AN ABSTRACT OF THE THESIS OF Nasim Adami for the degree of Master of Science in Civil Engineering presented on October 28, 213. Title: Development of an ACIP Pile-Specific Load-Displacement Model. Abstract approved: Armin W. Stuedlein Augered cast-in-place piles, also known as ACIP piles, have been used for more than seven decades in the United States and have gained in popularity due to their relatively quick installation and cost-effectiveness. Owing to the reduced impact on the neighboring environment as compared to some other deep foundation installation methods, ACIP piles are appropriate for use in urban areas. Although there has been an increase in application of ACIP piles, relatively little research on this type of pile has been performed as compared to similar deep foundations, such as drilled shafts. The insufficient experimental work on ACIP pile behavior and lack of ACIP pile specific loaddisplacement models have led practicing engineers to use the results and methodologies from drilled shafts. An example of this is the use of t-z and q-z based load transfer models
from drilled shaft-specific relationships to estimate the load-displacement behavior of ACIP piles. Such applications can result in an underestimation of shaft resistance and consequently disagreement between the predicted and measured load-displacement behavior of the ACIP piles. This thesis evaluates the ability of currently used load-displacement models to estimate the measured load-displacement behavior of ACIP piles. Also, a new empirically-based ACIP pile-specific t-z model is proposed that, in combination with the O Neill and Reese (1999) q-z model and ACIP pile-specific toe bearing resistance model, forms an ACIP pile specific load-displacement model. Experiments of instrumented ACIP piles installed in the granular soils of Western Washington were used to develop the ACIP pile specific t-z model. Comparison between the results from the currently used load-displacement models with the proposed model showed that the proposed model provides an improvement in the prediction of the load-displacement behavior of ACIP piles. Finally, an analysis of variability is performed using the Monte Carlo Simulation with the sample probability distributions of the uncertain variables in load-displacement model. These analyses result in provide a set of possible loads for a number of common service level displacements, which are reported as cumulative density function (CDF) curves. The CDF curves for loads corresponding to a displacement considered can be a useful tool in design procedure of ACIP piles.
Copyright by Nasim Adami October 28, 213 All Rights Reserved
Development of an ACIP Pile-Specific Load-Displacement Model by Nasim Adami A THESIS submitted to Oregon State University in partial fulfillment of the requirements for the degree of Master of Science Presented October 28, 213 Commencement June 214
Master of Science thesis of Nasim Adami presented on October 28, 213. APPROVED: Major Professor, representing Civil Engineering Head of the School of Civil and Construction Engineering Dean of the Graduate School I understand that my thesis will become part of the permanent collection of Oregon State University libraries. My signature below authorizes release of my thesis to any reader upon request. Nasim Adami, Author
ACKNOWLEDGEMENTS I would like to express my very great appreciation to my major advisor Dr. Armin Stuedlein for his guidance along this research and providing an environment for me to realize my abilities. I also want to thank my graduate committee members, Dr. Sessions, Dr. Mason, and Dr. Leshchinsky for their helpful suggestions. I would like to offer my special thanks to faculty and staff of the Oregon State University School of Civil and Construction Engineering for their dedication and support. Finally, I am particularly grateful for the support given by all my friends and graduate students of Geotechnical Engineering.
