD. This
higher figure is considered to be
realistic. The results of the analysis
of the test pile using the measured
strata levels (Table 1) and the
ground water level at -38mSD are
included in Table 4, using the same
limits on shaft friction and base
capacity as adopted in the design
calculations, with an alpha value
of 0.5. It is noted that the differing
ground water level records (-38mSD
or -46.5mSD) do not change the
results because of the use of the
limiting values of shaft and base
resistance.
The piles were designed for an
overall factor of safety of 2.25.
Working pile tests proved very
difficult to locate on the site due to
the congested nature of the site and
the presence of many existing piles.
It was therefore agreed that working
pile tests would not be carried out.
Instead, it was decided to design
for a factor of safety of 2.25 instead
of the normal 2.0 associated with
preliminary pile and working pile
tests or 2.5 for working piles tests
only, as defined by the London
District Surveyors Association
(1999). The pile design was also
checked for a factor of safety of
1.2 on the shaft friction alone as a
crude method of controlling pile
settlements.
The ultimate pile capacity was
calculated as 41,150kN. Applying
a factor of safety of 2.25 gave
a working load of 18,300kN.
However, as noted above, the pile
design was also checked for a factor
of safety of 1.2 on the shaft and
this reduced the pile capacity to
15,450kN. This was taken as the
Table 4: Test pile
design capacity
Capacity
Value kN
Overall ultimate
capacity
41,150
Shaft Capacity
18,530
Base Capacity
22,620
Working load (FOS
2.25 or 1.2 on shaft)
15,450
Table 3: Strata levels
Parameter
Nominal Diameter (m)
Length (m)
Toe Level (mSD)
Temporary casing – toe level (mSD)
Temporary casing – internal diameter (m)
Temporary casing – external diameter (m)
Founding Stratum
Penetration into Founding Stratum (m)
Value
1.2
53.8
-53.8
-19.5
1.23
1.25
Thanet Sand
8
Figure 4: Pile test frame
26
ground engineering january
2012
0 0
6 kN/m2
Stiffness
2
Stiffness
x1066x10
kN/m
2
6 kN/m
2
Stiffness
x10 x10
kN/m
Stiffness
Displacement
Displacement
(mm)(mm)
Displacement
(mm)(mm)
Displacement
50 50
00
10 10
5050
40 40
1010
20 20
4040
30 30
2020
30 30
3030
20 20
3030
40 40
2020
10 10
4040
50 50
5050
0 0
00
4 4
44
8 8
12 12
88
1212
16 16
20 20
1616
2020
24 24
2424
28 28
1010
0 0
0 0 100
100 200
200 300
300 400
400 500
500 600
600 700
700 800
800 900
900
-6 -6
00
Strain
Strain
x10
x10
100 200
200 300
300 400
400 500
500 600
600 700
700 800
800 900
900
0 0 100
-6 -6
Strainx10
x10
Strain
32 32
Load
Load
(MN)
(MN)
2828
3232
Load(MN)
(MN)
Load
Figure 5: Pile load v displacement
Figure 6: Measured concrete stiffness v strain
0 0
2
Stress
N/mm
2
Stress
N/mm
2
2
Stress
N/mm
Stress
N/mm
Depth
Depth
mSDmSD
Depth
mSDmSD
Depth
25 25
2525
20 20
00
10 10
1010
20 20
2020
15 15
2020
30 30
1515
10 10
1010
5 5
55
0 0
0 0
00
00
3030
40 40
BS8110
BS8110
Part
Part
1 1
Figure
Figure
2.12.1
2 12
fcu
f=cu42N/mm
= 42N/mm
BS8110
Part
BS8110
Part
1
Figure
2.1
Figure
Υm
Υm
= 1.0
=2.1
1.0
2 2
42N/mm
fcufcu= =42N/mm
Υm
1.0600 700
Υm
= =1.0
100
100 200
200 300
300 400
400 500
500
600
700 800
800 900
900
100 200
200
100
300 400
400
300
500
500
600
600
4040
50 50
5050
60 60
0 0
6060
00
-6 -6
Strain
x10
700 Strain
800x10
900
700
800
900
-6 -6
Strainx10
x10
Strain
5 5
55
10 10
1010
15 15
1515
20 20
2020
25 25
30 30
Load
Load
kNkN
(000s)
(000s)
2525
3030
LoadkN
kN(000s)
(000s)
Load
Figure 7: Measured stress-strain for pile concrete compared to
BS 8110 curve
Figure 8: Load distributions with depth
working load for the purposes of
assessment of the test results.
temperature compensation for the
strain gauges was not considered
necessary.
