Centrifuge modeling of single pile subjected to compression and tension in clay

CENTRIFUGE MODELING OF SINGLE PILE SUBJECTED TO COMPRESSION AND TENSION IN CLAY JIRASAK ARUNMONGKOL (B.Eng.(Hons.), CU) A THESIS SUBMITTED FOR THE DEGREE OF MASTER OF ENGINEERING DEPARTMENT OF CIVIL ENGINEERING NATIONAL UNIVERSITY OF SINGAPORE 2004 Acknowledgements ACKNOWLEDGEMENTS Firstly, I would like to express my deepest gratitude to my supervisors, Associate Professor Leung Chun Fai and Professor Chow Yean Khow for their advice, constant guidance and support throughout this research program. Thanks for their valuable time and efforts in shaping the framework of this thesis. Grateful thanks are also due to The National University of Singapore (NUS), for providing the research scholarship from July 2000 to July 2002 to conduct this research program and finance to laboratory research expenses. Without the funding, this research program would not have been accomplished. I would like to take this opportunity to thank Laboratory Professional officer Shen Rui Fu, Dr Retnamony Gnanaselvam Robinson and all the other Geotechnical Centrifuge Laboratory staffs, especially Mr Wong Chew Yuen and Mr Tan Lye Heng for giving useful advice, troubleshooting and solving technical problems. Further thanks to Mr. Foo Hee Ann, Mr. Loo Leong Huat, Mdm. Joyce Ang and Mdm. Jamilah for their assistance in fabricating model pile, sending out quotation forms and ordering equipment and transducers. Finally, I am grateful to my colleagues such as research assistants and research scholars in the Soft Ground Centre and Centrifuge Laboratory for their assistance, friendship and some invaluable support. i TABLE OF CONTENTS Acknowledgements i Table of contents ii Summary vi List of Tables viii List of Figures ix Chapter 1 Introduction 1.1 Background 1 1.2 Scope of research 4 Chapter 2 Literature review 2.1 Introduction 5 2.2 Behaviour of pile in cla y 6 2.2.1 Initial condition prior to pile installation 7 2.2.2 Pile installation effects and disturbances 8 2.2.3 Re-consolidation of clay around pile after 13 installation 2.2.4 Pile loading 2.3 Pile load test 14 15 2.3.1 Conventional pile load test 16 2.3.2 Osterberg pile load test 17 2.4 Estimation of pile bearing capacity in clay 2.4.1 Side resistance 18 19 2.4.1.1 Total stress approach, α method 19 2.4.1.2 Effective stress approach, β method 20 2.4.1.3 Mix approaches 21 2.4.2 Base resistance 2.5 Geotechnical centrifuge modelling 23 23 ii 2.5.1 Principle of centrifuge modelling 23 2.5.2 Centrifuge model tests on axially- loaded piles 25 2.6 Summary 28 Chapter 3 Experimental setup and material properties 3.1 Introduction 40 3.2 NUS geotechnical centrifuge 40 3.3 Experimental setup 42 3.3.1 Setup for pile load tests 42 3.3.2 Setup for base suction pile load test and in- flight 44 miniature cone penetration test 3.4 Transducers 45 3.4.1 Load cell 45 3.4.2 Linear potentiometer (POT) 46 3.4.3 Miniature pore pressure transducer (PPT) 46 3.5 Instrumented model piles 47 3.5.1 Strain gauge- instrumented model pile 47 3.5.1.1 Fabrication of model pile 47 3.5.1.2 Instrumentation of model pile 49 3.5.2 PPT-instrumented model pile 50 3.6 Clay sample preparation 51 3.7 Sign convention 53 3.8 In- flight miniature cone penetration test 54 3.9 Properties of clay 55 Chapter 4 Pile load test: effect of different loading methods 4.1 Introduction 71 4.2 Soil responses prior to pile installation 72 4.3 Pile installation 73 4.4 Dissipation of excess pore water pressures after 75 pile installation 4.5 Axial load transfer of the pile during and after 77 pile installation iii 4.6 Compression pile load test C1A 82 4.6.1 Results and discussion 82 4.6.2 Comparison with static pile design method 85 4.7 Ultimate tension pile load test T1B 87 4.8 O-cell pile load test T1C 91 4.9 Effects of different loading me thods on pile 93 bearing capacity 4.10 Base suction capacity test, test series D 4.10.1 Responses of pile and pore water pressure during 100 101 installation 4.10.2 Experimental procedure and results 4.11 Comparison between behaviour of piles installed in 101 103 sand and clay 4.12 Conclusion 105 Chapter 5 Pile load test: effects of loading history and maintained load 5.1 Introduc tion 136 5.2 Effect of loading history on pile performance 137 5.2.1 Test series A 137 5.2.2 Test series B 142 5.2.3 Test series C 145 5.2.4 Conclusion on the effect of loading history 147 5.3 5.4 Effect of maintained load on pile performance 149 5.3.1 Test series E 149 5.3.2 Test series F 153 5.3.3 Conclusion on the effect of maintained load 156 Comparison of the effect of loading history on single 157 pile installed in sand and clay Chapter 6 Conclusions 6.1 Conclusion remarks 179 6.1.1 179 Ground responses during and after pile installation iv 6.1.2 Residual stress 180 6.1.3 Compression between compression and uplift 181 behaviour of piles 6.1.4 Compression between conventional and 182 Osterberg pile load tests 6.1.5 Effect of loading history on pile behaviour 183 6.1.6 Effect of maintained load on the behaviour of 184 piles in clay 6.1.7 6.2 Behaviour of suction around tension piles in clay Recommendations on future study 184 185 References 186 Appendix 195 v Summary SUMMARY Chapter 5 This thesis describes an experimental investigation into the behaviour of pile subjected to compression and tension in clay. The investigation was carried out using the National University of Singapore geotechnical centrifuge. Several series of centrifuge model tests were carried out : (a) to compare the behaviour of a single pile, especially side resistance, when subjected to conventional compression, conventional tension and Osterberg pile load tests; (b) to investigate the effect of pile loading history and loading sequence on the pile performance; and (c) to investigate the effect of maintained dead load on pile behaviour in long term. The results indicate that positive excess pore water pressures are generated in clay during pile installation. A maximum excess pore water pressure as high as 4 times undrained shear strength of the clay can be generated at the pile base. Immediately after the completion of pile installation, significant residual stresses along the pile are not observed due to relatively low base resistance of the pile in soft to medium clay. However, increases in axial load along the pile due to dragged-down loads along the pile, which are caused by the consolidation of the clay after first compression loading, can be observed. A significant difference in the distribution of unit side resistance with depth between compression and tension piles in clay is observed in the present study. As compared to compression piles, the ultimate unit side resistance of tension piles is much lower around the upper part of the pile, but increases rapidly with depth until it is higher than that of compression piles at the pile base. As a result, the side resistance of tension piles is lower than that of compression piles. This can be attributed to the effect of different direction of pile loading. vi Summary It is found that the ultimate side resistance mobilized along the Osterberg piles is much lower than that of conventional compression piles. The side resistance characteristic of Osterberg piles is similar to that of conventional tension piles. However, the side resistance of Osterberg piles is slightly lower than that of conventional tension piles due to the effect of different load transfer mechanisms. The history of pile loading is found to affect the pile behaviour. When there is dissipation of positive excess pore water pressures before pile loading, the pile bearing capacity and load-displacement stiffness of the pile increase due to soil setup. However, the side capacity and load-displacement stiffness of the pile are found to deteriorate when the pile is loaded in compression-tension cycles. Finally, the pile under maintained dead load in long term is found to have a higher pile bearing capacity than that under short-term static pile load test. The gain in pile bearing capacity is caused by the soil set- up which is induced the dissipation of excess pore water pressures generated by the dead load. Keywords: pile load test, tension pile, Osterberg pile load test, centrifuge, residual stress, soil set- up, pile loading history vii LIST OF TABLES Table 2.1 Scaling relationship for geotechnical centrifuge modelling (After Leung et al., 1991) Table 3.1 Properties of Malaysian kaolin clay Table 3.2 Position of pore pressure transducers in clay Table 4.1 Summary of centrifuge model tests Table 4.2 Summary of pile load distribution at various stages during and after pile installation for test series A Table 4.3 Summary of pile load distribution at various stages during and after pile installation for test series A. (after pile weight adjustment) Table 4.4 Summary of results for load test C1A, T1B and T1C Table 4.5 Summary of results of test series D Table 5.1 Summary of observations on experimental results of test series A Table 5.2 Summary of experimental results for load test series A Table 5.3 Summary of observations on experimental results of test series B Table 5.4 Summary of experimental results for load test series B Table 5.5 Summary of observations on experimental results of test series C Table 5.6 Summary of experimental results for load test series C Table 5.7 Summary of observations on experimental results of test series E Table 5.8 Summary of experimental results for load test series E Table 5.9 Summary of observations on experimental results of test series F Table 5.10 Summary of experimental results for load test series F viii LIST OF FIGURES Figure 2.1 Ko of normally consolidated clay vs (a) friction angles (b) plasticity Index (After Ladd, 1977) Figure 2.2 Experimental results from CKo U direct simple shear tests on six clays by Ladd and Edgers (1972). (a) Undrained shear strength vs OCR (b) Relative increase in undrained shear strength ratio with OCR (After Ladd et al., 1977) Figure 2.3 Soil movement due to pile installation (After Randolph et al., 1982) Figure 2.4 Schematic diagram shows modeling of pile installation by using expanding cylindrical cavity theory in undrained soil with MohrCoulomb characteristics (After O’Neill, 2001) Figure 2.5 Stress distributions in clay around pile immediately after driving (a) OCR = 1 (b) OCR = 8 (After Randolph et al., 1979) Figure 2.6 Effect of residual loads after installation on the interpretation of the results of pile load tests. (After Briaud and Tucker, 1984) Figure 2.7 Time effect on undrained shear strength of clay around a displacement pile (a) variation of normalized undrained shear strength with time factor after driving at 1.15 pile radius from the pile center (b) variation of normalized shear strength at the end of consolidation with normalized radial distance. (After Randolph et al., 1979) ix Figure 2.8 Conventional pile load test setup using kentledge reaction. (After ASTM D1143) Figure 2.9 Load transfer mechanism for piles (After Das, 1998) Figure 2.10 Osterberg load test setup for bored piles. (Taken from the leaflet of Loadtest Inc.) Figure 2.11 Variation of adhesion factor α of piles with an embedded length less than 50 m in clay. (Adapted from Terzaghi et al., 1996) Figure 2.12 Correlation factor λ. (After Vijayvergiya & Focht, 1972) Figure 2.13 Design parameters α p and F proposed by Semple and Rigden (1984). (After Tomlinson, 1995) Figure 3.1 NUS geotechnical centrifuge Figure 3.2 Centrifuge model setup Figure 3.3 Setup for pile load test Figure 3.4 Schematic diagram of closed- loop servo-controlled actuator Figure 3.5 Setup for base suction capacity test and in- flight miniature cone penetration test Figure 3.6 Calibration for load cell SML-200 Figure 3.7 Schematic diagram of components of model pile Figure 3.8 Schematic diagram showing loading mechanism of conventional model pile subjected to (a) compression load at pile head (b) tension load at pile head. Figure 3.9 Schematic diagram showing loading mechanism of model O-cell pile when subjected to (a) tension load (b) compression load Figure 3.10 Location of strain gauges on model pile x Figure 3.11 Strain gauge instrumentation on pile (a) 4 gauge arrangement bridge circuitry (b) arrangement of strain gauges on pile surface. Figure 3.12 Schematic diagram of PPT- instrumented model pile Figure 3.13 Position of pore pressure transducers in clay Figure 3.14 Sign convention adopted in present study Figure 3.15 Schematic diagram showing structure of miniature cone penetrometer Figure 3.16 Physical properties of clay sample: (a) water content (b) bulk unit weight Figure 3.17 Shear strength properties of clay at 100g Figure 4.1 Pore water pressure responses registered by various PPTs in the clay during self- weight consolidation Figure 4.2 Average degree of self-weight soil consolidation with time after the centrifuge reaches 100g Figure 4.3 Pile installation process Figure 4.4 (a) Pile resistance-penetration response and (b) excess pore water pressure-pile penetration responses during pile installation of test series A Figure 4.5 Variation of hydrostatic pressure with time at 100g Figure 4.6 Variations of normalized excess pore water pressures with log time after the release of installation load for test series A Figure 4.7 Load distribution curves along the pile during and after pile installation for test series A Figure 4.8 Load distribution curves along the pile during and after pile installation for test series A. (after adjustment of pile weight) xi Figure 4.9 (a) Load-displacement curve (b) Excess pore water pressure responses during pile load test C1A Figure 4.10 Variations of normalized excess pore water pressure at depth of about 16.5 m with normalized radial distance for test C1A Figure 4.11 Variations of normalized excess pore water pressures with log time after the release of test load for test C1A Figure 4.12 Comparison between adjusted axial load distribution curves before and after linear interpolation of strain gauge readings with time of test C1A Figure 4.13 Compression of observed unit side resistance with depth at ultimate load of test C1A compared with prediction based on Randolph and Murphy (1985) method Figure 4.14 Variations of adhesion factor (α) with overconsolidation ratio (OCR) for load test C1A compared with that predicted by Randolph and Murphy (1985) method Figure 4.15 Excess pore water pressure-pile penetration responses during pile installation of test series B Figure 4.16 Variations of normalized excess pore water pressures with log time after the release of installation load for test series B Figure 4.17 (a) Load-displacement curve (b) Excess pore water pressure responses during pile load test T1B Figure 4.18 Variations of excess pore water pressure with ratio of radial distance to pile diameter at various applied loads of test T1B (at a depth of about 14.0 m) Figure 4.19 Variations of excess pore water pressures with time after the release of test load for load test T1B xii Figure 4.20 Adjusted axial load distribution curves along the pile during and after load test T1B Figure 4.21 Load distribution curves along pile during and after the release of installation load for test series C Figure 4.22 Load distribution curves along the pile at various applied loads during and after pile installation for test series C. (after adjustment of pile weight) Figure 4.23(a) Load-displacement curve (b) Excess pore water pressure responses during pile load test T1C Figure 4.24 Variations of excess pore water pressures with time after the release of test load for load test T1C Figure 4.25 Adjusted axial load distribution curves along the pile during and after load test T1C Figure 4.26 Load-displacement curves of tests C1A, T1B and T1C Figure 4.27 Soil resistance-displacement curves of tests C1A, T1B and T1C Figure 4.28 Variations of ultimate unit side resistance with depth at ultimate load for load tests C1A, T1B and T1C Figure 4.29 Stress path of a soil element close to pile wall since installation of displacement pile until loading in compression to failure for overconsolidated clay (After Ortigao, 1995) Figure 4.30 Idealized development of negative side resistance along pile when loaded at different locations Figure 4.31 Load-settlement curve and pore water responses during and after installation of pile of test series D xiii Figure 4.32 Variations of applied load and base suction force with time for load tests (a) T1D, (b) T2D and (c) T3D Figure 4.33 Variations of normalized suction pressure at the pile base (∆Uult /Cubase) with normalized pile displacement (δ/d) for tests T1D, T2D and T3D Figure 4.34 Load-displacement curves of test AC1, AT1, and JT1 performed in sand in the previous study (Adapted from Goh, 2000) Figure 4.35 Load distribution curves along the pile for test AC1 (Adapted from Goh (2000)) Figure 4.36 Load distribution curves along the pile for test AT1 (Adapted from Goh (2000)) Figure 4.37 Load distribution curves along the pile for test JT1 (Adapted from Goh (2000)) Figure 4.38 Variations of ultimate unit side resistance with depth at ultimate load for tests AC1, AT1 and JT1 performed in dry sand in previous study (Adapted from Goh, 2000) Figure 5.1 Load-relative pile displacement responses for load test series A. Figure 5.2 Load-displacement responses for load test series A Figure 5.3 Variations of unit side resistance with depth at the end of consolidation for test series A Figure 5.4 Schematic diagram illustrating the effect of existing side resistance before tension load test on load-pile displacement stiffness Figure 5.5 Adjusted axial load distribution curves of the pile at ultimate load for load test series A Figure 5.6 Variations of unit side resistance with depth at ultimate load for load test series A xiv Figure 5.7 Load-relative pile displacement responses for load test series B Figure 5.8 Variations of unit side resistance with depth at the end of consolidation for test series B Figure 5.9 Load-displacement responses for load test series B Figure 5.10 Adjusted axial load distribution curves of the pile at ultimate load for load test series B Figure 5.11 Variations of unit side resistance with depth at ultimate load for load test series B Figure 5.12 Load-relative pile displacement responses for load test series C Figure 5.13 Load-pile displacement responses for load test series C Figure 5.14 Variations of unit side resistance with depth at the end of consolidation for test series C Figure 5.15 Adjusted axial load distribution curves of the pile at ultimate load for load test series C Figure 5.16 Variations of unit side resistance with depth at ultimate load for load test series C Figure 5.17 Load-relative pile displacement responses for load test series E Figure 5.18 Load-displacement responses for load test series E Figure 5.19 Variations of unit side resistance with depth at the end of consolidation for test series E Figure 5.20 Adjusted axial load distribution curves of the pile at ultimate load for load test series E Figure 5.21 Variations of unit side resistance with depth at ultimate load for load test series E compared with the results of test series A Figure 5.22 Load-relative pile displacement curves for load test series F xv Figure 5.23 Load-displacement curves for load test series F Figure 5.24 Variations of unit side resistance with depth at the end of consolidation for test series F Figure 5.25 Adjusted axial load distribution curves of the pile at ultimate load for load test series F Figure 5.26 Variations of unit side resistance with depth at ultimate load for load test series F xvi Chapter 1 Introduction INTRODUCTION CHAPTER ONE Chapter 3 1.1 Background Piles are columnar structural load carrying members to facilitate the transfer of loads from the superstructures through weak compressible soil strata to stiff soil strata. The main materials used for piles are timber, steel and concrete. The piles can be driven, jacked or drilled into the ground and connected with a pile cap. In addition to carrying vertical compression loads, they may be required to resist uplift forces due to hydrostatic pressure or other loads. When used to support tall structures subjected to overturning forces due to wind or waves, piles may have to resist both compression and uplift forces but at different loading sequences depending on the direction of the overturning forces. Piles used to support marine structures, retaining walls, bridge piers and abutments and machinery foundations are required to resist combined vertical and horizontal loads. It is widely accepted that a pile transfers its load into its surround ing soil through two mechanisms. The first is the transfer of the load through the friction and adhesion along the pile shaft-soil interface. The rest of the load is then transferred to the pile base. Many attempts have been made to reliably predict the magnitude of these soil resistances. Unfortunately, owing to the complicated mechanism of pile-soil interaction, no theory is available to accurately predict the pile behaviour. The current pile design practice is still mainly based on empirical methods whose design parameters are often obtained from field pile load tests. 1 Chapter 1 Introduction Static pile load test on instrumented piles is the most direct way to evaluate the behaviour of piles at a site. The test involves in applying test load at the pile head at a certain rate against a reaction system which usually consists of dead weights on a kentledge or reaction piles. However, in certain circumstances, performing a pile load test may be cumbersome, hazardous or uneconomical, for example when a congested and inaccessible area is encountered or when a huge reaction system is required for a pile having very large capacity. To overcome such difficulties, Osterberg (1989) introduced an alternative pile load test method which uses a bellow- like device called ‘Osterberg cell’ or ‘O-cell’. The device was initially implemented with driven piles and subsequently with bored piles. The Osterberg cell, which is a bellow having top and bottom plates slightly smaller than the pile diameter, is installed at the bottom of a driven pile before installation or at the bottom of a bored pile before placing concrete. The load test is conducted by pumping a hydraulic pressure into the cell. This creates equivalent upward and downward forces to the pile shaft and pile base, respectively. As a result, the pile shaft moves upward and negative side resistance is mobilized along the shaft. Meanwhile, pile base resistance is mobilized due to an equivalent downward force acting at the top of the bottom plate. Based on the assumption that upward and downward side resistances of a pile are identical, an equivalent top-down loaddisplacement curve of the pile can be constructed from the load-displacement responses of the shaft and the bottom plate. The use of Osterberg cell becomes increasingly popular around the world as its implementation has been reported from many sites around the world, see for example Osterberg (1990), Schmertmann et. al. (1998), Fellenius et. al. (1999). The assumption of upward ultimate side resistance equal to ultimate downward shear resistance was investigated by field tests made by 2 Chapter 1 Introduction Ogura et. al. (1996). The piles were pushed upward with the O-cell and then pushed down after the O-cell was depressurized. It was found that the magnitudes of ultimate side resistance in both directions were the same. However, since the tests were performed consecutively on the same pile, the effect of preceding load test on subsequent load test may inherently be included in the results. Besides the costly expense of conducting full-scale pile load tests, the uncontrollable site condition during the test, for example fluctuation of ground water level, can cause difficulties and sometimes incompleteness of the test to study the pile behaviour. Physical modeling of pile foundation in laboratory is thus an alternative and economical way for the investigation of pile behaviour. Furthermore, a major benefit of physical modeling is the controllability of testing environment. As a result, the effect of each particular factor on the pile and soil behaviour can be distinctly investigated. However, the behaviour of soils is stress-dependent. Thus, the results obtained from a small model test may not represent prototype behaviour. Such limitation can be overcome by subjected to a forced acceleration field in the centrifuge such that the prototype stress levels can be reproduced. The scaling laws, which relate the results from the model test to the corresponding results in prototype scale, are presented by many researchers, see for example Leung et al. (1991). According to the scaling laws, the diffusion time for excess pore water pressure can be reduced by a factor of 1/N 2 , where N is number of times the Earth’s gravitational acceleration. Thus, the study of pile behaviour in clay in long term can be achieved within much shorter time than that of a full- scale test. The reliable and consistent results can be achieved by simulating the prototype condition and preparing the model carefully. Thus, centrifuge model 3 Chapter 1 Introduction testing is a robust tool in geotechnical engineering to study the behaviour of pile-soil interaction with lower cost than full- scale load testing. 