´ White, D. J. & Lehane, B. M. (2004). Geotechnique 54, No. 10, 645–658 Friction fatigue on displacement piles in sand D. J. W H I T E * a n d B. M . L E H A N E † Experiments with instrumented displacement piles have ´ Les essais avec des piles de deplacement instrumentees ´ shown that the ultimate shaft friction that can develop in ´ ont montre que la friction d’arbre ultime qui peut a given sand horizon decreases as the pile tip penetrates ˆ apparaıtre dans un horizon sableux donne diminue a ´ ` to deeper levels. This phenomenon, which is now com- ´ ` mesure que la pointe de la pile penetre a des niveaux` monly referred to as friction fatigue, is investigated here ´ ´ ` plus bas. Nous etudions ici ce phenomene qui est au- using centrifuge model piles equipped with lateral stress ´ ´ jourd’hui appele communement fatigue de friction, en sensors, and by drawing on other experimental data from ` utilisant des piles en modeles centrifuges equipees de ´ ´ the laboratory and the ﬁeld. It is shown that the primary ´ ´ detecteurs de contrainte lateraux et en puisant dans mechanism controlling friction fatigue is the cyclic his- ´ ´ d’autres resultats experimentaux en laboratoire et sur le tory imparted during pile installation to soil elements at ´ terrain. Nous montrons que le mecanisme primaire con- the pile–sand interface. For a given installation method ˆ trolant la fatigue de friction est l’histoire cyclique impar- the stationary lateral stress acting at any given level on a ´´ tie pendant l’installation des piles sur les elements de sol displacement pile can be described as a relatively unique ` ´ a l’interface pile-sable. Pour une methode d’installation function of the cone penetration test end resistance and ´ ´ donnee, la contrainte laterale stationnaire agissant a un ` the number of cycles imposed during installation. The ´ ´ niveau donne sur une pile de deplacement peut etre ˆ strong inﬂuence of cycling, which is also seen in cyclic ´ decrite comme une fonction relativement unique de la constant normal stiffness interface shear tests, is attribu- ´ resistance ﬁnale CPT et du nombre de cycles imposes ´ ted to contraction of a narrow shear zone at the shaft– pendant l’installation. La forte inﬂuence du cyclage qui soil interface that is surrounded by soil with a relatively ´ ´ est egalement notee dans les essais de cisaillement sur high lateral stiffness. ` ´ ` une interface a rigidite normale a constante cyclique est ´ ` attribuee a la contraction dans une zone de cisaillement ´ ` ´ etroite a l’interface arbre-sol qui est entouree par un sol KEYWORDS: piles; sands ´ ´ ayant une rigidite laterale relativement elevee. ´ ´ INTRODUCTION techniques are to become more widely accepted, research is There are signiﬁcant uncertainties associated with the pre- needed to investigate the inﬂuence of the installation method diction of the axial capacity of displacement piles in sand. on pile behaviour. Since conventional design methods have Recent experiments involving ﬁeld-scale instrumented piles been developed based on historic experience with dynami- (e.g. Lehane, 1992; Chow, 1997) have successfully reduced cally installed piles, caution is required when these methods the level of this uncertainty, leading to the development of are applied to jacked piles. improved approaches such as those proposed by Lehane & Jardine (1994), Randolph et al. (1994) and Jardine & Chow (1996). These methods have a greater prediction reliability BACKGROUND primarily because of the improved understanding of shaft Field measurements of the distribution of horizontal and friction characteristics obtained through measurement of shear stress acting on a pile shaft have highlighted the horizontal stresses acting on the pile shaft (Poulos et al., deﬁciencies of conventional ‘earth pressure’ design ap- 2001). However, a number of signiﬁcant facets require proaches in which the in situ vertical effective stress proﬁle further investigation to enable formulation of more generally is multiplied by assumed earth pressure and interface friction accepted design approaches with reduced reliance on empiri- coefﬁcients to estimate a distribution of the available ulti- cal correlations. mate shaft shear stress (ôf ) with depth; ôf values derived in A second motivation for research into shaft friction on this way are often restricted to limiting maximum values displacement piles lies in the development of new pile (e.g. API, 2000). The ‘earth pressure’ approach for displace- installation techniques that involve jacking rather than con- ment piles in sand has been largely replaced in the onshore ventional dynamic techniques, which are becoming less environment by methods that relate ôf to an in situ test acceptable in urban areas because of noise, vibration and parameter such as the cone penetration test (CPT) end emission restrictions. White et al. (2002) and Lehane et al. resistance (qc : e.g. Bustamante & Gianeselli, 1982). The (2003) describe two novel pile-jacking systems with capaci- design method of Jardine & Chow (1996) also assumes a ties greater than 4 MN for the installation of large pre- direct relationship between ôf and qc but is shown to have formed displacement piles. These machines offer an alter- greater reliability primarily because it incorporates the trend, native to bored piling for the construction of deep founda- seen in ﬁeld experiments, for the available shaft friction at a tions in urban areas. However, if these new jacking given soil horizon to decrease with increasing penetration of the pile tip (e.g. Lehane et al., 1993). Heerema (1980) referred to this phenomenon as friction fatigue and demon- Manuscript received 17 February 2004; revised manuscript accepted strated its inﬂuence on pile driveability. 20 September 2004. Discussion on this paper closes on 1 June 2005, for further details The Randolph et al. (1994) and Jardine & Chow (1996) see p. ii. design methods account for friction fatigue using relation- * Department of Engineering, University of Cambridge, UK. ships that can be divided into (a) the calculation of a † School of Civil and Resource Engineering, University of Western maximum available shaft friction (ôf ) from a sand horizon Australia, Crawley, Western Australia. just above the pile tip and (b) the calculation of a degrada- 645 646 WHITE AND LEHANE tion, or fatigue, of ôf as the pile tip penetrates further. gests a possible inﬂuence of differences between the respec- Randolph et al. (1994) propose a reduction in lateral stress tive jacking procedures. For example, the ICP was installed (and hence shaft friction) towards active earth pressure in jacking strokes with a typical length of 0.23 m (2.3D), conditions in proportion to eÀì h=D , where h is the distance and therefore the number of cycles encountered at z ¼ 3 m above the pile tip, D is the pile diameter, and ì is an prior to arrival of the three instruments was about 2, 6 and empirical degradation factor. They note that ì is likely to be 17 respectively. A 40% reduction in ó hc (or ôf ) occurs 9 inﬂuenced by a number of factors, including soil compressi- between h/D of 4 and 37 for this case, whereas no friction bility, pile roughness and incremental driving energy. Jardine fatigue is measured between h/D of 4 and 35 for the CPT, & Chow (1996) refer to friction fatigue as the ‘h/R effect’ for which the number of cycles experienced at a soil horizon (R being pile radius), and propose that the horizontal stress less than 1 m