Some observations of the effects of time on the capacity of piles driven in sand R. J. JARDINE, J. R. STANDING and F. C. CHOW†

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´
Jardine, R. J., Standing, J. R. & Chow, F. C. (2006). Geotechnique 56, No. 4, 227–244

Some observations of the effects of time on the capacity of piles driven
in sand
R . J. JA R D I N E * , J. R . S TA N D I N G * a n d F. C . C H OW †
´
Les piles enfoncees par battage dans le sable montrent de
´
remarquables augmentations de la capacite de leur arbre
axial dans les mois qui suivent l’installation. De nombreux
´
´
avantages pratiques en decoulent si les capacites de service
´
´
depassent les niveaux prouves dans les essais sur le site,
´
`
essais qui sont normalement effectues quelques jours apres
´
le battage. Cet expose rapporte les conclusions d’un programme d’essais de tension sur des piles en tuyaux d’acier

`
dans un sable dense a Dunkerque dans le nord de la France.
´
´
Ces essais ont montre un accroissement plus marque que
´
´
prevu de la capacite d’arbre au fil du temps. Les piles plus
´ ´
´
vieilles ont montre egalement des modes de defaillance
´
´
´
etonnamment fragiles ; des essais prealables avaient de´
´
´
grade la capacite et modifie le processus de vieillissement,
´
donnant des traces capacite-temps non monotones de l’ar´
bre, tombant bien en dessous de la caracteristique de
´
vieillissement intacte (IAC) definie par les essais sur des
´
´
´
piles neuves, n’ayant montre aucune defaillance anterieure.
´
´ ´
Les tendances capacite-temps derivees des essais multiples

´
peuvent donner des resultats trompeurs. Ces nouvelles con´
clusions permettent de caracteriser plus clairement les
´´
effets de vieillissement et permettent une reevaluation des
´
`
´
bases de donnees existantes relatives a divers types enfonces
dans toute une gamme de limons, de sables et de graviers.
Nous en tirons des conclusions pratiques importantes du
´
point de vue de la conception des piles et de l’interpretation
des essais de charge sur le terrain.

Piles driven in sand can show remarkable increases in
their axial shaft capacities in the months that follow
installation. Many practical benefits follow if service
capacities can be relied upon to exceed the levels proven
in site tests, which are usually performed within a few
days of driving. This paper reports findings from a
programme of mainly tension tests on steel pipe piles
performed in dense sand at Dunkirk, northern France.
The tests demonstrated more marked shaft capacity
growth with time than expected. The aged piles also
showed surprisingly brittle failure modes; prior testing to
failure both degraded capacity and modified the ageing
processes, leading to non-monotonic shaft capacity–time
traces that fall far below the intact ageing characteristic
(IAC) defined by tests on fresh, previously unfailed, piles.

Capacity–time trends inferred from repeatedly tested
piles can give misleading results. The new findings allow
the ageing effects to be characterised more clearly, and
permit a re-evaluation of existing databases involving
piles of various types driven in a range of silts, sands and
gravels. Important practical conclusions are drawn regarding pile design and the interpretation of field pile
load tests.

KEYWORDS: creep; piles; sands; time dependence

INTRODUCTION
As part of the research described by Chow (1995, 1997), in
1994 a team from Imperial College, London, performed
tension retests on open-ended, 324 mm diameter, steel pipe
piles driven at a sand research site near the Gravelines
Power Station complex at Dunkirk, north-west France. Surprisingly strong effects of time were found, which were
reported by Chow et al. (1997) along with a database of
comparable measurements made elsewhere, and a tentative
exploration of the possible root causes. This paper reports a
subsequent systematic investigation of the same topic involving a suite of new piles driven at the same test site and a
reappraisal of existing pile test databases.

Chow had been driven in 1988 by the French CLAROM
research group (Brucy et al., 1991). Fig. 1 shows the
location of these tests and the other instrumented pile
experiments conducted by Chow in 1994.
The key results from the 1988–1994 studies regarding
ageing are illustrated in Fig. 2 through two sets of tests to
failure performed on the 11 m long (strain-gauged) CS pile.
Pile CS was subjected to a first restrike five months after

installation. One tension test to failure, followed by a compression test, was performed six months after driving, and a
second set of similar tests was conducted three months later
after a second restrike following partial removal of the sand
plug inside the pile. The CLAROM group found only minor
differences between the pairs of tests performed in 1989,
giving no indication of any strong effect of pile age. However, the tension capacity was 85% higher when Chow
(1997) retested the same pile, and others, five years later.
The 1994 retest results prompted Chow (1997) to research
and assemble a database from the literature and practitioners’ files of tests on ‘aged’ steel, concrete and timber driven
piles, drawing on the sources detailed in Appendix 1, which
comprised a mix of first-time tests on ‘fresh piles’, restrikes
and static retests of previously failed piles. The broadly
scattered dataset of mixed compression and tension tests is
presented in Fig. 3, covering the total compressive capacities
Qt and the shaft capacities Qs (where these could be
isolated) in Figs 3(a) and 3(b) respectively. The capacities
developed at times t are divided by Qt (EOID) or Qs (EOID),
the values assessed at the end of initial driving (EOID) by

Summary of earlier pile ageing tests at Dunkirk and their
interpretation
The Dunkirk experimental area was provided by the Port
Autonome de Dunkerque and is located close to the Port
Ouest Industrial Zone (ZIP), about 200 m south of the
Institut Pasteur laboratories. The open-ended piles tested by
Manuscript received 14 December 2004; revised manuscript
accepted 25 January 2006.
Discussion on this paper closes on 1 November 2006, for further
details see p ii.
* Imperial College, London, UK.

† WorleyParsons Pty Ltd, Australia; formerly Imperial College
London.

227


JARDINE, STANDING AND CHOW

228

N
Avant Port
Ouest

The Channel

TEST SITE

Centre
Aquacole

Bassin de
l’Atlantique

Gravelines
Power Station

Grand
Fort-Philippe
0

5 km

c
Tra

k
P1

BH1

N
line
Reference

T post
to Borne H

To look-out
tower

60 m

40 m

LS

CL
P2

CLAROM tests
CL 1 CS 5 11.7 m long piles
LL 1 LS 5 22.4 m long piles
P1 1 P2 5 cone penetration
P1 1 P2 5 tests
BH1 5 borehole

LL
CS

re

o
rdc

ck

tra

DPH2
DPH1

Ha

DPM15/1

DK1
DK2

IC container

DK2b

DK3

Shed

CPT1/SC1
DMT1

IC pile tests

DMT2
CPT2/SC2
0

10 m

BRE In situ tests
CPT 5 cone penetration test
SC 5 seismic cone test
DMT 5 dilatometer test
DPH ỹ
ý 5 dynamic penetration tests
DPM
ỵ

Fig. 1. Site plan showing position of earlier pile test locations (taken from IFP Plan No.
FM89.04.04.13) (Chow, 1997)

EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND
1600

Pile CS

T¢89a & C¢89a, soil plug intact
T¢89b & C¢89b, soil plug cored out

Total load: kN

1200

T¢94, five years later

C¢89a

800
Compression

Base load
C¢89b

C¢89b

C¢89a

400
0
260

2400

240

220

T¢89a
T¢89b

20
40
Cumulative pile head
displacement: mm

60

Tension

T¢94

2800

0

Faster rate of displacement

Fig. 2. Load–displacement curves for 11 m CS pile from
CLAROM research programme (Chow, 1997)
5·0
4·5

Qt(t)/Qt (t 5 1)

4·0
3·5
3·0
2·5

Samson & Authier (1986)
Seidel et al. (1988)
Astedt et al. (1992)
Dunkirk CS
Dunkirk CL
Dunkirk LS
Skov & Denver (1988)
Tavenas & Audy (1972)
York et al. (1994)
Svinkin et al. (1994)
Bullock & Schmertmann (1995)
Holm (1992)
Tomlinson (1996)

2·0
1·5
1·0
0·5
0
0·1

1

10
100
Time after driving: days
(a)

1000

10 000

5·0
4·5

QS(t)/QS (t 5 1)

4·0
3·5

Samson & Authier (1986)
Seidel et al. (1988)
Astedt et al. (1992)
Dunkirk CL
Dunkirk LS
Dunkirk CS
Tomlinson (1996)

Jardine & Chow (1996)
trendline

3·0

2·5
2·0
1·5
1·0

Range of data from
Bullock et al. (2005a)

0·5
0
0·1

1

10
100
Time after driving: days
(b)