A
fluctuation
of
approximately 25 micro-strains
occurred in the range 0-80°C which
was not considered to compromise
the reliability of the strain readings.
The strain recorded in the strain
gauges at -1.5mSD was considered
to correspond to the load applied
to the pile, due the proximity of
Pile test data
The test was carried out in
accordance
with
the
ICE
Specification for Piling and
Embedded Retaining Walls (2007).
The pile was loaded (Figure 3) using
a load cell and a reaction system
consisting of a load frame and
four 1,500mm diameter anchors to
provide the reaction for the test as
illustrated in Figure 4. The test pile
was subjected to three cycles of load
to maximum loads of 100, 150 and
222% of the originally specified
working load of 13,500kN. The
pile settlement in relation to load is
shown on Figure 5. The maximum
settlement at ground level recorded
at 222% of the original working
load (30,000kN) was 52.18mm. The
maximum relative movement of the
base recorded by the extensometers
for the maximum load was
23.36mm, giving a maximum base
movement of 28.82mm.
The raw data from the pile test
was assessed initially and any
unusual or inconsistent readings
obtained from the strain gauges
were discounted. Errors in the
ground engineering january
2012
strain gauges could be a result of
erroneous or incorrectly installed
gauges or damage to the gauges
during construction of the pile. The
results from three strain gauges,
one each at -51mSD, -35mSD and
-1.5mSD were discounted at this
stage and the average strain at these
levels calculated using the remaining
three gauges. Temperature was
measured for the gauges. However,
the strain gauges to the head of
the pile and point of application
of the load (note that the diameter
of the concrete through the casing
is 1,230mm and this figure has
been used for the assessment at
-1.5mSD and at -19.5mSD). The
stiffness of the concrete in the pile
was estimated from the strain gauge
at -1.5mSD for each load increment
and the varying stiffness was
Table 5: Load from strain gauges
Jack load
(kN)
Load (kN) at each level (mSD)
-51.0
-44.5
-35.0
-28.0
-19.5
-1.5
3,374
134
367
995
1,625
2,834
3,115
6,833
486
1,131
2,591
3,954
6,383
6,787
10,135
972
2,206
4,488
6,455
9,910
10,189
13,565
1,825
3,723
6,782
9,369
13,644
13,761
16,879
2,519
5,099
9,125
12,059
16,909
17,070
20,254
3,980
6,953
12,098
15,008
20,014
20,239
23,626
6,200
9,456
15,140
18,067
23,185
23,530
27,000
8,586
12,023
18,084
21,059
26,429
26,800
30,005
10,788
14,419
20,731
23,712
29,161
29,638
27
technical paper
used to obtain the loads at all
other strain gauges as described
below.
The stiffness of the concrete
during the test period was found to
vary with the strain as shown on the
graph presented in Figure 6, from a
maximum value of 42x106kN/m2
at 60x10-6 strain to 26x106kN/m2
at 780x10-6 strain. This is compared
to the graph presented in Figure 2.1
of BS 8110 Part 1 – as shown on
Figure 7. A similar stiffness curve
is found if a concrete cube strength
of 42N/mm2 is adopted, setting the
concrete material factor to 1.0 for
this comparison.
The strain gauge data for each
load increment at each level was
averaged and was converted to
a corresponding load, using the
varying concrete stiffness as noted
above. The formula below was used
to obtain the load at each strain
gauge.
Fy = εyAG EG
where: εy is the average strain
reading of the four (or three)
strain gauges placed at a distance
y from the surface, AG is the cross
sectional area of the pile and EG
is the gross stiffness of the pile.
The pile gross stiffness allows for
the pile reinforcement (34 H40s
as compression reinforcement) in
addition to the pile concrete. The
stiffness of the casing has been
ignored in this assessment because
the strain gauges at -1.5mSD and
-19.5mSD are close to the top and
bottom of the casing and therefore
the casing does not appear to change
the pile stiffness.
AGEG = (π x 1.22 / 4 – 34 x π x
0.042 / 4) x Ec + 34 x π x 0.042 / 4 x
210 x 106 kN/m2
ie “Concrete area” + “Steel Area”
where: Ec is the stiffness of the
concrete relative to the strain at each
strain gauge.