1.2 Scope of research The research work discussed in this thesis is the extension of the work done by Goh (2000) who studied the behaviour of single pile in sand. It covers the investigation of the pile behaviour in clay as follows: Ø To study the behaviour of a single pile when subjected to compression and tension load and its surrounding clay using centrifuge modeling. Ø To compare the behaviour of side resistance of a pile subjected to tension load at the pile base (similar to Osterberg pile load test setup) with that of a pile subjected to conventional compression and tension loads at the pile head. Ø To study the effect of preceding loads on the pile in subsequent loading. 4 Chapter 2 Literature review LITERATURE R EVIEW CHAPTER TWO 4 2.1 Introduction A pile is a long slender structural member employed to transfer load from superstructure to a firm/stiff stratum beneath the ground surface. Pile foundation can be classified into many categories according to: Ø Method of installation: bored/augered, continuous augered, driven, screw- in, jacked- in, jetting, etc. Ø Degree of displacement: full-displacement, partial-displacement, non- displacement Ø Nature of load transfer: friction and end-bearing Ø Pile material: concrete, steel, timber, composite Ø Shape of pile section: rectangular, circular, hollow, H-pile, octagonal, triangular piles etc. Ø Types of ground that the piles are installed in: cohesionless, cohesive, and mixed soil types Ø Manufacture: in-situ, precast or combined Ø Support during pile installation: no support, temporary casing, permanent casing, drilling mud, soil/concrete/grout Ø Enlarged base: with or without From engineering point of view, the load bearing capacity and loaddisplacement characteristics of pile are the major concerns. Tomlinson (1970) reported that soil stratification, pile material and shape, and time after pile driving can affect the load bearing capacity of piles. Meyerhof (1976) stated that the behaviour of 5 Chapter 2 Literature review pile-soil interaction is not only dependent on the nature of soil, pile dimensions, pile layout, but also the method of installation. Azzouz et al. (1990) summarized that the complicated behaviour of pile-soil interaction is dependent on site conditions (stratigraphy, soil properties, water table etc.), pile characteristics (diameter, length, material, surface roughness etc.), installation methods (closed versus open-ended, time history of driving versus jacking), and loading conditions (set-up time after driving, rate of axial loading etc.). In this chapter, the literatures relating to the behaviour of a pile in clay, especially displacement pile, is reviewed. Since pile performance can be greatly affected by the installation method of the pile, the literatures on the behaviour of pile and its surrounding clay prior to, during and after pile installation are firstly reviewed. The widely- used prediction methods of pile bearing capacity are then reported, especially on the methods for predicting the ultimate side resistance. To verify actual pile performance, full-scale static pile load tests are usually conducted. Thus, various methods for conducting full- scale static pile load tests are presented followed by the review of previous centrifuge modeling studies on pile foundation. 2.2 Behaviour of pile in clay The performance of pile is mainly affected by the characteristic s of the surrounding soils. For a pile installed in clay, its responses can be divided into 4 stages as follows: 1) Initial condition of clay prior to pile installation, 2) Pile installation effects and disturbances 3) Re-consolidation of clay around pile after installation, 4) Pile loading 6 Chapter 2 Literature review 2.2.1 Initial condition prior to pile installation It is well known that the stress-strain and strength characteristics of soils can be explained using the well kno wn concept of effective stress as firstly proposed by Terzaghi (1923). The initial state of effective stress and stress history of saturated cohesive soils under geostatic state of stress can be expressed by 3 parameters which are the in-situ vertical effective stress, σ’vo , the coefficient of lateral earth pressure at rest, Ko , and the overconsolidation ratio, OCR. As σ’vo = σvo - uo , where σvo = total vertical stress and uo = pore water pressure, it can be determined with the highest degree of confidence among the three parameters. Ko is usually estimated from the correlation with soil parameters obtained from standard laboratory tests. Ladd et al. (1977) compiled the data of Ko measured in laboratory of normally consolidated remolded and undisturbed clays plotted with effective friction angle of soil, φ’, and plasticity index, PI. Figure 2.1 reveals that Ko of normally consolidated clay is approximately equal to 1-sin φ’. Brooker and Ireland (1965), and Campanella and Vaid (1972) reported that the change of Ko with consolidation pressure is generally small. For overconsolidated clay, Schmidt (1966) proposed an empirical equation: K o ,OC K o , NC = OCR m (Eq. 2.1) where Ko,NC, Ko,OC = the coefficient of lateral earth pressure at rest of normally consolidated and overconsolidated clays, respectively. OCR = overconsolidation ratio = σ’p / σ’vo σ’p = maximum past vertical effective stress to which the soil element has been subjected m = 1.2 sin φ’ 7 Chapter 2 Literature review Equation 2.1 clearly shows the influence of stress history on the in-situ horizontal effective stress. Mayne and Kulhawy (1982) also proposed a similar relationship except m = sin φ’. Note that this relationship is valid only with overconsolidated clay due to monotonic unloading. Besides having the influence on the in situ effective horizontal stress, stress history also plays an important role on the stress-strain-strength properties of cohesive soils. Based on the results of CKo U direct simple shear tests on six clays by Ladd and Edgers (1972) as shown in Figure 2.2, Ladd and Foott (1974) proposed a reasonably well-defined expression as follow: C ( u ) OC σ 'vc = OCR m Cu ( ) NC σ 'vc (Eq. 2.2) where Cu = undrained shear strength of clays obtained from CKo U direct simple shear tests σ’vc = vertical consolidation stress of a soil element m = 0.8, better fit can obtained if m is decreased from 0.85 to 0.75 with increasing OCR for this set of data Further results of plane strain and triaxial tests on the same soils also shows the relative increase of Cu/σ’vc with OCR similar to the data obtained from direct simple shear tests. 2.2.2 Pile installation effects and disturbances When a pile is driven or jacked into cohesive soil deposits, there is minimal migration of pore water pressure within surrounding soil mass due to low permeability of soils. Hence, during installation, the volume of soil mass displaced 8 Chapter 2 Literature review must be equal to the volume of pile in undrained condition. The early research conducted by Housel and Burkey (1948) and Cummings et al. (1950) suggested that pile driving can cause extensive distortion and fully remo lding to the soil mass adjacent to the pile shaft. However, minor effects were observed in the soil at approximately 2 times pile diameter from the pile. A similar observation was also reported by Tomlinson (1970). The soil specimens taken after pile driving revealed that during pile penetration, the surrounding soil had been carried down with the pile shaft extending to a short distance away from the pile. This implies that a zone of soil around the pile has been intensely sheared and remolded to failure. The observation of vertical ground movement s during pile jacking by Cooke (1979) confirms the observation made by Tomlinson (1970). Besides the disturbance of soil during installation, Cooke and Price (1973), Randolph et al. (1979), Hwang et al. (2001), and Pestana et al. (2002) reported the existence of outward radial deformation of soil around the pile. According to the data of extensive instrumented field measurement of driven piles, Hwang et al. (2001) concluded that lateral displacement of ground caused by pile driving decreases with increasing distance from the pile. Hagerty and Peck (1971) and Bozozuk et al. (1978) observed ground heave during pile driving. Cooke and Price (1973) mentioned that at a penetration up to about 10 times the radius of pile, heaving of ground surface could occur. At greater depths, the soil moves predominantly in radial direction. It is well known that the behaviour of soil mass is controlled by its state of effective stress. In order to apply the effective stress concept to pile foundation, the initial and change of state of effective stress around the pile shaft must be fully understood. Thus, many researchers paid efforts on monitoring total radial stresses and pore water responses of soil mass during and after pile installation; see for 9 Chapter 2 Literature review example Bjerrum and Johannessen (1960), Soderman and Milligan (1961), D’Appolonia and Lambe (1971), Azzouz and Morrison (1988), Eigenbrod and Issigonis (1996), Hwang et al. (2001), Pestana et al. (2002). Generally, the above researchers reported that significant excess pore pressures are generated during pile installation. The maximum excess pore water pressure is at the pile shaft and decreases as a function of invert of square root of radius away from the pile-soil interface. Moreover, the generated excess pore water pressures in a zone close to the pile-soil interface were reported to exceed the effective overburden pressure (Hwang et al., 2001, and Pestana et al., 2002). Figure 2.3 shows an idealized schematic diagram of soil movement due to pile installation. Based on field observations, the displaced soils predominantly move outward in radial direction for most pile length. Initial attempts to model the installation of pile analytically were thus based on the expansion of cylindrical cavity with the final radius equal to the radius of pile (Soderberg, 1962, Butterfield and Banerjee, 1970, Vesic, 1972, Leifer et al., 1979, Randolph et al., 1979, Matlock et al., 1982, Heydinger, 1982, Hydinger and O’Neill, 1986). Esrig et al. (1977) made a major step in understanding the effect of soil parameters on side resistance by proposing an analytical model termed “general effective stress method” (GESM). The modeling of pile installation process is based on undrained plain strain cylindrical cavity expansion theory in an idealized elastic-plastic soil. Further development of this model was done by Kirby and Esrig (1979) to enhance the capability of coping with a general stress-strain relationship. This model can provide an estimation of the state of stress and pore water pressures developed in soil mass dur ing and after pile installation Figure 2.4 presents a schematic diagram showing the idealization of pile installation using cylindrical cavity expansion theory by Vesic (1972). However, 10 Chapter 2 Literature review GESM tends to overestimate the unit side resistance as compared to field data, as pointed out by O’Neill (2001). Some other models which have different attributes from GESM were also deve loped and applied to cylindrical cavity expansion theory; see for example, CAMFE (Randolph et al., 1979), CASH (Matlock et al., 1982), and VECONS/AXIPLN (Heydinger, 1982, Heydinger and O’Neill, 1986). Figure 2.5 shows the results of an analytical solution based on cylindrical cavity expansion theory conducted by Randolph et al. (1979). Immediately after pile installation, positive excess pore water pressures are generated with the maximum value at the pile-soil interface. In addition, the clay at the pile-soil interface has the highest effective stress in the radial direction, σ’r, no matter the clay is normally consolidated (OCR = 1) or overconsolidated (OCR = 8). Nevertheless, the assumption of all these methods mentioned above is based on undrained plain strain condition. There is not treatment for the heave at ground surface and soil movement around the advancing pile tip. Moreover, the variation of undrained Young’s modulus, Eu, in the yield plastic region must be assumed. A significant better prediction was proposed by Azzouz et al. (1990) by coupling strain path method (Ba ligh, 1985) with MIT-E3 constitutive model (Whittle, 1987). In this method, there is no need to assume the variation of Eu in the yield region. Furthermore, the end condition at the pile base is also taken into account. Recently, Whittle (1999) improved the strain path model to include the effect of free surface. Whitaker and Cooke (1966) reported that upon unloading of a test load, each pile shaft remained in compression under the reaction of a residual load at the pile base and a balancing negative side resistance along the pile shaft. Hunter and Davission (1969) conducted a full pile test program in medium sand including compression, tension, and lateral load tests on instrumented steel pipe and H-piles. 11 Chapter 2 Literature review From the test results, they observed large apparent tension forces at the pile base when compressive loading tests were followed by tension tests. Since the presence of tension forces at the base in sand was not possible, they concluded that unexpected residual compression forces had not been taken into account. Cooke (1979) presented the results of a jacked instrumented steel tube pile in clay having horizontal inclinometer instrumented at different elevations around the pile. The results suggested that residual forces in displacement piles arise because of differences in the rates of mobilization of resistance on the shaft and the base as a pile is displaced, either during installation or under compressive loading. During unloading, the soil under pile base tries to push the pile back up while the pile decompresses. If the base resistance is high enough, the rebound could create enough upward movement to mobilize downward pile-soil resistance along the shaft. Finally, equilibrium is reached when the mobilized downward resistance is equal to the pushing force at the pile base. Hunter and Davisson (1969) reported that residual loads at the pile base could be as high as 80 percent of anticipated base load. However, residual base loads were negligible if the piles were installed by a vibratory hammer. Since the vibration of the hammer is effectively minimizing side resistance during installation, inadequate negative side resistance could be mobilized to oppose the residual compression load at the base. They further reported that ignoring residual loads after installation could cause serious errors in the division of load between side and base resistances. However, the total load bearing capacity could not be affected. Figure 2.6 illustrates that ignoring residual loads after pile installation could lead to overestimation of ultimate side resistance and underestimation of ultimate base resistance for compression piles and vice versa for tension piles. From the results of model tests of long extensively instrumented piles in deep beds of sand, Hanna and Tan (1973) 12 Chapter 2 Literature review concluded that the shape and magnitude of the load-settlement curve of the piles were influenced by residual stresses and thus the behaviour of the piles is dependent on its previous load history. This conclusion is supported by the analytical solutions performed by Poulos (1987). According to the analyzes, the load-settlement stiffness of the pile head of a driven or jacked- in pile is generally larger in compression than in tension due to the effect of residual stresses. The residual stresses and their effects are most significant for piles in sand, but do not significantly influence the response of piles in clay. 2.2.3 Re-consolidation of clay around pile after installation Cummings, Kerkhoff and Peck (1950) measured the water content of soft clay at various radii from a large cluster of piles over a 1- year period after pile driving. It was found that the re was horizontal migration of pore water initiated by driving. The change in water content was roughly constant with depth and varied with time. From a test pile in California, Seed and Reese (1957) showed that the bearing capacity of a pile increased whereas the water content in the clay decreased. Field measurements made by Holtz and Lowitz (1965), and Fellenius and Samson (1976) suggested a decrease of undrained shear strength of clay within 1.5-2.0 times pile diameter immediately after pile installation and the decrease is largely recovered at the end of consolidation. These observations suggest that once the installation of a displacement pile is complete, the built- up excess pore water pressures around the pile will gradually dissipate, allowing the surrounding clay to consolidate. Consequently, the effective stresses around the pile, especially lateral effective stresses, increase with time after driving as supported by field measurement obtained by Soares and Dias 13 Chapter 2 Literature review (1989). The increase in lateral effective stresses results in an increase of side resistance and this effect phenomenon is commonly called as ‘set- up’ effect. The distribution of excess pore water pressures around a driven pile has been investigated by many researchers (Bjerrum and Johannessen, 1960, Lo and Stermac, 1965, Koizumi and Ito, 1967, etc.). It is found that if a pile is very long compared with its diameter, the excess pore water pressures would generally dissipate in a radial direction, except at the ground surface and pile base where there is also vertical flow at these regions. Lehane and Jardine (1994) found that the dissipation of excess pore water pressures at the pile base is faster than those around the pile shaft due to threedimensional drainage path near the base. With regards to these observations, most numerical analyses model the consolidation after cavity expansion in only radial direction. Figure 2.7 shows the results of a numerical analysis conducted by Randolph et al. (1979) which modeled pile installation as the undrained expansion of a cylindrical cavity in work-hardening elasto-plastic soil. Figure 2.7(a) shows that the undrained shear strength of surrounding clay increases with time after driving. At the end of consolidation, the undrained shear strength is highest at the pile shaft and reduces with increasing radial distance from the pile as shown in Figure 2.7(b). Randolph et al. (1979) showed that the prediction agrees well with field measurements made by Seed and Reese (1957) and Eide et al. (1961). 2.2.4 Pile loading For a vertical rigid cylindrical friction pile with a very large embedment ratio, the effect of ground surface and pile base on the side resistance is negligible. Thus, the soil elements adjacent to the pile surface are subjected to only straining due to axial loading. Thus, many investigators (Esrig et al., 1977, Butterfield and Banerjee, 14 Chapter 2 Literature review 1970, Azzouz et al., 1990, etc.) studied the effect of pile loading analytically by assuming that the soil elements are subjected to pure shear. Direct measurement by Esrig et al. (1977) has suggested that the excess pore water pressures generated at the pile-soil interface during pile loading are small, typically 0-25% of the undrained shear strength of clay. Esrig and Kirby (1979) proposed that such low value is possible due to the presence of residual stresses. Beyond unloading, the stress paths of soil elements adjacent to pile shaft move inwards from the yield surface to the elastic region. Since no excess pore water pressure is generated due to pure shear in a soil behaving elastic behaviour, only excess pore water pressure generated after soil yielded could be obtained during pile loading. Azzouz et al. (1990) suggested that the shear surfaces of failure where slippage occurs can be either soil-pile or soil-soil slippages depending on the relative magnitude of the roughness of pile surface and the size of soil particles. The soil-soil slippage could occur if the pile is rough enough resulting in higher peak-shearing resistance. However, the skin friction ratio, which is the ultimate side resistance over effective horizontal stress acting on the shaft after full consolidation of soil caused by installation, is not sensitive to in-situ overconsolidation ratio of the soil. Based on the results of a series of pile load tests of intensively instrumented closed-ended steel pile installed in heavily overconsolidated London clay, Bond and Jardine (1995) proved that the shaft capacity of displacement piles in London clay is governed by an effective stress interface sliding criterion, not by the failure of soil continuum. 2.3 Pile load test In practice, static pile load test is usually conducted to determine the load bearing capacity and the load-settlement responses of a pile. The test is also used to 15 Chapter 2 Literature review evaluate whether the designer’s estimation of pile capacity and length is appropriate. The current commonly used methods for pile load tests are summarized below. 