1000

10 000

229

ultimately reach equilibrium capacities after long durations:
the interpretation of EOID capacities is open to considerable
potential uncertainty.
Chow et al. (1997) considered possible explanations for
the time-dependent processes. Some, including corrosion,

were discounted as probable major causes (because the
process is common to steel, timber and concrete driven piles
and develops most rapidly below the active corrosion zone).
The dominant process was thought to be gains in the radial
effective stresses acting on the pile shafts resulting from the
relaxation, through creep, of circumferential arching established around the pile shafts during installation.1 Increases
in sand shear strength and stiffness with time (ageing)
involving the reorientation of sand grains and possible cementing or micro-interlocking processes were also potential
contributing factors resulting in enhanced interface dilation
and possibly interface friction angles. The relative contributions of these causes were uncertain, but the creep-induced
stress redistribution process was thought by Chow et al. to
have the greatest influence.
Research by Axelsson (2000) has broadly reinforced these
preliminary conclusions, reporting large rises with time in
the horizontal effective stresses measured (in the field) on
the sides of a 235 mm square concrete pile. Axelsson also
observed significant gains in the positive effective stress
changes developed through restrained dilation during successive load tests, and noted that similar gains applied to
smaller rods driven into sand. Bullock et al. (2005b) argue
that the time-related increases are due either to a more
dilatant response to loading, or to gains in interface friction
angle ä. Bowman (2002) investigated the microscopic and
macroscopic behaviour of dense sands under high stress
ratios with a view to understanding the mechanisms of pile
set-up. She suggests an alternate hypothesis for pile set-up
based on dilatant creep and ageing, arguing that the creep
volume changes (initiated by the intense shearing imposed
during installation) are initially contractant as the soil grains
rearrange themselves to redistribute stresses. However, Bowman argues that the creep straining gradually changes to
become dilatant, both microscopically and macroscopically.

As the kinematic restraint provided by the pile would inhibit
expansion of the soil, any such dilation would lead to
increased radial stresses.
The potential roles that sand state and type might play in
any of the above processes are open to speculation. Stiffness
and the ability to sustain arching are likely to increase with
in situ density, as would the level of prestressing imposed by
pile driving. However, creep can also be expected to become
more marked in elements subjected to high stress levels, and
in sediments having high initial void ratios or angular
particles. Equally, corrosion processes would depend on the
specific chemical interactions between the pile material, sand
mineralogy and groundwater ionic concentrations.

Fig. 3. Original database of pile capacity against time in terms
of: (a) total pile capacity; (b) shaft resistance alone

pile-driving analysis or other means. Bullock et al. (2005a)
have recently used a similar format to report multiple retests
on 457 mm wide square concrete piles driven at five sites in
Florida, two of which consist mainly of (possibly calcareous)
sands as described in Appendix 1. Their wide range of ‘side
shear set-up’ results is also indicated in Fig. 3(b). Base
capacity was interpreted as remaining relatively unaffected
by time in the above studies.
In the same way as Skov & Denver (1988) and others,
Jardine & Chow (1996) drew a tentative semi-logarithmic
trendline through the shaft resistance scatter, but it was not
clear whether the ageing process was continuous, or if its
start was delayed over the first few days, or if piles would

Application to other test programmes
The 1994 Dunkirk test results were communicated to the
EURIPIDES group, which was researching the behaviour of
large steel (760 mm diameter, up to 46.8 m penetration) pipe
piles driven in very dense sand at Eemshaven in Holland
(Zuidberg & Vergobbi, 1996; CUR, 2001; Fugro, 2004). The
information was passed through the same route to the team
responsible for tests on steel pipe piles 762 mm in diameter
and up to 75 m penetration, which were being driven in
1 ˚
Astedt et al. (1992) included in their list of potential causes a
related concept of radial variations in density combined with
unstable circumferential ‘vaulting’, without referring to creep.


JARDINE, STANDING AND CHOW

230

looser mica sand in order to check the design of the still
larger main Jamuna Bridge piles in Bangladesh (Tomlinson,
pers. comm., 1996; CUR, 2001; Fugro, 2004). Retests
performed on these piles after rest periods of around
six months at both the EURIPIDES and Jamuna Bridge sites
showed increases in capacity of between 70% and 90% over
six months, confirming the trends seen in Fig. 3. Other
experiments conducted on concrete-driven piles showed even
more marked gains in capacity with time.

sea wall that had been formed by hydraulically placed
dredged marine sand. The hydraulic fill, which was placed
between 1972 and 1975, is around 3 m thick and is underlain by Flandrian marine sand that was deposited between
2100 and 900 bp. The Eocene Ypresienne Clay found
´
beneath the sand is part of a major stratum that extends
beneath the North Sea, known as London Clay in the UK.
In the late 1980s the CLAROM group commissioned two
cone penetration tests (CPTs), a 26 m deep sampled borehole, and a programme of laboratory testing. The work
conducted by the Imperial College group in the early 1990s
(summarised by Chow, 1997) included mineralogical and
index tests and direct and interface shear, triaxial stress path
and resonant column torsional shear experiments (Connolly,
1996; Kuwano, 1999). Additional fieldwork was performed
by the Building Research Establishment (BRE), including
CPT, seismic cone, Marchetti dilatometer and Rayleigh wave
testing.
Figure 4 shows the typical site profile for the CLAROM/
Imperial College test area, and Fig. 5 shows the range of
particle size distributions from the CLAROM borehole. The
grains are sub-rounded to rounded and are composed (on
average) of 84% quartz, 8% albite and microcline, and 8%
coarse shell fragments (CaCO3 ). The relative density fluctuates with depth, and Chow (1997) interpreted values close to
100% over the first 2 m of depth, tending to an average of
around 75% at greater depth. A much looser organic layer
was located at around 7–8 m depth. Direct shear and triaxial
compression testing indicated soil–soil peak ö9 values of
35–408 and critical state values of around 328, and interface
shear tests against steel surfaces similar to those of the
driven piles showed critical state ä9 values of around 278.

Other experiments gave details of the sand’s anisotropy,
very-small-strain elastic parameters, non-linear stiffness
characteristics and creep behaviour (Kuwano, 1999; Jardine
& Standing, 2000; Kuwano & Jardine, 2002; Jardine et al.,
2004).
An advance series of new CPT tests was performed in
1998 at the intended GOPAL piling position. However, the
pile driving had to be relocated at a late stage, and seven

TEST PROGRAMMES AT DUNKIRK 1998–1999
GOPAL project and test site
The EU-funded GOPAL project2 described by Parker et
al. (1999) provided an opportunity to investigate the timerelated field behaviour of piles driven in sand systematically.
The Port Autonome de Dunkerque made the same test site
(shown in Fig. 1) available to the GOPAL team. Multiple
experiments on fresh and pretested plain driven piles were
incorporated into the research programme, the primary purpose of which was to study the behaviour of piles enhanced
by jet-grouting. The aim of the tests on the plain driven
piles listed in Table 1 was to investigate
(a) the rate of capacity gain in ‘fresh’ piles
(b) whether the ageing processes are subject to any initial
delay
(c) whether the gains stabilise over a medium-term timescale
(d) whether pretesting to failure affects the time–capacity
relationships
(e) how cyclic loading might affect the ageing trends.
The GOPAL pile tests were conducted within 70 m of the
earlier Dunkirk instrumented pile tests, in essentially similar
ground conditions, and so contributed to a common dataset.
Ground conditions

The test sites employed by the CLAROM, Imperial College and GOPAL groups lie in a flat area behind the main
Table 1. Details of first-time tests on driven pipe piles at Dunkirk
Pile

Outside diameter (o.d.)
and penetration

Wall thicknesses

Date of
driving

First test type
(age after driving in days)

Maximum load and
displacement

R1

457 mm o.d.
19.31 m
457 mm o.d.
18.85 m
457 mm o.d.
19.24 m
457 mm o.d.
19.37 m
457 mm o.d.
19.05 m

457 mm o.d.
18.90 m
457 mm o.d.
10.02 m

20 mm over top 2.5 m,
13.5 mm to base
As above

24/08/98

As above

20/08/98

As above

24/08/98

As above

25/08/98

As above

21/08/98

As above

25/08/98

Tension to failure
(9 days)
Tension to failure
(235 days)
Unfailed tension proof
load test (85 days)
Unfailed tension, proof
load test (74 days)
Unfailed tension, proof
load test (86 days)
Tension to failure
(80 days)
Compression to
failure (68 days)

1450 kN
at 24mm
3210 kN
at 34 mm
2000 kN
at 10 mm
2000 kN
at 9 mm
2000 kN
at 9 mm
2400 kN
at 30 mm
2800 kN
at 33 mm

R2
R3
R4
R5
R6
C1

21/08/98

Pile-head loads shown have not been adjusted for pile and plug weights.

2
The EU-funded GOPAL project involved Bachy Soletanche
(France), D’Appolonia (Italy) and Imperial College London (UK).
It investigated jet-grouting as a means of enhancing the base
capacity of piles driven in sand. The results have been detailed in
reports prepared by D’Appolonia for the EU; they are also
summarised by Parker et al. (1999) and Jardine et al. (2001b).


EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND
CPT qc: MPa
0
10
20

Borehole log
0
2

4

CPT fc: kPa
0 100 200

40

300

400

Very dense, light brown, uniform, fine to medium,
subrounded SAND with occasional shell fragments
(Hydraulic fill)
GWL

6

Dense with shell fragments
(Flandrian Sand)

8

Organic layer

10
Depth: m

30

231

12

Dense, green-brown and grey-brown, uniform
fine to medium, subrounded SAND with some
shell fragments
(Flandrian Sand)

14
16
18
20

Becoming very dense

22
24

Fig. 4. Typical site profile for CLAROM/Imperial College test site (Chow, 1997)
Silt

Percentage finer by weight

100

Fine

Sand
Medium

Gravel

N

Coarse

90
80
70
60
50
40

C2: 0·80–1·12 m
C6: 3·51–3·73 m
C8: 5·19–5·55 m
C9: 6·20–6·35 m
C15: 11·5–11·9 m
CLAROM envelope of results

30
20
10
0
0·01

0·1

Particle size: mm

1

R3

10
A

JP1

Fig. 5. Range of particle size distribution curves from CLAROM borehole (Chow, 1997)

PS.R23

further CPT tests were performed within the eventual
GOPAL pile test area, some months after pile driving and
jet-grouting. As indicated in Fig. 6, the soundings were
typically conducted 5–6 m (around 12 pile diameters) away
from the six reaction piles (R1 to R6) described below.
Three further soundings were made 1.5 m from the axes of
the two compression test piles C1 and JP1. Summary logs
are given in Fig. 7, which also shows the synthesis made by
Parker (2000) of the local stratification. More detailed
traces3 are given in Fig. 16 in Appendix 2. As with Chow’s
profile (Fig. 4) for the nearby CLAROM site, the qc traces
fluctuate with depth and typically fall between 10 and
35 MPa. A significant band of sand containing organic
matter was identified at 7–8 m in Chow’s profile, which
divided the sand into upper and lower units. Although this

Comparing the qc traces obtained around 1.5 m from the
compression piles’ axes with those from more remote locations
indicates that pile installation can have influenced the C1 and JP1
traces only marginally, and is unlikely to have affected the other
CPT measurements significantly.

3

R6

B
PS.R12

R1

2

4

A

R5

R2

Institut
Pasteur

0

PS.GP1A
PS.R56
PS.GP1B

PS.C1
C1

B
PS.R45

R4

8

12 m

Fig. 6. Plan showing layout of test and reaction piles and CPTs
from GOPAL project, Dunkirk (sections A–A and B–B relate
to the CPT profiles shown in Fig. 7)


JARDINE, STANDING AND CHOW

232
Section A–A
GP1.B
qc: MPa
20 40

R2-3

qc: MPa
20 40

R5-6
qc: MPa
20 40

z (m)

5

10

15

R1-2
qc: MPa
20 40

Section B–B
C1
qc: MPa
20 40

R4-5
qc: MPa
20 40

10

z (m)

5

15

Fig. 6, to provide the necessary reaction capacity, giving a
minimum spacing (s) to diameter (D) ratio % 15. The piles
were driven by simple drop weight hammering. The ram and
helmet weighed 4.7 t, and these fell through a variable (but
recorded) height of 1–4 m. Fig. 8 shows the site blow count
records in which the blows/m penetration rate has been
multiplied by the drop height, giving an equivalent number
of blows normalised to a 1 m drop. Piles C1, R1, R4 and R5
required the hardest driving, whereas JP1, R3 and R6 drove
more easily, with their driving performance correlating well
with the local CPT qc profiles. Measurements made inside
each pile indicated partial plugging, with the internal soil
columns rising to 7.65–10.0 m below ground level and
occupying about 60% of the pile length; similar core heights
had been recorded in the 22 m long CLAROM piles. The
effective weight of the (19 m long, 457 mm OD) GOPAL
steel piles was about 35 kN with their cores, whereas that of
the (22 m long, 324 mm OD) CLAROM piles was about
20 kN. The equivalent weights of the 10 m long GOPAL
compression piles C1 and JP1 were around 18 kN.
One of the two 10 m long compression piles, JP1, was
modified, after all driving had been completed, by forming a
deep jet-grout cylinder, 3.5 m in diameter, 5 m beneath its
base. The primary GOPAL research aim was to investigate

the behaviour of the jet-grouted pile after a suitable curing
period, and to contrast this with that of the plain driven
‘control’ pile C1. The GOPAL testing did not involve
loading any of the reaction piles (termed R1 to R6) to
beyond about 1350 kN, less than 60% of the tension capacities expected at that age. This allowed the six piles to be
considered as un-failed, or ‘fresh’, and to be used for other
experiments. A programme of first-time tension static and
cyclic loading tests was conducted over eight months that
involved multiple tests on both ‘fresh’ and pre-failed piles.
Additional data were provided by one incomplete static
0

2
R1
R2
R3
R4
R5
R6

Hydraulic fill
Flandrian Sands

4

Organic layer
Interbedded

6

Horizontal scale
1

2

3

4

5m

Fig. 7. CPT profiles with depth and interpreted soil profile
(see Fig. 6 for locations of sections A–A and B–B)

8

Penetration: m

0

10

12

organic layer was less well developed in the GOPAL area,
two comparable bands (with lower qc and higher friction
ratio, fs /qc ) were found between 14 and 20 m depth: these
are more extensive in the region of the jet-grouted pile
(section A–A) than around the control pile (section B–B).
The CPT programme allowed these variations in ground

conditions to be accounted for in the test interpretation.

14

16

18

Testing programme and procedures
The GOPAL project involved primary compression tests
on two steel-driven pipe piles 10 m long, 457 mm outside
diameter, and with 13–20 mm wall thickness (C1 and JP1).
Six piles with the same section were driven to penetrations
of 18.9–19.4 m during August 1998, in the grid shown in

20

0

100
200
300
Normalised blow count: blows/m

400

Fig. 8. Normalised blow count against average penetration
depth for reaction pile driving

EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND

233

around ten months of their installation. The make-up and test
histories of the CLAROM piles are detailed in Table 2.

tension test and two ‘rapid’ pullout tests conducted by Chow
in 1994 on 22 m long steel pipe piles (LL and LS) that had
been driven five to six years before as part of the 1988
CLAROM research programme (Chow, 1997). Both piles
had been restruck within six months of their installation.
Despite their five-year ‘recovery’ period these piles could
not be considered as being truly ‘fresh’. Use was also made
of tests conducted on the two shorter CLAROM piles (CS
and CL) that had been subjected to more intensive pretesting
in 1989. Both CS and CL had sustained restrikes and two
sets of tests to failure in tension and compression within

Pile testing procedures
The GOPAL field testing was conducted in association
with Precision Monitoring and Control (PMC) Ltd, which
provided the equipment and key site staff. A slow, loadcontrolled procedure was specified that used an automated
hydraulic system and the beam arrangements illustrated in
Figs 9(a) and (b); the compression, tension and two-way

Table 2. Details of tests on CLAROM driven pipe piles at Dunkirk
Pile

Outside diameter (o.d.)

and penetration

Wall thickness

Date of
driving

First test type
(age after initial driving)

Maximum shaft load
and displacement

CS

324 mm
11.1 m
11.3 m after restriking

12.7 mm, 19.1 mm over
0.64 m toe length

15/12/88

End of initial driving
(EOID)
First restrike
(160 days)
Tension test
(188 days)

Compression test (189 days)
Soil plug removal
Second restrike
(259 days)
Tension test
(272 days)
Compression test
(273 days)
Tension test
(1991 days)
Pile extraction

233 kN

24/05/89
21/06/89
22/06/89
30/08/89
31/08/89

11.6 m after restriking

13/09/89
14/09/89
29/05/94
27/08/94
CL

324 mm o.d.
11.1 m

11.3 m after restriking

12.7 mm over entire
length

13/12/88
23/05/89
07/06/89
08/06/89
30/08/89
01/09/89

11.6 m after restriking

27/09/89
28/09/89
31/08/94
LS

324 mm o.d.
22.0 m
22.0 m after restriking

12.7 mm, 19.1 mm over
0.64 m toe length

16/12/88
24/05/89
30/08/89
31/08/89

22.1 m after restriking

27/05/94
LL

324 mm o.d.
22.0 m
22.0 m after restriking
22.1 m after restriking

12.7 mm over entire
length

30/08/94
12–14/12/88
23/05/89
30/08/89
1/09/89
1/09/94

End of initial driving
(EOID)
First restrike
(161 days)
Tension test
(176 days)
Compression test
(177 days)
Soil plug removal

Second restrike
(262 days)
Tension test
(288 days)
Compression test
(289 days)
Pile extraction
(2087 days)
End of initial driving
(EOID)
First restrike
(159 days)
Soil plug removal
Second restrike
(258 days)
Tension test*
(1988 days)
Pile extraction
End of initial driving
(EOID)
First restrike
Soil plug removal
Second restrike
Pile extraction
(2089 days)

630 kN assuming
Qb ¼ Qb (EOID)
395 kN at 12 mm
623 kN at 90 mm

No Qs measurement
435 kN at 12 mm
535 kN at 82 mm
750 kN at 9.5 mm
Not measured,
assumed to be
same as CL
No information
No information
458 kN at 20 mm
692 kN at 68 mm
No information
548 kN at 20 mm
676 kN at 71 mm
810 kN
1280 kN
2560 kN assuming
Qb ¼ Qb (EOID)
No Qs measurement
3200 kN at 25 mm
No information
No information
No information
3100 kN

Pile-head loads shown have not been adjusted for pile and plug weights.
*First test on CLAROM LS involved loading to 2500 kN without failure. Its capacity was projected on the basis of observed creep rate
characteristics and on the subsequent rapid pile extraction loads, which indicated a required head load .3100 kN at 2115 days’ age.