Load transfer along shaft
of pile
Table 5 summarises the jack load
and the load at each level from
the strain gauge results.
The
assessment of the load at each level
based on the strain gauge results
and the assessment of the concrete
stiffness do not give exact figures.
The measured loads at -1.5mSD
do not correspond precisely to the
recorded jack load for the first load
increment. This may be because
the casing had some effect on the
stiffness at this depth for the first
load increment. The bottom of the
casing was at -19.5mSD and was
unlikely to affect the measured
strain at -19.5mSD. The comparison
between the recorded jack load and
28
Table 6: Measured and design shaft friction in different strata (kN/m2)
Jack load
(kN)
Debonded
length
London Clay
London/Lambeth
Clay
Lambeth Clay
Thanet Sand
/ mSD
-1.5 to -19.0
-19.0 to -28.0
-28.0 to -35.0
-35.0 to -44.5
-44.5 to -51.0
3,374
4
36
24
18
9
6,833
6
72
52
41
26
10,135
4
102
75
64
50
13,565
2
126
98
85
77
16,879
2
143
111
112
105
20,254
3
148
110
144
121
23,626
5
151
111
159
133
27,000
6
158
113
169
140
30,005
7
161
113
176
148
140
200
Values assumed in design
–
0
140
140
Table 7: Measured pile settlement
Total applied load
from strain gauge
at -1.5m (kN)
Shaft friction
measured over
debonded length (kN)
Load applied at
pile cut off level
(kN)
Measured
settlement at top
of pile (mm)
Settlement at
pile cut off level
(mm)
0
0
0
0
3,115
281
2,834
1.62
0.52
6,787
403
6,383
3.77
1.24
10,189
279
9,910
6.25
2.25
13,761
117
13,644
9.39
3.70
17,070
161
16,909
13.11
5.76
20,239
226
20,014
21.91
12.84
23,530
345
23,185
28.96
17.96
26,800
371
26,429
39.70
26.58
29,638
477
29,161
52.18
37.09
0
the measured results at -1.5mSD are
within 1% except for the first load
increment.
These results are plotted in Figure
8. Average values for the ultimate
shaft friction in the various strata
at each major load increment are
presented in Table 6. The bond
stresses in each stratum have been
compared with the design values
and are discussed as follows:
n The bond stress over the
debonded length was small, with
7kN/m2 maximum for the final load
increment.
n The maximum measured average
shaft friction in the London Clay
was 161kN/m2 which was slightly
above the 140kN/m2 limit applied
in the design and the delayed
concreting of the pile appears
to have had a minimal effect.
(Alternatively it could be considered
that the shaft friction in the London
Clay would have been higher if
the pile construction had not been
delayed.)
n The zone between 28 and 35m
depth was the lower section of
the London Clay and the upper
Lambeth Group and the peak shaft
friction of 113N/m2 was slightly less
than the 140kN/m2 limit applied in
the design (16% reduction). The
lower part of the London Clay often
contains sandier bands as noted on
the borehole logs and these zones
may have deteriorated during the
extended construction period. The
Upper Lambeth Group was referred
to in borehole BH1 as “silt” which
similarly could have reduced the
shaft friction.
n The shaft friction in the lower
Lambeth Clay appeared to be
increasing with each load increment
and it is not clear if the peak was
reached. The maximum recorded
value of 176kN/m2 was more than
the 140kN/m2 limit applied in the
design.
n The shaft friction in the Thanet
Sand appeared to be increasing with
each load increment and it was not
clear if the peak was been reached.
The maximum recorded value of
148kN/m2 was less than the 200kN/
m2 limit applied in the design.
The base capacity was not
mobilised in this test, with 10,788kN
recorded as the maximum load
applied at some 3m above the toe
level of the pile, equivalent to a
stress of 9,540kN/m2.
Performance of the Pile
The settlements at pile cut off
level were calculated based on
the average measured strain at
ground engineering january
2012
Load/ultimate load (%)
80
70
60
50
40
30
References
20
10
0
0
2
4
6
8
10
12
Settlement/diameter (%)
Pile shaft diameter (DS) = 1.2m
Pile base diameter (DB) = 1.2m
Deformation modulus below base (EB)
= 300,000kN/m2
Young's modulus of concrete (EC) = 50x106kN/m2
Friction length coefficient (KE) = 0.4
Upper pile length carrying no load (LO) = 0m
Pile length transferring load by friction (LF) = 34.3m
Flexibility factor (MS) = 0.00015
Pile design load (PT) = 16,680kN
Ultimate shaft friction load (US) = 18,000kN
Ultimate pile base load (UB) = 45,000kN
Figure 9 Fleming analysis of the pile test
-1.5mSD and -19.5mSD, over
the debonded length. The values
(Table 7) indicated that the pile
settlement at cut off level for the
design working load of 15,450kN
was approximately 6mm, which
includes the axial shortening of
the pile below pile cut off level.