behind the cone tip is either zero or one, and acting on the pile shaft reduces in proportion to (h/R)À r , between 1 m and 2 m is either one or two. where an r value of 0.38 was found to provide a best ﬁt to a The ﬁnal example of friction fatigue is from a full-scale large database of high-quality load tests. instrumented tubular pile installed in dense sand for the The mechanisms governing friction fatigue are unclear, EURIPIDES project, described by Fugro (1996). A heavily and there is no particular behavioural model that provides a instrumented 762 mm diameter open-ended tubular pile was basis for the derivation of factors such as ì and r. Further- load-tested at three penetrations between 30 and 47 m below more, while both of these friction fatigue approaches choose ground level. The test site comprised dense sand below 22 m to normalise the distance h by pile diameter (or radius), this depth, with CPT qc values in the range 45–70 MPa. The normalisation remains to be justiﬁed by a governing me- pile was driven using a 90 kJ hammer reaching a blowcount chanism. Few ﬁeld records include measurements of local of $400 blows/m at the ﬁnal embedded depth. shaft friction distribution, and these are often unreliable Figure 1(c) shows the distribution of external ôf values at owing to residual loads, instrument damage during driving, a pile head settlement of D/10 calculated from the axial and the difﬁculty of separating internal and external shaft force distribution in the pile and ignoring internal shaft friction on open-ended piles. friction (which Randolph et al., 1991, and others, show to To illustrate friction fatigue, proﬁles of ultimate shaft be negligible at more than about 4D above the tip). Some friction from three different types of pile or penetrometer scatter is evident, as would be expected considering the are shown in Fig. 1. These have been selected to demon- demands placed on the instrumentation during such hard strate the extreme cases of low and high friction degradation driving. It is also noted that the external shaft frictions during installation, and show that the rate of degradation is inferred over the lower two diameters are inﬂuenced by high strongly inﬂuenced by the method of installation. internal shaft friction mobilised on the plug (shown dotted), The results from four soundings using a CPT instrument as evidenced from internal horizontal stress measurements. equipped with four friction sleeves located above the penet- However, there is a clear trend for signiﬁcant friction rometer tip are shown in Fig. 1(a) (DeJong, 2001; DeJong & fatigue. For example, over the depth range 30–35 m, the Frost, 2002). Three conﬁgurations of the CPT allow meas- mean ôf value was 330 kPa during the load test when the urements of sleeve friction at distances of between 167 mm pile tip was at 38.7 m. After a further 8.2 m (11D) of and 1517 mm behind the shoulder of the cone tip. During driving, a subsequent re-test indicated that this value of ôf the soundings, each of which involved jacking strokes with a reduced to 130 kPa, which is only 40% of the original value. length of 1 m (or 23D), the second and third sleeve indicate During the period of driving between load tests, approxi- frictions that are $8% below and above those registered by mately 3000 hammer blows were applied to the pile head. the ﬁrst and last sleeves respectively. Post-test measurements The three examples described above suggest that friction of the sleeve diameters revealed the second and third sleeves fatigue does not occur in the absence of loading cycles, and to be 0.05 mm smaller and larger than the ﬁrst and last, that a greater number of cycles imposed during installation explaining the observed consistent minor difference in re- leads to a larger reduction in shaft friction at a given soil corded friction. Aside from these minor discrepancies, the horizon. This observation suggests that the distribution of three soundings indicate no reduction in shaft friction be- available shaft friction after pile installation might be better tween h/D ¼ 4 and 35 (or between 167 and 1517 mm behind predicted by linking friction fatigue to the loading cycles the penetrometer shoulder). induced by the installation method rather than employing an Measurements of stationary horizontal effective stress, ó hc , 9 empirical relationship with h/D. acting on the shaft of the 101 mm diameter, Imperial College Prompted by trends such as those inferred from the data instrumented model pile (ICP) at three stages during installa- in Fig. 1, this paper describes a systematic investigation into tion in dense sand are shown in Fig. 1(b) (Chow, 1997; test the effects of the installation method on friction fatigue and DK2). The term ‘stationary’ is employed as these values were shaft capacity of displacement piles in sand. The investiga- recorded between jacking strokes when the pile was under tion was conducted at reduced scale using instrumented zero head load. Four horizontal stress transducers located model piles in a geotechnical drum centrifuge so that a large along the pile shaft provide a series of observations at a given number of pile installations could be performed in a uniform soil horizon as each instrument passes that depth, z. A clear soil sample at a cost that was approximately two orders of trend for ó hc measured at a given z to reduce with increasing 9 magnitude lower than that of an equivalent series of ﬁeld- penetration is evident. For example, close to z ¼ 3 m, ó hc is 9 scale tests. The paper focuses on the distribution of lateral recorded as 169 kPa when the ﬁrst instrument, located stresses acting on the centrifuge piles during installation and 406 mm (4D) behind the pile tip, passes that level. A 36% subsequent cyclic load tests to assist understanding of the reduction occurs prior to the arrival of the second instrument friction fatigue mechanism. after a further 965 mm (9.5D) of penetration. On arrival of the fourth instrument, located 3760 mm (37D) behind the pile tip, the registered ó hc value had fallen to 98 kPa, which is 9 DESCRIPTION OF EXPERIMENTS less than 60% of the value measured at h ¼ 4D. Drum centrifuge Measured ôf values showed a virtually identical rate of This investigation comprised 18 pile installations con- degradation with h to that of the ó hc values (see Chow, 9 ducted in the drum centrifuge at the University of Western 1997). This degradation is in marked contrast to that shown Australia (UWA). The ring channel of this machine has an in the multiple friction sleeve CPT experiments, which sug- outer diameter of 1.2 m, an inner diameter of 0.8 m, and a FRICTION FATIGUE ON DISPLACEMENT PILES IN SAND 647 Stationary horizontal stress, σ′ hc 0 50 100 150 200 40 0 0 Normalised instrument distance behind cone shoulder, h/D S1 1600 35 S2 Instrument distance behind cone shoulder: mm 1 10 S3 1400 30 S4 1200 Normalised instrument depth, z/D 25 2 20 Instrument depth, z: m 1000 20 800 3 30 15 600 4 40 10 400 5 Conventional f riction 200 5 50 sleeve z tip 3·26 m = 32 D measurements z tip 4·36 m = 43 D 0 0 0 25 50 75 100 z tip 5·77 m = 57 D Mean shear stress, τf: kPa 6 60 (a) (b) Ultimate shaft friction, τf: kPa 0 200 400 600 800 1000 20 z tip 30·45 m 30 z tip 38·70 m 25 z tip 46·90 m 35 30 40 Normalised depth, z/D Depth, z: m 45 35 50 40 55 45 60 50 65 (c) Fig. 1. Observed distributions of shaft friction on: (a) a multi-sleeve CPT instrument (D 43.7 mm) (DeJong, 2001); (b) a jacked instrumented model