JARDINE, STANDING AND CHOW

234

PMC
tension
head
914 3 419
5·3 m long
I-beam

400 mm

500 t
hydraulic
jack and
load cell

400 mm

914 mm

914 mm

6 m by 1 m
ground
panel
Levelled
sand

1400 mm
Displacement
and reference
system

200 mm
500 mm

6m
(a)

Tension
head

Loadspreading
beams
914 3 419
5·3 m long
I-beam

6m

Displacement
and
reference
system

6m
(b)

Fig. 9. Details of rig used for testing reaction piles in tension (not to scale): (a) elevation; (b) plan

cycling experiments required different configurations. The
6 m long reaction beam transferred compressive loads to the
ground through steel frames that sat on steel spreader mats
located at least 2 m from the pile. The pile-head movements
were measured by four independent transducers attached to

long reference beams that were supported at points at least
2 m away from both the piles and the reaction pads.
The tension loads were applied in increments separated by
creep pause periods. The load increments diminished and
the pause periods extended as failure approached. The initial


EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND
load increments were each 100 kN, applied at a steady rate
over about 100 s, and the initial pause periods were each
30 min long. Creep rates were monitored continuously, and
generally reduced with time during each pause period. The
load increments were reduced to 50 kN once the 30 min
creep rates exceeded 0.01 mm/min, then to 20 kN, and
finally to 10 kN if the 30 min creep rates exceeded 0.08 mm/
min. From this point the pause periods were extended to
60 min, and failure was considered imminent. In some cases
the procedures were modified slightly for operational reasons. Overall, most tests to failure involved 15–30 load
increments and took 10–20 h. The time-dependent movements accumulated during pause periods grew from being
practically negligible at loads below a ‘threshold load’ of
600–1000 kN to being dominant in the final test stages—by
which stage most of the observed movements were accumulating as creep during the extended pause periods. Failure

was defined in terms of critical creep rates (0.5 mm/min over
the first 10 min or 0.1 mm/min after 30 min). Some piles
demonstrated a runaway failure mode in tension, with capacity reducing after developing a peak value. The pile-head
displacements required to reach peak tension capacity are all
less than 34 mm, or 7% of the pile diameter. The compression test on pile C1 had to terminate while the load was still
climbing, and before settlement reached 10% of the pile
diameter (45.6 mm). A load of 2849 kN was projected for
C1 at the intended 10% diameter settlement limit. The
instrumentation installed on the piles did not allow the base
and shaft components to be separated reliably, and the
distribution had to be interpreted by other means.
Additional fast loading tests were performed at the end of
cyclic loading tests on some reaction piles: these involved
applying a constant rate of loading so that failure developed
within a period of several minutes, rather than hours. Regarding the 22 m long CLAROM piles, a single load-controlled
test was performed on pile LS in May 1994. Although the
maximum load that could be applied with the site equipment
(2500 kN) was insufficient to fail the pile, a locally calibrated

235

projection based on the piles’ creep characteristics indicated a
static capacity of $3200 kN. The CLAROM piles were then
extracted by hydraulic jacking, with relatively low-resolution
pressure gauge measurements being made of the pile-head
loads. Check tests involving the shorter CLAROM piles
indicated that the ‘pullout’ capacities measured in this way
are $8% higher than the equivalent static capacities. Taken
together, these additional data indicate static tension capacities of around 3150 kN for the 22 m CLAROM piles at
around 2000–2100 days; Chow (1997) gives further details.

Test sequences and main results
Table 1 gives the key features of the four ‘first-time’ static
tests to failure on piles R1 to R6, C1, and Table 2 gives
details of the tests on the four CLAROM piles. Note that
the tabulated loads were as-measured at the pile heads,
without any modification for self-weight. The first tests on
R3, R4 and R5 involved tension proof-loading to 2000 kN,
well beyond their previous maximum loads but at least 15%
below the tension capacities proven by testing R6 at a
similar age. The proof-tested piles developed maximum
displacements of 8–10 mm, but without large creep components and with most (60–70%) of the pile-head movement
recovering on unloading. These unfailed piles were subsequently used for cyclic loading experiments and tension
tests, as were R1, R2 and R6. The programme of reaction
pile ‘proof’ loading and CPT testing pattern allowed an
assessment to be made of how (a) variations in soil profile
and (b) site activities including the jet-grouting might have
affected the reaction piles’ behaviour.
Table 3 gives details of static retests that were performed
on all of the piles listed in Table 1. The associated notes
indicate which piles were subjected to intermediate stages of
significant cyclic loading, and where relatively rapid-capacity
check tests were performed in place of the standard slow
maintained load procedures. The tabulated pile-head loads
do not account for pile self-weight.

Table 3. Retests on Dunkirk reaction piles and compression pile C1
Pile

Test 2
(days after driving)

R1

Tension failure
1500 kN
(57 days)
at 8 mm
1700 kN
Cyclic failure, then
at 8 mm
rapid tension
failure (236 days)
Cyclic failure, then 1650 kN (ultimate)
at 30 mm
tension failure
(87 days)
Cyclic failure, then 1650 kN (average)
at 15 mm
rapid tension
failure (82 days)

R2
R3
R4

R5
R6
C1

Cyclic failure, then

rapid tension
failure (88 days)
Cyclic failure, then
tension failure
(82 days)
Tension failure
(69 days)

Maximum static
load and
displacement

1300 kN (average)
at 10 mm
Static 1585 kN
at 7 mm
821 kN
at 33 mm

Test 3
(days after driving)

Maximum static
load and
displacement

Tension failure
(239 days)
N/A

1646 kN
at 8 mm
N/A

Tension failure
(250 days)

Test 4
Maximum load
(days after driving) and displacement

N/A

N/A

N/A

N/A

2000 kN
at 10 mm

N/A

N/A

Low-level cyclic test,
then rapid tension to
failure
(236 days)

Tension failure
(226 days)

2491 kN
at 12 mm

N/A

N/A

1794 kN
at 9 mm

N/A

N/A

Cyclic failure, then
rapid tension failure
(83 days)
Two-way cyclic
failure, then tension
failure (72 days)

1300 kN (average)
at 10 mm

Cyclic failure, then
rapid tension
failure (238 days)

N/A

1300 kN
at 15 mm

450 kN
at 20 mm

N/A

Pile-head loads have not been adjusted for pile and plug weights; details of the cyclic tests are given by Jardine & Standing (2000).
Note: the initial ‘virgin’ tension capacities of piles R3, R4 and R5 at 80 to 88 days’ age assumed to be around 1.9 times ICP capacity, as
indicated by the test on R6.


236

JARDINE, STANDING AND CHOW

INTERPRETATION
Factors that were addressed in the interpretation of the pile
load tests include the potential variations in ground conditions
and stress states between the test piles, the crucial differences
between first-time tests and retests, and the relationship between the new information and Chow’s original database.
Potential effects of variations in soil properties, pile details
and site operations
As noted earlier, piles C1, R4 and R5 experienced harder
driving than R3, R6 and JP1, which were located in the
opposite corner (see Fig. 8), in keeping with the slightly
different layering indicated by the CPT testing (see Fig. 7).

