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Cọc lái xe trong cát có thể thấy tăng đáng kể trong axial năng lực trục của họ trong những tháng tiếp theo cài đặt. Nhiều lợi ích thiết thực làm theo nếu dịch vụ năng lực có thể được dựa vào để vượt quá mức đã được chứng minh trong các thử nghiệm trang web, thường được thực hiện trong vòng một vài ngày của lái xe. Bài báo này báo cáo kết luận từ một Chương trình chủ yếu là các bài kiểm tra căng thẳng trên các cọc ống thép thực hiện trong cát dày đặc tại Dunkirk, miền bắc nước Pháp. Các cuộc thử nghiệm đã chứng minh năng lực trục đánh dấu nhiều hơn tăng trưởng với thời gian hơn dự kiến. Các cọc niên cũng cho thấy chế độ thất bại đáng ngạc nhiên dễ gãy; kiểm tra trước khi thất bại cả hai khả năng suy thoái và biến đổi sự lão hóa quá trình, dẫn đến không đơn điệu trục công suất thời gian dấu vết mà rơi xa dưới đặc trưng lão hóa còn nguyên vẹn (IAC) được xác định bằng cách kiểm tra trên tươi, unfailed trước, cọc. Xu hướng công suất thời gian suy ra từ thử nghiệm nhiều lần cọc có thể cho kết quả sai lệch. Những phát hiện mới cho phép các tác động lão hóa được đặc trưng rõ ràng hơn, và cho phép đánh giá lại cơ sở dữ liệu hiện có liên quan đến cọc các loại điều khiển trong một phạm vi của bùn, cát và sỏi. Kết luận quan trọng thực tiễn được rút ra về thiết kế cọc và việc giải thích của đống trường kiểm tra tải
See discussions, stats, and author profiles for this publication at: http://www.researchgate.net/publication/245411222 Time-related increases in the shaft capacities of driven piles in sand ARTICLE in GÉOTECHNIQUE · JANUARY 2001 Impact Factor: 1.67 · DOI: 10.1680/geot.51.5.475.39977 CITATIONS DOWNLOADS VIEWS 20 69 92 4 AUTHORS, INCLUDING: Richard Jardine Jean-Franỗois Nauroy Imperial College London IFP Energies nouvelles 103 PUBLICATIONS 1,367 CITATIONS 24 PUBLICATIONS 170 CITATIONS SEE PROFILE SEE PROFILE Available from: Richard Jardine Retrieved on: 10 August 2015 ´ 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 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 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 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 lAtlantique Gravelines Power Station Grand Fort-Philippe 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 CS 11.7 m long piles LL LS 22.4 m long piles P1 P2 cone penetration P1 P2 tests BH1 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 10 m BRE In situ tests CPT cone penetration test SC seismic cone test DMT dilatometer test DPH ü ý dynamic penetration tests DPM ỵ Fig 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 400 260 2400 240 220 T¢89a T¢89b 20 40 Cumulative pile head displacement: mm 60 Tension T¢94 2800 Faster rate of displacement Fig Load–displacement curves for 11 m CS pile from CLAROM research programme (Chow, 1997) 5·0 4·5 Qt(t)/Qt (t 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 0·1 10 100 Time after driving: days (a) 1000 10 000 5·0 4·5 QS(t)/QS (t 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 0·1 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 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 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 ˚ 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 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 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 shows the typical site profile for the CLAROM/ Imperial College test area, and Fig 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 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 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 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 mm 2000 kN at 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 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 10 20 Borehole log CPT fc: kPa 100 200 40 300 400 Very dense, light brown, uniform, fine to medium, subrounded SAND with occasional shell fragments (Hydraulic fill) GWL Dense with shell fragments (Flandrian Sand) 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 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: 0·801·12 m C6: 3·513·73 m C8: 5·195·55 m C9: 6·206·35 m C15: 11·511·9 m CLAROM envelope of results 30 20 10 0·01 0·1 Particle size: mm R3 10 A JP1 Fig 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 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 R6 B PS.R12 R1 A R5 R2 Institut Pasteur PS.GP1A PS.R56 PS.GP1B PS.C1 C1 B PS.R45 R4 12 m Fig 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 AA GP1.B qc: MPa 20 40 R2-3 qc: MPa 20 40 R5-6 qc: MPa 20 40 z (m) 10 15 R1-2 qc: MPa 20 40 Section BB C1 qc: MPa 20 40 R4-5 qc: MPa 20 40 10 z (m) 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 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 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, 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 R1 R2 R3 R4 R5 R6 Hydraulic fill Flandrian Sands Organic layer Interbedded Horizontal scale 5m Fig CPT profiles with depth and interpreted soil profile (see Fig for locations of sections A–A and B–B) Penetration: m 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 20 100 200 300 Normalised blow count: blows/m 400 Fig 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 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 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 914 419 5·3 m long I-beam 400 mm 500 t hydraulic jack and load cell 400 mm 914 mm 914 mm m by m ground panel Levelled sand 1400 mm