TABLE OF CONTENTS Page 1 Introduction... 1 1.1 Problem Statement... 1 1.2 Research Outline... 1 2 Literature Review... 3 2.1 Introduction... 3 2.1.1 Drilled Shaft Foundations... 5 2.1.2 Augered Cast-in-Place Piles... 5 2.2 Bearing Capacity of Drilled Deep Foundations... 7 2.2.1 Estimation of Bearing Capacity of Drilled Shaft Foundations... in Granular Soils... 8 2.2.2 Estimation of Bearing Capacity of Augered Cast-in-Place Piles... 16 2.3 Prediction Models of Pile Settlement in Granular Soils... 21 2.3.1 Vesic (1977) Settlement Method... 21 2.3.2 Introduction to t-z and q-z Models... 24 2.4 Summary... 34 3 Research Objectives... 35 4 Database of Loading Test Cases... 37 4.1 Static Pile Load Test Data for 1997 Test Series... 37 4.1.1 General... 37
TABLE OF CONTENTS (Continued) Page 4.1.2 Development of Soil Profile and Soil Parameters for the 1997 Test Series... 4 4.2 Load Transfer and Static Load Test Results for 29 Test Series... 45 4.2.1 General... 45 4.2.2 Development of Soil Profiles and Soil Parameters for the 29 Test Series.. 46 4.2.3 Description of the instrumentation of Piles in 29 Test Series... 51 4.2.4 Determination of Composite Modulus... 51 4.2.5 Load Transfer... 53 4.3 Summary... 57 5 Development of a Load Transfer Model for Augered Cast-in Place Piles... 58 5.1 Introduction... 58 5.2 Development of the t-z Model... 58 5.2.1 Calculation of Pile Shear Stress, t... 59 5.2.2 Calculation the Relative Soil-Pile Movement, z... 61 5.2.3 Normalizing Parameters... 63 5.2.4 Proposed t-z Model Parameters... 64 5.3 Selected q-z Model... 72 5.4 Evaluation of the Proposed Load-Transfer Model...72 5.5 Comparison of Results to 29 and 1997 Test Series... 74 5.6 Summary... 87
TABLE OF CONTENTS (Continued) Page 6 Analysis of the Proposed t-z Model by Use of Monte Carlo Simulations... 88 6.1 Introduction... 88 6.2 Monte Carlo Simulation in Geotechnical Engineering... 89 6.3 Input Variables and Probability Distribution of Uncertain Parameters... 91 6.3.1 Uncertain Variables of the Proposed Prediction Model... 91 6.3.2 Probability Distribution of Uncertain Variables... 92 6.3.3 Monte Carlo Simulation of Random Variables... 97 6.4 Monte Carlo Simulation Correction... 11 6.5 Discussion of the Results... 18 6.6 Summary... 112 7 Summary and Conclusion... 113 7.1 Summary... 113 7.2 Conclusions... 115 7.3 Suggestions for Future Study... 117 References... 118
TABLE OF CONTENTS (Continued) Page APPENDICES... 127 Appendix A: MATLAB Code for Generating the Random Variables... 128 Appendix B: MATLAB Code for Calculating the Load-Displacement Response of ACIP Piles Using t-z and q-z Models... 131 Appendix C: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation... 135 Appendix D: Detailed Monte Carlo Simulation Results... 147
LIST OF FIGURES Figure Page Figure 2.1: Situations in which deep foundations may be needed. (a) weak soil layer with a dense stratum within a reasonable depth, (b) weak soil layer without a dense stratum, (c) uplift forces, (d) and (f) horizontal force resist by a single pile, (e) horizontal forces and moments resist by pile group,(g) erosion around the footing, (h) liquefaction susceptible layer, (i) deep foundation acting as a fender system, (j) probability of future excavation in adjacent region, (k) swelling soil layer(fhwa 26, after Vesic 1977).... 4 Figure 2.2: ACIP construction steps, (1) drilling, (2) concrete injection, (3) inserting the reinforcements (after Mainwinch 213).... 6 Figure 2.3: A conceptual sketch of resisting parameters (Das 24).... 8 Figure 2.4: Bearing capacity factor from Prakash and Sharmad (199).... 1 Figure 2.5: The relationship between SPT N values and the toe bearing resistance, after Neely 1991.... 17 Figure 2.6: The relationship between SPT N values and the toe bearing resistance, with comparison between the Stuedlein et al. (212) expression with FHWA method, after Stuedlein et al. (212).... 18 Figure 2.7: Average β-coefficient for use in Equation 2.15 (Neely 1991).... 19 Figure 2.8: The relationship between depth of embedment and β coefficient. The data points consist of experiments conducted in the Stuedlein et al. (212) and Neely (1991) experiment. after Stuedlein et al. (212).... 2 Figure 2.9: The load distribution factor for different typical loading cases: (a) uniform distribution, α =.5, (b) extreme case of linear distribution, α =.67, (c) extreme case of linear distribution, α =.33. After Vesic (1977).... 23 Figure 2.1: Schematic concept used in t-z method modeling (after FHWA 213).... 25 Figure 2.11: Concentric cylinder model for settlement analysis of axially loaded piles (modified from Randolph and Wroth 1978).... 26