2.3.1 Conventional pile load test The equipment and procedures for conducting pile load tests have been developed and refined for many years and set out as standard specification such as ASTM D1143 and ASTM D3689. The conventional or Slow Maintained Load (SML) test is currently widely used procedure. In a SML test, the static load is applied in increments, usually by hydraulic jack via a load cell. The settlement of the pile head is measured by dial gauges mounted with a reference beam. A new load increment is applied once the rate of pile settlement is very small. As the expected bearing capacity is approached, the size of load increment is reduced to obtain a more accurate indication of bearing capacity. In principle, the ultimate load bearing capacity of a pile is defined as a load for which rapid movement occurs under maintained or slight increase of applied load, i.e. the pile plunges. However, in most circumstances, such distinct plunging is not always obtained. Several mathematical procedures were thus proposed to estimate the ultimate bearing capacity. Some of the widely used methods are summarized by Fellenius (1980). According to specifications, the SML test requires several days to complete and hence may not be economical. The duration of the load test can be reduced by conducting either the Constant Rate of Penetration (CRP) or Quick Maintained Load (QML) tests. In CRP test, the pile is jacked down at a constant rate of penetration until no further increase in load is registered. In QML test, the pile is loaded at intervals of 2.5 minutes in increments of about 15% of design load until further penetration is required to maintain the applied load. To perform a conventional load 16 Chapter 2 Literature review test, the load is applied by jacking against a reaction system. The reaction system can be either dead weights on kentledge or anchor piles. However, if the pile capacity is very large, very large dead weights or size of anchor piles are required. This may lead to safety hazards and a cumbersome test. Figure 2.8 illustrates the setup for a conventional pile load test using the kentledge reaction system. The load transfer mechanism from a pile to supporting soil can be investigated during a pile load test by instrument ing either telltales or strain gauges along the pile shaft. The axial loads along the pile can then be obtained from the measured local displacements or strains at various elevations along the pile. To obtain the true separation of side and base resistance, all strain gauges should be zeroed prior to installation. Figure 2.9 shows the load transfer mechanism for a pile under test load. A pile loaded at the pile head with load Q(z=0) has side (Q 1 ) and end (Q 2 ) resistances as shown in Figure 2.9(a). The development of axial load distribution along the shaft is shown in Figure 2.9(b). From the figure, the applied load is gradually transferred to the supporting soil along the side of the pile. The difference in axial loads in the pile between any two elevations is thus the side resistance, ∆Q(s) . By dividing the side resistance with corresponding circumferential area of the pile (p), the unit side resistance along the given pile shaft segment (∆Q(s) /(p.z)) can be obtained. At the pile base, the rest of axial load is taken by the soil beneath the base (end bearing resistance, Q2 ). Similarly, the unit end bearing resistance can be obtained by normalizing Q2 with pile base area, Ap . 2.3.2 Osterberg pile load test Osterberg (1989) proposed a new method of static pile load test which can overcome some disadvantages of conventional pile load test. The method can be 17 Chapter 2 Literature review applicable to both precast and bored piles. In the test, the pile to be tested is installed with an ‘Osterberg load cell’ or ‘O-cell’, which is a hydraulic jack or bellow-like device, as part of the pile to the desired depth prior to pile installation. The load test is conducted by pumping fluid into the cell. The cell is then activated to push the lower part of the pile downward whereas the upper part is pushed upward with the equivalent applied load. The upper and lower parts of the pile act as a reaction to each other. The forces and movements at the pile head and base are measured independently. The load increments are applied following the QML method until either the upper or lower part of the pile reaches ultimate condition, whichever is earlier. Thus, the load-displacement responses of the pile shaft and base can be obtained separately. The equivalent top-down equivalent load-displacement responses can be obtained. By adding the side and base resistances at the same measured movement, a point on the equivalent top-down load-displacement curve can be obtained. Figure 2.10 illustrates the setup of Osterberg pile load test for bored piles. 2.4 Estimation of pile bearing capacity in clay In general, the bearing capacity of a pile is the sum of the pile base resis tance and the total side resistance around the pile shaft. When no strong end-bearing soil layer exists, the base resistance of a pile is relatively small and the major pile resistance is derived from the side resistance. Therefore, the primary concern of estimating the bearing capacity of a floating in clay is to estimate the ultimate unit side resistance. 18 Chapter 2 Literature review 2.4.1 Side resistance 2.4.1.1 Total stress approach, α method In this approach, the ultimate unit side resistance is correlated with the in-situ undrained shear strength of the clay. Different correlations have been proposed by various researchers (Tomlinson, 1957; Peck, 1958; Woodward et al., 1961; Kerisel, 1965; etc.). The ultimate unit side resistance, fs, is given by f s = αCu (Eq. 2.3) where α is adhesion factor and Cu is the undrained shear strength of clay. Figure 2.11 shows the correlation between α with average undrained shear strengt h based on the field measurements from Dennis and Olson (1983), and Stas and Kulhawy (1984). The results of driven and bored piles in compression and tension are plotted in the figure which shows that the adhesion factor decreases with increasing average undrained shear strength. Since the effects of pile installation, time effect, pile geometry and length are collapsed into only one factor, a poor correlation is noted in Figure 2.11. This may lead to the conclusion that the unit side resistance of compression piles equals to that of tension piles. Olson (1984) analyzed eleven cases of compression and tension pile load tests in clay from 5 different sites in USA by using the API (American Petroleum Institute) design guideline. In 10 out of the 11 cases, the ratio of calculated/measured ultimate side resistance in tension is higher than that in compression. In 9 out of 10 cases, the tension tests followed the compression tests. The results indicate that the assumption of unit side resistance in tension equal to compression overpredicts the tension capacity by 18%. If the pile base resistance is insignificant and the side resistance does not reduce in the second test, it may be possible that the reversal of stresses reduces the ultimate side resistance for the first cycle of reversal, the ultimate side 19 Chapter 2 Literature review resistance in tension is lower than that in compression, or other factors. Although the method does not consider the state of effective stress around the pile which governs the pile-soil behaviour, it continues to be commonly used due to the ease of obtaining soil parameter. 2.4.1.2 Effective stress approach, β method The effective stress concept was firstly applied to calculate the side resistance by Zeevaert (1959). The method correlates ultimate unit side resistance, fs, with initial effective overburden stress, σ’vo , through a β factor. The ultimate unit side resistance can be estimated by f s = K c σ ' vo tan φ ' = βσ ' vo (Eq. 2.4) where Kc is the coefficient of lateral earth pressure after excess pore water pressure generated by pile driving had fully dissipated and φ’ is the effective stress friction angle. The main difficulty in applying the effective stress approach is estimating the radial effective stress on the pile at failure. By assuming that (1) cohesion intercept of remolding soil after pile driving is zero (2) horizontal effective stress after dissipation of excess pore water pressure generated by pile driving, σ’hc, is at least equal to initial horizontal effective stress, σ’ho , and (3) excess pore water pressures ge nerated by shear distortion during pile loading can dissipate rapidly, Burland (1973) proposed that Kc is approximately equal to Ko . For normally consolidated clay, Ko is roughly equal to (1-sin φ’). Based on laboratory or field tests, Meyerhof (1976) suggested that Ko is roughly equal to (1-sin φ’) OCR0.5. For driven pile in overconsolidated clay, Kc is about 1.5Ko according to field pile load tests in London clay. 20 Chapter 2 Literature review 2.4.1.3 Mixed approaches The ultimate unit side resistance estimated from these approaches is correlated to both Cu and σ’vo . Based on data of 42 pile load tests, Vijayvergiya and Focht (1972) correlated the ultimate unit side resistance to the average effective overburden stress and undrained shear strength over the embedded length of driven piles using an empirically determined factor λ. f s = λ (σ ' vo +2Cu ) (Eq. 2.5) where the variation of correlation factor λ plotted with depth is shown in Figure 2.12. Thus, the method is so-called the ‘λ method’. From the comparison between predicted and measured ultimate unit side resistance of pile load tests, Flatte and Selnes (1977) reported that the λ method generally overpredicts the ultimate unit side resistance with poor correlation. Esrig and Kirby (1979) pointed out that this is because most of pile load tests reported by Flatte and Selnes (1979) were conducted in normally consolidated or lightly overconsolidated clay with pile length less than 50 ft. In fact, the data points having pile penetration less than 50 ft in Figure 2.12 were obtained from pile load tests of driven piles in heavily overconsolidated clay whereas piles having penetration longer than 50 ft were performed in normally consolidated clay. Esrig and Kirby (1979) observed that the decrease of λ factor in normally consolidated clay may indicate the decrease of ultimate side resistance with increasing pile penetration. The observation is similar to the data showing the decrease of β factor with pile penetration presented by Meyerhof (1976). Poulos (1982) presented a theoretical study showing that this is due to the progressive failure along the pile-soil interface due to pile length effect. Based on the results of driven pile load tests from American Petroleum Institute (API) database, Semple and Rigden (1984) correlated the average ultimate 21 Chapter 2 Literature review unit side resistance to the average shear strength by taking into account the effect of ambient effective stress level and embedded length of pile. The recommended equation for estimating average unit side resistance at failure is f s = Fα p C u (Eq. 2.6) where F = length correction factor (a function of the pile aspect ratio, l/d), α p = peak value of α (a function of the strength ratio, Cu/σ’v ). Figure 2.13 presents the charts for αp and F. From the figure, the strength ratio of normally consolidated of 0.35 together with minimum α p of 0.5 were assumed. The length effect is treated by using l/d rather than pile flexibility. An alternative correlation of the same database was proposed by Randolph and Murphy (1985). The method makes use of the simplicity of the α method by correlating the strength ratio with fs,max/Cu which is in turn the α factor. The ultimate side resistance at any depth along the pile can be estimated from C f s = ( u ) 0nc. 5 C u0.5σ '0v .5 σ 'v for Cu ≤1 σ 'v (Eq. 2.7) C f s = ( u ) 0nc. 5 Cu0. 75σ ' 0v. 25 σ 'v for Cu >1 σ 'v (Eq. 2.8) where the value of (C u /σ’v )nc is the strength ratio for remoulded normally consolidated clay. The method assumes that the α value equals to 1 for normally consolidated clay. The assumption is supported by the results of field measurements reported by Cox et al. (1979), and Pelletier and Doyle (1982). Since the strength ratio can best reflect the effect of past stress history of the clay (Randolph and Wroth, 1982), there is no need to derive a profile of OCR from laboratory tests. 22 Chapter 2 Literature review 2.4.2 Base resistance For a pile in saturated homogeneous clay, the ultimate unit base resistance in undrained condition, q b, can be estimated from q b = Cu N c (Eq. 2.9) where Nc is the bearing capacity factor with respect to cohesion at pile base. The value of Nc factor depends on the sensitivity and characteristics of the clay. Roy et al. (1974) reported a value of about 5 for very sensitive brittle normally consolidated clay. For insensitive stiff overconsolidated clay, Meyerhof (1951) and Skempton (1951) reported a Nc value of about 10. However, the value of 9 is frequently used in the estimation of ultimate base resistance of driven and bored piles. 2.5 Geotechnical centrifuge modeling 2.5.1 Principle of centrifuge modeling The behaviour of soil is well-known to be a function of stress level and stress history. The study of soil behaviour from field tests is rare and often incomplete. The results of laboratory model tests at normal gravity condition may not be representative of the prototype behaviour because of low overburden stresses at various points in the models. Geotechnical centrifuge modeling can replicate stress level at a point in the model identical to the corresponding point in the prototype by increasing gravitational field of the model. To conduct centrifuge modeling, soil models placed on the end of a centrifuge arm is spun so that it is subjected to a centrifugal acceleration at many times stronger than the Earth’s gravity. For a centrifuge model, the increase in overburden stresses with depth is dependent on the soil density and the magnitude of centrifugal 23 Chapter 2 Literature review acceleration. The results of a model of scale N accelerated to N times Earth’s gravity can be related to the prototype scale by applying appropriate scaling laws. Basic scaling laws can be derived by means of either using dimensional analysis or considering the governing differential equations. Some of them which relates to the present study had been presented by many researchers, e.g. Leung et al. (1991). Table 2.1 summarizes the scaling relations relevant to the present study. In centrifuge modeling, some limitations of the method should be realized. In practice, the Earth’s gravity is uniform. However, there is variation of acceleration field with the radius to any element in the centrifuge soil model. To minimize the error due to non- linear stress distribution, Schofield (1980) suggested that the effective centrifuge radius should be measured from the central axis to the upper onethird of the soil depth. In addition, to keep the maximum error in the stress profile to less than 3% of the prototype stress, the ratio of depth of soil sample to effective centrifuge radius should be less than 0.2. Feld et al. (1991) investigated the effect of size of centrifuge model container on geostatic stresses by conducting a series of 16 geostatic stress tests at 10 to 50g. Comparison of the measured results and expected geostatic stresses suggested the order of importance of the factors was: gravity, width of container, void ratio and whether or not the container wall was lubricated. However, the height of the sample did not affect the results. It is also found that the widest container as possible should be used. From the results of centrifuge tests on a single pile in compacted sand by varying pile diameter and g levels, Craig and Sabagh (1994) found tha t the effect of container base boundary effect may be insignificant for piles in dense sand having distance between the container base to the pile base larger than 5 times pile diameter at low stress levels, but only 3 times pile diameter at higher stress levels. The lateral 24 Chapter 2 Literature review boundary constraints also show similar trend to be less critical at higher stress levels. It was found that a container width of 35 times pile diameter showed no lateral restraint effect even at low stress levels. Zeng and Lim (2002) studied the effect of centrifuge radius and boundary effect imposed by model container analytically by using the Original Cam-Clay model. The soil in the model is assumed to be uniform, homogeneous and normally consolidated clay. Comparison of geostatic stresses in the soil elements between models subjected to gravitatio nal and centrifuge acceleration at various g leve ls and container sizes suggested that the difference in vertical stresses is small but more significant for horizontal and shear stresses. The influences of centrifugal acceleration, radius of centrifuge acceleration and boundary conditions due to model container vary from problem to problem and from parameter to parameter. 2.5.2 Centrifuge model tests on axially-loaded piles Among the early studies of centrifuge modeling on pile foundation, Mikasa et al. (1973) conducted tests on model piles installed at 1g but load tested at 100g. The observed load transfer mechanisms along the pile were not consistent with field measurement s. The model tests of jacked- in single pile conducted at 70g by Ko et al. (1984) shows that serious error in bearing capacity of up to 40% could occur if the pile is installed at 1g. Thus, pile installation should be performed in the stress condition comparable to a field prototype. The effect of the acceleration level was studied in more detail by Craig (1984). He concluded that installation of pile at 1g could lead to reduction in pile capacity as high as 50% due to lack of modeling stress regime around the pile during installation. 25 Chapter 2 Literature review A system using pneumatic jack and electric control units to jack in a pile and subsequent cyclic loading on the pile was developed by Cook and Lewis (1980). A similar system using a double acting cylinder was developed by Ko et al. (1984). In an attempt to overcome the drawback of pneumatic jack which has less precision than a hydraulic jack, Clegg (1981) developed a loading system which used the self- weight of the pile and pile cap to overcome the soil resistance during installation and subsequent load tests by using a flexible cable, a set of pulleys and a loading system. By using such displacement-controlled system, the range of rate of penetration of the pile that the system can make is wider tha n a single jacking system. Applying lateral load on the pile is possible by varying the eccentricity of the cable. An additional advantage is that the pile will align itself with centrifugal acceleration. Thus, it can be ensured that the pile can be installed vertically. The system was later modified by Nunez et al. (1988) to accommodate a double-acting pneumatic hammer to drive a pile in- flight with frequency up to 2.5 Hz. Further development of driving system was done by Cyran et al. (1991) which allows a pile to be driven in- flight with a frequency of 27.7 Hz at 25g. Nicole and Randolph (1994) developed a miniature pile driving actuator which can produce a maximum blow frequency of 30 Hz. However, such low frequency is still not high enough to drive a pile in undrained condition. To simulate the pile driving properly, Randolph (1979) showed that a driving rate of over 1000 blows per second may be required. Thus, it is suggested that an alternative way of installation by jacking a pile to a full depth within a time which no significant excess pore water pressure can dissipate may be used. Comparison between static and dynamic (a frequency of 14 Hz) driving of piles conducted by Craig (1985) suggested that installation methods do 26 Chapter 2 Literature review not significantly affect the behaviour of piles if the total installation time is about the same. The application of centrifuge modeling on studying pile behaviour has been performed by many researchers. Ko et al. (1984) reported a direct comparison between full-scaled load and centrifuge static load tests of a single driven pile in clay. The comparison of load-displacement curves, load transfer curves and pile capacity shows that the centrifuge results can produce good agreement with the field tests. The effect of residual stresses in pile jacked- in slightly overconsolidated kaolin clay at 100g was reported by Thomas et al. (1996). It is shown that neglecting residual stresses can cause significant errors in the estimation of load transfer in the pile both in compression and tension. It is shown that the residual stresses cause pile head stiffness of a compression pile higher than that of a tension pile. In addition, the pile head load and stiffness seem insensitive to the different constant rate of penetration ranging from 15-300 mm/min. In many occasions, the load and bending moment distribution along a model pile may be required. Such information may be obtained by installing load cells at several sections of a model pile (Cook and Lewis, 1980, Nunez and Randolph, 1984, and Lee et al., 1998). The more common instrumentation is installing a load cell near the pile head and instrumenting a series of bridges of strain gauges on the pile shaft. The methods of strain gauging a model pile adopted so far by many researchers can be categorized into two methods: internal (Fioravatnte et al., 1994) and external instrumentation. For internal instrumentation, the model pile needs to be a closed-end pile for protecting the instrumentation on the internal surface of the pile. In addition, the model pile needs to be formed from segments which would affect the uniformity of stiffness of the model pile. Nunez et al. (1988) attached strain gauges on recessed 27 Chapter 2 Literature review cross-sections along the pile. To prevent the damage of strain gauges from pile-soil abrasion and water intrusion, a protective epoxy coating is applied uniformly on the pile surface. This method makes the modeling of open-ended pile possible. Nunez and Randolph (1984) further concluded that external instrumentation is more efficient in terms of providing less fabrication time, more economic, higher accuracy of instrumentation and greater number of measurement locations. 