The CPT soundings, which were made after driving, grouting and curing had been completed, proved lower qc values
around the jet-grouted pile JP1 than were found around the
plain driven pile C1. The site operations may have added to
the effects of spatial variations, especially at locations PS
C1 and PS GP1B. Whereas plain driving is likely to elevate
the local qc values, jet-grouting creates a column of soil–
cement slurry. This soft inclusion, along with shrinkage
caused by water bleeding and grout de-airing, allows local
stress relief that may reduce local qc values.
One way of gauging the overall effects of local variations
in ground conditions on capacity is to apply the CPT-based
capacity calculation procedures set out by Jardine & Chow
(1996) and Jardine et al. (2005) for ‘short-term’ shaft
resistance, adopting the nearest CPT profiles (see Fig. 6).
These capacity estimates are termed the ICP, or Imperial
College Pile, predictions. When checked against databases
of field load tests the ICP approach gives far better predictions than conventional methods, and Chow (1997) showed
that it applied well to her tests at Dunkirk. The main
features of the shaft capacity approach are as follows.
(a) Shaft capacity is found by integrating shear stress over
the embedded area:
ð
(1)
Qs ¼ ðD ôf dz
(b) Under tension loading,4 the local maximum shaft shear
stress expected at any given depth on the shaft, and
height h above the pile tip, is
9
9
(2)

ụf ẳ 0:90:8ú rc ỵ ˜ó rd Þ tan äf
(c) The pre-loading radial effective stress is
 0:13  À0:38
ó v0
9
h
ó rc ¼ 0:029qc
9
Pa
RÃ

(3)

(d) ó v0 is the free-field vertical effective stress, and Pa , the
9
atmospheric pressure, is introduced to make the term
Ã
non-dimensional; R is found from the pile’s inner and
outer radii as
À
Á0:5
Ã
(4)
R ¼ R2 À R2
outer
inner
(e) The dilatant component of radial effective stress change
is ˜ó rd ¼ 2Gd r =Router , where ˜ r is the pile peak to
9
trough surface roughness, and the operational secant

shear stiffness G may be estimated from the CPT
profile, as recommended by Jardine & Chow (1996)
and Jardine et al. (2005).
Table 4 summarises the ICP predictions, taking account of
the actual pile lengths as detailed in Table 1. The capacities
4

The equivalent expression applying to compressive loading is
ụf ẳ (ú rc ỵ ú r )tan äf .
9
9

calculated for the 19 m long, 457 mm o.d. plain piles are
highest at locations where driving was relatively hard (C1,
R4, R1) and are lowest where driving was easier (R6, JP1
and R3). The potential impact of local CPT variations
amounts to approximately Ỉ15%, and the good correlation
with driving records suggests that the variation was primarily
natural rather than induced by installation activities. The
tabulated values indicate that the potential effects of driving
and grouting on the CPTs taken within 2 m of C1’s and
JP1’s axes are not resolvable against the background of the
site’s natural soil variability. The availability of driving
records and CPT profiles located near to each pile reduces
the scope for variability to cause errors in test interpretation.
The four field tests made on ‘fresh’ piles 81 to 90 days
after driving allow another way of checking potential variations between the reaction piles. Fig. 10 presents the plots
for the test to failure on R6 along with the proof load tests
on R3, R4 and R5 (see Table 1). Piles R6 and R3 developed
more creep and their pile-head movements were around 20%

greater at the 2000 kN stage than for the other two piles,
supporting the hierarchy of calculated capacities given in
Table 4 and the driving records. Such local variations are
allowed for in the assessment that follows by normalising all
the measured capacities by the ICP ‘short-term’ predictions.
ICP tension capacities of 481–503 kN and 1251 kN were
assessed for the short (CS and CL) and long (LS and LL)
CLAROM piles, depending on the relevant pile-tip depth
and based on a single CPT test (CPT1) performed by BRE
around 25 m from the test piles (see Fig. 1). Naturally, local
variations in CPT resistance could affect the individual piles.
Tests on ‘fresh’ piles
Figure 11 presents the overall load–displacement curves
developed during first-time tension tests conducted to failure
on piles R1, R6 and R2 after 9, 81 and 235 days respectively. The three tests are hard to distinguish up to 1000 kN,
after which the curves spread—each following an approximately smooth trend until reaching its particular limiting
tension capacity at $30 mm displacement. The main feature
is the overwhelming effect of pile age on (tension) shaft
capacity, and this is the main focus of the paper. However, it
is useful to consider whether the pre-failure load–displacement behaviour is unusual in any respect. Detailed finiteelement analyses have been undertaken for selected tests
with the code ICFEP, in which the sands’ non-linear and
pressure-dependent stiffness characteristics were derived
from detailed laboratory studies and input through appropriate ‘small-strain’ formulations into generalised Mohr–
Coulomb models of the sand mass that accounted for (a) the
effective stress regime expected around the piles and (b) the
sand-interface shear behaviour seen in laboratory tests
(Jardine & Kovacevic, 2000; Jardine et al., 2004). Best
estimates for the operational soil parameters led to good
matches for the ultimate capacities and the first two thirds
of the loading curves, provided account was taken of the

probable effects of age on the radial effective stresses acting
near the pile shaft. This finding suggests that the ground’s
stiffness response was neither unusual, nor greatly affected
at working load levels by ageing. However, the measured
pile-head movements exceeded the predicted values as failure was approached; as noted earlier, (unmodelled) rate
processes, including creep, had an important influence at
high loading levels.
Shaft capacity–time trend for fresh piles
Returning to the question of shaft capacity, the peak
tension loads are corrected for self-weight and plotted as


EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND

237

Table 4. Short-term capacities calculated for GOPAL piles following ‘ICP’ procedures
Pile

CPT profile used for calculation

Calculated ICP capacity: kN

R1
R2
R3
R4
R5
R6
C1

R1–R2
Mean R1–R2 and R2–R3
R2–R3
R4–R5
Mean R4–R5 and R5–R6
R5–R6
C1
JP1B

1500 (shaft: tension)
1390 (shaft: tension)
1430 (shaft: tension)
1700 (shaft: tension)
1420 (shaft: tension)
1270 (shaft: tension)
910 (shaft: compression)
673 (shaft: tension)
753 (base)
1290 (shaft: tension)

C1

1720 (shaft: tension)

3500

3·0

R3—1st test 13/11/1998 (85 days)*

R4—1st test 16/11/1998 (85 days)*
R5—1st test 19/11/1998 (90 days)*
R6—1st test 09/11/1998 (81 days)

3000

*Tests curtailed at a maximum rig load of 2000 kN

2500
2000
1500

2·0

C1

500

0

0

5

10
15
20
25
Pile head displacement: mm

30

3500

Jardine & Chow
(1996) trendline

2500
2000
1500
1000
500
0

0

5

10
15
20
25
Pile head displacement: mm

30

1

10

100
1000
Time after driving: days

10000

35

R1—1st test 02/09/1998 (9 days)
R2—1st test 17/04/1999 (235 days)
R6—1st test 09/11/1998 (81 days)

3000

CLAROM piles LL and LS
22 m long (restruck)

R1

1·0
0·5

0

R6

?

1·5

1000

Fig. 10. Load–displacement curves from first-time tests on
reaction piles R3, R4, R5 and R6

Force applied to pile head–tension positive: kN

Intact ageing characteristic (IAC)
R2

2·5

QS(t)/QSICP

Force applied to pile head–tension positive: kN

19 m long pile close
to JP1
19 m long pile close
to C1

35

Fig. 11. Overall load–displacement curves from first-time tension testing to failure of piles R1, R2 and R6

non-dimensional ratios against logarithmic time in Fig. 12.
The mean of the 2020- to 2100-day capacities of the
CLAROM LS and LL piles described in Table 2 (normalised
by the ICP tension shaft capacity corresponding to the BRE
CPT1 cone profile) is added to provide a minimum estimate

for the capacity available to ‘fresh’ piles after 5.5 years. (LL

Fig. 12. Normalised shaft capacities against time for first-time
tests on 19 m long reaction piles R1, R2, R6; 22 m long
CLAROM piles (all in tension) and 10 m long pile C1 (in
compression)

and LS had been restruck within six months of its installation: see Table 2.) The initial Qs (EOID) capacity point,
shown as Qs (t)=QICP , was evaluated on the basis of the
s
CLAROM group’s dynamic measurement of an end of initial
driving (EOID) compressive shaft resistance of 1280 kN for
their 22 m long LS pile, factored as recommended by the
ICP procedures to 896 kN to predict the initial tension
capacity. Given the problems of defining EOID reliably, the
initial EOID point and early age trend should be considered
tentative; data from other sand sites have indicated dynamic
Qs (EOID)=QICP ratios that may be higher (Overy, pers.
s
comm., 2000).
The relevant tension data points are joined together to
form an intact ageing characteristic (IAC) for the family of
‘fresh’, previously unfailed, 19 m long steel pipe piles driven
at Dunkirk and tested in tension. The compression shaft
capacity assessed from the first test on the 10 m long pile
C1 is also shown. Base capacity measurements made in the
well-instrumented CLAROM program were combined with
local CPT measurements to estimate C1’s base load as
1040 kN (Ỉ25%) at the 10% diameter settlement limit,
leading to a shaft capacity of 1780 kN (Ỉ14%). Applying

the ICP compression shaft capacity given in Table 4 leads to
the plotted range for Qs (t)=QICP, with the mean plotting
s
close to the IAC established for the tension piles, suggesting
that the ‘compression’ IAC for C1 may be similar to that of
the longer tension piles—at least at the age considered. It is
obvious that the ‘fresh’ Dunkirk piles gained capacity much