LIST OF FIGURES (Continued) Figure Page Figure 2.12: Linear t-z curve obtained using Randolph and Wroth (1978) (after FHWA 213).... 28 Figure 2.13: (a) Normalized load transfer curves for clay proposed by Coyle and Reese 1966 and (b) Normalized skin friction curves for sand proposed by Coyle and Solaiman 1967.... 29 Figure 2.14: Normalized skin friction curve for clay and sand proposed by Vijayverjia (1977).... 3 Figure 2.15: Normalized curves of load transfer in side friction vs. settlement for drilled shafts in sand (O'Neill and Reese 1999).... 31 Figure 2.16: Normalized q-z curve for clay and sand (Vijayverjia 1977).... 32 Figure 2.17: Normalized curves of load transfer in toe bearing vs. settlement for drilled shafts in sand (after O'Neill and Reese 1999).... 33 Figure 4.1: Comparison between actual load-displacement and predications from the stability plot method, along with the ultimate loads proposed by Gurtowski (1997)... 39 Figure 4.2: Subsurface information for test pile 1997-1: (a) SPT N-value, (b) unit weight, (c) friction angle, (d) soil profile.... 41 Figure 4.3: Subsurface information for test pile 1997-2: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 42 Figure 4.4: Subsurface information for test pile 1997-3: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 42 Figure 4.5: Subsurface information for test pile 1997-4: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 43
LIST OF FIGURES (Continued) Figure Page Figure 4.6: Subsurface information for test pile 1997-5: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 43 Figure 4.7: Subsurface information for test pile 1997-6: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 44 Figure 4.8: Comparison between actual load-displacement and the ultimate loads proposed by Gurtowski (29)... 47 Figure 4.9: Subsurface information for test pile 29-1: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 48 Figure 4.1: Subsurface information for test pile 29-2: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 48 Figure 4.11: Subsurface information for test pile 29-3: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 49 Figure 4.12: Subsurface information for test pile 29-4: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 49 Figure 4.13: Subsurface information for test pile 29-5: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 5 Figure 4.14: Subsurface information for test pile 29-6: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.... 5 Figure 4.15: Calculations of tangentmodulus for all strain gages on pile 29-4 and demonstration of their approach to the lower limit presented by strin gages near pile head, after Stuedlein and Gurtowski (212).... 52 Figure 4.16: Tangent modulus slope for 29 Test Series, adopted from Stuedlein and Gurtowski (212).... 53
LIST OF FIGURES (Continued) Figure Page Figure 4.17: An example of the interpolation of pile 29-6: (a) as observed, (b) after interpolation.... 54 Figure 4.18: Interpolated load transfer curve for test pile 29-1.... 54 Figure 4.19: Interpolated load transfer curve for test pile 29-2.... 55 Figure 4.2: Interpolated load transfer curve for test pile 29-3.... 55 Figure 4.21: Interpolated load transfer curve for test pile 29-4.... 56 Figure 4.22: Interpolated load transfer curve for test pile 29-5.... 56 Figure 4.23: Interpolated load transfer curve for test pile 29-6.... 57 Figure 5.1: Schematic illustration showing definitions of depths and corresponding loads from load transfer curves in the calculation procedure.... 61 Figure 5.2: A Schematic illustration of P d,i and L d,i product, as an area on a load transfer curve.... 63 Figure 5.3: Comparison of experimental and fitted t-z curves for pile 29-1... 66 Figure 5.4: Comparison of experimental and fitted t-z curves for pile 29-2... 67 Figure 5.5: Comparison of experimental and fitted t-z curves for pile 29-4... 69 Figure 5.6: Comparison of experimental and fitted t-z curves for pile 29-5... 69 Figure 5.7: Comparison between the experimentally-derived coefficient a and the proposed relationship.... 71 Figure 5.8: Comparison between the experimentally-derived coefficient b and the proposed relationship.... 71
LIST OF FIGURES (Continued) Figure Page Figure 5.9: Comparison between the measured load transfers and those approximated by the proposed t-z model for piles, (a) 29-1, (b) 29-2, (c) 29-4, (d) 29-5.......73 Figure 5.1: Comparison between field measurements and predictions of proposed model: (a)test pile 29-1, (b) Test pile 29-2, (c) Test pile 29-4, (d) Test pile 29-5.... 75 Figure 5.11: Comparison between field measurements and predictions of proposed model for: (a) Test pile 1997-1, (b) Test pile 1997-2.... 76 Figure 5.12: Comparison between field measurements and predictions of proposed model for driven grout piles: (a) Test pile 29-3, (b) Test pile 29-6.... 78 Figure 5.13: Point by point comparison of bias with respect to normalized head displacement, for the 29 Test Series... 79 Figure 5.14: Point by point comparison of bias with respect to normalized head displacement, for the 1997 Test Series... 