pile (D 101.6 mm) (Chow, 1997); (c) a long offshore pile (D 762 mm) (Fugro, 1996) channel height of 0.3 m. The key advantage of a drum A key feature of the centrifuge is its independently centrifuge over a beam centrifuge is the increased plan area rotating central shaft and tool table, which can be driven of the soil sample, which permits a large number of tests to relative to the ring channel by a hollow stepper motor, and be conducted in a uniform soil model. Compared with the brought to a halt independently of the channel. An actuator 1.8 m radius beam centrifuge at UWA, a sample in the drum is mounted on the tool table, onto which instrumented tools centrifuge has three times more plan area, and so is ideally can be attached and controlled. Twin stepper motors allow suited to parametric studies of the kind reported in this precise vertical and radial movement of the tool, while a paper. further stepper motor controls a counterbalance. Two per- 648 WHITE AND LEHANE sonal computers (PCs) are mounted on the tool table: one pile, after which the slots were backﬁlled with clear epoxy controls the actuator stepper motors, and the other is for (Fig. 2). The head of the pile was attached to a cap that was data acquisition. The actuator control PC receives commands connected via a 10 kN axial load cell to the tool table. The from a PC in the control room via a serial link across the wiring for the pressure cells passed up the body of the pile slip rings. Feedback between the onboard data acquisition in a 6 mm diameter hole, giving the pile a net cross- PC and the control room PC allows the actuator to be sectional area of 52.7 mm2 . Noting that the ratio of the operated in a load-controlled mode. A full technical descrip- Young’s modulus of steel to that of concrete is $7, the axial tion of the facility is provided by Stewart et al. (1998). stiffness of the model pile is about 4.5 times higher than that of a solid concrete pile. After construction of the pile, the pressure cells were Model pile exercised under air in a pressurised chamber. No hysteresis To measure the horizontal stress acting on the shaft, a was evident in the range 0–500 kPa, and linear calibration model pile was fabricated with miniature total pressure cells factors, including those to correct for a small cross-sensitiv- (Kyowa PS-5KA) embedded on the surface (Fig. 2). A ity to axial load in the range 10–20 kPa/kN, were estab- square section (9 mm 3 9 mm) pile was used so that these lished for each sensor. The possibility of cell action effects ﬂat circular cells could be mounted ﬂush with the pile inﬂuencing the operative calibration factor in soil was con- surface. The pile was fabricated from stainless steel and sidered, although it will be shown later that cell action machined to a mean centreline surface roughness, RCLA , of effects were, at worst, consistent between instruments, per- 0.55 ìm. mitting comparison of horizontal stress measurements at Six pressure cells were mounted at four locations between different locations. 9 and 108 mm behind the pile tip (i.e. h/B ¼ 1, 3, 6 and 12, where h is the height above the pile tip and B is the pile width). By instrumenting both sides of the pile close to the Soil properties tip, a degree of redundancy was incorporated and the repeat- The centrifuge tests employed uniform rounded ﬁne silica ability of the measurements could be checked. The pressure sand, with properties summarised in Table 1. A series of 30 cells were glued into slots machined into the face of the constant normal load direct interface shear box tests were performed at a range of relative densities in excess of $50% and vertical stresses between 50 and 250 kPa. Although the interface would normally be considered smooth, having a relative roughness, Rn % 2RCLA /D50 of 0.006, all samples dilated by between 0.01 and 0.02 mm and registered a peak strength. The average peak and constant-volume friction angles were 168 and 128 respectively, showing no signiﬁcant variation with initial density or conﬁning stress over the range investigated. Soil model preparation A single bed of sand was used for the entire testing programme. This bed was prepared in three stages. First, the channel was ﬁlled to the maximum depth of 200 mm with dry sand using an automatic pourer while the ring channel Cell was spinning at 20g. The sand was then saturated with water B4 and left to drain overnight through the base, leaving the material slightly damp and with sufﬁcient suction to hold 108 mm the sample in position when the ring channel was halted. (12B) Finally, the channel was brought to a halt and the surface 185 mm was screeded ﬂat to a nominal sample depth of 180 mm using a rotating cutter attached to the tool table. The actual variation in sample depth around the channel, as shown by the touchdown position of each pile test, was less than Æ 1 mm. The channel was then accelerated to 50g, and Cell remained spinning for 9 days while the test programme was B3 conducted. The bottom channel drain remained open throughout testing, and no outﬂow of water was evident after 54 mm the ﬁrst 24 h of spinning. However, during emptying of the 6B channel after testing, the lower 50 mm of the sand bed Cell remained slightly damp, indicating that completely dry con- Cell F2 B2 ditions had not been achieved. 27 mm Cell Cell (3B) Table 1. Soil properties F1 B1 Property Value 9 mm (B) D10 particle size: mm 0.100 D50 particle size: mm 0.180 Maximum voids ratio, emax 0.762 9 mm (B) 9 mm (B) Minimum voids ratio, emin 0.493 Speciﬁc gravity, Gs 2.65 Fig. 2. Schematic diagram of instrumented model pile FRICTION FATIGUE ON DISPLACEMENT PILES IN SAND 649 SOIL PROFILE CHARACTERISATION distribution of Dr may be attributed to the reducing ‘fall A total of 12 CPT soundings were conducted to character- height’ of the sand during deposition. ise the sand bed, using a 6 mm diameter CPT probe inserted into the soil at a rate of 1 mm/s. These tests were conducted at different positions around the sand bed to assess the TEST PROGRAMME homogeneity of the sample, and after each change in g-level. The full test programme included tests at acceleration The CPT soundings conducted during the initial phase of levels of 50g and 150g. This paper will consider only the testing at 50g are shown in Fig. 3. Good agreement is found initial series of tests conducted at 50g (tests T1–T13, Table between all tests in the upper 80 mm of the sample, with 2). The tests were located at 108 intervals around the drum deviation of up to Æ15% from the average evident between perimeter, and staggered between the upper and lower parts 80 and 120 mm. The average of soundings 1A, 1B, 2A and of the channel. This arrangement was selected to maximise 2B has been used as the ‘design’ proﬁle throughout the the separation of the test locations and the channel walls. back-analysis described in the paper. Adjacent tests were located 108 mm (12B) apart at the sand Using the Lunne & Christofferson (1983) correlation be- surface, increasing to 123 mm (13.7B) apart at the ﬁnal tween qc , vertical effective stress, ó v , and relative density, 9 installation depth of 120 mm owing to the radial direction of Dr, it is estimated that the sample’s relative density (Dr ) installation. increases with depth from $20% at the sample surface to $80% at a depth of 60 mm and $90% at 120 mm. This Installation methods Cone resistance, qc: MPa The pile was installed to a ﬁnal depth, L, of 120 mm in all 18 pile installations performed. Installation was generally 0 10 20 30 40 0 paused after 60 mm, L/2, of penetration so that a static compression test could be performed before installation continued to the full depth (Table 2). After reaching full Average depth, the pile was unloaded before undergoing a sequence 20 CPT 1A of static and/or cyclic load tests. Three techniques were used to install the test pile in order to investigate the inﬂuence of CPT 1B cycling during penetration: CPT 2A 40 CPT 2B (a) monotonic installation, comprising a monotonic push at 0.2 mm/s to half of the ﬁnal pile depth followed by another monotonic push to the ﬁnal pile length 60 (b) jacked installation, comprising cycles of ﬁxed down- ward displacement (2 mm at 0.2 mm/s, i.e. one ‘jack- Depth: mm stroke’) followed by unloading to zero head load at 0.005 mm/s 80 (c) ‘pseudo-dynamic’ installation, comprising two-way cycles of ﬁxed downward (2 mm at 0.2 mm/s) and upward (1.5 mm at 0.2 mm/s) displacement. 