238

JARDINE, STANDING AND CHOW

more quickly than expected from Jardine & Chow’s tentative
trendline, which had been drawn through a mixed dataset of
first tests, retests and restrikes involving a variety of piles
and different sands. The fresh piles developed their ICP
capacities around ten days after driving, climbing to values
around 2.3 times higher over the following eight months.
The most conservative interpretation that can be made for
the long-term IAC would be to assume that a plateau developed at some point after eight months that passed through
the long-term LL and LS CLAROM piles’ capacities (with
an average Qs (t)=QICP ¼ 2:48), effectively assuming that
s
these two piles had been able to recover fully over five years
from their earlier restrikes. It is more likely that their 1994
shaft capacities would have been higher if they had been left
untested since their installation in 1988.
The very strong effect on pile capacity implies that all
time-independent calculation methods will be subject to considerable scatter unless they are restricted to predicting a

closely specified age range. Considering the time relationships suggested by Jardine & Chow (1996), the age at which
piles can be expected to match the ICP predictions should
now be reduced to around ten days, and the trend suggested
for subsequent capacity growth now appears over-conservative for ‘fresh’ piles.
Retests on previously failed piles
A surprising feature of the recent Dunkirk tests was the
brittle response of the aged piles. This brittleness is not
apparent in the load–displacement curves shown in Fig. 11.
But, as detailed in Table 3, retests performed soon after
unloading from an earlier test were unable to achieve the
same capacities. Figs 13(a), (b) and (c) present traces for
each of the GOPAL piles, showing how the (normalised and
pile-weight corrected for tension) capacity of pre-failed piles
varied with time. The fresh piles’ IAC and the Jardine &
Chow trendline are also shown. As detailed in the notes
accompanying Fig. 13, the traces sketched between the
known data points had to involve some elements of interpretation, as not all paths could be followed continuously or
precisely. It is difficult to investigate and separate all of the
potential processes with uninstrumented field-scale piles. For
practical reasons repeat loading tests often had to follow
hours or days after the preceding failure, so the immediate
reductions in shaft capacity could not always be quantified.
The normalised tension shaft capacity trends interpreted
from the dynamic and static tests5 on the CLAROM piles
are presented in Fig. 13(d). Pile LS gave a first restrike
capacity between the IAC and Jardine & Chow trendline,
whereas the shorter CS and CL piles appear to fall anomalously below the latter. The restrike tests on the short (CS,
CL) piles involved driving on to greater incremental penetrations (160–400 mm) than the 60–110 mm increments developed by the longer LL and LS piles. Their subsequent
histories were also more complex (see Table 2), implying
heavier ‘damage’ levels that are consistent with their far

lower long-term normalised capacities. Although the potential errors in interpreting the dynamic tests, and the possible
spatial variations of the CPT profiles within the CLAROM
test area, should be borne in mind, the CLAROM pile retest
traces scatter (like their GOPAL equivalents) between the
IAC and the possible EOID lower limit.

5
Note that tensile capacities have been plotted from dynamic tests
by applying the tension–compression capacity ratio given by
Jardine & Chow (1996), and that data from second restrikes have
been omitted because the base capacities could not be isolated
reliably after tension failures and soil plug removal.

Figures 14(a) and (b) present examples of the corresponding load–displacement curves for the GOPAL piles, and the
following observations are offered on the combined dataset.
(a) Any restrike or load cycle that causes failure degrades
capacity, despite the apparently ductile ‘first-time’
loading curves.
(b) Pre-failed piles have much lower capacities than fresh
piles (over the age range considered), which scatter
sporadically around the trendline suggested by Jardine
& Chow.
(c) It is not clear whether any lower bound applies to the
Qs (t)=QICP values of repeatedly failed piles. However,
s
none of the available static test data fall below the ratio
(% 0.72) applying at the EOID.
(d) Pre-failed piles develop more distinctly ‘brittle’ load–
displacement curves, although their peaks are hard to
capture in essentially load-controlled tests.

(e) Repeated cyclic or static loading to failure can cause
piles to degrade from the IAC towards the EOID
capacity. The potential losses of capacity become more
marked with time over the first months of the pile’s
life.
( f ) Pre-failed piles follow non-monotonic capacity–time
trends that depend on the timing and severity of their
prior loading episodes. Although their capacities recover with time, their capacity growth rates are
generally slower than those of fresh piles.
The last point may be emphasised by considering the
CLAROM piles. The CS pile had been restruck twice before
being subjected (in 1989) to pairs of tension and compression tests to failure, giving the results shown in Fig. 2. The
‘damage’ inflicted by the restrikes counteracted the otherwise beneficial effects of time, and the shaft capacity developed in the first tension test (T’89a conducted 180 days
after driving) fell below the ten-day ICP capacity with the
ratio Qs (t)=QICP ¼ 0:78; the retest capacity measured three
s
months later recovered to a marginally higher ratio of 0.87.
More significant gains occurred when the pile was allowed
to recover for a further 4.6 years, but the long-term
Qs (t)=QICP ratio amounted to 1.52, well below the minimum
s
ratio of around 2.3 suggested by Fig. 12 for fresh piles at
ages exceeding eight months. The single restriking of piles
LS and LL (involving 60 mm of incremental penetration)
caused less ‘damage’ than the more comprehensive testing
of the short piles, leading to Qs (t)=QICP ¼ 2:46 after the
s
same period. The long-term capacities of LS and LL might
well have been higher but for the earlier restrike tests.
It seems probable that much of the ‘damage’ associated

with the testing to failure takes place during the post-peak
and unloading stages. Building from the hypothesis of Chow
et al. (1997) that the key process is the changing efficiency
of circumferential stress arching action around the shaft, it is
suggested that the arching action is strengthened and the
radial stresses are reduced as the pile fails and is unloaded.
Cyclic loading experiments with the Imperial College Instrumented Pile (Lehane, 1992; Chow, 1997) in sands at Labenne and Dunkirk show that radial effective stresses fall on
unloading. This is interpreted as being due to contraction
taking place within the shear zone close to the shaft, which
allows the relatively stiff circumferential arch to carry more
load, effectively shielding the pile shaft from the high
ambient radial stresses. Creep is likely to reduce the efficiency of the arch gradually with time, leading to higher
shaft radial effective stresses and shaft capacity recovery.
An experiment was conducted on pile R4 to assess
whether low-level cyclic loading could accelerate the postulated arching creep processes and enhance capacity growth.
As summarised in Table 3, a rapid test performed in


EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND
3·0

3·0

Dunkirk
IAC

2·0

R2
C1

?

2·5

R2(1)
R2(2)

1·5

C1(2)

C1(1)

1·0

R1(1)

R1

R1(2)

C1(3)
R1(5)

R1(4)
R1(3)

R1(6)

Jardine & Chow
(1996) trendline

QS(t)/QSICP

QS(t)/QSICP

2·5

C1(4)

0·5
0

239

1

10

2·0

R3(1) & R4(1)

1·5

R4(2)

1·0

100
1000
Time after driving: days

0

10 000

R3(2)

R3(4) & R4(5)

R3(3)
R4(4)

Jardine & Chow
(1996) trendline

R4(3)

R3(1) & R4(1)

0·5

Line representing
capacity at end of driving

?

Dunkirk

IAC

Line representing
capacity at end of driving
1

C1(1): virgin path for C1 (end point proven by first test—in compression)
C1(2): decrease in capacity from compression test
(end point proven by second test)
C1(3): further decrease in capacity from three phases of cycling
(end point proven from third test)
C1(4): presumed regain in capacity with tim

10

100
1000
Time after driving: days

10000

R3(1): virgin path for pile R3 (end point estimated to be similar to that for R6)
R3(2): decrease in capacity from two phases of cycling testing
(end point proven by second test)
R3(3): increase in capacity (end point proven by third test)
R3(4): decrease in capacity indicted by brittle response of third test to failure
R4(1): virgin path for pile R4 (end point estimated to be similar to that for R6)
R4(2): decrease in capacity from two phases of high-level cycling
(end point proven by second test)
R4(3): increase in capacity (end point estimated by drawing line parallel to that

proven for R3(3))
R4(4): increase in capacity from low-level cycling
(end point proven by third test)
R4(5): decrease in capacity indicated by brittle response of third test to failure

R1(1): virgin path for pile R1 (end point proven by first test)
R1(2): decrease in capacity following fit test (projected from other field tests)
R1(3): increase in capacity (end point proven by second test)
R1(4): decrease in capacity following second test (projected from R1(5))
R1(5): increase in capacity (end point proven by third test)
line drawn parallel to that proven for R3(3)
R1(6): presumed decrease in capacity following third test to failure
R2(1): virgin path for pile R2 (end point proven by first test)
R2(2): decrease in capacity from one phase of cyclic testing to failure
(end point proven by second test)

(a)

QS(t)/QSICP

2·5

?