8 Figure 5.15: Comparison between the load-displacements produced by the proposed model with models proposed by Vijayvergia (1977) and O Neill and Reese (1999): (a) ACIP pile 29-1, (b) ACIP pile 29-2, (c) DG pile 29-3.... 83 Figure 5.16: Comparison between the load-displacements produced by the proposed model with models proposed by Vijayvergia (1977) and O Neill and Reese (1999): (a) ACIP pile 29-4, (b) ACIP pile 29-5, (c) DG piles 29-6.... 84 Figure 5.17: Comparison between the load-displacements produced by the proposed model with models proposed by Vijayvergia (1977) and O Neill and Reese (1999), ACIP piles: (a) 1997-1, (b) 1997-2, (c) 1997-3.... 85 Figure 5.18: Comparison between the load-displacements produced by the proposed model with models proposed by Vijayvergia (1977) and O Neill and Reese (1999), ACIP piles: (a) 1997-4, (b) 1997-5.... 86
LIST OF FIGURES (Continued) Figure Page Figure 6.1: Cumulative distribution function for the bias of the a coefficient.... 94 Figure 6.2: Cumulative distribution function for the bias of the b coefficient.... 95 Figure 6.3: Cumulative distribution function for the bias of the.... 96 Figure 6.4: The load displacement predictions produced by MCS and the proposed prediction model for 29 Test Series... 98 Figure 6.5: The load displacement predictions produced by MCS and the proposed prediction model for 1997 Test Series... 1 Figure 6.6: The load displacement predictions produced by corrected MCS and the proposed model for 29 Test Series... 15 Figure 6.7: The load displacement predictions produced by corrected MCS and the proposed model for 1997 Test Series... 17 Figure 6.8: Cumulative density functions of loads for four displacements of 5, 1, 15 and 25mm of 29 Test Series... 11 Figure 6.9: Cumulative density functions of loads for four displacements of 5, 1, 15 and 25 mm of 1997 Test Series... 111
LIST OF TABLES Table Page Table 2.1: Approximate φ' f φ' values for the interface between deep foundations and soil (after Kulhawy et al. 1983 and Kulhawy 1991).... 13 Table 2.2: Approximate ratio of coefficient of lateral earth pressure after construction to that before construction (after Kulhawy et ai. 1983 and Kulhawy 1991).... 14 Table 2.3: Values for C p coefficient (after Vesic 1977).... 23 Table 4.1: Geometric details and locations of test piles for the 1997 Test Series, after Gurtowski (1997)... 37 Table 4.2: Criteria for the estimation of unit weight, adopted from ODT Geotechnical Design Manual (29).... 41 Table 4.3: Geometric details and locations of test piles for the 29 Test Series, after Gurtowski (29)... 45 Table 5.1: Comparison between predicted ultimate loads with those by Gurtowski (1997, 29).... 81 Table 5.2: A comparison between the load mean bias obtained from the proposed model by those created by O Neill and Reese (1999) and Vijayvergia (1977) models.... 82 Table 6.1: Comparison of number of realization from Equations 6.1 and 6.2, from Phoon (28).... 9 Table 6.2: Statistical characteristics of distribution fitting for the a coefficient.... 94 Table 6.3: Statistical characteristics of distribution fitting for the b coefficient.... 95 Table 6.4: Statistical characteristics of distribution fitting for the.... 96 Table 6.5: Statistical characteristics of each actual displacement for test pile 29-6... 12
LIST OF TABLES (Continued) Table Page Table 6.6: Statistical characteristics of corrected MCS results.... 13 Table 7.1: The comparison of the results obtained using Vijayvergia (1977) and O Neill and Reese (1999) models.... 115 Table 7.2: The average mean bias in load and COV in load bias results from the proposed load-displacement model.... 116
LIST OF APPENDIX TABLES Table Page Table C. 1: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 29-1.... 136 Table C. 2: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 29-2.... 137 Table C. 3: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 29-3.... 138 Table C. 4: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 29-4.... 139 Table C. 5: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 29-5.... 14 Table C. 6: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 29-6.... 141 Table C. 7: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 1997-1.... 142 Table C. 8: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 1997-2.... 143 Table C. 9: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 1997-3.... 144 Table C. 1: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 1997-4.... 145 Table C. 11: Comparison of Deterministic Coefficients Results with Results from Random Variable Generation, pile 1997-5.... 146