100 Load tests 120 Each test included a sequence of load tests as shown in Table 2. The procedure for each type of load test presented in this paper is as follows: 140 (a) compression load test, where the pile was jacked downwards at 0.005 mm/s for a distance of at least Fig. 3. Cone resistance against depth 1.5 mm Table 2. Summary of test programme Test number T1 T2 T3 T4 T5 T6 T7 T8 T9 T10 T11 T12 T13* Installation method Monotonic installation Jacked installation Pseudo-dynamic Load testing Compression load test at L/2 Compression load test at L Tension test on extraction L 1-way cyclic compression test 2-way cyclic compression/tension test 1-way cyclic tension test * Test halted after compression load test at 60 mm penetration. 650 WHITE AND LEHANE (b) cyclic compression load test, where the pile head load installation on the density, stress history, and hence stiffness was cycled between a nominally zero pile head load of of the soil beneath the pile tip. 50 N and a maximum of 750 N at a velocity of It should be noted that the relative stiffness of the shaft 0.01 mm/s and the base response of model piles does not mimic ﬁeld (c) cyclic compression-tension load test, where the pile piles. Base response scales primarily with diameter, since it head load was cycled between –50 N and 750 N at a is a continuum failure mechanism, whereas shaft response velocity of 0.01 mm/s. does not scale, since it is governed partly by the shear stress–displacement response of the interface and partly by continuum downdrag of the soil surrounding the pile, as required to generate this shear stress in the far ﬁeld. As a PILE HEAD LOAD MEASUREMENTS result, the settlement required to mobilise ultimate base Although the primary aim of this paper is to examine resistance during model tests scales correctly as a fraction of mechanisms controlling friction fatigue for displacement pile diameter, typically D/5 or D/10. In contrast, the settle- piles in sand, it is instructive to ﬁrst examine trends indi- ment to mobilise shaft resistance does not scale, and remains cated by the pile head load measurements obtained during similar to the absolute ﬁeld values of a few millimetres, pile installation and subsequent load testing. rather than reducing in proportion to the model scale to a fraction of a millimetre. Therefore, although the extraction strokes (1.5 mm) represent a realistic rebound and reversal Installation of shaft resistance for modelling dynamic installation, the The pile head loads recorded during each installation base rebound is very high, at D/6. This is an unrealistic mode are plotted in Fig. 4. The monotonic installation data model to examine the effect of installation method on base are collated in Fig. 4(a) and show excellent repeatability, response, and therefore the resulting differences in base with a maximum deviation between the four tests of Æ10%. stiffness (and hence head stiffness) should not be considered Although the pile was not equipped with a base load cell, representative of the ﬁeld case. some estimate of the base resistance can be made by subtracting the pullout load, which was typically 200 N, or 10% of the ﬁnal installation resistance. Since the head load HORIZONTAL STRESS DURING MONOTONIC is dominated by base resistance, any difference between INSTALLATION compressive and tensile shaft capacity can be overlooked in The six total stress cells provide measurements of hori- an approximate estimate of base resistance. The mean ﬁnal zontal stress acting on the pile shaft at four different installation force of 2100 N therefore indicates a mean unit locations behind the pile tip (h/B ¼ 1, 3, 6 and 9). The base resistance, qb , of about 23.5 MPa, which is slightly measurements recorded during monotonic installation (test below the mean CPT resistance at the same depth (Fig. 3). T10) are shown in Fig. 6. These are seen to mirror the CPT The pile head loads recorded during jacked installation at qc proﬁle, apart from interruptions corresponding to a tip the end of each 2 mm jacking stroke are shown in Fig. 4(b). depth of 60 mm (when the pile head was unloaded prior to a Between strokes, the pile was extracted until the head load static compression load test) and the ﬁnal tip depth of reduced to a nominal zero value (of 50 N) to mimic the one- 120 mm (when the pile was unloaded prior to further load way cycling at the pile head imposed by pile jacking. The testing). pseudo-dynamic installation data are shown in Fig. 4(c) and, Stresses recorded when the pile is moving are referred to although showing greater scatter than the other two installa- as ó hm . All ó hm data recorded during the four monotonic 9 9 tion methods, exhibit a deviation of less than 15% between installations are normalised by the corresponding qc values piles. It is clear that the 1.5 mm of extraction during each and plotted against instrument depth in Fig. 7; the error bars cycle mobilises signiﬁcant negative shaft friction, particu- shown in this ﬁgure indicate one standard deviation above larly as the ﬁnal embedment is approached. This two-way and below the average value. It is evident that: cycling of shaft friction mimics the conditions during dy- namic pile installation in an approximate manner. (a) The normalised horizontal stress, ó hm /qc , remains 9 The maximum pile head loads for each installation mode approximately constant throughout installation at are very similar and, since the head load is dominated by 0.016, with possibly a very slight decrease with base resistance, this close agreement indicates that the increasing depth (or stress level). This value corre- installation procedure has minimal inﬂuence on the ultimate sponds to a friction ratio, ôf /qc , of 0.34%, using the base resistance. constant-volume friction angle of 128 recorded during interface testing. This ratio is at the lower end of the range typically measured during CPTs in dense uniform sands (Lunne et al., 1997), reﬂecting the higher Monotonic compression load tests roughness and therefore interface friction angle of a The pile head load–displacement data recorded for each CPT (e.g. Rcla ¼ 0.18–6.85 ìm; DeJong et al., 2001) monotonic compression test conducted after installation to compared with the model pile. It may therefore be 120 mm are shown in Fig. 5 for the three installation meth- assumed that no gross error exists in the horizontal ods employed. Although plunging failure is not reached stress measurements. during these tests, all curves approach similar maximum (b) ó