2.6 Summary Although piles have been used as foundation to support superstructure for a long time, a reliable theory to predict their responses when subjected to compression and tension throughout their lifetime is rather difficult due to sophisticated pile-soil interaction. Pile load test is thus the most reliable measure to obtain the actual pilesoil response at any particular sites when subjected to applied loads. Several empirical design methods have been developed based on the data base of pile load tests. The limitation of such methods is the reliability of the prediction extrapolated out of the database. The attempts to overcome such limitation have been proposed by using analytical theories based on effective stress analysis. However, the validation of these analytical results with field measurements is still required. In pile load test, controlling site condition, e.g. soil stratigraphy and ground water level, which could affect the test results, is very difficult. Furthermore, the cost of conducting full-scale pile load tests is expensive. Centrifuge modeling technique is an effective and economical way to investigate the behaviour of pile and its surrounding clay. Many techniques and theories have been developed to simulate prototype condition as close as possible. Geotechnical centrifuge modeling has been widely adopted to investigate of the behaviour of pile in clay. 28 Chapter 2 Literature Review Table 2.1 Scaling relationship for geotechnical centrifuge modeling (After Leung et al., 1991). Parameter Prototype Centrifuge model at Ng Linear dimension 1 1/N Area 1 1/N2 Volume 1 1/N3 Density 1 1 Mass 1 1/N3 Displacement 1 1/N Strain 1 1 Stress 1 1 Force 1 1/N2 Time (dynamic) 1 1/N Time (consolidation) 1 1/N2 29 Chapter 2 Literature Review Figure 2.1 Ko of normally consolidated clay vs (a) friction angles (b) plasticity index. (After Ladd, 1977) 30 Chapter 2 Literature Review (a) Undrained shear strength vs OCR (b) Relative increase in undrained shear strength ratio with OCR Figure 2.2 Experimental results from CKoU direct simple shear tests on six clays by Ladd and Edgers (1972). (After Ladd et al., 1977) 31 Chapter 2 Literature Review Figure 2.3 Soil movement due to pile installation (After Randolph et al., 1982) Figure 2.4 Schematic diagram shows modeling of pile installation by using expanding cylindrical cavity theory in undrained soil with Mohr-Coulomb characteristics (After O’Neill, 2001). 32 Chapter 2 Literature Review (a) (b) Figure 2.5 Stress distributions in clay around pile immediately after driving (a) OCR = 1 (b) OCR = 8 (After Randolph et al., 1979). 33 Chapter 2 Literature Review Figure 2.6 Effect of residual loads after installation on the interpretation of the results of pile load tests. (After Briaud and Tucker, 1984). 34 Chapter 2 Literature Review (a) variation of normalized undrained shear strength with time factor after driving at1.15 pile radius from the pile center (b) variation of normalized shear strength at the end of consolidation with normalized radial distance Figure 2.7 Time effect on undrained shear strength of clay around a displacement pile. (After Randolph et al., 1979) 35 Chapter 2 Literature Review Figure 2.8 Conventional pile load test setup using kentledge reaction. (After ASTM D1143). Figure 2.9 Load transfer mechanism for piles (After Das, 1998) 36 Chapter 2 Literature Review α Figure 2.10 Osterberg load test setup for bored piles. (Taken from the leaflet of Loadtest Inc.) Cu averaged over the embedded length of pile (kPa) Figure 2.11 Variation of adhesion factor α of piles with an embedded length less than 50 m in clay. (Adapted from Terzaghi et al., 1996) 37 Chapter 2 Literature Review Figure 2.12 Correlation factor λ. (After Vijayvergiya & Focht, 1972) 38 Chapter 2 Literature Review Figure 2.13 Design parameters αp and F proposed by Semple and Rigden (1984). (After Tomlinson, 1995). 39 Chapter 3 Experimental setupand material properties EXPERIMENTAL SETUP AND MATERIAL PROPERTIES CHAPTER THREE Chapter 3 3.1 Introduction This chapter describes the NUS geotechnical centrifuge facilities, the model setup and instrumentation, instrumented model piles and the control system used in the present study. The procedure of clay preparation is then presented. At the end of this chapter, details of in- flight miniature cone penetration test and clay properties are described. The results obtained from these tests on clay are also included in this chapter. 3.2 NUS geotechnical centrifuge In the present study, all centrifuge experiments were performed on the NUS geotechnical centrifuge shown in Figure 3.1(a). The centrifuge has a twin-swingplatform configuration with both platforms fully swung up during flight. The radial distance from the center of rotation to the base of the model container is 1.872 m. Figure 3.1(b) shows the position of the swing platforms during centrifuge operation. Normally, one of the platforms is placed with model package while the other is loaded with appropriate counterweight to provide moment equilibrium for the centrifuge arm during spinning. The maximum payload mass that the centrifuge can operate is 200 kg at the maximum g- level of 200g. The payload capacity of the centrifuge is hence 40gtonnes. Each platform has a working area of about 750x700 mm with a clear headroom of 1.2 m to allow equipment to be mounted on top of the model container. 40 Chapter 3 Experimental setupand material properties The centrifuge is driven by a 56-kW hydraulic drive system. A hydraulic pump is driven by an AC-motor which is located outside the centrifuge enclosure. DC power supply and signal outputs are transmitted from the control room to the centrifuge, and vice versa, via the multi-way connectors, electrical slip rings, and a junction box which is placed on the platform nearby the container. Of the total 100 copper-graphite slip rings, 90 low noise slip rings are used for signal transmission. For the remaining 10 rings, three are used for 230VAR, two are used for video signals from on-board closed-circuit TV camera, and five are used for transmitting electrical power with a maximum rating of 15 A. All the signal cables are screened and earthed to minimize electromagnetic and ground noise pick- up. The detail description of the centrifuge can be obtained from Lee (1991) and Lee et al. (1991). The control room accommodates the main control system of the centrifuge machine and data acquisition system. The data acquisition system consists of a group of NEC amplifiers, DAP3000a/111 analog to digital converter cards, and several micro-computers. After output signals are transmitted from the slip rings to the control room, all data signals are amplified and filtered by the NEC amplifiers. Amplification factor of each transducer is selected in a manner that the amplified output signals are within a range of -10 to 10 V. For load cell and pore pressure transducers, an amplification factor of 100 is used while the signals of potentiometers are not amplified. Thereafter, the amplified data are converted to digital signal by the DAP3000a/111 analog to digital converter cards prior to be recorded by Dasylab-1 and -4 microcomputers. For the control system, Dasylab-2 microcomputer is used to generate command signals. The command signals are then converted to analog data via DAP3000a/111 digital to analog converter and sent to the centrifuge room via the amplifier. No amplification is applied to the command system. 41 Chapter 3 Experimental setupand material properties 3.3 Experimental setup This section describes the model package used in the experiments. The setup used for performing pile load tests is first described followed by detail of the setup for measur ing suction pressure at the base of a tension pile and conducting an in- flight miniature cone penetration test. The experimental setup consists of model container, loading frame and load actuator, as shown in Figure 3.2. 3.3.1 Setup for pile load tests The experiment setup developed by Yet (1998) was modified and used in the present study. The detail description of the initial designed experimental setup is given by Yet (1998) and Goh (2000). Figure 3.3 also shows the main components of the present load test setup which comprises model container, loading frame and load actuator. The model container is a cylindrical stainless steel tub flanged at the top to allow the loading frame to be bolted on it. The container has an internal diameter of 500 mm and a height of 400 mm. A valve was installed close to the base of the container to facilitate double-drainage during consolidation of clay samples at 1g. A miniature video camera can be fixed on a mounting bracket at the side of the container to capture a side view of the model pile during centrifuge tests. The loading frame is a stainless steel frame structure laterally braced by an intermediate mounting plate. The hydraulic actuator is bolted with the top of the frame which is stiffened by stainless steel plates. The other end of the actuator is mounted with an intermediate mounting plate. Vertical guide rod is fixed below the intermediate mounting plate to prevent downward movement of the pile from lateral sway and to provide verticality to the model pile. A 250- mm potentiometer is secured 42 Chapter 3 Experimental setupand material properties in a perspex holder resting on the intermediate mounting plate with its slider rod bearing on the movable bar. The movable bar is mounted in-between the threaded end of the piston rod of the actuator and load cell. As the piston rod is moving either upward or downward, the movable bar, the sliding rod, model pile, and load cell would move with the piston. This allows the penetration depth of the model pile to be monitored. The installation of the model pile and the subsequent load tests were performed using the servo-controlled hydraulic actuator. The actuator consists of a hydraulic cylinder and piston rod, a servo valve, and an electro-hydraulic closed- loop control system. The hydraulic pressure in the hydraulic system is generated by a 3.2 kW electric pump which is driven by a constant speed 3-phase motor. The oil from a 76-liter tank is drawn by the pump at a flow rate of 12.5 liter a minute resulting in a maximum pressure of 21 MPa. However, the pressure in the system is limited to 7.5 MPa to prevent damage to the hydraulic union. The pressure is maintained constant by a 10-liter nitrogen filled accumulator. From the tank, the oil flows through a directional valve, a pressure reducing valve, a filter, and a rotating hydraulic union before reaching the servo- valve which is secured on board. The hydraulic cylinder is a double-acting single-rod cylinder having a bore diameter of 50 mm and a stroke length of 190 mm. The diameter of the piston rod is 36 mm resulting in the ratio of piston area to annulus area of 2. A 6-mm thick rubber mat between the hydraulic cylinder and the intermediate mounting plate is used to damp the oscillation induced by dithering of the servo-valve. The closed- loop servo-controlled system can be operated in either load or displacement control mode. The performance of the servo-valve is controlled by MOOG servo-controller. Figure 3.4 shows a schematic diagram of the closed- loop 43 Chapter 3 Experimental setupand material properties servo-controlled actuator. The command signal fed to the servo-controller is generated from a computer using the triangular wave signal generator of Dasylab v.3.0 computer program. The generated digital signals are then converted to analog data signals before reaching the controller. As illustrated in Figure 3.4, the command signal is fed to the servo- valve controller to activate the spool movement inside the hydraulic servo-controlled valve. By comparing the feedback signal and the command signal, the controller will send command signal to the servo-valve to activate the spool movement which in turn adjust the pressure in the hydraulic system. The change of hydraulic pressure will activate the movement of the hydraulic cylinder in the direction which will result in the summation of the feedback and command signals to be zero. In the mean time, the registered signals of potentiometer and load cell are sent to the acquisition system in the control room. In the present study, all experiments were conducted using displacement control since the feedback of the load cell fluctuates too much resulting in severe vibration of the piston rod inside the hydraulic cylinder. The output signal of the 250-mm potentiometer is fed to servovalve controller as the feedback signal of the system. 3.3.2 Setup for base suction pile load test and in-flight miniature cone penetration test Both pile load test for investigating suction at the base of tension pile and inflight miniature cone penetration test were conducted in a rectangular model container having internal dimension of 540 mm in length, 200 mm in width and 470 mm in height. A steel frame having two load actuators on the top was fixed on the top of the container as shown in Figure 3.5. The actuators are located 140 mm away from each other. By turning the frame around and re- fixing it, a maximum of 3 tests can be 44 Chapter 3 Experimental setupand material properties performed in one clay sample. A moveable aluminium plate was fixed in between the end of piston rod of the load actuator and load cell. The movement of load cell, model pile and cone penetrometer was thus measured by placing the 250- mm potentiometer on the top of the plate. The control system of the load actuators is identical to that described in the previous section. 3.4 Transducers 3.4.1 Load cell An Interface SML-200 load cell was used to measure the applied load on the pile. The load cell has a stainless steel diaphragm construction with temperature compensated strain gauges. The load cell is designed to have 0.05% accuracy with a maximum load capacity of 200 lbf in tension and compression. It is designed with an overload safety factor of 2. The thread size at the dead (top) and live (bottom) ends of the cell is 3/8-24 UNF-2B with a depth of 0.38 in. According to the manufacturer’s calibration certificate, the calibration factor of the load cell is 42.34 N/mV. After being used for a few months, calibration was performed to check the accuracy of the calibration factor. It is done by connecting two thread adapters on both ends of the cell. The cell was then placed on a level floor and loaded on the live end by applying dead weights on the live end of the cell. The calibration chart is presented in Figure 3.6. The calibration factor is determined to be 40.725 N/mV for loading and 40.602 N/mV for unloading. 45 Chapter 3 Experimental setupand material properties 3.4.2 Linear potentiometer (POT) Five potentiometers with different stroke lengths were utilized in the present study. The excitation and maximum output voltage of these potentiometers are 10 VDC. The movement of pile head connector, which was fixed with load cell and moveable bar, was measured by the Midouri 250- mm potentiometer. The relative movement between pile head connector and pile head was measured by two 50- mm potentiometers. To monitor the settlement of the clay sample, two 100- mm potentiometers were mounted on an aluminium tube as shown in Figure 3.3. All potentiometers were calibrated using a digital vernia caliper. The calibration factors of 250-mm potentiometer and two 100- mm potentiometers are 24.85 mm/V, and 10.13 and 10.08 mm/V, respectively. For two 50- mm potentiometers, the factors are 5.06 and 5.08 mm/V. Re-calibration of the potentiometers reveal that the factors do not significantly change with time. 3.4.3 Miniature pore pressure transducer (PPT) Druck PDCR 81 miniature pore pressure transducer is widely used in centrifuge modeling due to its tiny size. The transducer comprises two major parts. The first part is a single crystal silicon diaphragm with a fully active bridge of strain gauges on its surface. The second part is a porous filter stone placed in front of the cell to prevent soil particles from contacting the diaphragm. Prior to installing the transducer in the clay, the porous stone was saturated by deairing the transducer with a vacuum pump. All transducers were calibrated using a calibrator manufactured specially for this purpose. The calibration was performed by applying air pressure to the 46 Chapter 3 Experimental setupand material properties diaphragm. The corresponding output signal was then recorded against the pressure applied. 3.5 Instrumented model piles Two model piles were fabricated in the present stud y. The first model pile was extensive ly instrumented by strain gauges and used for conducting pile load tests. The second mode pile was specially designed with a PPT at the pile base. 3.5.1 Strain gauge-instrumented model pile To investigate the effect of different loading methods on the pile behaviour, the model pile was designed to be loaded either at the pile head or at the pile base. Strain gauges were placed along the pile shaft to monitor the load transfer characteristics during load tests. 3.5.1.1 Fabrication of model pile The model pile shaft was fabricated from a 12- mm aluminium hollow pipe with a wall thickness of 1 mm. The pipe was cut to 255 mm long with a 10- mm sliding slot near the pile head, as shown in Figure 3.7. At the base of the model pile, a 12-mm hollow brass tube was threaded with M4 female thread size in the inner side of the tube and inserted into the aluminium tube. Four 2-mm diameter holes were drilled through the aluminium and brass tubes and fixed together with 4 2- mm diameter brass pins. The bottom of the model pile was then closed using a brass pile tip shown in Figure 3.7. The pile shaft can be utilized as a conventional test pile, i.e. compression and tension loads are applied at the pile head, by using brass connector 2 and brass shear 47 Chapter 3 Experimental setupand material properties pin 2 shown in Figure 3.7. Brass connector 2 was drilled with a 4- mm diameter hole at the lower end and threaded with 3/8-24UNF thread size at the other end. To connect the pile head with load cell, brass connector 2 was inserted into the model pile at the top end. Brass shear pin 2 was inserted through the sliding slot of the shaft and the hole of the connector to guide the movement of the connector along the slot. The brass connector was then connected to the live end of load cell by thread. Using a brass connector with a sliding slot in loading the pile head has advantages over fixing the pile head directly to load cell. Firstly, it can be ensured that after installation, the model pile is normal to the soil surface since the pile must hang on the shear pin and follow the direction of centrifugal acceleration. Secondly, the load applied to the pile head can be completely released by moving the shear pin to the middle of the sliding slot. Thus, pile load tests using displacement control can be performed. One of drawbacks of fixing model pile head to load cell is that only loadcontrolled mode of testing can be conducted. To release applied load, the command signal must be set to zero applied load datum value. The servo-controller will then adjust hydraulic pressure in the system until feedback signal of load cell is equal to the command signal. Thus, downdrag force along the pile shaft due to consolidation of surrounding clay will be mobilized along the shaft first resulting in tension load registered by the load cell. Subsequently, the model pile is moved down to the clay to release the tension load. Thus, residual stresses left along the pile after installation may be lost. For the current design, it is clear that by moving the shear pin to the middle of the sliding slot allows the model pile to settle freely with the clay-water system. Figures 3.8(a) and (b) show the loading mechanism of conventional test pile loaded in compression and tension, respectively. As shown in Figure 3.8(a), 48 Chapter 3 Experimental setupand material properties compression load on pile head is obtained by pushing the connector against the pile head. By pulling the pile head connector out, double shear occurring in the brass shear pin causes a tension force at the pile head, see Figure 3.8(b) To simulate the loading method of O-cell pile load test, brass connector 1, internal aluminium tube and internal aluminium loading rod were fabricated to facilitate application of load at the pile base. A schematic dia gram of these components is shown in Figure 3.7. By pulling brass connector 1 out, applied tension load is transferred to internal aluminium tube via brass pin 1. It is then transferred to the internal aluminium loading rod by mean of a 10x10x1.5 mm brass key. Once the lower end of the 30x1.5 mm slot of the aluminium tube hits the key, the key pulls the rod up. As a result, the applied load acts directly at the pile base. Figure 3.9(a) shows a schematic diagram of the model O-cell pile. When the brass connector is subjected to compression, the pile head is directly pushed by the connector since the 10x10x1.5 mm shear key is not in contact with either end of the slot of the aluminium tube. Thus, the model O-cell pile, installation and compression loads are applied at the pile head whereas tension load is applied at the pile base. 3.5.1.2 Instrumentation of model pile To monitor the axial load distribution in the pile, the pile shaft is instrumented with 10 bridges of strain gauges. The location of the bridges is shown in Figure 3.10. A bridge comprises 4 strain gauges bonded on the external surface of the pile shaft and wired together to form a complete wheatstone bridge. Kyowa KFG 1-DO-C1-23 uniaxial strain gauges were used. To bond a strain gauge on the model pile, the pile is first cleaned to remove dirt, paint or oil and then polished with sand paper. Oil and grease along the pile are removed using acetone. CN adhesive is applied on the back of strain gauge prior to 49 Chapter 3 Experimental setupand material properties attaching it at appropriate location. Strong thumb-pressure is applied to the gauge via a polythene sheet for about a minute. After the adhesive has hardened, gauge- lead is carefully soldered to a gauge terminal by using tweezers to prevent the sensitive gauges from damage. Insulated copper wires of 0.1 mm diameter are used to join the terminal of each gauge to form a full bridge. Another 4 insulated copper wires are then connected to the terminals. Two gauges are used to transmit bridge excitation from the pile head while the other two are used to feedback output signals. The wires are then connected to strain gauge terminals at the pile head. By connecting another set of insulated wires to the terminals at the pile head, output signals from strain gauges can be sent to a strain meter via 2 sets of 25-pin parallel connector and vice versa for the excitation from the strain meter. Figures 3.11(a) and (b) show a wheatstone bridge circuit of four strain gauges. Such circuit enhances the measurement accuracy by temperature compensation and elimination of effects of bending. Strain gauges are bonded in the manner as shown in Figure 3.11(b). After installation of strain gauges, a thin layer of rubber putty-type coating agent is used to cover all the strain gauges and terminals. Thereafter, a 1.5 mm thick layer of epoxy is coated on the pile shaft to provide waterproof to the strain gauges. 