LIST OF APPENDIX TABLES (Continued) Table Page Table D. 1: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 29-1. 148 Table D. 2: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 29-2.... 149 Table D. 3: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 29-3.... 15 Table D. 4: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 29-4.... 151 Table D. 5: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 29-5.... 152 Table D. 6: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 29-6.... 152 Table D. 7: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 1997-1.... 153 Table D. 8: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 1997-2.... 153 Table D. 9: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 1997-3.... 153 Table D. 1: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 1997-4.... 154 Table D. 11: The detailed results of the corrected Monte Carlo Simulation for all the head displacements of test pile 1997-5.... 154
LIST OF SYMBOLS Applied load, constant Load resisted by shaft Load resisted by toe Area of the pile Vertical effective stress Horizontal effective stress Bearing capacity factor Shape factor Depth factor D t K K The diameter of the pile toe Ultimate unit toe bearing capacity SPT N value corrected for 6 percent hammer efficiency Unit shaft resistance Effective soil-foundation friction angle The lateral earth pressure coefficient after installation The lateral earth pressure coefficient prior to the installation Pile length Settlement of the pile head t z q The mobilized shaft resistance, shear stress along the shaft Relative displacement between the pile and soil Mobilized toe resistance Tangent modulus Strain Area of steel
Circumference of pile H P COV Z i i N CDF Observed pile head displacement Load, changing value along the pile Relative soil-pile movement Bias Mean bias Standard deviation Coefficient of variation Probability of occurrence Standard normal variate Probability density function Sample ranking Number of realizations Cumulative density function Mean value Tolerance margins of the target probability Level of confidence
1 1 Introduction Augered cast-in-place piles, also known as ACIP piles, have been used for more than 7 years in the United States. Augered cast-in-place piles have gained in popularity because they are relatively fast to install and are economical for a wide range of projects. Additionally, the installation of ACIP piles poses less impact on the neighboring environment than some other deep foundation installation methods. 1.1 Problem Statement Although there has been an increase in application of ACIP piles, relatively little research on this type of pile has been performed as compared to drilled shafts or driven piles. The insufficient experimental work on ACIP piles behavior and lack of an ACIP pile specific load-displacement model has led practicing engineers to use the results and methodologies from similar cases, such as drilled shafts. An example of this is the use of load transfer models, known as t-z curves when shaft resistance is concerned. Implementation of drilled shaft specific t-z models results in an underestimation of the shaft resistance of ACIP piles and consequently a disagreement between the predicted and measured load-displacement behavior of the ACIP piles. 1.2 Research Outline The goal of this study is to develop an ACIP pile specific t-z model to generate loaddisplacement estimates that are more accurate than currently used models. Chapter 2
2 includes a detailed literature review on the previously proposed and currently used loaddisplacement models for ACIP piles. Chapter 3 provides the objectives of this research. Chapter 4 describes the experimental database that is used in this study along with the procedure necessary to generate the experimental load transfer curves. Chapter 5 presents the methodology for developing the individual t-z curves from the experimental datasets and the proposed load-displacement model. Also, the comparisons between the actual field measurements and existing load-displacement models are included with the results obtained from the proposed model. Chapter 6 provides a statistical analysis using Monte Carlo Simulation to estimate the variability associated with the proposed load-displacement model. Finally, Chapter 7 provides a summary and conclusions derived from this research. A complete list of references and four appendices follow Chapter 7. Appendix A presents the code developed in MATLAB to generate the random variables, a, b and, parameters required for calculation of the load-displacement. Appendix B presents the code developed in MATLAB to calculate the load-displacement behavior of the piles based on the proposed model. Appendix C contains the details of comparison between values of the random variables, a, b and calculated from deterministic formulation with those obtained from random variable generation and Appendix D provides a point by point comparison of the bias between the measured and simulated load-displacement curves of all piles.