Dunkirk
IAC

2·0

2·5

R6
R5(1) & R6(1)

1·5

R6(2)

1·0

R6(3)

R5(1) & R6(1)

0·5
0

R5(2)
R6(4)
R5(3)

R6(5)

Jardine & Chow
(1996) trendline

10

100
1000

Time after driving: days

10 000

2·0
1·5

R5(1): virgin path for pile R5 (end point estimated to be similar to that for R6)
R5(2): decrease in capacity from two phases of cycling testing
(end point proven by second test)
R5(3): increase in capacity (end point proven by third test)
R6(1): virgin path for pile R6 (end point proven by first test)
R6(2): decrease in capacity from one phase of cycling
(end point proven by second test)
R6(3): further decrease in capacity following another phase of cyclic loading
(end point proven by third test)
R6(4): increase in capacity (end point estimated by drawing line parallel to that
proven for R5(5))
R6(5): decrease in capacity from cyclic failure (end point proven by fourth test)

LL(1)

LS(3)

CL(2)
LS(2)

LS(1)

CL(1)

1·0

0

?

Dunkirk
IAC

CS(1)

0·5

Line representing
capacity at end of driving
1

(b)

3·0

QS(t)/QSICP

3·0

CS(2)

Jardine & Chow
(1996) trendline

1

10

CS(4)

CS(3)

Line representing
capacity at end of driving

100
1000
Time after driving: days

10000

CS(1): path for pile CS with first restrike 160 days after EOID
(end point determined from restrike driving analysis)
CS(2): decrease in capacity from first restrike
(end point proven by tension test performed at 188 days)
CS(3): increase in capacity (end point proven by tension test at 272 days)
CS(4): increase in capacity (end point proven by tension test at 1991 days)
CL(1): increase in capacity of CL (start point proven by tension test at
176 days and end point proven by tension test at 288 days)
CL(2): increase in capacity (end point proven by pile extraction at 2087 days)
LS(1): path for pile LS with first restrike 159 days after EOID
(end point determined from restrike driving analysis)
LS(2): decrease in capacity from first restrike
(end point determined from estimated shaft capacity at 180 days)

LS(3): increase in capacity (end point proven by tension test at 1988 days)
LL(1): increase in capacity (end point proven by pile extraction at 2089 days)

(c)

(d)

Fig. 13. Normalised pile capacities against time for first-time and pre-failed tension tests for: (a) control pile C1 and reaction piles
R1 and R2; (b) reaction piles R3 and R4; (c) reaction piles R5 and R6; (d) CLAROM piles CS, CL, LS and LL

November 1998 indicated a capacity of around 1625 kN. In
April 1999 1000 regular tension cycles were applied over a
16 h period, with load maxima and minima of 805 kN and
5 kN respectively, giving a cyclic load amplitude of around
20% of the shaft capacity estimated for the time of testing.
The pile-head displacement amplitude maintained constant at
Ỉ1.25 mm under this relatively low-level cycling, and very
little permanent displacement developed. A tension test to
failure performed on the following day indicated a capacity
of 2491 kN, 53% greater than five months earlier. A parallel
pair of static tests were performed on a similarly pre-failed

pile (R3) over a comparable set of test dates (November
1998 and April 1999), but without applying the cyclic
loading. R3 developed a far more modest recovery (17%) in
capacity. Noting that gentle vibration accelerates creep in
granular media (Jardine et al., 2001a), the response of piles
R3 and R4 is compatible with the hypothesis that the
observed ageing and pretesting trends originate from a
circumferential arching action that (a) becomes more marked

after each extreme load cycle (involving slip) associated
with driving or testing, and (b) weakens with time through
creep.


JARDINE, STANDING AND CHOW

Force applied to pile head–tension positive: kN

Force applied to pile head–tension positive: kN

240
3500

R1: 1st test 02/09/1998 (9 days)
R1: 2nd test 28/10/1998 (57 days)
R1: 3rd test 26/04/1999 (239 days)
R2: 1st test 17/04/1999 (235 days)
R3: 2nd test 20/04/1999 (85 days*)
* Test performed same day as first,
but after episide of cyclic loading

3000
2500
2000
1500
1000
500
0

0

3500

5

10
15
20
25
Pile head displacement: mm
(a)

30

35

SUMMARY AND CONCLUSIONS
A programme of first-time loading and retest experiments
has been performed on ten field-scale open-ended steel pipe
piles driven in a predominantly dense silica marine sand at
Dunkirk, northern France. A careful interpretation has been
made that takes account of differences in pile dimensions,
local spatial variations across the test area, and a re-evaluation of time-effect studies performed by others. The following main conclusions follow from the programme described.

R1: 1st test 02/09/1998 (9 days)
R2: 1st test 17/04/1999 (235 days)
R5: 2nd test 15/04/1999 (234 days)
R6: 2nd test 11/11/1998 (244 days)

3000

sands and gravels) of various mineral compositions,7 in
states ranging from very loose to very dense. The tests
involved dynamic, static and Osterberg cell procedures. The
24 shaft data points on Fig. 15(a(ii)) scatter around the
Dunkirk IAC, even though some of the piles considered had
been restruck. The only unambiguous set of first-time static
tests in the database is the Jamuna Bridge series involving
mica sands (Fig. 15(b(ii))), and these fall relatively close to
the IAC. More research is required before drawing definitive
conclusions, and this should include reliable calculations or
measurements of the ten–day static shaft capacities of fresh
piles to allow the normalisation adopted in Figs 12 and 13.
However, the available information suggests that the Dunkirk
intact ageing characteristic may not be unduly site specific
or dependent on pile details.

2500
2000
1500

Synthesis with Chow’s database
The new test results have prompted a reassessment of the
database, assembled by Chow, of tests by others on aged
piles driven in sand. The CPT test profile and other parameters needed to run ICP capacity estimates were absent for
most of these cases, so the axes were kept as in Fig. 3, with
capacity expressed as a multiple of the EOID resistance
rather than the ICP values. The resulting total and shaft
capacities ratios are plotted in Figs 15(a) to (b). Fig. 15(a)(i)

contains the combined tension and compression database for
all cases (including restrikes) where there has been no prior
static failure. Fig. 15(a)(ii) presents the subset of cases
where it was possible to isolate shaft capacity,6 and contains
mostly compression tests. Figs 15(b)(i) and 15(b)(ii) show
the equivalent datasets after removing any cases where the
piles had been restruck prior to static load testing, along
with the Dunkirk IAC (noting that the relative position of
the Dunkirk IAC depends on the interpreted Qs (EOID)/QICP
s
ratio). These diagrams represent the wide range of pile and
soil types listed in Appendix 1, with concrete, timber and
steel piles of widths/diameters from 0.27 m to 1.03 m driven
to lengths 8.75–78 m in granular soils (ranging from silts to

(a) The ‘fresh’ Dunkirk piles developed substantial increases in their tension shaft capacities during the
weeks and months after driving, defining the piles’
intact ageing characteristic (IAC). A single check on
pile C1 indicated that compression shaft capacity grew
in a similar same way. Base resistance is not thought to
vary significantly with time.
(b) The beneficial ageing processes appeared to commence
within a few days of driving, although this early period
requires further investigation.
(c) Shaft capacities rose over eight months to more than
double those seen in load tests conducted a few days
after driving, or expected from calculation procedures
designed to match short-term test capacities.
(d) Further checks are required on completely undisturbed
aged piles to confirm longer-term trends. Tests

performed on two piles five years after they had been
restruck indicated ‘damaged’ normalised capacities
greater than those seen with ‘fresh’ piles after
eight months.
(e) The Dunkirk piles showed brittle responses after failure
and unloading. Extreme loading cycles, including
pretesting to failure and subsequent unloading, degraded shaft capacity and disrupted the growth of
capacity with time.
( f ) The capacity of pre-failed piles recovered with time,
but at relatively modest rates, giving non-monotonic
time–capacity traces that plot well below the ‘fresh’
piles’ IAC. Their capacity–time traces could appear to
be steeply upward, flat or negatively inclined—depending on the sequence and intensity of prior testing. The
EOID tension shaft capacity may provide a lower bound
to the shaft capacity of piles subjected to multiple static
or cyclic tests to failure.
(g) Whereas high-level cycling caused damage, low-level
one-way load cycling (with an amplitude around 20%

6
The capacities plotted in Fig. 15 from Tomlinson, pers. comm.
(1996) and Bullock et al. (2005a) represent only the shaft
capacities determined from Chin’s (1970) analyses of compression
tests, tension load tests and Osterberg cell tests.

7
Note that the capacities of piles driven in mica and calcareous
sands may be poorly predicted by standard pile capacity calculations; modified procedures are needed for these soils (Jardine et al.,
2005).