Displacement and reference system 200 mm 500 mm 6m (a) Tension head Loadspreading beams 914 419 5·3 m long I-beam 6m Displacement and reference system 6m (b) Fig Details of rig used for testing reaction piles in tension (not to scale): (a) elevation; (b) plan cycling experiments required different configurations The m long reaction beam transferred compressive loads to the ground through steel frames that sat on steel spreader mats located at least 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 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 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 creep rates exceeded 0.01 mm/min, then to 20 kN, and finally to 10 kN if the 30 creep rates exceeded 0.08 mm/ 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 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 gives the key features of the four ‘first-time’ static tests to failure on piles R1 to R6, C1, and Table 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 gives details of static retests that were performed on all of the piles listed in Table 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 not account for pile self-weight Table Retests on Dunkirk reaction piles and compression pile C1 Pile Test (days after driving) R1 Tension failure 1500 kN (57 days) at mm 1700 kN Cyclic failure, then at 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 mm 821 kN at 33 mm Test (days after driving) Maximum static load and displacement Tension failure (239 days) N/A 1646 kN at mm N/A Tension failure (250 days) Test 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 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 (2) ụf ẳ 0:90:8ú rc ỵ ú rd Þ tan äf (c) The pre-loading radial effective stress is 0:13 À0:38 ó v0 h ó rc ¼ 0:029qc Pa Rà (3) (d) ó v0 is the free-field vertical effective stress, and Pa , the atmospheric pressure, is introduced to make the term à non-dimensional; R is found from the pile’s inner and outer radii as À Á0:5 à (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 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 summarises the ICP predictions, taking account of the actual pile lengths as detailed in Table The capacities The equivalent expression applying to compressive loading is ơf ¼ (ó rc þ ˜ó r )tan äf 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 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 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 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 R31st test 13/11/1998 (85 days)* R41st test 16/11/1998 (85 days)* R51st test 19/11/1998 (90 days)* R61st 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 10 15 20 25 Pile head displacement: mm 30 3500 Jardine & Chow (1996) trendline 2500 2000 1500 1000 500 0 10 15 20 25 Pile head displacement: mm 30 10 100 1000 Time after driving: days 10000 35 R11st test 02/09/1998 (9 days) R21st test 17/04/1999 (235 days) R61st test 09/11/1998 (81 days) 3000 CLAROM piles LL and LS 22 m long (restruck) R1 1·0 0·5 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 headtension positive: kN Intact ageing characteristic (IAC) R2 2·5 QS(t)/QSICP Force applied to pile headtension 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 (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 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 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 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) 0·5 239 10 2·0 R3(1) & R4(1) 1·5 R4(2) 1·0 100 1000 Time after driving: days 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 C1(1): virgin path for C1 (end point proven by first testin 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 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 ? Dunkirk IAC CS(1) 0·5 Line representing capacity at end of driving (b) 3·0 QS(t)/QSICP 3·0 CS(2) Jardine & Chow (1996) trendline 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 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 headtension positive: kN Force applied to pile headtension 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 3500 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% 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 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 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 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 5·0 4·5 2·5 2·0 0·1 10000 QS(t)/QS(EOID) 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 0·1 10 100 Time after driving: days (ii) (a) 1000 0·1 10000 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 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 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 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 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 0 R23 R56 GP1A GP1B R23 R56 GP1A GP1B 5 10 10 10 Depth: m Depth: m Depth: m R23 R56 GP1A GP1B 15 15 15 20 20 20 25 10 20 30 40 50 Cone end resistance, qc: MPa 25 60 25 100 200 300 400 500 600 700 Sleeve friction, fS: kPa (a) 0 Friction ratio: % R12 R45 C1 R12 R45 C1 R12 R45 C1 5 10 10 10 Depth: m Depth: m Depth: m 15 15 15 20 20 20 25 10 20 30 40 50 Cone end resistance, qc: MPa 60 25 100 200 300 400 500 600 700 Sleeve friction, fS: kPa (b) 25 Friction ratio: % 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 REFERENCES ˚ Astedt, B., Weiner, L & Holm, G (1992) Increase in bearing capacity with time for friction 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