3.5.2 PPT-instrumented model pile To measure suction pressure generated under the base of a tension pile, another model pile was fabricated with a pore pressure transducer at the model pile base, as shown in Figure 3.12. The model pile was fabricated from a 270-mm long aluminium hollow pipe having a diameter of 12.7 mm and a wall thickness of 1.1 mm. At the bottom end of the pipe, two pieces of brass adapters were machined to hold the 50 Chapter 3 Experimental setupand material properties transducer in place. The first adapter is a hollow brass tube having outside diameter of 12.7 mm with 2.1 mm thick wall. The upper end of the tube is fixed with the aluminium pipe by using four steel pins. The other end is threaded to connect with the other adapter. The other adapter is used to hold the head of the transducer and the pile shaft. The bottom end of the adapter is the base of model pile. A 10 mm hole was bored and threaded at the bottom end to accommodate a threaded perspex plate. The plate was drilled with several 1- mm diameter holes to allow water flow to the transducer. A 1.5 mm deep groove was then cut at the bottom of the plate allowing it to be tightened or removed by a screw driver. Leakage of water from the pile base into the pile body is prevented by a rubber O-ring. By tightening the perspex plate, the head of the transducer will compress the O-ring resulting in air and water tight. At the pile head, an 8- mm slot was cut to connect the cable of the transducer with the junction box placed on the centrifuge platform. A 4- mm hole was drilled so that a shear pin can be inserted for connecting the model pile with brass connector 2. The location of the slot and the hole at the pile head are shown in Figure 3.12. The model pile is finally coated with a layer of epoxy to achieve the same diameter as the strain gauge- instrumented model pile. 3.6 Clay sample preparation All clay samples in the present study were prepared from whitish Malaysian kaolin clay powder. Clay slurry having water content of 120% (1.5 times liquid limit) is thoroughly mixed in a mixer under vacuum condition for at least 3 hours. The clay slurry is then laid on top of a 2.5-cm thick layer of coarse sand covered by a sheet of geotextile in the model container. To reduce interface friction between the clay and 51 Chapter 3 Experimental setupand material properties container wall, a layer of grease is applied to the wall. To avoid air bubbles from being trapped in the clay, the slurry is laid in the container under water. Once the clay slurry reaches the required level, a sheet of geotextile is laid on it prior to covering it with a stainless steel cover plate. Thereafter, the clay slurry is gradually loaded using a pneumatic jack. Once the applied load has been fully consolidated under 100 kPa pressure which is about two-third of the maximum required pressure, the applied load and cover plate are removed to install pore pressure transducers in the clay. Firstly, the location of each PPT was marked on the clay surface by using a plumb bob. To install PPTs at predetermined depths, a half-cut 6- mm hollow pipe is used to make holes at the marked locations on the clay surface. To make a hole, the pipe is penetrated into the clay with a predetermined length. By turning the pipe around many rounds, the cut clay can be removed with the pipe resulting in a hole. The depth of each hole depends on the predetermined position of each PPT inside the clay. PPTs are then inserted into the cut holes prior to filling up all voids in the holes with clay slurry. After the installation of pore pressure transducers, another layer of kaolin slurry is poured on the top of first clay layer to get the required final thickness of preconsolidated clay. A sheet of geotextile and the cover plate are put back and gradually reloaded by pneumatic jack. The final pre-consolidation pressure applied on the sample is 150 kPa. It should be noted that the benefit of installing pore pressure transducers at about two-third of the maximum load is that there would be smaller change of the transducers’ elevations after the final pressure is applied. Moreover, the stress history of the clay around the holes drilled for installing PPTs will be the same as the other part of the clay. Hence, the uniformity of the clay throughout the sample 52 Chapter 3 Experimental setupand material properties would be expected. The position of PPTs in the clay sample is shown in Figure 3.13 and Table 3.2. After the clay had been fully consolidated, the applied load is removed. The loading frame for conducting pile load tests is bolted on the top of the container. The typical thickness of clay bed is about 24 cm. The container is then placed on the centrifuge platform. Two short-distance potentiometers mounted on an aluminium tube for monitoring the degree of self-weight consolidation are fixed on the top of the container. The connectors of all transducers are connected to the junction box. The model pile is subsequently fixed to the pile top connector which connects to the load cell. Lubricant oil is then applied to the sliding slot to reduce sliding friction. Finally, an aluminium pile head clamp is fixed to the pile top and the sample is ready to be tested. 3.7 Sign convention The sign convention adopted in the present study is illustrated in Figure 3.14. The r- z reference axes for positive values of dimension in length are shown in the figure. For the sign convention of force, any forces resulting in axial contraction of the model pile are positive and vice versa. Therefore, positive sign is given to the following force: downward applied force at the pile head, upward side and base resistances of pile, downward dead weight of pile and upward uplift force applied at the pile base. 53 Chapter 3 Experimental setupand material properties 3.8 In-flight miniature cone penetration test A miniature in-flight cone penetration test (Tani and Craig, 1995) was carried out at 100g to determine the undrained shear strength profile of the clay sample used in the experiments. The model setup and sample preparation are described in Sections 3.3.2 and 3.6, respectively. Figure 3.15 depicts structure of the miniature cone penetrometer used in the present study. The cone has a diameter of 10 mm with a tip angle of 60o . A miniature load cell and a pore pressure transducer are instrumented at the cone tip as shown in the figure. Hence, only resistance and pore water pressure at the tip can be measured. To ensure the accuracy of measured tip resistance, a calibration for influence of stresses due to pore water pressure and seal filling acting on the rear of the cone tip was performed. Details of calibration have been described by Tani and Craig (1995). The undrained shear strength of clay can be calculated as follows: Cu = qc − σ v Nk q c = qm + (1 − α ).U b (Eq. 3.1) (Eq. 3.2) where Cu = undrained shear strength of clay, qm = measured cone resistance, qc = corrected cone resistance for ambient total overburden pressure of the clay around the cone tip, σv = total overburden pressure of clay at cone tip, Nk = cone factor, α = cone area ratio, which is approximately equal to the ratio of cross-sectional area of shaft connected to load cell divided by the projected area of cone, and Ub = measured pore water pressure at the cone tip. According to the calibration of the cone in a closed chamber under various water pressures, the α factor obtained is 0.51. To obtain an accurate measurement of pore water pressure, the pore pressure transducer was de-aired with silicone oil. The 54 Chapter 3 Experimental setupand material properties use of silicone oil for de-airing pore pressure transducer is to prevent the holes which connect the cone tip to the transducer from being clogged by clay. Once the clay has consolidated under its self-weight at 100g, the cone was penetrated down at a standard rate of 5 mm/sec using a closed- loop servo-controlled loading system described in Section 3.3.1. The data of measured tip resistance and pore water pressure at the cone tip were recorded with a sampling rate of 50 data/sec. After the centrifuge was spun down, a layer of water above the soil surface was removed as soon as possible to prevent clay swelling. Another cone and a laboratory vane shear tests were then conducted at 1g to find out the cone factor, Nk . By calibrating the result of cone penetration test with the vane shear strength of the clay at 1g, a cone factor, Nk , of 11.81 is obtained. 3.9 Properties of clay The basic engineering properties of the kaolin clay are summarized in Table 3.1. Figures 3.16(a) and (b) show the distribution of water content and total unit weight of the clay samples, respectively, collected with in 1 hour after spinning down of the centrifuge. It should be noted that the depth values in the figures are shown in prototype scale. Figure 3.16(a) shows that the water content around the clay surface and bottom of the samples are relatively high. This is the result of clay swelling during spinning down. Since the middle portion of the sample has the longest drainage path, the average water content of 61% of this region can best represent the in- flight water content of the clay. An average bulk unit weight of about 16 kN/m3 can be observed from Figure 3.16(b). Figure 3.17(a) shows the distribution of undrained shear strength of the clay sample with depth obtained by in- flight miniature cone penetration test. Majority of 55 Chapter 3 Experimental setupand material properties the sample is medium clay overlaid by a layer of 5-m thick soft clay. Based on the bulk unit weight of the clay of 16 kN/m3 , the distribution of overconsolidation ratio (OCR) of the clay sample with depth can be determined and plotted as shown in Figure 3.17(b). According to the Eq. 2.2, the relationship between Cu /σ’vo and OCR of the clay can be correlated as follows: Cu / σ ' vo = 0.267OCR 0.65 (Eq. 3.3) 56 Chapter 3 Experimental setupand material properties Table 3.1 Properties of Malaysian kaolin clay Properties Specific Gravity (Gs ) Values 2.60 Sieve Analysis Clay (%) 87 Silt (%) 13 Sand (%) 0 Atterberg’s limits Liquid Limit (LL) (%) 80 Plastic Limit (PL) (%) 40 Compression Index (Cc) 0.55 Swelling Index (Cs ) 0.14 57 Chapter 3 Experimental setupand material properties Table 3.2 Position of pore pressure transducers in clay. Test series A B C D r, m Depth/pile diameter z/d Radial distance/pile diameter r/d 3.0 5.6 2.0 3.7 2 10.0 4.9 6.7 3.2 3 11.8 5.3 7.9 3.5 4 17.0 4.0 11.3 2.7 5 11.5 7.5 7.7 5.0 7 12.1 10.9 8.1 7.3 8 13.0 14.5 8.7 9.7 2 9.8 4.6 6.5 3.1 3 13.8 5.5 9.2 3.7 4 19.5 4.6 13.0 3.1 5 15.2 7.2 10.1 4.8 6 18.3 8.9 12.2 5.9 7 14.9 10.6 9.9 7.1 2 9.5 5.8 6.3 3.9 3 14.5 3.0 9.7 2.0 4 18.9 3.0 12.6 2.0 5 14.1 7.5 9.4 5.0 6 18.9 8.2 12.6 5.5 7 14.5 10.1 9.7 6.7 1 5.0 6.8 3.3 4.5 2 11.0 6.8 7.3 4.5 3 18.0 6.8 12.0 4.5 Depth Radial distance z, m 1 PPT No. 58 Chapter 3 Experimental setup and material properties Servo-valve Slip rings On-board camera Strain meter Miniature VDO camera Counterweigh Model package Swing platform (a) Photograph of NUS geotechnical centrifuge and experimental setup. (b) Position of the swing platform during centrifuge operation. Figure 3.1 NUS geotechnical centrifuge 59 Figure 3.2 Centrifuge model setup Chapter 3 Experimental setup and material properties 60 Chapter 3 Experimental setup and material properties Load actuator 250-mm potentiometer Intermediate plate Vertical guide rod 50-mm potentiometer Movable bar Pile Head Clamp 100-mm potentiometer Instrumented model pile Miniature VDO camera Figure 3.3 Setup for pile load test. Figure 3.4 Schematic diagram of closed-loop servo-controlled actuator. 61 Chapter 3 Experimental setup and material properties 250-mm potentiometer Load actuators Model container Junction box Figure 3.5 Setup for base suction capacity test and in-flight miniature cone penetration test. 200 180 y = 40.725x - 12.962 R2 = 1 160 Load, N 140 120 y = 40.602x - 12.547 R2 = 1 100 80 60 40 Loading 20 Unloading 0 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 Load cell readings, mV Figure 3.6 Calibration for load cell SML-200. 62 Chapter 3 Experimental setup and material properties Figure 3.7 Schematic diagram of components of model pile. 63 Chapter 3 Experimental setup and material properties (a) (b) Figure 3.8 Schematic diagram showing loading mechanism of conventional model pile subjected to (a) compression load at pile head (b) tension load at pile head. (a) (b) Figure 3.9 Schematic diagram showing loading mechanism of model O-cell pile when subjected to (a) tension load (b) compression load. Figure 3.10 Location of strain gauges on model pile. 64 (b) Figure 3.11 Strain gauge instrumentation on pile: (a) 4 gauge arrangement bridge circuitry (b) arrangement of strain gauges on pile surface. (a) Chapter 3 Experimental setup and material properties 65 Figure 3.12 Schematic diagram of PPT-instrumented model pile. Chapter 3 Experimental setup and material properties 66 Chapter 3 Experimental setup and material properties Figure 3.13 Position of pore pressure transducers in clay. Figure 3.14 Sign convention adopted in present study. 67 Chapter 3 Experimental setup and material properties Figure 3.15 Schematic diagram showing structure of miniature cone penetrometer. 68 Chapter 3 Experimental setup and material properties Water content (%) 58 60 62 64 66 68 70 Depth below soil surface (z), m 0 5 10 T est series A T est series B T est series C 15 T est series D T est series E T est series F 20 (a) 3 Bulk unit weight (kN/m ) 14 15 16 17 18 19 20 Depth below soil surface (z), m 0 5 10 T est series A T est series B T est series C 15 T est series D T est series E T est series F 20 (b) Figure 3.16 Physical properties of clay sample: (a) water content (b) bulk unit weight. 69 Chapter 3 Experimental setup and material properties Cu, kPa 0 10 20 30 40 50 0 Cone penetrometer data Depth below soil surface, m_ Best Fit curve 5 R2 = 0.99 10 15 20 25 (a) Overconsolidation ratio 0 5 10 15 20 Depth below soil surface (z), m 0 5 10 15 20 25 (b) Figure 3.17 Shear strength properties of clay at 100g. 70 Chapter 4 Pile load test: Effect of different loading methods P ILE LOAD TEST: EFFECT OF DIFFERENT LOADING METHODS CHAPTER FOUR Chapter 4 4.1 Introduction To verify the load carrying capacity of piles, static pile load test is commonly carried out. Traditionally, the test load may be applied in compression or tension modes at the pile head. Osterberg (1989) proposed a new pile loading method in which the test load is applied via a bellow-like cell, so-called Osterberg cell or O-cell, at the pile base to pus h the pile shaft upward by the reaction force of the pile base. Several series of centrifuge experiments were conducted at 100g to investigate the behaviour of a jack- in, closed-ended pile in clay under displacement-controlled mode as summarized in Table 4.1. It should be noted that the first character of the name for each load test given in the table represents the type of loading method (Ccompressive, T-tensile). The second character indicates the number of time that a particular type of loading has been tested. Finally, the third character indicates the name of test series (A-G). For example, T3A means the third tensile pile load test of load test series A and C2B means the second compressive pile load test of load test series B. The different loading sequences are designed to investigate the influence of loading direction and loading history on the pile bearing capacity. After the in- flight installation of pile, all subsequent pile load tests were conducted in- flight at 100g without stopping the centrifuge. The effective radius of the spinning is taken as the upper one-third height of the clay specimen, as suggested by Schofield (1980). It is important to note that since the area of plunger inside the hydraulic cylinder on the piston rod side is smaller, the applicable uplift force is smaller than the compression force. As a result, the rate of penetration of tension pile 71 Chapter 4 Pile load test: Effect of different loading methods load tests is lower than that of compression pile load tests. However, the difference of penetration rate between the compression and tension pile load tests is only about 10% which is considered relatively insignificant and both rates are still within the range recommended by ASTM D 1143 and ASTM D 3689. Unless otherwise stated, all test results are presented in prototype scale. In this chapter, only the effect of different loading methods on pile beha viour is presented. This involves the first pile load tests of each test series which are tests C1A (conventional compression pile test), T1B (conventional tension pile test), and T1C (O-cell pile test). The remaining results are presented in Chapter 5. The preparation method of clay sample and its properties as well as the position of five PPTs installed in the samples of test series A, B, and C are described in Chapter 3. Pile installation for test series A and the results of test C1A are described in detail as an illustrative example. 4.2 Soil responses prior to pile installation Prior to tests, all readings of strain gauges, potentiometers, and PPTs were set to zero. While the centrifugal acceleration escalates, the weight of the model pile and clay sample keep increasing with increase in g level. Since the clay sample has a low coefficient of permeability, the increase in clay self- weight generates positive excess pore water pressures in the clay. The excess pore water pressures in clay increases with g level until the g level reaches 100g as shown in Figure 4.1. Thereafter, the generated excess pore water pressures dissipate with time under approximately onedimensional consolidation condition. Figure 4.1 shows the responses of pore water pressures registered by various PPTs in the clay during self-weight consolidation. The corresponding settlement of the clay surface was measured by two 10-cm 72 Chapter 4 Pile load test: Effect of different loading methods potentiometers. By using the hyperbolic method proposed by Tan et al. (1991), the average degree of self- weight consolidation can be calculated from the predicted final soil settlement. Figure 4.2 plots the degree of self-weight consolidation with time after the centrifuge reaches 100g. As shown in Figures 4.1 and 4.2, after the clay has consolidated for about 8 years, all registered pore water pressures in the clay almost equalize to hydrostatic pressures and the clay has an average degree of consolidation of about 92%. At this stage, the magnitude of the remaining excess pore water pressures is small and the clay can be taken as essentially fully consolidated and is ready for pile installation. 4.3 Pile installation Once the average degree of self- weight consolidation of the clay sample exceeds 90%, pile installation is performed by jacking the model pile into the clay. For all tests, the installation of pile and subsequent load tests were conducted using displacement-controlled mode. The typical rate of penetration for pile installation is 3 mm/min. For subsequent load tests, the rates of penetration or uplift are given in Table 4.1. The rate of recording data via the data acquisitioning system during pile installation and all subsequent load tests is 0.04 sec/reading. Typically, the installation process comprises 4 stages. Figure 4.3 illustrates pile installation process in each stage. In stage I, the pile shaft is hanged above the clay surface on the shear pin which is connected to the pile head connector. Hence, the load cell connected to the pile head connector registers a tensile load equal to the pile weight or the effective pile weight if it is submerged into water. Figure 4.4(a) presents the load-displacement curve during pile installation of test series A. During stage I, the registered reaction force at the pile head is -0.45 MN. Since the pile shaft 73 Chapter 4 Pile load test: Effect of different loading methods is submerged into water for 3m, the total pile weight of 0.53 MN can be calculated by back analysis. Stage II of the pile installation commences when the pile is lowered down at a constant rate of 3 mm/min. The registered tensile load at the pile head decreases with increasing pile penetration until zero load is measured at a pile embedment depth of about 2.4 m as shown in Figure 4.4(a). This reveals that during stage II, the pile penetrates into the clay by its self-weight. At the end of stage II, the resistance from the clay against the weight of the pile is high enough to reduce the pile velocity. Consequently, the shear pin which is moving at a constant speed with the pile head connector starts sliding apart from the top end of the sliding slot to the lower end, as illustrated in stage II of Figure 4.3. At this moment, zero pile head load is registered by the load cell. It is found that the shear pin displaces by about 13 mm (model scale) to reach the other end whereas the slot length is only 10 mm (model scale). This indicates that the pile shaft penetrates into the clay for another 0.3 m to deplete its momentum after the shear pin begins to slide down. The total weight of pile is then fully supported by the initial mobilized side (Q si) and base (Q bi ) resistances due to pile self-weight. In stage III, the shear pin reaches the lower end of the sliding slot. The pile is then jacked into the clay at a constant rate of 3mm/min. The corresponding compressive installation force increases with pile penetration as shown in Figure 4.4(a). The maximum installation load and the final pile penetration are 1.42 MN and 16.17 m, respectively. Subsequently, the pile head is fully unloaded by moving the shear pin up to the middle of the sliding slot at stage IV. 74 Chapter 4 Pile load test: Effect of different loading methods During pile installation, as the pile base is being advanced into the clay by the installation load, the clay around the pile is displaced and severely disturbed. Consequently, normal and shear stresses are induced in the clay. The excess pore water pressures in the clay generated by the installation load are shown in Figure 4.4(b). The position of PPTs measured after the completion of experiment is also shown in the figure. It should be noted that PPT 2 and PPT 4 installed in the clay in test series A had malfunctioned. Thus, the ir readings could not be shown in the figure. As the installation load increases, the clay dilates initially due to increase in shear stresses resulting in negative excess pore water pressures as shown in the figure up to a pile penetration of 7 m. As the pile penetrates deeper, the total stresses in the soil mass around the pile increase with the installation load. Once the soil is sheared to the critical state and cannot resist any further increase in shear stresses, the increase in total stress has to be entirely taken by pore water. This results in positive pore water pressures registered for pile penetration higher than 7 m. It can be seen that the positive excess pore water pressures in the clay increase with increasing installation load. Comparing the excess pore water pressures at different radii from the pile at a given depth, the excess pore water pressure tends to be highest right at the pile shaft and decreases with increasing radial distance from the pile. 4.4 Dissipation of excess pore water pressures after pile installation Once the pile is installed to the desired embedment length, the shear pin is moved up to the middle of the sliding slot hence releasing the applied load on the pile 75 Chapter 4 Pile load test: Effect of different loading methods head. The clay sample is then allowed to consolidate for about 420 days before the subsequent load test. Figure 4.5 shows the variation of hydrostatic pressure with time from a supplementary centrifuge test. In this test, the container has a PPT fixed at the bottom of the container which is filled only with water to a level similar to that for the pile load tests. It is evident that there is a gradual reduction in static pore water pressure with time due to evaporation of water causing a lowering of water level. The observed reduction rate of hydrostatic pressure is 0.00407 kPa/day. Typically, there are 6 pile load tests in a test series. Between successive load tests, the clay is allowed to consolidate for 420 days. Thus, the hydrostatic pressure can be reduced by as high as 12 kPa for the duration of a test series. To account for the water evaporation phenomenon, the pore water pressure readings in the present study have been corrected using the above observed reduction rate. Since the magnitude of excess pore water pressures at different pile radii are different, the change of pore water pressures around the pile after pile installation is presented in terms of normalized excess pore water pressure with log time in Figure 4.6 for comparison purpose. The term ‘normalized excess pore water pressure’ is defined as the ratio of excess pore water pressure at any time (U) to the initial excess pore water pressure (Ui). It can be observed from the figure that the excess pore water pressures around the pile dissipate immediately after the pile was unloaded. At any given time, the soil closer to the pile has a higher degree of consolidation, i.e. 1-U/Ui, than farther points. Figure 4.6 also reveals that while PPT 3 recorded decreasing excess pore water pressure with time, the excess pore water pressure recorded by PPT 5 remains constant for some time before it decays. PPT 7, which is farthest from the pile, even recorded an increase in excess pore water pressure until 2 days after 76 Chapter 4 Pile load test: Effect of different loading methods releasing pile installation load. This phenomenon happens due to the threedimensional consolidation of the clay which is so-called ‘Mandel-Cryer effect’ (Mandel, 1953 and Cryer, 1963). During pile installation, higher excess pore water pressure is generated at a point closer to the pile. Once the consolidation process starts after releasing of installation load, excess pore water pressure at every point in the clay starts dissipates outward from the pile. At a given time after pile installation, when higher excess pore water pressure dissipated from a point closer to the pile reaches the point of consideration, it superimposes to the excess pore water pressure at the point of consideration which still has not fully dissipated. Consequently, an increase in excess pore water pressure at the early stage of consolidation at a point farther away from the pile can be observed from the result of PPT 7. After 420 days, all excess pore water pressures readings are less than 10% of their respective initial values. In other words, the soil around the pile shaft has consolidated by at least 90%. At this stage, the cla y can be treated as almost fully consolidated and the first load test is performed. However, it should be noted that in practice, piles are normally tested between 14 and 28 days after installation where excess pore water pressure may not have substantially dissipated after pile installation. 