At 1.5 x working load, 23,175kN,
the settlement at pile cut off level
was approximately 23mm, again
including axial shortening of the
pile below cut off level.
These values were less than the
specified maximum values of 12mm
and 36mm. It should be noted that
these values have been recorded for
a pile that was left open for two days
between boring and concreting.
Ultimate Capacity of the Pile
The pile test results were used to
carry out load/settlement analysis
ground engineering january
2012
for the pile to extrapolate to an
ultimate capacity at a settlement
equivalent to 10% of the pile base
diameter, namely 120mm total
settlement. The data was analysed
in accordance with the Fleming
Method (1992) (Figure 9). The
input parameters were manipulated
to best fit the data from the pile test
and a prediction for the ultimate
load was estimated from the
extrapolated load/settlement curve.
The base capacity can only be
estimated approximately given the
limited load applied to the pile in
comparison to the ultimate capacity.
For a pile deflection of 10%
of the pile diameter (120mm) the
ultimate loading on the pile was
estimated conservatively to be
43,690kN. Based on a SWL of
15,450kN this provided an overall
FOS of 2.8 for the ultimate capacity
1. Burland J B, Standing J R and
Jardine F M: Building response
to tunnelling – Case studies from
construction of the JLE – CIRIA
Special Publication 200; 2001
A and Yeow H C: Advances in
understanding of base grouted pile
performance in very dense sand.
Tunnel Construction and Piling,
London 8-10 Sept. 1999 pp57-69
2. ICE Specification for Piling
and Embedded Retaining Walls
(ICE 2007) published by Thomas
Telford, London
6. Guidance notes for the
design of straight shafted bored
piles in London Clay. London
District Surveyor’s Association
Publications No 1, 1999
3. Suckling T P and Eager D: Nonbase Grouted Piled Foundations
in Thanet Sand for a Project
in East India Dock, London
Underground
Construction
Symposium. London. 2001
4. Troughton V M and Platis A:
The effects of changes in effective
stress on base grouted pile in sand.
Proc. Int. Conf. Piling and Deep
Foundations, London 1989 pp445453
5. Chapman T J P, Connolly
M L, Nicholson D P, Raison C
of the pile. Alternatively, for a FOS
of 2.25, the SWL of the pile was
19,400kN. Whilst the base capacity
in particular is a very approximate
estimate, the figures suggest that the
base capacity is about 22,500kN/m2
at 120mm pile settlement, similar to
the 20,000kN/m2 limit.
Conclusions
The test confirmed an overall factor
of safety in excess of 2.25. The pile
performance was also satisfactory,
giving less than 12mm settlement at
pile cut off level at working load and
less than 36mm settlement at 1.5 x
working load. These values were
satisfied in spite of leaving the pile
open for two days between boring
and concreting.
The test measured a maximum
average shaft friction of 161kN/
m2 in London Clay. The lower 3m
7. Fleming W G K: A new method
for single pile settlement prediction
and analysis, Geotechnique 42,
No 3, Sept 1992 pp411-425
8. BS 8110 Structural Use of
Concrete - Part 1: Code of Practice
for Design and Construction
British Standards 1997
9. Nicholson D P, Chapman T J P,
Morrison P. Pressuremeter proves
its worth in London’s Docklands.
Ground Engineering, March 2002
pp32-34.
of London Clay and the upper 4m
of the Lambeth Beds recorded a
reduced shaft friction of 113kN/
m2. These values are considered to
represent peak values.
The maximum recorded shaft
frictions in the lower Lambeth
Group and the Thanet Sand
of 176kN/m2 and 148kN/m2
respectively are not considered to
represent the peak values because
the values continued to increase
with each load increment.
Similarly, the maximum base
capacity in Thanet Sand was not
mobilised, although extrapolation
of the data indicates very
approximately that the base capacity
is similar to the 20,000 kN/m2 limit.
29