hm /qc ratios are independent of the distance behind the 9 loads of 2100 Æ 300 N after 2.5–3 mm pile head displace- pile tip: that is, at a given soil horizon, approximately ment. However, the initial stiffness of the load–displacement equal horizontal stress is recorded by each instrument response varies with the installation method. Both the mono- as it passes that point. tonic and jacked piles show an initial constant pile head stiffness of about 2000 N/mm, whereas the pseudo-dynamic The independence of ó hm /qc during monotonic installation 9 piles, which are load tested after ending installation with a from the distance h is emphasised in Fig. 8, which plots the 1.5 mm extraction stage, show a lower initial stiffness of ratio of the normalised horizontal stress recorded at h/B ¼ 1 1500 N/mm. The difference in stiffness of the monotonic (i.e. the instrument level closest to the pile tip) to that for and jacked piles compared with the pseudo-dynamic piles each of the other three instrument levels. No systematic must arise primarily from the inﬂuence of the ﬁnal stage of friction fatigue is evident between h/B ¼ 1 and 9. This FRICTION FATIGUE ON DISPLACEMENT PILES IN SAND 651 Pile-head load: N 0 500 1000 1500 2000 2500 0 Pile-head load at end of jacking stroke: N 0 500 1000 1500 2000 2500 0 20 T1 20 T4 40 T2 T5 T6 T10 40 T7 Pile-tip depth: mm 60 T1 1 Pile-tip depth: mm 60 80 80 100 100 120 120 140 140 (a) (b) Pile-head load at each end of ‘blow’: N 500 0 500 1000 1500 2000 2500 0 T3 20 T8 T9 T12 40 Pile-tip depth: mm 60 80 100 120 140 (c) Fig. 4. Pile-head load against tip depth during installation: (a) monotonic installations; (b) jacked installations; (c) pseudo-dynamic installations result conﬁrms the observation shown in Fig. 1(a) through incompatible with these observations. The observations con- direct measurement of the horizontal stress acting on a pile. trast with those of Klotz & Coop (2001) and Vesic (1970), The proposition that friction fatigue is associated with relief who show maximum ó hm values at between 5 and 10 9 from the highly stressed region close to the pile tip is diameters behind the pile tip. Fellenius & Altaee (1995) 652 WHITE AND LEHANE Pile-head load: N Horizontal stress,σ′hm: kPa 0 500 1000 1500 2000 2500 0 100 200 300 400 500 0 0 0·5 Pile-head settlement: mm T5 1·0 T1 T4 20 1·5 h/B 1 h/B 3 2·0 T10 h/B 6 2·5 40 h/B 9 3·0 Instrument depth: mm (a) 60 Pile-head load: N 0 500 1000 1500 2000 2500 0 0·5 Pile-head settlement: mm T2 80 1·0 T6 1·5 T7 T1 1 2·0 100 2·5 3·0 (b) 120 Fig. 6. Horizontal stress measurement during monotonic instal- Pile-head load: N lation (test T10) 0 500 1000 1500 2000 2500 0·0 0·5 Pile-head settlement: mm Normalised horizontal stress, σ′hm/qc 1·0 0 0·01 0·02 0·03 0·04 T3 0 T8 1·5 T12 2·0 T9 20 2·5 3·0 (c) 40 Instrument depth: mm Fig. 5. Pile-head load–settlement during compression load tests: (a) monotonic installations; (b) jacked installations; (c) pseudo- dynamic installations 60 suggest that residual loads, ignored during the interpretation, may have led Vesic to falsely conclude that the maximum value of unit shaft resistance occurs some distance above the 80 pile base, although Klotz & Coop’s observations cannot be explained in this way. h/B 1 HORIZONTAL STRESS DURING CYCLIC 100 h/B 3 INSTALLATION h/B 6 The jack stroke length was insufﬁcient to mobilise full friction during the pseudo-dynamic and jacked cyclic instal- lation methods. This became evident from the results of 120 static compression tests, which showed that lateral stresses and pile head loads continued to increase until the pile head Fig. 7. Normalised horizontal stress during monotonic installa- displacement reached between 5 and 8 mm. The ó hm data9 tion (mean of all four tests, error bars 1 std dev.) FRICTION FATIGUE ON DISPLACEMENT PILES IN SAND 653 Horizontal stress reduction: σ′hm, h nB /σ′ hm, h B stress, ó hc , recorded during each installation cycle. For 9 0 0·5 1·0 1·5 2·0 jacked installation, this corresponds to the value acting when 0 the pile is unloaded to nominally zero head load (actually 50 N). For pseudo-dynamic installation, this is the minimum value recorded during each cycle, and occurs close to the 10 moment of zero pile head load. The two unloading stages at pile tip penetrations of 60 and 120 mm provide two meas- 20 urements of ó hc for the monotonic installation method. 9 The proﬁles of ó hc with depth for each installation 9 method are shown in Fig. 9, grouped by instrument position 30 and averaged over four tests for each method. Friction fatigue is clearly evident in these data: that is, at a given depth, ó hc decreases as each instrument passes. The reduc- 9 40 tion in ó hc due to two-way cycling during installation is 9 demonstrated by the very low values recorded on the Depth: mm pseudo-dynamic piles compared with the monotonic installa- 50 tions. The upper half of the pseudo-dynamically installed piles has typically only 10% of the stationary horizontal stress of the monotonic case. The one-way cycling induced 60 by jacking leads to ó hc values that lie between those devel- 9 oped during pseudodynamic and monotonic installation. n 3 The reduction in ó hc at a given depth with increasing 9 70 distance from the pile tip (h) is quantiﬁed in Fig. 10. The n 6 measured values of ó hc have been normalised by qc and 9 80 averaged over the entire installation (the error bars show Æ1 n 9 standard deviation). Two methods of presenting these nor- malised data have been used. In Fig. 10(a), the decay in 90 ó hc /qc behind the pile tip is plotted against the distance h. 9 The high lateral stress close to the pile tip matches the trends observed in the ﬁeld (Lehane, 1992; Chow, 1997) but, 100 evidently, the data from the two installation methods do not overlie each other. As seen in Fig. 10(b), by plotting ó hc /qc 9 Fig. 8. Reduction in horizontal stress between instruments against number of cycles at that soil horizon, a better during monotonic installation (mean of all four tests, error agreement between the two sets of data is found. In a given bars 1 std dev.) soil horizon, the number of cycles (N) experienced at each lateral stress sensor position is simply 2h for the pseudo- dynamic installation since the net penetration is 0.5 mm per recorded during these installation methods cannot, therefore, cycle. For jacked installation, the value of N required to be compared directly with the ó hm measurements recorded 9 bring a given sensor to a speciﬁc soil horizon increases during monotonic installation (i.e. those in Figs 6 and 7). slightly with depth since the rebound during unloading in- Instead, the inﬂuence of installation method on shaft friction creases as the pile head load increases. The mean set per can be examined by considering the stationary horizontal jacking cycle is approximately 1.5 mm, and error bars show- Stationary horizontal stress, σ′hc: kPa Stationary horizontal stress, σ′ : kPa hc Stationary horizontal stress, σ′ : kPa hc 0 20 40 60 80 100 120 140 160 0 10 20 30 40 50 60 0 2 4 6 8 10 0 0 0 Jacked Jacked Jacked Pseudo- Pseudo- Pseudo- 20 dynamic 20 dynamic 20 dynamic Monotonic Monotonic Monotonic 40 40 40 Instrument depth: mm Instrument depth: mm Instrument depth: mm 60 60 60 80 80 80 100 100 100 120 120 120 (a) (b) (c) Fig. 9. Variation of stationary horizontal stress with installation method: (a) h/B 1; (b) h/B 3; (c) h/B 6 654 WHITE AND LEHANE 70 140 Jacked Jacked 60 120 N 108 Pseudo-dynamic Pseudo-dynamic Nmean 35 h/B 6 