3 2 Literature Review 2.1 Introduction Deep foundations are considered to be one of the most commonly implemented foundation types in geotechnical engineering practices (Prezzi 27). A summary of the conditions for which implementation of piles could be a solution is presented by Vesic (1977). These typical situations are illustrated in Figure 2.1. Some of the reasons that might influence an engineer s decision to use a deep foundation over a shallow foundation include the following (Coduto 21): 1. The upper soil layers do not provide sufficient strength for use of shallow foundations. 2. High loads that would require large and impractical dimensions of shallow foundations. 3. Possibility of scour or undermining that could result in failure of the shallow foundation. 4. Constraints and obstacles, such as water, for accessing the ground surface. 5. Need for uplift support. 6. Requirement to resist large lateral loads. Shallow foundations and deep foundations can be distinguished by their depth of embedment. A deep foundation is defined as a foundation that transfers the applied loads to the soil layers below the ground surface to the depths in range of 15 to 45 m or more.
4 Figure 2.1: Situations in which deep foundations may be needed. (a) weak soil layer with a dense stratum within a reasonable depth, (b) weak soil layer without a dense stratum, (c) uplift forces, (d) and (f) horizontal force resist by a single pile, (e) horizontal forces and moments resist by pile group,(g) erosion around the footing, (h) liquefaction susceptible layer, (i) deep foundation acting as a fender system, (j) probability of future excavation in adjacent region, (k) swelling soil layer(fhwa 26, after Vesic 1977). The decision regarding the type of deep foundation to use is based on several factors such as the type and magnitude of loading, soil properties, and construction site constraints. Among all types of deep foundations, the focus of this thesis is mainly on augered cast-in-place piles, but in some cases where no ACIP pile specific relationships are available, expressions were adopted from a comparable kind of deep foundation,
5 drilled shafts. Therefore, these two deep foundation types are described in the following sections. 2.1.1 Drilled Shaft Foundations Drilled shafts are a common type of deep foundation that is usually constructed using a cast-in-place procedure, installed using a drill rig. Drilled shafts may be constructed up to 6 m in diameter and up to 45 m long, or more, and are able to support a variety of loading modes. The construction procedure is considered more appropriate in urban areas due to the production of lower noise levels and vibrations than other types of installation methods like pile driving. Additionally, if soils that were not anticipated during design are observed, engineers are able to change the dimensions of the shaft during construction. As with any engineering tool there are some disadvantages to drilled shaft foundations. Human proficiency plays a critical role in the quality of the foundation; therefore construction is very dependent on the contractor s skill. Because drilled shafts are constructed using excavation, they may have less unit side and toe bearing resistance than driven foundations (Coduto 21). 2.1.2 Augered Cast-in-Place Piles Augered cast-in-place piles (ACIP piles) also known as Continuous Flight Auger piles (CFA) have been used in the United States since 194s (Neate 1989). The typical
6 diameter of ACIP piles ranges from.3 to 1. m (Brown 25) and their lengths can extend more than 3 to 35 m (Brettmann and NeSmith 25, Mandolini et al. 22). ACIP piles are installed by drilling a plugged hollow-stem, continuous-flight auger into the ground to the depth of interest, during which soil is removed by auger flights. After reaching the depth of interest, concrete or grout is pumped into the hollow stem and auger slowly removed under a head of grout. The steel reinforcement cage is then inserted and tied off at the surface. The construction technique is illustrated in Figure 2.2 (Prezzi 27). Augered cast-in-place piles are more economical than other type of piles, like driven piles, although they typically do not provide as high capacities. In addition to sharing the advantages of drilled shafts, the ACIP pile construction technique is more effective for caving soils and is most often used in granular soils (Coduto 21). Figure 2.2: ACIP construction steps, (1) drilling, (2) concrete injection, (3) inserting the reinforcements (after Mainwinch 213).