4.5 Axial load transfer of the pile during and after pile installation The axial load distribution along a pile can be derived from the responses of the strain gauges placed along the pile shaft. When a pile is loaded, the corresponding 77 Chapter 4 Pile load test: Effect of different loading methods change of axial strains along the pile is registered by the strain gauges. Subsequently, the change of axial loads in the pile can be obtained from the change of axial strains if the stiffness of the pile is known. The difference in changes of axial loads between any two points along the pile is the load transfer to the surrounding soil via the pile shaft. This transferred load is termed ‘side resistance’ or ‘shaft resistance’ (Q s). The remaining applied load which has not been transferred to the supporting soil around the shaft has to be taken by the soil beneath the pile base. Such resistance from the soil is commonly termed ‘base resistance’ (Q b ). Prior to spinning up the centrifuge (i.e. stage I), all strain gauge readings were set to zero. The strain meter was set to record strain gauge data at a maximum rate of 2 second/set of readings. At 1g, the weight of the model pile is insignificant and can be neglected. However, the weight of the pile increases substantially with centrifugal acceleration. Figure 4.7 present s the axial load distribution curves of the pile at various applied loads during and after pile installation for test series A. Prior to pile installation, tensile forces were recorded by the load cell at the pile head and by strain gauges along the pile shaft. At this stage, the axial load registered by each strain gauge represents the weight of pile segment below that particular strain gauge. Since it is impracticable to instrument strain gauges right at the pile base, the axial loads at the pile base shown in this and all subsequent figures are linearly interpolated from the two bottommost strain gauges. The gradient of the load transfer curve before pile installation indicates that the weight of pile per unit length is 26.7 kN/m. It is noted that the back-calculated pile unit weight of 15.2 kN/m3 is approximately the same as the unit weight of the clay sample. At stage II of pile installation (Figure 4.3), there is no applied load on the pile head yet. The side and base resistances are thus mobilized to support only the weight 78 Chapter 4 Pile load test: Effect of different loading methods of the pile. It is noted that the axial load recorded at the ground level is the weight of the pile segment above the clay surface. At stage III, as the pile is jacked into the clay, the axial load recorded at the pile head increases due to increasing clay resistance. It is noted that since the centrifuge machine cannot perfectly maintain its rotating speed to be constant, the kinking of the axial load distribution curves is observed in Figure 4.7. The magnitude of the kinking depends on the sensitivity of the strain gauges used and the workmanship of strain gauging. Therefore, different strain gauges show different magnitudes of kinking. Immediately after the installation load was removed, Figure 4.7 reveals that little axial load was measured at the pile base. Consequently, the majority of pile weight is resisted by positive side resistance. As the time goes on, the measured axial load in the pile gradually increases until the axial load at the pile base approximately equals to the weight of the clay displaced by the pile at 420 days after pile installation. The change of axial loads in pile with time after pile installation is believed to be due to the different magnitude and dissipation rate of excess pore water pressures around the pile. In the previous section, it is shown that higher excess pore water pressure is generated at a point closer to the pile. It will be shown later by subsequent test results that the highest excess pore water pressure is generated at the pile base and its magnitude can be as high as effective overburden stress of the clay. This would mean that the effective stress of clay around the pile base is almost zero during pile installation. Once the installation load is unloaded, the excess pore water pressures in the clay close to the pile, especially around the pile base, dissipate much faster than that at a farther point due to Mandel-Cryer effect. In other words, the rate of volume change, i.e. densification, of the clay beneath the pile base is much higher than that 79 Chapter 4 Pile load test: Effect of different loading methods close to the pile shaft. Since the total volume of the clay around the pile shaft is much more than that around the pile base, it is believed that the densification of the clay beneath the pile base does not significantly cause settlement to the above clay layer. Therefore, there is a tendency that the pile base would separate (but may not actually separate) from the clay beneath. As a result, axial load at the pile base less tha n the weight of displaced clay was measured in the experiment. As the consolidation process progresses, the clay around the pile shaft gradually settle. At this stage the resistance of the clay beneath the pile base is relatively high enough to support the pile. As a consequence, the settlement of the clay around the pile shaft is higher than that of the pile. This causes drag-down loads along the shaft which gradually increases the axial load in the pile as shown in Figure 4.7. At 420 days after the release of applied load, the clay around the pile is almost fully consolidated. The axial load in the pile recorded by any particular strain gauge represents the sum of the side resistance on the pile segment above that strain gauge and the weight of the pile segment. Table 4.2 summarizes the results from the axial load distribution curves shown in Figure 4.7 at various stages during and after pile installation for test series A. The resultant downward and upward forces acting on the pile are also shown in the table. From the table, it is found that the force equilibrium of the pile is not satisfied. Since the weight of the pile and the applied load at the pile head measured directly by the load cell are reliable, it is evident that the traditional method of determining side and base resistances from axial load distribution curves may not be suitable for this case. This is believed to be due to the negligence of pile weight in determining the side resistance. 80 Chapter 4 Pile load test: Effect of different loading methods Figure 4.8 shows the axial load distribution curves during pile installation of test series A after adjustment of pile weight along the shaft. At any elevation along the pile, the axial load is obtained by subtracting the weight of pile segment above that point from the measured axial load and adding the result with Wp -σv Ap . Appendix describes the principle of the adjustment in which the axial load in the pile at ground level is equal to the sum of applied load and (Wp -σv Ap ). Table 4.3 tabulates a summary of axial load distribution shown in Figure 4.8. It is noted that the resultant upward and downward forces acting on the pile are comparable. This means that the axial load distribution curves plotted in Figure 4.8 provide a more meaningful interpretation. It is noted that, at 420 days after the pile head is unloaded, no significant residual load remains in the pile. There is only a positive side resistance of 0.03 MN acting along the shaft to resist extra load of the pile weight in excess of the weight of the clay displaced by the pile, Wp -σv Ap . No significant residual load in the pile is observed after pile installation. This agrees with the results of a simplified boundary element analysis in elastic continuum conducted by Poulos (1987) who pointed out that the residual force generally increases as the pile stiffness decreases. The maximum residual load in the pile increases while the end bearing capacity increases. The analysis of a 1-m diameter pile in soft clay with an axial pile stiffness of 19.6 GPa.m2 revealed that the residual load is small and generally well below the limiting shaft resistance values. In the present study, the Young’s modulus of the aluminiummade model pile is 72 GPa. The axial stiffness of the model pile works out to be 24.9 GPa.m2 , which is higher than the value in the above analysis. Thus, the residual load in the pile in the present study should be small. 81 Chapter 4 Pile load test: Effect of different loading methods 4.6 Compression pile load test C1A 4.6.1 Results and Discussion Test C1A refers to the first compression pile load test of test series A. The test was conducted once the clay was allowed to fully consolidate for 420 days after pile installation. Figure 4.9(a) presents the load-displacement curve of pile load test C1A. The characteristic of this curve can be divided into three stages. For the first stage, the load-displacement curve shows the stiffest response. The applied load increases linearly with the pile displacement until a load of about 1.17 MN. Thereafter, the rate of increase in load reduces rapidly for the second stage. Once the pile has displaced by 3.65 cm, the pile shaft plunges into the clay denoting the third stage. In general, the ultimate load carrying capacity of a pile is defined as the load at which rapid pile movement occurs under maintained or slight increase of load, i.e. the pile plunges. In several occasions, the distinct plunging at the ultimate load is not always clearly defined. Various procedures defining the ultimate pile capacity have been proposed, as reported by Fellenius (1980). Since the load-displacement curve of the pile in Figure 4.9(a) clearly sho ws the pile plunging process. The ultimate load can be determined as 1.80 MN at a pile head displacement of 3.65 cm. The welldefined load-displacement curve of Figure 4.9(a) confirms that the constant rate of penetration (CRP) method is more appropriate for determining the true ultimate pile capacity in clays than the maintained load test method, as suggested by Fleming et al. (1992). In addition, it is found that the ultimate pile capacity of load test C1A is about 26.7% higher than the pile installation load (1.42 MN). This is attributed to the set-up of the clay around the pile shaft. After the full dissipation of excess pore water 82 Chapter 4 Pile load test: Effect of different loading methods pressures due to pile installation, the shear strength and effective lateral stress around the pile increase as described in Chapter 2. Figure 4.9(b) shows the responses of excess pore water pressures in the clay with pile displacement during test C1A. It should be noted that all PPTs were installed in the soil at about the same depth but at different radial distances from the pile. The excess pore water pressures are noted to increase with applied load even after the ultimate load has been reached. Figure 4.10 plots the variation of normalized excess pore water pressure (∆U/σ’v ) at depth of about 16.5 m with radial distance normalized by pile diameter (r/d) at various applied loads for test C1A. The trend lines interpolated from the data points as a function of 1/[(r/d)-a]2 are also plotted in the figure, where a is a constant. It is evident that the excess pore water pressure decreases with increasing radial distance from the pile until it practically vanishes at a radial distance ratio (r/d) of about 8.6. Figure 4.11 shows the variation of normalized excess pore water pressures with time after unloading for test C1A. Note that immediately after the shear pin stopped pushing the pile head, the compression load at the pile head dropped by certain extent. At the third day after the pile was partially unloaded, the applied load was completely removed by moving the shear pin to the middle of the sliding slot. This is because the hydraulic cylinder-controlling computer requires several seconds (model scale) to change the movement direction of the shear pin. This explains why there is a sudden drop of excess pore water pressures at the third day. The excess pore water pressures in the soil show similar responses to those after pile installation as described earlier. The soil closer to the pile shows a higher degree of consolidation at a given time. The Mandel-Cryer effect can be observed from the responses of PPT 5 83 Chapter 4 Pile load test: Effect of different loading methods and 7. The time delay for reaching peak pore pressure of the soil located farther away indicates radial consolidation in the clay. For example, PPT 7 in the figure records peak excess pore pressure at about 1.5 days later than that of PPT 5. This suggests that excess pore water pressures dissipate in radial direction outward from the pile, as suggested by Randolph et al. (1979). As stated earlier, the maximum rate of recording strain gauge data was 2 second/set of readings. According to the strain meter manual, the strain meter requires 80 msec to record one reading. Since the total number of strain gauges instrumented on the pile is 10 gauges, the readings of the bottom bridge are thus recorded at 0.72 sec (model scale) earlier than the one at the top of the pile. Unfortunately, the time required to mobilize the ultimate pile capacity is only about half a second. Hence, the axial load responses along the pile prior to ultimate load cannot be investigated. Figure 4.12 plots the load distribution curves of the pile during and after load test C1A. To investigate the time-delayed recording effect on the load transfer curves, another set of corrected load distribution curves which are linearly interpolated with time from the raw data points are also shown in the figure. It is evident that at ultimate condition, the load distribution curves plotted from raw data are approximately the same as the ones plotted from the interpolated data. This implies that the measured axial loads along the pile at ultimate condition only change slightly with time. The load distribution obtained from raw data can provide sufficient accuracy without linear interpolation of the data with time. Therefore, the data points of load distribution curves for all subsequent load tests in the present study will be plotted without any adjustment with time. From Figure 4.12, the ultimate side and net base resistances of 1.54 MN and 0.53 MN can be observed at an applied load of 2.07 MN. Immediately after the pile 84 Chapter 4 Pile load test: Effect of different loading methods head is unloaded, axial loads along the pile reduce accordingly. A positive side resistance of about 0.30 MN and net base resistance of -0.30 MN are observed from the figure. As the clay consolidates, axial loads along the pile increase gradually with time due to drag-down load along the pile shaft. At 420 days after the pile head is unloaded, residual compression loads are observed along the pile. A net base resistance of 0.10 MN is developed with negative side resistance mobilized along the pile to maintain the pile in equilibrium. This implies that residual loads along the pile can increase after the pile is unloaded due to consolidation of the surrounding clay. 4.6.2 Comparison with static pile design method In this section, the ultimate side and base resistances obtained from the experiment are compared with a widely- used static pile design method. The undrained shear strength used in the calculation is obtained from the in- flight CPT test described in Chapter 3. In the calculation of in-situ effective overburden stress of the clay, the bulk unit weight of 16 kN/m3 is adopted as described in Chapter 3. To estimate the end bearing capacity, a commonly- used bearing capacity factor (N c) of 9 proposed by Skempton (1951) is adopted. For a pile having a diameter of 1.50 m and embedment length of 16.18 m, the theoretical net base resistance is determined to be 0.55 MN. This agrees reasonably well with the ultimate net base resistance of 0.53 MN observed from load test C1A. The prediction of ultimate unit side resistance is more complex since there are many factors affecting the side resistance of a pile as described in Chapter 2. Many pile design methods are empirical in nature. One of the short-comings of these methods is that all effects are dealt with using a single parameter (for example, α factor). In this section, the ultimate side resistances predicted by an approach 85 Chapter 4 Pile load test: Effect of different loading methods proposed by Randolph and Murphy (1985) are used to compare with the experimental results. The approach is adopted as the current design method suggested by American Petroleum Institute. This approach has several advantages over conventional α (Tomlinson, 1957, Peck, 1958, Woodward et al., 1961, Flaate, 1972 and others) and β (Zeevaert, 1959, Chandler, 1968, Burland, 1973, Mayerhof, 1976 and others) methods that it takes into account of ambient effective stress level and overconsolidation ratio. The approach gives a simple computation which links the principles of soil mechanics with the database of American Petroleum Institute (API). Detail of this design method has been described in Chapter 2. Figure 4.13 plots the variation of ultimate unit side resistance (fs) with depth along the pile together with the one calculated by the Randolph and Murphy approach. The experimental results agree well with the predicted values. The ultimate unit side resistance increases with depth and this is mainly due to the increase in ambient effective stress level. The approach proposed by Randolph and Murphy (1985) is based on the observation made by Wroth (1972). Wroth (1972) noted that adhesion factor (α) decreases with increasing overconsolidation ratio (OCR). Furthermore, Wroth (1972) and Semple (1980) suggested that α factor should be correlated with OCR which can be represented by strength ratio (C u /σ’vo ). This may lead to the conclusion that α factor would have a value of one for normally consolidated clay. The results of field load tests performed in stiff normally consolidated clay by Cox et al. (1979), and Pelletier and Doyle (1982) support this conclusion. Figure 4.14 plots the variation of α factor with OCR for load test C1A. The experimental results agree well with the α values predicted by Randolph and Murphy (1985) method for OCR up to 5. For higher OCR values, the prediction slightly 86 Chapter 4 Pile load test: Effect of different loading methods underestimates the α factor. It is evident from the experimental result s that α decreases with increasing OCR, as reported by Wroth (1972). 4.7 Ultimate tension pile load test T1B Test T1B is the first load test after pile installation for test series B. The model pile and testing method used in test series B are the same as test series A except for the sequence of load tests. Table 4.1 shows the test sequence and the rate of loading for each test in test series B. Five PPTs were installed in the clay sample at predetermined positions as described in Chapter 3. Unfortunately, PPT 3 malfunctioned during the self- weight consolidation process. For test series B, the maximum installation load of 1.69 MN is recorded once the pile embedment length reaches 15.98 m. Figure 4.15 plots the excess pore water responses in the clay during pile installation. PPTs 4, 5 and 7 show similar trends to those for test series A. The maximum excess pore water pressures (∆Umax) were recorded at maximum installation load. However, PPT 2 shows a reduction in excess pore water pressures after the pile base advances past the depth of about 10 m, although the installation load is still increasing. This indicates that the ∆Umax generated in the clay during pile installation depends on the location of the pile base rather than the installation load. By comparing the pile embedment length that ∆Umax was registered with the elevation of PPT 2, it is likely that ∆Umax is generated when the pile base reaches the elevation of PPT 2. This is because ∆Umax requires some time to migrate from the pile shaft in a radial direction to PPT 2 located at 5.3 m away. Once the pile base travels past PPT 2, excess pore water pressure (∆U) starts to dissipate. 