50 100 Distance above pile tip, h: mm Number of cycles, N 40 80 30 N 54 60 Nmean 20 h/B 3 h/B 6 20 40 Nmean 6 h/B 3 10 N 18 20 h/B 1 h/B 1 N: Number of cycles during installation Mean value shown for jacked installation 0 0 0 0·002 0·004 0·006 0·008 0 0·002 0·004 0·006 0·008 Normalised stationary horizontal stress, σ′ /qc hc Normalised stationary horizontal stress, σ′ /qc hc (a) (b) Fig. 10. Inﬂuence of loading cycles during installation on stationary horizontal stress: (a) normalised horizontal stress against distance above pile tip; (b) normalised horizontal stress against number of cycles ing the variation throughout installation are shown in Fig. of the loading type imposed during installation. All four 10(b). piles were ﬁrst subjected to one cycle of loading in the form A comparison of Fig. 10(a) with Fig. 10(b) indicates that of a static compression test to a pile settlement of 1.5 mm the variation of ó hc along the pile shaft is better related to N 9 before initiation of the scheduled component of the cyclic than to h (or h/B). A direct comparison is possible between tests. the instrument located at h/B ¼ 3 during jacked installation The variation of horizontal stresses at h/B ¼ 1 during and h/B ¼ 1 during pseudo-dynamic installation (Fig. 10(b)). selected cycles throughout tests T2 and T8 is presented in These instruments encounter 18 and 20 cycles respectively, Figs 11 and 12. The increase in stress during loading and register approximately equal normalised stationary hor- followed by a sharp reduction after changing direction is a izontal stresses. In this case, although jacked installation characteristic of interface shear under conditions of constant involves one-way cycling of the pile head load and pseudo- normal stiffness or constrained dilation. Comparable patterns dynamic installation involves two-way cycling, the degrada- of normal stress–shear displacement are widely observed in tion after around 20 cycles is comparable. It is shown later, interface shear box testing under constant normal stiffness from cyclic load tests, that over a greater number of cycles, (CNS) conditions (e.g. Airey et al., 1992; Fakharian & two-way head loading leads to greater degradation. It is also Evgin, 1997; DeJong et al., 2003). observed that one-way loading of the pile head leads to a Under one-way pile head loading there is some evidence degree of two-way loading along the pile shaft due to of two-way shear stress cycling at h/B ¼ 1 since the mini- rebound. mum horizontal stress does not coincide with the minimum pile head load. For interpretation of these cyclic load test data, ó hc is deﬁned as the minimum value within each cycle. 9 HORIZONTAL STRESS DURING CYCLIC LOAD This deﬁnition differs slightly from that used by Lehane TESTING (1992) and Chow (1997), who measure ó hc under zero head 9 The effects on shaft friction of cyclic loading during load, and is arguably more fundamental since it is the installation and a cyclic ‘working load’ may be compared by minimum value of ó hc within a particular loading cycle, 9 examining the horizontal stress measurements during the rather than the value under an arbitrary amount of residual cyclic load tests. Four tests will be highlighted (see Table shear stress. The reduction in ó hc due to cycling is clearly 9 2): evident in both Fig. 11 and Fig. 12. In the case of one-way cycling (of the pile head load), a limiting value of ó hc is 9 (a) the cyclic compression test following jacked installation reached after 30 cycles whereas, under more arduous two- of pile T2 way cycling, a progressive reduction towards zero is re- (b) the cyclic compression–tension test following pseudo- corded. dynamic installation of pile T8 The progressive reductions of ó hc observed at h/B ¼ 1 9 (c) the cyclic compression test following monotonic throughout the cyclic load tests are plotted in Fig. 13. It installation of pile T1 should be noted that if ó hc was deﬁned during the cyclic 9 (d) the cyclic compression–tension test following mono- load tests as the value under zero head load, a lower tonic installation of pile T4. degradation would be evident. Figs 11 and 12 show a small The cyclic tests on piles T2 and T8 represent an extension but increasing discrepancy between ó hc at Phead ¼ 0 and the 9 FRICTION FATIGUE ON DISPLACEMENT PILES IN SAND 655 160 Phead 750 N End of cycle 140 Cycle 1 Phead 50 N Start of cycle Phead 50 N 120 σ′ hc Cycle 5 Horizontal stress, σ′h: kPa 100 Cycle 30 Cycle 55 Cycle 100 80 60 40 20 0 119·6 119·7 119·8 119·9 120·0 120·1 120·2 120·3 120·4 120·5 120·6 Pile tip depth: mm Fig. 11. Horizontal stress degradation during one-way cyclic compression load test (test T2, jacked installation, h/B 1) 160 Phead 50 N Phead 750 N Start of cycle 140 Phead 0 N 120 σ′hc End of cycle Horizontal stress, σ′ : kPa 100 Cycle 1 Phead 0 N Cycle 5 h Cycle 30 80 Cycle 100 60 40 20 0 122·0 122·2 122·4 122·6 122·8 123·0 123·2 123·4 Pile tip depth: mm Fig. 12. Horizontal stress degradation during two-way cyclic compression-tension load test (Test T8, pseudo-dynamic installation, h/B 1) minimum value. Regardless of the chosen deﬁnition there is Hence the starting offsets of 11 and 18 cycles adopted for a reducing trend in ó hc with number of cycles. 9 the jacked and pseudo-dynamically installed piles (T2 and The one-way cyclic load tests on the monotonic and T8) enable direct comparisons with the monotonic base jacked installations (T1 and T2) are compared in Fig. 13(a), cases (T1 and T4). This representation leads to closely and the two-way cyclic load tests on the monotonic and comparable ó hc variations with N for tests T1 and T2 (in 9 pseudo-dynamic installations (T4 and T8) are compared in Fig. 13(a)) and for tests T4 and T8 (in Fig. 13(b)), suggest- Fig. 13(b). The number of cycles includes those experienced ing that the cycling is the only mechanism leading to during installation (and the extra cycle due to the subsequent degradation of ó hc . 9 static compression tests) in addition to the load test cycles. It is apparent in Fig. 13(a) that the relatively high values 656 WHITE AND LEHANE Normalised stationary horizontal stress, σ′ /qc hc Normalised stationary horizontal stress, σ′ /qc hc 0 0·002 0·004 0·006 0·008 0 0·002 0·004 0·006 0·008 140 140 T4 (monotonic) two-way load test (σ′hc) T1 (monotonic) one-way load test (σ′ ) hc T1 (jacked) one-way load test (σ′hc) (Fig. 11) T8 (Pseudo-dynamic) two-way load test (σ′hc) (Fig. 12) 120 Jacked installation (σ′hc/qc, Fig. 10) 120 Pseudo-dynamic installation (σ′hc/qc, Fig. 10) 100 100 Number of cycles, N Number of cycles, N 80 80 60 60 40 40 20 20 0 0 0 25 50 75 100 125 150 175 200 0 25 50 75 100 125 150 175 200 Stationary horizontal stress, σ′hc: kPa Stationary horizontal stress, σ′hc: kPa (a) (b) Fig. 13. Degradation of stationary horizontal stress with cycling at h/B 1 during load tests: (a) one-way compression load test; (b) two-way compression-tension load test of ó hc that exist at h/B ¼ 1 after a low number of cycles 9 the number of cycles and their amplitude on friction fatigue. reduce to a constant value of about 50 kPa after 30 one-way Such sensitivity was also observed by Kelly (2001), who cycles. A similar degradation of ó hc with number of cycles 9 showed in CNS interface shear tests that larger-amplitude is apparent in Fig. 13(b) for two-way cycling, and ó hc 9 cycles lead to a higher rate of degradation. continues to reduce beyond 30 cycles to a value close to zero. This reduction in ó hc does