1000
500
0

0

5

10
15
20
25
Pile head displacement: mm
(b)

30

35

Fig. 14. Selection of load–displacement curves for pre-failed
reaction piles with youngest (R1) and oldest (R2) first time tests
also shown: (a) R1 to R3; (b) R5 and R6


EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND
4·5

Qt(t)/Qt(EOID)

4·0

3·5
3·0
2·5

5·0

Samson & Authier (1986)
Seidel et al. (1988)
Astedt et al. (1992)
Dunkirk CS
Dunkirk CL
Dunkirk LS
Skov & Denver (1988)
Tavenas & Audy (1972)
York et al. (1994)
Svinkin et al. (1994)
Tomlinson (1996)
Bullock et al. (2005a) BKM
Bullock et al. (2005a) VLE

4·0

2·0

3·5
3·0
2·5
2·0
1·5

1·5
1·0

1·0

0·5

0·5

5·0
4·5
4·0

1

10
100
Time after driving: days
(i)

Samson & Authier (1986)
Seidel et al. (1988)
Astedt et al. (1992)
Tomlinson (1996)
Bullock et al. (2005a) BKM
Bullock et al. (2005a) VLE

3·5

1000

10
100
Time after driving: days
(i)

1000

10 000

Tomlinson (1996)

4·0

?

3·0
Dunkirk
IAC

1·5

1

5·0
4·5

2·5
2·0

0
0·1

10000

QS(t)/QS(EOID)

0
0·1

QS(t)/QS(EOID)

Skov & Denver (1988)
Tavenas & Audy (1972)
York et al. (1994)
Tomlinson (1996)

4·5

Qt(t)/Qt(EOID)

5·0

241

?

3·5
3·0
2·5

2·0

Dunkirk
IAC

1·5

1·0

1·0

0·5

0·5

0
0·1

1

10
100
Time after driving: days
(ii)
(a)

1000

0
0·1

10000

1

10
100
Time after driving: days
(ii)
(b)

1000

10 000

Fig. 15. Data relating to: (a) restrike tests or first-time static tests; (b) first-time static tests. Selected from original database;
expressed in terms of: (i) total pile capacity and (ii) shaft resistance alone

of shaft capacity) accelerated the beneficial ageing
processes at Dunkirk.
(h) It is not certain how scale, or sand type and state,
affect shaft capacity growth with time. However, a reevaluation of the database summarised in Appendix 1
indicates that piles formed from concrete, steel or
timber, with diameters between 0.26 and 1.03 m, driven
in various sands types and densities, all show considerable gains in shaft capacity within six months of
driving, following trends that are at least broadly
comparable to the recent Dunkirk experiments. Other
workers have also reported substantial gains with time
in tests on smaller-scale rods or SPT samplers.
(i) It is essential to separate tests on ‘fresh’ and pretested

piles when studying ageing effects. Eliminating restrikes and retests from Chow’s dataset leads to a subset
of ‘first-time’ tests on ‘fresh’ piles that conforms with
the ‘fresh’ pile IAC relationship found at Dunkirk.
( j) Pile capacity calculation procedures that take no
account of time will be subject to considerable error
unless they consider only a tightly specified age range.
In the case of the ICP procedures, the standard
calculation is most likely to match field capacities in
tests conducted around ten days after installation.
(k) The capacity–time relationship suggested earlier by
Jardine & Chow (1996), and those offered by most
other workers, can now be understood as being
arbitrarily affected by pre-failures. The retests described
in this paper give shaft capacities that scatter

(l)

sporadically around the initially suggested trendline,
while the ‘first-time’ IAC lies far above it.
The field data described above are consistent with the
previously offered explanation for the time dependence
of shaft capacity for piles driven in sands: that the
radial stresses developed on the shaft grow through a
relaxation (with time) of a circumferential arching
stress field. Changes in the degree of restrained dilation
that develop as the shaft is loaded to failure may also
occur as the piles ‘age’, and physico-chemical processes such as corrosion may also contribute. The
degree of radial expansion required for sand grains to
unlock from the pile shaft, and the stiffness of the
restraining soil mass, may increase with time.

The field observations have many practical repercussions and
implications, especially with regard to test timing and
matching service loading requirements. The ageing processes
offer potential practical benefits if piles have been, or can
be, driven months or years before any critical loading events
can occur—as in carefully staged construction, or when
reusing pre-installed aged foundations. Equally, it is clear
that piles driven in sand are damaged by failure and cannot
carry the same loads as intact aged piles. There are therefore
clear problems in performing multiple tests on the same pile
to assess differences between tension and compression behaviour, length–depth or ageing effects. Further research
into the underlying processes, their quantification and possible practical exploitation is required urgently.


JARDINE, STANDING AND CHOW

242

ACKNOWLEDGEMENTS
The above research was funded by the EU (through the
GOPAL project) and the UK Health and Safety Executive
(HSE), and their support is gratefully acknowledged. The
authors thank Mr Eric Parker of D’Appolonia, Genova, for his
major contribution to the project, the Port Autonome de Dunkerque for their generous loan of the site, and Mme Francoise
¸

Brucy and M. Jean-Francois Nauroy for providing data from the
¸
CLAROM tests. The field testing was performed in conjunction

with Precision Monitoring and Control (PMC) of Teeside (UK);
much of the laboratory work at Imperial College was conducted
by Dr Reiko Kuwano and Mr Tim Connolly; the in situ soil
testing was performed by the Building Research Establishment
(Garston, UK) and Simecsol of Dunkerque, France.

APPENDIX 1
Reference

Test location

Soil description

Tavenas & Audy (1972)

Medium dense, fine
sand; k ¼ 10À4 m/s
Sand and gravel
Medium dense
hydraulic sand fill
Sand and silt
Very loose to very
dense sand. Piles
founded 2 m into
limestone
Sweden, various Silts and sands,
sites
insignificant
carbonate contents
Orsa, Sweden

Loose to dense,
fine to medium sand
JFK Int.
Medium dense,
Airport, USA
medium fine sand.
2 m thick clay and
peat layer near
surface
Alabama, USA
Silty sand
Jamuna Bridge,
Loose to medium
Bangladesh
dense, silty, medium
fine, micaceous sand
Buckman Bridge Dense fine sand
(BKM), Florida,
USA
Dense fine sand
Vilano Bridge
East (VLE),
Florida, USA

Pile type

Equivalent
diameter: m

Average

length: m

(S)tatic/
(D)ynamic
testing

Max.
time:
days

11

S

56

14
9.1

D & S
S

51
100

8.75
14.7

D & S
D & S

23
535

D & S

300

S

64

St Charles
River, Quebec
Samson & Authier (1986) Jasper, Canada
Ng et al. (1988)
Hunters Point,
San Francisco
Skov & Denver (1988)
MBB, Hamburg
Seidel et al. (1988)
Barwon Bridge,
Australia

Concrete
0.320
hexagonal
Steel H
HP 310X79
Concrete square

0.344

Holm (1992); Holm &
˚
Astedt, pers. comm.
(1995)
˚
Astedt et al. (1992)

Concrete square 0.265–0.305

Various

Concrete square
0.305
10:1 inclined
0.200–0.355
Monotube,
timber and steel

26.3
20

D & S

224

Concrete square 0.516–1.032
Concrete square 0.451–0.508
Steel tubular

0.762
inclined
Concrete square
0.516

21.5
25
78

D & S
D & S

23
86
270

9.16

D & S
(O-cell)

268

10.68

D & S
(O-cell)

77

York et al. (1994)

Svinkin et al. (1994)
Tomlinson, pers. comm.
(1996)
Bullock et al. (2005a)
Bullock et al. (2005a)

Concrete square
Concrete square
with tapered toe

Concrete square

0.395
0.508

0.516


EFFECTS OF TIME ON CAPACITY OF PILES DRIVEN IN SAND

243

APPENDIX 2
0

0

0

R2–3
R5–6
GP1A
GP1B

R2–3
R5–6
GP1A
GP1B
5

5

10

10

10

Depth: m

Depth: m

Depth: m

5

R2–3
R5–6

GP1A
GP1B

15

15

15

20

20

20

25

0

10
20
30
40
50
Cone end resistance, qc: MPa

25

60

0

25

100 200 300 400 500 600 700
Sleeve friction, fS: kPa
(a)

0

0

0

1

2
3
Friction ratio: %

4

5

0
R1–2
R4–5
C1

R1–2

R4–5
C1

R1–2
R4–5
C1

5

5

10

10

10

Depth: m

Depth: m

Depth: m

5

15

15

15

20

20

20

25

0

10
20
30
40
50
Cone end resistance, qc: MPa

60

25

0

100 200 300 400 500 600 700
Sleeve friction, fS: kPa
(b)

25

0

1

2
3
Friction ratio: %

4

5

Fig. 16. Cone penetration test data from test site at Dunkirk in terms of cone end resistance, sleeve friction and
friction ratio against depth for (a) Section A–A and (b) Section B–B, as given in Fig. 6

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