87 Chapter 4 Pile load test: Effect of different loading methods Figure 4.16 shows the dissipation of ∆U (in terms of normalized ∆U) with time after pile installation of test series B. It can be observed that at an early stage, the clay at PPT 2 has a normalized excess pore water pressure lower than 100%. This is because the soil at PPT 2 starts to dissipate earlier prior to the completion of pile installation. Although the clay at PPT 2 has a shorter drainage path to the clay surface than that of PPT 4, after 4 days of unloading, the normalized excess pore water pressure recorded by PPT 4 is lower than that of PPT 2. In other words, the clay at PPT 2 has a lower degree of consolidation than that at PPT 4 at a given time. This would mean that the rate of consolidation of the clay at PPT 2 decreases due to excess pore water pressure migrating from elsewhere to PPT 2. Figure 4.15 shows that, at the same radial distance from the pile, higher excess pore water pressures are generated only at deeper depth. This probably indicates that, for a pile having embedment depth as high as 10.7 times pile diameter, excess pore water pressure due to pile installation does not dissipate only in radial direction as many analytical solutions assume (Kirby et al, 1983, Randolph et al., 1979, Matlock et al., 1982, and others). Figure 4.17(a) shows the load-displacement curve of load test T1B. The ultimate tensile load of 1.80 MN is well defined at a pile displacement of 4.37 cm. In practice, it is normally assumed that the base resistance of a tension pile is negligible. Thus, the tension pile capacity minus the pile weight (1.80-0.53 = 1.27 MN) is the ultimate downward or negative side resistance. The corresponding excess pore water pressures are shown in Figure 4.17(b). Unlike the soil response in compression load test, negative excess pore water pressures are generated in the clay around the pile. The highest negative excess pore water pressure is registered by PPT 4 which is closest to the shaft. However, PPT 2, which is at about the same radial distance as PPT 4, recorded much smaller values. This suggests that the maximum negative ∆U is 88 Chapter 4 Pile load test: Effect of different loading methods generated at the base of a tension pile. The pile base resistance due to suction pressure is thus expected to exist. The behaviour of suction pressure beneath the base of a tension pile will be further investigated in test series D. Figure 4.18 plots the variation of excess pore water pressure with radial distance to pile diameter ratio at various applied loads for PPTs 4, 5, and 7. It should be noted that PPTs 5 and 7 are 4 m above the pile base while PPT 4 is at about the same level as the pile base. The readings of PPT 4 suggest that negative excess pore water pressure decreases with increasing r/d ratio similar to that observed from test C1A. However, at high applied loads, e.g. at ultimate load, positive excess pore water pressures are registered at r/d ratio = 8.4 which reveals that excess pore water pressure generated at r/d = 8.4 is not influenced by the suction pressure at the pile base. At high applied load, suction pressure developed underneath the pile base may be high enough to cause lateral movement of the surrounding soil towards it. If this is the case, the soil close to the pile base is being sheared under compression unloading mode. This is supported by the ∆U response with pile displacement of PPT 7 which is similar to that of soil in triaxial test under compression unloading mode. Figure 4.19 shows the variation of excess pore water pressures with time after release of test load for load test T1B. It can be seen that once the applied load si completely removed, sudden changes of excess pore water pressure from negative to positive value can be observed for every PPT. This reveals that negative ∆Us generated during tension pile load test exist only in short term and diminish rapidly after unloading. The positive excess pore water pressures generated after unloading indicates that the clay around the pile has been loaded in compression mode. Figure 4.20 presents the load distribution curves along the pile during and after load test T1B. As the tension load increases, the self-weight of the pile is 89 Chapter 4 Pile load test: Effect of different loading methods progressively overcome. The increase of tension loads along the pile corresponding to increasing tension at the pile head are then recorded until the ultimate tension capacity is reached. At an applied load of -1.83 MN, the axial load at the ground surface and at the pile base is -1.83 MN and -0.82 MN, respectively. The ultimate side resistance is thus –1.01 MN. It is noted that the negative sign of side resistance indicates downward direction. Since a tension load of 0.53 MN is necessary to lift the pile weight the actual base resistance is therefore equa l to -0.82+0.53 = -0.29 MN. The existence of suction force at the pile base agrees with the suction pressures in the surrounding clay recorded by PPTs, as shown earlier. Immediately after the pile head is unloaded, Figure 4.20 reveals that the net base resistance is -0.50 MN which is nearly equal to σv Ap . This means that there is no force acting at the pile base. In other words, suction pressure at the pile base vanishes rapidly once the test load is released similarly to the observations of PPTs. As a result, a side resistance of 0.53 kN is mobilized to support the pile weight, as shown in Figure 4.20. The shearing due to the mobilization of the side resistance would result in positive excess pore water pressures in the clay adjacent to the pile shaft, as recorded by PPTs and shown in Figure 4.19. At 420 days after the pile head is unloaded, no significant change in load transfer is observed. From the test result, the suction force at the pile base is (0.29/1.01)x100 = 28.7% of the ultimate side resistance. This implies that the interpretation of ultimate side resistance of a tension pile in clay by subtracting the pile weight from the ultimate tension capacity would give a higher value than the actual value. For a very long pile, side resistance usually contributes as the major part of pile capacity. Thus, the error due to such interpretation method would be insignificant. 90 Chapter 4 Pile load test: Effect of different loading methods 4.8 O-cell pile load test T1C Test T1C was performed to simulate the loading mechanism similar to O-cell pile load test. A model O-cell pile was used in this test. The loading mechanism of this model pile allows the pile head being jacked into the clay during installation and the pile base being pulled up during load test. The detail of the model O-cell pile has been described in detail in Chapter 3. Five PPTs were installed in the clay and their positions are also given in Chapter 3. Unfortunately, two PPTs did not work during the test. The test sequence and rate of loading for test series C are provided in Table 4.1. The responses of load-displacement curve and pore water pressures in the clay during installation are similar to those of tests C1A and T1B. Hence, these will not be presented herein. The pile was installed to an embedment depth of 16.90 m at an applied load of 1.45 MN. After installation, the pile was unloaded and left for about 420 days to allow the clay to fully consolidate. Figure 4.21 presents the axial load distribution curves of the model O-cell pile during and after installation for test series C. Prior to pile installation (stage I), the pile is held at its base via the internal rod connected to the pile head connector. Therefore, the weight of the pile segment above each strain gauge is registered at this stage. Meanwhile, the load cell registers a tension load of 0.60 MN. It is noted that the total weight of the pile is 0.74 MN due to the weight of internal rod. The lower weight of the pile registered by the load cell at this stage is due to hydrostatic pressure at the pile base. At stage II, the pile stops penetrating into the clay by its self- weight. Subsequently, the pile head is pushed down by the pile head connector at stage III. Immediately after the release of installation load, the axial load distribution curves are similar to those of test series A. 91 Chapter 4 Pile load test: Effect of different loading methods Figure 4.22 shows the load distribution curves during and after installation of test series C after adjustment of pile weight according to the method described in Appendix. It should be noted again that the axial load observed at the ground surface is equal to the sum of applied force and Wp -σv Ap . The ultimate side and net base resistances of 1.40 MN and 0.25 MN are observed at an applied load of 1.45 MN. At 420 days after pile head is unloaded, a positive side resistance and a net base resistance of 0.19 MN and 0.06 MN are mobilized to support the weight of pile in excess of σv Ap . Figures 4.23(a) and (b) show the load-displacement curve and excess pore water pressure responses registered by various PPTs during pile load test T1C, respectively. The ultimate tension capacity of 1.77 MN was mobilized at the pile displacement of 4.72 cm which is larger than the pile displacement at ultimate load of test C1A. Similarly to the response of test T1B, the pile-soil stiffness decreases at a higher rate than that of test C1A after the applied load exceeds about half the ultimate load. Figure 4.23(b) shows that negative excess pore water pressures are generated during the test. At very large pile displacement, PPT 7 registers positive excess pore water pressure which is caused by shearing in compression unloading mode similar to that observed from load test T1B. Figure 4.24 shows the variation of excess pore water pressures with time after release of test load for load test T1C. The sudden change of excess pore water pressures from negative to positive values occurs after the pile base is completely unloaded at 2.5 days. This confirms that negative excess pore water pressures due to tension tests exist only in short term. After 420 days, the clay around the pile completely consolidates. 92 Chapter 4 Pile load test: Effect of different loading methods The load distribution curves corresponding to the test load pulling the pile base during and after test T1C are shown in Figure 4.25. As the pile base is pulled up, negative side resistance is mobilized along the pile-soil interface to resist the applied load. As a result, the pile experiences compression loads as revealed by the strain gauges while tension loads are recorded in the pile for conventional tension load test. It is noted that an applied tension load of 0.74 MN must be used to overcome the total weight of the pile prior to mobilizing shaft resistance. Therefore, an axial load recorded by the bottom strain gauge is equal to the difference in the magnitude of applied load and the pile weight. From Figure 4.25, the ultimate negative side resistance of 0.80 MN is observed at the applied load of -1.93 MN. The net base resistance can be back calculated from the force equilibrium of the pile as -1.93 + (0.74+0.80) = -0.39 MN. Immediately after release of test load, the load transfer is similar to that observed from test T1B. The negative net base resistance having a magnitude of about σv Ap is recorded. A positive side resistance of 0.74 MN is developed to hold the pile in equilibrium. No significant change of the load transfer is observed. 4.9 Effects of different loading methods on pile bearing capacity In this section, the results of pile load tests C1A, T1B and T1C whic h were conducted by different loading methods are compared and discussed. For load test C1A, the pile head is pushed down into the clay causing upward side resistance acting along the pile shaft. As a result, the pile experiences compression loads over the whole length. In contrast, the upward movement of the pile due to tension load for test 93 Chapter 4 Pile load test: Effect of different loading methods T1B mobilizes downward side resistance along the shaft. Therefore, tension loads are recorded along the pile. For test T1C, the test load applied directly to the pile base brings about the directions of pile movement and side resistance same as those of load test T1B. However, the pile is in compression for this test. Figure 4.26 compares the load-displacement characteristics of load tests C1A, T1B and T1C. It is shown in the figure that the compression pile load test (test C1A) requires less pile movement to mobilize ultimate pile bearing capacity than the tension tests (tests T1B and T1C) at a comparable pile embedment depth. The observed ultimate pile capacity of tests C1A, T1B and T1C from the figure are 1.80, 1.80 and -1.77 MN, respectively. Since the tension pile capacity of tests T1B and T1C is partly contributed by the pile weight, the load-displacement curves in Figure 4.26 are adjusted to eliminate the effect of pile weight. For tension tests T1B and T1C, the curve is adjusted by removing pile weight from the applied tension load. For compression test C1A, no adjustment needs to be carried out since the unit weight of the pile is approximately the same as the unit weight of the displaced clay. As a result, the soil resistance-pile displacement responses of tests C1A, T1B and T1C can be plotted as shown in Figure 4.27. It is noted that the piles in tests T1B and T1C weigh 0.53 and 0.74 MN, respectively. From Figure 4.27, the observed ultimate soil resistance of tests C1A, T1B and T1C are 1.80, -1.26 and -1.05 MN, respectively. It is thus shown that the soil resistance of compression pile C1A is significantly greater than that of tension piles T1B and T1C. Supposing that the ultimate base resistance for compression piles is equal to 9Cu, the ultimate side resistance for test C1A is calculated to be 1.25 MN can be calculated. If tension piles are assumed to have no base resistance, it is apparent that the downward side resistance of tension piles is the same as that of upward 94 Chapter 4 Pile load test: Effect of different loading methods resistance of compression piles. However, pore water pressure measurement around the pile of test T1B shows that suction pressure, which contributes to the base resistance to tension piles, is likely to be generated at the pile base. Therefore, a more accurate magnitude of side and base resistances should be obtained from axial load distribution curves of the piles. It is noted that the positive soil resistance at relative small pile displacement of tests T1B and T1C are actually the weight of pile. The figure shows that the weight of tension pile has to be overcome first before the mobilization of soil resistance. The cause that makes test T1C having a lower tension capacity than that of test T1B will be discussed later in this section. In the current design practice, it is usually assumed that the downward side resistance of a tension pile in clay is equal to the upward side resistance of an equivalent compression pile (Vesic, 1977, NAVFAC DM 7.2, 1986, Fleming et al., 1992, Tomlinson, 1994). However, the direct comparison between compression and tension pile load tests conducted at a comparable condition has been rarely found in literature. It is likely that this assumption is made on the comparison of a single empirical parameter reflecting an average unit side resistance of a pile, for example the one shown in Figure 2.11. Figure 2.11 shows the variation of adhesion factor α with average undrained shear strength of clay over the pile embedment depth. The data points are obtained from conventional compression and tension load tests of steel pipe piles, untapered precast concrete piles and cast- in-place drilled shafts tested at many sites. Unfortunately, the data points of the tension load tests are not clearly specified in the figure. Although the data ha ve a high degree of scattering, it is assumed that the upward and downward side resistances are identical. Therefore, a single correlation 95 Chapter 4 Pile load test: Effect of different loading methods curve for design purpose is drawn and shown in the figure. Since the adhesion factor α includes all factors affecting side resistance of piles which were tested at different site conditions, it may not be appropriate to make such conclusion. Table 4.4 reveals that on average, the downward side resistance of tests T1B and T1C are about 66% and 52% of the upward side resistance of test C1A. Figure 4.28 shows the variation of ultimate unit side resistance with depth for load tests C1A, T1B and T1C. The negative side resistances of load tests T1B and T1C are mirrored to the y-axis and re-plotted in the figure for comparison. It can be seen that the unit side resistance increases with depth for all types of load tests due to increasing overburden stress. However, the difference in the variation of unit side resistance with depth between compression pile and tension piles can be clearly observed from the figure. As compared to the results of tension piles, much greater magnitude of ultimate positive side resistance of test C1A is mobilized around the upper part of the pile shaft. In contrast, the ultimate negative unit side resistance of tension piles increases rapidly towards the pile base. Significantly higher magnitude of ultimate negative unit side resistance than that of ultimate positive unit side resistance at a given depth around the lower part of the shaft can be observed from Figure 4.28. It will be discussed later that the difference in the variation of ultimate unit side resistance with depth between compression and tension piles is due to the effect of loading direction. The rapid increase in unit negative side resistance towards the pile base of tension tests T1B and T1C could be attributed to the increase in effective stresses in the clay due to suction pressure created around the pile base. The similarity of the variation of unit side resistance with depth of test T1C to that of T1B reveals that loading direction controls the characteristic of the side resistance of pile. 96 Chapter 4 Pile load test: Effect of different loading methods Based on the results of an analytical study, De Nicola and Randolph (1994) reported that, for a relatively slender pile in sand, a significant lower side resistance of a pile under tensile loading than that of compressive loading would be obtained. This is attributed to the change in total stress field due to the direction of loading and the effect of Poisson’s ratio of the pile, which causes radial contraction and expansion of the pile when subjected to tension and compression, respectively. For such a pile, the tendency of the pile to contract radially leads to lower radial stresses, and thus lower side resistance. However, such an effect on side resistance of a pile in clay is not discussed. It may be prudent to investigate the effect of Poisson’s ratio effect on the pile in the present study by calculating the maximum possible radial expansion or contraction of the pile. For the pile having an axial stiffness of 9.596x103 MN, its compressive or tensile strain under an applied load of 2.00 MN (maximum test load in this study) is 2.08x10-2 %. Suppose the Poisson’s ratio of the pile material is 0.3, the maximum radial contraction or expansion of the pile is 0.047 mm. Thus, it is unlikely that such a tiny radial movement can affect the side resistance along the pile. Hence, it is possible that the direction of loading or pile movement governs the characteristic of side resistance of the pile. Figure 4.29 illustrates the stress paths of an element of overconsolidated clay adjacent to the pile wall prior to pile installation to failure under compression loading. It is noted that the t and s’ axes in the figure represent deviatoric and effective mean stresses, respectively. From the figure, a clay element at point A is loaded under extension mode due to radial expansion of cavity of the clay to failure during pile installation to critical state. The corresponding stress path is the path AB drawn from point A to point B on the critical state line (CSL). After installation, the mean effective stress of the clay increases as it consolidates. 97 Chapter 4 Pile load test: Effect of different loading methods Consequently, the void ratio of the clay decreases and the clay shear strength increases. The corresponding total stress path moves from point B towards point C at the end of consolidation (or point B’ to point C’ for effective stress). It is noted that negative deviatoric stress of point C indicates that the stress of the clay element in radial direction is higher than that in vertical direction, as confirmed by the analytical study by Randolph et al. (1979). When the pile is subjected to compression loading, the mobilized upward shear stress at the shaft causes the stress path of the clay element to move along the path CD or C’D’. It can be observed tha t there is rotation of principal stress as the stress path moves towards point D or D’ on the critical state line, i.e. change of the sign of deviatoric stress. On the other hand, for a tension pile, the mobilized downward shear stress will cause the deviatoric stress of the element to be more negative. As a result, the stress path will move towards the critical state line by not having the rotation of principal stress. Thus, it is evident in the theoretical point of view that the magnitude of ultimate upward side resistance for a compression pile is higher than the downward side resistance for an equivalent tension pile. This is confirmed by the experimental results of the present study. For Osterberg pile load test, it is usually assumed that the characteristic of upward and downward side resistances of a pile in clay are identical (Osterberg, 1989). However, it is shown by the experimental result that the characteristic of the side resistance obtained from Osterberg pile load test is similar to that of conventional tension test rather than that of conventional compression test. This is mainly due to the direction of loading as discussed earlier. Thus, the ultimate side resistance estimated from such an assumption would be significantly higher than the actual capacity as evidenced by the experimental results. 