not, however, mean that the 9 available shaft friction is negligible. For example, it may be DISCUSSION inferred from Fig. 12 that dilation at the interface can lead Implications for design to an increase in normal stress at ultimate conditions to well The key observation from this investigation is that friction in excess of 100 kPa. This surprising recovery of normal fatigue arises from cycles of loading. During continuous stress is also observed in CNS interface shear testing (e.g. penetration there is no reduction in horizontal stress at a Airey et al., 1992; Shahrour et al., 1999). given soil horizon as the pile penetrates deeper: that is, The agreement seen in Fig. 13 between the degradation of lateral stresses are independent of distance, h. Friction ó hc during each cyclic load test, when modiﬁed to account 9 fatigue cannot therefore be attributed to the departure of a for the cycles induced during installation, supports the link zone of high stress around the pile tip as the pile penetrates between friction distribution and cyclic history. To examine deeper. Instead, it has been found that the reduction in this link more closely, in addition to the two cyclic load test stationary horizontal stress at a given soil horizon is better curves shown in each of Figs 13(a) and 13(b), the normal- linked to the cyclic history. On-pile measurements of hor- ised stationary horizontal stresses, ó hc /qc , recorded at 9 izontal stress have shown trends of behaviour that agree with h/B ¼ 1, 3 and 6 during jacked and pseudo-dynamic instal- observations during constant normal stiffness (CNS) cyclic lation (i.e. as shown in Fig. 10) are also plotted in Figs interface shear testing (e.g. Tabucanon, 1997; Kelly, 2001; 13(a) and 13(b). There is reasonable agreement between the DeJong et al., 2003). A mechanism of net contraction with installation data and the cyclic load test data, supporting the cycling of a thin interface layer that is conﬁned by the far link between cycling and friction fatigue, which is indepen- ﬁeld soil explains the behaviour observed in both cases. dent of the distance h. For the pseudo-dynamic installation, These results suggest that, for design, an appropriate non- in which the cyclic load test amplitude (Fig. 12) is compar- dimensional quantity to govern the reduction in horizontal able to the amplitude of the installation cycles, the agree- effective stress from an initial ‘unfatigued’ value behind the ment is better than that seen for jacked installation, possibly pile tip with continued pile penetration is number of cycles, because the jacked installation procedure involves larger N, rather than h/D. This observation is supported by the cycles of displacement than that induced in the one-way similar relationship between ó hc /qc and N during both 9 cyclic tests. This discrepancy highlights the inﬂuence of both installation and load testing. However, the centrifuge tests FRICTION FATIGUE ON DISPLACEMENT PILES IN SAND 657 support inferences made from interface shear tests by de- qc: MPa Jacking force: kN monstrating that the rate of degradation depends not only on 0 20 40 0 2500 5000 N, but also on the mode and amplitude of cycling. Two-way 0 cycling leads to a greater degradation than one-way cycling during both installation and load testing. Predrilled Although this investigation has demonstrated that friction 1 for CPT fatigue is not linked per se to h or to the normalised distance, h/D, some dependence on diameter may remain, Sand since the conﬁning stiffness, which governs the reduction in 2 ó hc for a given contraction of the interface layer, depends on 9 Fine to 4G/D (G being the operational shear stiffness of the soil medium surrounding the installed pile). 3 The relative success of the Jardine & Chow (1996) design method, which relates friction fatigue to h/D, is likely to be primarily because the chosen power law for the h/R effect 4 reﬂects the degradation in friction caused by contraction of Depth: m the interface during the number of installation cycles of typical driven piles (forming the database used to validate 5 Jardine & Chow’s design method), under the corresponding conﬁning stiffness. The examples highlighted in Fig. 1 warn against extrapolation beyond this range. The sharp reduc- 6 tions in ó hc with cycling seen during this investigation may 9 be attenuated by the lower conﬁning stiffness (/ 1/D) around ﬁeld piles; for a given interface contraction a lower drop in ó h would result. However, the mechanism of behav- 9 7 iour is scale-independent, and cycling would still lead to friction fatigue. Finally, it is noted that this paper does not address the 8 inﬂuence of cycling, and of ó hc reduction, on the subsequent 9 increase in horizontal effective stress observed during static loading to failure. Any increase in ó h during loading 9 9 contributes an additional component of normal stress at failure and hence increases ôf . Therefore any reduction in Fig. 14. Jacking record for precast concrete pile in sand ó hc may not cause a proportional reduction in ôf . However, 9 CNS interface shear box tests, which can be considered analogous to elements of a pile–soil interface, show that any pleting installation with far fewer cycles than dynamic cyclic contraction and loss of normal stress is not fully methods. Furthermore, optimisation of dynamic methods can recovered during subsequent shearing to failure (Ghionna et reduce the total blowcount during installation. It should be al., 2004), so any drop in ó hc will have an inﬂuence on ôf . 9 noted, however, that any additional friction on a jacked pile may degrade more quickly under a cyclic working load than on a ‘pre-degraded’ dynamically installed pile. Implications for construction Considerable further research is needed in order to capture This investigation has demonstrated that the stationary this behaviour in a prediction method. However, since site- horizontal stress acting on a pile shaft is strongly inﬂuenced speciﬁc pile load testing is usually required for major by the number of installation cycles. Although some recov- projects, the value of these ﬁndings lies equally in the ery in lateral stress is possible during subsequent loading suggestion that optimisation of the pile installation process (e.g. Figs 11 and 12), signiﬁcant load cycles reduce the can yield higher shaft friction. available pile shaft friction. This behaviour is illustrated in Fig. 14, which plots the jacking record of a 350 mm square precast concrete pile installed using a jacking machine of CONCLUSIONS the type described by Lehane et al. (2003). The site con- Drum centrifuge tests have been conducted to examine sisted of medium dense sand with a CPT qc value of more closely the distribution of horizontal stress acting on 20 MPa between depths of 3 and 9 m. For experimental the pile shaft during installation and subsequent cyclic purposes, the contractor jacked the pile in 1.5 m increments loading. During monotonic installation no friction fatigue to a depth of 6.3 m and then completed the installation to a was recorded, which is in agreement with ﬁeld data from ﬁnal pile tip depth of 8 m using 22 no. 75 mm jacking multi-sleeve CPT installation. The variation of available increments. It can be assumed that the base resistance shaft friction with depth followed the CPT proﬁle, indicating (following the qc proﬁle) was unchanged between 6.3 m and that the normalisation of lateral stresses by qc provides a 8 m, and therefore that the