98 Chapter 4 Pile load test: Effect of different loading methods Besides loading direction, the position of load application on a pile can also affect the side resistance of the pile as observed by the lower tension capacity and ultimate negative side resistance of test T1C than that of test T1B. This may be due to the difference in the load transfer mechanisms between tests T1B and T1C. Unfortunately, the sampling rate of the strain meter used in the experiments is not fast enough to reveal the load transfer mechanism of the pile at small applied load. Thus, the development of side resistance along the pile with applied load cannot be investigated. It is believed that when a pile is loaded, the applied load is progressively transferred to the surrounding soil via side resistance from the point that the load is applied. Thus, the pile segment in the soil closer to the point of load application will be subject to higher load than the segment farther away. The pile segment closer to the point of load application will have a greater pile-soil movement due to greater pile elongation or contraction. As a result, the side resistance along the pile closer to the point of load application will be mobilized first. As the applied load increases, the side resistance is mobilized until its ultimate and residual resistances are reached. The side resistance along the pile farther from the point of load application is then progressively mobilized to resist the increasing applied load. The ultimate pile capacity is reached once the ultimate side resistance has been fully mobilized throughout the pile length. Figure 4.30 shows the idealized development of negative side resistance with applied tension load when the pile is loaded at the pile head and at the pile base, respectively. Figure 4.30(a) illustrates that side resistance is initially mobilized at the top part of the pile for a tension pile loaded at the pile head. At ultimate load, ultimate side resistance around the bottom part of the pile is mobilized together with large 99 Chapter 4 Pile load test: Effect of different loading methods negative excess pore water pressure generated in the clay. Owing to relatively larger pile-soil displacement, the side resistance around the top part of the pile is mobilized to its residual value. In contrast, Figure 4.30(b) shows that, for a tension pile pulled out at the pile base, the side resistance around the pile base reaches its residual value at ultimate load while that around the top part of the pile reaches its ultimate value. Since the side resistance at ultimate value has a higher magnitude than that at residual value, the tension pile pulled out at the pile base has a lower side resistance around the pile base than the pile pulled out at the pile head. This is evidenced by the lower side resistance at ultimate load of test T1C than that of test T1B as shown in Figure 4.28. Consequently, test T1C has a lower soil resistance than that of test T1B as shown in Figure 4.27. It is important to note that the experimental results of the present study are limited to only a full displacement pile. For a non-displacement pile such as bored pile, the radial stresses of the soil adjacent to the pile may be brought back close to their original values once concrete has been placed. The effect of rotation of princip al stress during loading may not be pronounced. The suggestion given by Osterberg (1989) may be valid for such a pile. This should be investigated further in future studies. 4.10 Base suction capacity test, test series D Test series D was performed on a model pile instrumented with a PPT at the pile base. The detail of this model pile is described in Chapter 3. The purpose of this test series is to investigate the significance of suction pressure at the pile base towards the load carrying capacity in view of the observed made in test T1B. Three tension 100 Chapter 4 Pile load test: Effect of different loading methods pile load tests were conducted at different pile embedment depths at the same rate of loading as tests T1B and T1C. 4.10.1 Responses of pile and pore water pressure during installation The pile was penetrated into the clay at a rate of 3.0 mm/sec until a pile embedment depth of 6.52 m. The total weight of the pile was 0.62 MN. The maximum installation load was 0.16 MN, as observed from the load-displacement curve shown in Figure 4.31(a). Figure 4.31(b) shows that, at the end of installation, a maximum pore water pressure of 215 kPa was measured at the pile base. After the excess pore water pressures due to pile installation have fully dissipated, the measured hydrostatic pore water pressure at the pile base is 115 kPa as shown in Figure 4.31(c). Therefore, the maximum excess pore water pressure at the pile base is determined to be 215-115 = 100 kPa. This is 3.98 times the undisturbed undrained shear strength of clay at the level of pile base. Since the in-situ effective overburden stress at 6.52 m deep is 40.36 kPa, the maximum excess pore water pressure at the pile base is found to be 2.48 times higher than the in-situ effective overburden pressure. This conforms to the field measurement made by Hwang et al. (2001). 4.10.2 Experimental procedure and results Once the clay has fully consolidated, tension load test T1D was conducted at 0.42 mm/min (model scale) till a pile displacement of half pile diameter was reached. Subsequently, the test load is partially unloaded and sustained for about four days prior to being fully released. Then, the pile was jacked to a deeper depth of 9.58 m 101 Chapter 4 Pile load test: Effect of different loading methods and load test T2D was conducted following the same procedure as test T1D. The test procedure was then repeated for load test T3D at a pile embedment depth of 17.02 m. Figures 4.32(a) to (c) present the variations of applied load at the pile head and base suction force which is the measured suction pressure times the pile area with time for load tests T1D, T2D and T3D, respectively. Assuming force equilibrium on the pile, the variation of the load resistance provided by pile weight and side resistance with time are calculated and plotted in the figures. It can be established that after the ultimate load is reached, the side resistance of a tension pile generally remains constant while the applied load still increases gradually due to increasing suction pressure at the pile base. Once the test load is partially unloaded and sustained, the suction force at the pile base for each lo ad test dissipates rapidly within a few days. Table 4.5 summarizes the results of pile load tests for test series D. The ultimate side resistance for load tests T1D and T2D, which have an embedment depth/pile diameter ratio (D/d) lower tha n 10, is quite little. This is attributed to the relatively low ultimate unit side resistance for a tension pile having a shallow embedment depth as shown in Figure 4.28. Therefore, the test load is mainly mobilized to overcome the pile weight. Table 4.5 reveals that the ultimate side resistance of load test T3D increases significantly beyond that of load test T2D. This is attributed to the increase in pile embedment depth and hence the increase in average radial stress on the pile. Table 4.5 also shows that at ultimate load, the suction force at the pile base increases with increasing pile embedment depth. It is likely that the normalized suction pressure with undrained shear strength of clay at the pile base (∆Uult /Cub) increases towards a value of -3.00 when the D/d ratio of the pile is larger than 10. 102 Chapter 4 Pile load test: Effect of different loading methods Hence, it can be deduced that the magnitude of suction pressure at the base of a tension pile is about 3 times the undrained shear strength of the clay at the pile base. Figure 4.33 plots the variation of normalized suction pressure at the pile base (∆Uult /Cubase) with normalized pile displacement (δ/d) for tests T1D, T2D and T3D. It can be observed that the ∆Uult /Cubase ratios of all load tests reach their maximum value at a δ/d value of approximately -0.3. At such a large pile displacement, the observed maximum ∆Uult /Cubase ratios are within the range of 5.5-6.9 times Cubase. This observation agrees with the finding of Bemben and Kupferman (1975) who reported that a suction pressure at the pile base of 6 times Cubase is a significant holding component of an anchor fluke in clay at ultimate condition. 4.11 Comparison between behaviour of piles installed in sand and clay In a previous study by Goh (2000), the behaviour of a single pile in sand under similar loading conditions was investigated. The installation and load testing were conducted using a load-controlled loading method. The properties of sand used in the experiment and the detail of experimental results are presented by Yet (1998) and Goh (2000), respectively. In this section, the results of conventional compression (AC1), conventional tension (AT1), and O-cell (JT1) pile load tests in sand are compared with those in clay obtained from the present study. Figure 4.34 compares the load-displacement curves of load tests AC1, AT1 and JT1. Similarly to a pile in clay, it is observed that the tension pile in sand reaches its ultimate load at higher pile displacement than that of the compression pile. Furthermore, it is shown in Figure 4.34 that the load carrying capacity of the 103 Chapter 4 Pile load test: Effect of different loading methods compression pile is much higher than that of the tension pile. This is mainly due to the contribution of large base resistance (5.13 MN) of the compression pile in sand as shown the load distribution curves along the pile fo r test AC1 in Figure 4.35. It is noted that the axial loads registered by load cell is excessive because of the weight of brass connector at the pile head. After pile installation, it is found that there are residual loads left in the pile. Upon unloading the test load, the residual loads along the pile increase. In contrast, for a floating pile in clay, Figure 4.12 reveals that the pile derives its load capacity mainly from the side resistance. No residual loads along the pile are observed after pile installation. The results of these tests support the conclusion of the analytical study carried out by Poulos (1987) that the higher the base resistance, the greater residual stresses would occur after installation. However, the results of the present study sho w that, upon the release of first compression test load, residual compression loads along the pile may occur and increase as the surrounding clay consolidates. Figures 4.36 and 4.37 present the load distribution curves of the pile for load tests AT1 and JT1, respectively. Similarly to the response of the pile in clay, the pile in test AT1 experiences tension while that of test JT1 is subjected to compression. A linear interpolation of the axial load distribution curves to the pile base of test AT1 reveals that the pile has no base resistance. This implies that suction pressures could not be generated around tension piles in sand due to the high permeability of sand. Figure 4.38 plots the variation of ultimate unit side resistance with depth for tests AC1, AT1 and JT1. The mirrored variations of tests AT1 and JT1 with the y-axis are also plotted in the figure for comparison with that of test AC1. Figure 4.38 shows that the unit side resistance of conventional compression test AC1 is higher than that 104 Chapter 4 Pile load test: Effect of different loading methods of tests AT1 and JT1, especially at greater depths. However, the variation pattern of unit side resistance with depth for test AC1 is similar to that for tension tests AT1 and JT1, unlike that for test C1A in clay. This may be due to the difference in behaviour of sand from that of clay with interaction between solid and liquid phases during shear. The side resistance of tension pile AT1 is found to be only about 58% of compression pile AC1. Similarly, the experimental results of the present study shows that the tens ion pile T1B in clay has an ultimate side resistance about 66% of that of the compression pile C1A. Thus, it is suggested by the experimental results that the negative side resistance of a jacked-in pile in clay or sand is about 60-70% of that of the positive side resistance. This is attributed to the effect of loading direction as discussed earlier. It is also shown in Figure 4.38 that the unit side resistance of test JT1 is slightly lower than that of test AT1 although the pile in test JT1 is subjected to compression. This implies that the effect of Poisson’s ratio of the pile does not play an important role on the unit side resistance of a pile in sand as well as that in clay. Instead, this is caused by the different load transfer mechanisms due to different points of loading application on the pile as discussed earlier. 4.12 Conclusion Several centrifuge model tests were conducted at 100g to simulate the loading mechanisms of conventional compression (C1A), conventional tension (T1B), and partially simulated O-cell pile (T1C). An additional model test (test series D) was performed to examine the presence of suction pressure at the pile base. Based on the results of the previous study in sand (Goh, 2000) and the present study in clay, several findings were made and summarized as follows: 105 Chapter 4 Pile load test: Effect of different loading methods Ø During installation of pile in clay, the clay close to the pile is severely disturbed. Excess pore water pressures are generated in the clay by shear stresses mobilized during pile installation. The excess pore water pressure is found to be highest at the pile shaft and decreases rapidly with increasing radial distance from the pile. During installation, the excess pore water pressure is highest at the pile base and can be as high as 4 times the undrained shear strength of the clay at pile base. Such high pressure could be higher than the in-situ effective overburden pressure of the clay at the pile base. Ø The excess pore water pressures generated during installation or load test dissipate outward from the pile in both radial and vertical directions for a pile having pile embedment depth/pile diameter ratio up to about 11. At any given time after the pile is unloaded, the clay element closer to the pile base has a higher degree of consolidation. Ø For a pile in soft clay, the results of this study show that no significant residual loads along the pile are observed due to the relatively low base resistance of the pile. However, residual loads due to drag-down load along the pile-soil interface are observed as the clay consolidates after releasing the first compression test load. In contrast, a pile installed in sand shows significant residual loads immediately after unloading the pile. Ignoring residual loads in the interpretation could result in an overestimation of side resistance for a compression pile and vice versa for a tension pile. Ø For a compression pile load test, the observed ultimate unit side resistance agrees well with the prediction proposed by Randolph and Murphy (1985). Ø For a tension pile load test, the pile displacement at ultimate load is higher than that of a compression pile. 106 Chapter 4 Pile load test: Effect of different loading methods Ø The variation pattern of ultimate unit side resistance along a tension pile in clay is found to be different from that of a compression pile. However, the variation pattern of ultimate unit side resistance of a compression pile in sand is found to be similar to that of a tension pile. The results of the present study suggest that the average unit side resistance of a tension pile in clay is about 50 to 65% of that of a compression pile while Goh (2000) found that it is about 58% for a pile in dry dense sand. This is attributed to the difference in stress paths of soil elements around the pile when subjected to loading due to different loading directions. Ø The results of an additional model test suggest that when a tension pile is subjected to short-term loading, suction pressures generated around the pile considerably contributes to the base resistance of the pile, especially for a pile having embedment depth/pile diameter lower than 11. The base resis tance of a tension pile at ultimate load may be determined as equal to 3 times the undrained shear strength of the clay at the pile base. However, it is found that the suction pressures can fully dissipate within a few days upon the release of the tension load. In the interpretation of the result of a tension pile load test in clay, the ultimate side resistance greater than the actual side capacity would be a result if the presence of suction pressure at the pile base is ignored. Ø For the O-cell pile load test, the responses of the pile and surrounding clay are similar to those of a conventional tension pile load test although the pile shaft is subjected to compression like that of a compression pile. This confirms that, for a displacement pile, the direction of loading on a pile controls the characteristic of side resistance rather than the Poisson’s ratio of pile. Based on the assumption that the downward and upward side resistances of a full- 107 Chapter 4 Pile load test: Effect of different loading methods displacement pile in clay are identical, the interpretation of an O-cell pile load test may give a lower side resistance than that of an equivalent compression pile. 108 Chapter 4 Pile load test: Effect of different loading methods Table 4.1 Summary of centrifuge model tests. Rate of penetration or uplift (4) Test Sequence(1) Test s eries mm/min Compression test Tension test A C1A-C2A-T1A-T2A-C3A -T3A 0.48 0.42 B T1B-T2B-C1B-C2B 0.48 0.42 C(2) T1C-T2C-C1C-T3C 0.48 0.42 D(3) T1D-T2D-T3D - 0.42 E C1E-C2E-C3E-T1E-T2E-T3E 0.60 0.48 F T1F-T2F-T3F-C1F-C2F-C3F 0.60 0.48 (1) The notation of each load test comprises 3 letters. The first letter stands for type of load tests, C = Compression load test and T = Tension load test. The second letter identifies the number of times of a particular loading type. The last letter identifies the test series. For example, T1A is the first tension load test in test series A. (2) Partially-simulated O-cell test. (3) Pile base suction capacity test. (4) the rates are within the range suggested by ASTM D 1143 and ASTM D 3689 standards Table 4.2 Summary of pile load distribution at various stages during and after pile installation for test series A. MN Resultant downward force MN Resultant upward force MN 0.53 -0.47 0.06 0.05 0.40 0.53 0.00 0.53 0.50 0.88 0.68 0.53 1.43 1.96 1.56 at maximum installation load -0.45 0.55 0.53 0.00 0.53 0.10 at 420 days after pile head unloaded Qs Qb Wp P MN MN MN I 0.00 0.05 II 0.10 III IV Stage Remark 109 Chapter 4 Pile load test: Effect of different loading methods Table 4.3 Summary of pile load distribution at various stages during and after pile installation for test series A. (after pile weight adjustment) MN Resultant downward force MN Resultant upward force MN 0.53 -0.47 0.06 0.05 0.12 0.53 0.00 0.53 0.56 0.18 0.51 0.53 1.46 1.99 1.97 at maximum installation load 0.00 0.51 0.53 0.00 0.53 0.54 at 420 days after pile head unloaded Qs Qb,net σv Ap Wp P MN MN MN MN I 0.00 0.00 0.05 II 0.16 0.28 III 1.28 IV 0.03 Stage Remark Table 4.4 Summary of results for load test C1A, T1B and T1C. Embedment depth of pile Pile weight m MN Ultimate load observed from load distribution curve MN C1A 16.2 0.53 T1B 16.0 T1C 16.9 Test Ultimate side resistance Ultimate net base resistance MN MN 2.07 1.54 0.53 0.53 -1.83 -1.01 -0.29 0.74 -1.93 -0.80 -0.39 110 Chapter 4 Pile load test: Effect of different loading methods Table 4.5 Summary of results of test series D. Test No. D/d Pult MN Qb,net MN Qs MN Qb,net/Qs % ∆Uult kPa Cu b kPa ∆Uult /Cu b T1D 4.43 -0.73 -0.12 0.00 - -62.7 25.2 -2.48 T2D 6.39 -0.84 -0.14 -0.07 193.8 -81.4 28.7 -2.83 T3D 11.35 -1.50 -0.18 -0.70 25.8 -102.8 35.2 -2.92 Note : D = embedment depth of pile, d = pile diameter, Pult = ultimate tension load, Qb,net = measured suction pressure at the pile base at ultimate load x pile cross sectional area, Qs = ultimate side resistance of pile, ∆Uult = excess pore water pressure at pile base at ultimate load, Cub = undrained shear strength of clay at pile base. 111 [...]... subjected to compression and tension load and its surrounding clay using centrifuge modeling Ø To compare the behaviour of side resistance of a pile subjected to tension load at the pile base (similar to Osterberg pile load test setup) with that of a pile subjected to conventional compression and tension loads at the pile head Ø To study the effect of preceding loads on the pile in subsequent loading... previous centrifuge modeling studies on pile foundation 2.2 Behaviour of pile in clay The performance of pile is mainly affected by the characteristic s of the surrounding soils For a pile installed in clay, its responses can be divided into 4 stages as follows: 1) Initial condition of clay prior to pile installation, 2) Pile installation effects and disturbances 3) Re-consolidation of clay around pile. .. history of driving versus jacking), and loading conditions (set-up time after driving, rate of axial loading etc.) In this chapter, the literatures relating to the behaviour of a pile in clay, especially displacement pile, is reviewed Since pile performance can be greatly affected by the installation method of the pile, the literatures on the behaviour of pile and its surrounding clay prior to, during... robust tool in geotechnical engineering to study the behaviour of pile- soil interaction with lower cost than full- scale load testing 1.2 Scope of research The research work discussed in this thesis is the extension of the work done by Goh (2000) who studied the behaviour of single pile in sand It covers the investigation of the pile behaviour in clay as follows: Ø To study the behaviour of a single pile. .. diagram of components of model pile Figure 3.8 Schematic diagram showing loading mechanism of conventional model pile subjected to (a) compression load at pile head (b) tension load at pile head Figure 3.9 Schematic diagram showing loading mechanism of model O-cell pile when subjected to (a) tension load (b) compression load Figure 3.10 Location of strain gauges on model pile x Figure 3.11 Strain gauge instrumentation... stresses and their effects are most significant for piles in sand, but do not significantly influence the response of piles in clay 2.2.3 Re-consolidation of clay around pile after installation Cummings, Kerkhoff and Peck (1950) measured the water content of soft clay at various radii from a large cluster of piles over a 1- year period after pile driving It was found that the re was horizontal migration of. .. undrained shear strengt h based on the field measurements from Dennis and Olson (1983), and Stas and Kulhawy (1984) The results of driven and bored piles in compression and tension are plotted in the figure which shows that the adhesion factor decreases with increasing average undrained shear strength Since the effects of pile installation, time effect, pile geometry and length are collapsed into only... only one factor, a poor correlation is noted in Figure 2.11 This may lead to the conclusion that the unit side resistance of compression piles equals to that of tension piles Olson (1984) analyzed eleven cases of compression and tension pile load tests in clay from 5 different sites in USA by using the API (American Petroleum Institute) design guideline In 10 out of the 11 cases, the ratio of calculated/measured... after pile installation could lead to overestimation of ultimate side resistance and underestimation of ultimate base resistance for compression piles and vice versa for tension piles From the results of model tests of long extensively instrumented piles in deep beds of sand, Hanna and Tan (1973) 12 Chapter 2 Literature review concluded that the shape and magnitude of the load-settlement curve of the piles... soil caused by installation, is not sensitive to in- situ overconsolidation ratio of the soil Based on the results of a series of pile load tests of intensively instrumented closed-ended steel pile installed in heavily overconsolidated London clay, Bond and Jardine (1995) proved that the shaft capacity of displacement piles in London clay is governed by an effective stress interface sliding criterion, ... behaviour of single pile in sand It covers the investigation of the pile behaviour in clay as follows: Ø To study the behaviour of a single pile when subjected to compression and tension load and its... Osterberg pile load tests; (b) to investigate the effect of pile loading history and loading sequence on the pile performance; and (c) to investigate the effect of maintained dead load on pile behaviour... experimental investigation into the behaviour of pile subjected to compression and tension in clay The investigation was carried out using the National University of Singapore geotechnical centrifuge