mobilisation of a constant jacking useful basis for design. resistance over this depth interval, as seen in Fig. 14, was Cyclic installation methods have been seen to cause sig- due to friction fatigue induced by cyclic loading. Clearly, in niﬁcant degradation of shaft friction. One-way and two-way this example the additional time spent during installation installation methods lead to differing proﬁles of stationary from 6.3 m to 8 m and the additional cost of pile length was 9 lateral effective stress (ó hc ), when plotted against the relative entirely negated by the additional loading cycles. position of the pile tip, h. The centrifuge cyclic load tests This link between the cycles during installation and the also showed that, for a given installation method, the degra- shaft friction during ﬁrst loading suggests that opportunities dation of ó hc during cyclic loading followed the same decay 9 may exist for improving pile capacity, and increasing design pattern as that during installation. These measurements agree efﬁciency, if the loading cycles during installation are mini- with trends indicated in CNS interface shear tests. A mised. Pile-jacking machines offer the possibility of com- mechanism of net contraction with cycling of a thin inter- 658 WHITE AND LEHANE face layer that is conﬁned by the far ﬁeld soil explains the came into being and why it does not exist. Proc. Inst. Civ. behaviour observed in both cases, and provides a rational Engrs Geotech. Engng 113, No. 2, 107–111. basis for improved design. Fugro (1996). EURIPIDES database report, Vols 1–5. Leidschen- Based on these centrifuge test data, and corroboratory dam, The Netherlands: Fugro BV. evidence from ﬁeld-scale test results, it is concluded that the Ghionna, V. H., Mortara, G. & Vita, G. P. (2004). Sand-structure interface behaviour under cyclic loading from constant normal degradation of available shaft friction at a given soil horizon stiffness direct shear tests. Proceedings of the international during installation and subsequent cyclic loading is better symposium on deformation characteristics of geomaterials, characterised by the number of cycles experienced at that Lyon, pp. 231–238. point, than by the non-dimensional distance from the pile Heerema, E. P. (1980). Predicting pile driveability: heather as tip, h/D. This conclusion has implications for construction, an illustration of the friction fatigue theory. Ground Engng 13, as well as for design. Modern installation techniques of pile 15–37. jacking involve reduced cycling, and may therefore yield Jardine, R. J. & Chow, F. C. (1996). New design methods for higher shaft friction than conventional dynamic installation offshore piles, MTD96/103. London: Marine Technology Direc- methods. torate. Kelly, R. (2001). Development of a large diameter ring shear apparatus and its use for interface testing. PhD thesis, Univer- sity of Sydney. ACKNOWLEDGEMENTS Klotz, E. U. & Coop, M. R. (2001). An investigation of the effect The support provided by the Australian Research Council ´ of soil state on the capacity of driven piles in sands. Geotech- (ARC) for this research project is gratefully acknowledged. nique 51, No. 9, 733–751. Lehane, B. M. (1992). Experimental investigations of pile behaviour The ARC Centre for Offshore Foundation Systems at UWA using instrumented ﬁeld piles. PhD thesis, Imperial College, also provided support. The authors also acknowledge the University of London. excellent technical assistance provided by Mr Bart Thomp- Lehane, B. M. & Jardine, R. J. (1994). Shaft capacity of driven son and the staff of the UWA civil engineering workshop. piles in sand: a new design method. Proc. 7th Int. Conf. on the We also thank Dr Peter Mitchell (formerly of PPK Interna- Behaviour of Offshore Structures, Boston 1, 23–36. tional) for his permission to present the data shown in Fig. Lehane, B. M., Jardine, R. J, Bond, A. J. & Frank, R. (1993). 14. Mechanisms of shaft friction in sand from instrumented pile tests. ASCE J. Geotech. Engng 119, No. GT1, 19–35. Lehane, B. M., Pennington, D. & Clark, S. (2003). Jacked end- bearing piles in the soft alluvial sediments of Perth. Aust. REFERENCES Geomech. J. 38, No. 3, 123–134. Airey, D. W., Al-Douri, R. & Poulos, H. G. (1992). Estimation of Lunne, T. & Christofferson, H. P. (1983). Cone penetrometer pile friction degradation from shearbox tests. ASTM Geotech. interpretation for offshore sands. Proc. Offshore Technology Test. J. 15, No. 4, 388–392. Conf., Houston, OTC4464, 181–192. API (2000). RP2A: Recommended practice of planning, designing Lunne, T., Robertson, P. K. & Powell, J. J. M. (1997). Cone and constructing ﬁxed offshore platforms: Working stress design, penetration testing in geotechnical practice. London: Blackie 21st edn, pp. 59–61. Washington: American Petroleum Institute. Academic & Professional. Bustamante, M. & Gianeselli, L. (1982). Pile bearing capacity Poulos, H. G., Carter, J. P. & Small, J. C. (2001). Foundations and prediction by means of static penetrometer CPT. Proc. 2nd Eur. retaining structures: research and practice. Proc. 15th Int. Conf. Symp. on Penetration Testing, Amsterdam, 1, 493–500. Soil Mech. Found. Engng, Istanbul 4, 2527–2606. Chow, F. C. (1997). Investigations into the behaviour of displace- Randolph, M. F., Dolwin, J. & Beck, R. (1994). Design of driven ment piles for offshore foundations. PhD thesis, Imperial Col- ´ piles in sand. Geotechnique 44, No. 3, 427–448. lege, University of London. Randolph, M. F., Leong, E. C. & Houlsby, G. T. (1991). One- DeJong, J. T. (2001). Investigation of particulate-continuum inter- ´ dimensional analysis of soil plugs in pipe piles. Geotechnique face mechanisms and their assessment through a multi-friction 41, No. 4, 587–598 sleeve penetrometer attachment. PhD thesis, Georgia Institute of Shahrour, I., Rezaie, F. & Nauroy, J.-F. (1999). Experimental study Technology, Atlanta, GA. of the behaviour of calcareous sand: structure interface. Proc. DeJong, J. T. & Frost J. D. (2002). A multi-friction sleeve attach- 2nd Int. Conf. on Engineering for Calcareous Sediments, ment for the cone penetrometer. ASTM Geotech. Test. J. 25, No. Bahrain, 69–77. 2, 111–127. Stewart, D. P., Boyle, R. S. & Randolph, M. F. (1998). Experience DeJong, J. T., Frost, J. D. & Cargill, P. E. (2001). Effect of surface with a new drum centrifuge. Proc. Int. Conf. Centrifuge ’98, texturing on CPT friction sleeve measurements. ASCE J. Geo- Tokyo, 1, 35–40. tech. Geoenviron. Engng 127, No. 2, 158–168. Tabucanon, J. T. (1997). Shaft resistance of piles in sand. PhD DeJong, J. T., Randolph, M. F. & White, D. J. (2003). Interface load dissertation, University of Sydney, Australia. transfer degradation during cyclic loading: a microscale investi- Vesic, A. S. (1970). Tests on instrumented piles, Ogeechee River gation. Soils Found. 43, No. 4, 81–93. site. J. Soil Mech. Found. Div. ASCE 96, No. SM2, 561–584 Fakharian, K. & Evgin, E. (1997). Cyclic simple-shear behaviour of White, D. J., Finlay, T. C. R., Bolton, M. D. & Bearss, G. (2002). sand-steel interfaces under constant normal stiffness condition. Press-in piling: ground vibration and noise during pile installa- ASCE J. Geotech. Geoenviron. Engng 123, No. 12, 1096–1105. tion. Proc. Int. Deep Foundations Congress, Orlando, ASCE Fellenius, B. H. & Altaee, A. A. (1995). Critical depth: how it Special Publication 116, 363–371.
Pages to are hidden for
"Friction fatigue on displacement piles in sand"Please download to view full document