7 Similar to drilled shaft foundations, ACIP piles are highly dependent on contractor s skill. Placement of steel caging into the grout is another difficulty associated with construction of ACIP s and it can be problematic if heavy steel cages are required in squeezing soils. Augered cast-in-place piles are inappropriate for use in soils that contains cobbles or boulders. Additionally, if the stratigraphy includes a soil layer with high compressibility, such as organic deposits, use of ACIP piles may not be appropriate. 2.2 Bearing Capacity of Drilled Deep Foundations The controlling factor for the bearing capacity of piles subjected to axial loads is the load transfer mechanism by which the load is transferred from the pile shaft and toe to the surrounding soil. The basic concept of load transfer is the same for all types of piles, although the rate of load transfer can vary significantly between different deep foundation and soil types. The applied load, Q, on a pile with length of L, is resisted by shaft friction that is mobilized along the length of the shaft (Q s ), and resistance of the soil below the toe of the pile (Q t ), as demonstrated in Figure 2.3.
8 Figure 2.3: A conceptual sketch of resisting parameters (Das 24). The ultimate shaft resistance is usually completely mobilized at displacements relatively lower than those for the ultimate toe resistance. The amount of displacement for full mobilization of shaft resistance is about 5 to 1 mm, whereas the value of displacement required to mobilize the ultimate toe bearing capacity is about 1% to 25% of the footing diameter (Vesic 1977). The methods to calculate these ultimate values are presented in following sections. 2.2.1 Estimation of Bearing Capacity of Drilled Shaft Foundations in Granular Soils 2.2.1.1 Toe Bearing Capacity of Drilled Shafts Since most of the bearing capacity design models in geotechnical engineering have their roots in empiricism, there is usually more than one single approach for each
9 parameter of interest. In fact, most of these approaches have shown to provide results within the same order, with small differences due to the assumptions used. One of the first relationships for the ultimate toe bearing capacity of drilled shaft piles in granular soils, provided by Meyerhof (1963) is: [ ] where is the area of the toe and equals whereas D t is the diameter of the pile toe, is effective vertical stress at the base of the shaft, is the bearing capacity factor and equals, is a shape factor that equals, and is depth factor that could be obtained by. Chen and Kulhawy (1994) suggested the use of the critical rigidity index and reduced rigidity index factors to account for soil compressibility effects. Prakash and Sharmad (199) proposed a relationship for calculating the ultimate toe bearing capacity as: where is the ultimate unit toe bearing capacity, is the bearing capacity factor which is a function of the friction angle as shown in Figure 2.4, and equals vertical effective stress at the pile toe.
Nq 1 8 7 6 5 Proposed data points Power fit 8 8 3 4 3 2 1 1 2 3 4 5 φ (Degrees) Figure 2.4: Bearing capacity factor from Prakash and Sharmad (199). One of the criteria that should be considered in calculation of the ultimate toe bearing capacity is the displacement of the pile itself. O Neill and Reese (1999), recommend: for the ultimate toe bearing capacity of drilled shafts based on settlement equal to 5 percent of the toe diameter, where equals the ultimate unit toe bearing capacity, is the mean SPT N-value between the pile toe and a depth of 2D t below the toe, and D t is the toe diameter of drilled shafts. This relationship is valid for N 6 5, where N 6 is mean SPT N-value, corrected for 6% hummer efficiency, between the toe and a depth of 2D t below the toe. O Neill and Reese (1999) considered soils with N 6 5 as an intermediate geomaterial, and further soil testing may be necessary to determine the ultimate toe bearing capacity of drilled shafts in these soils or rock conditions.
11 Considerable displacement is required to develop significant toe resistance, and serviceability of the supported structure must be taken into account. Typically, 25 mm is considered as the maximum allowable settlement although in reality, the critical maximum displacement depends on the structure. There are two approaches for incorporating the relationship between displacement and resistance; first, one can reduce the magnitude of used in the ultimate resistance calculation, one can perform a settlement analysis as described in Section 2.3.1 and evaluate the ultimate toe bearing capacity accordingly. As described earlier, Equation 2.4 is only valid for the cases where N 6 5. In the cases where N 6 5 and the soil is non-cohesive, O Neill and Reese (1999) recommended: [ ] 8 where is the ultimate unit toe bearing capacity, defines as corrected SPT N-value, and equals the vertical effective stress at toe of the drilled shaft. The O Neill and Reese (1999) relationship was modified by Brown et al. (27) for determination of the ultimate unit toe resistance of ACIP piles as: where represents the SPT N values corrected for 6 percent hammer efficiency.
12 2.2.1.2 Shaft Resistance of Drilled Shafts The majority of load in normal service loading cases is transferred through the shaft by mobilized frictional resistance, since the ultimate values of shaft friction could be mobilized under smaller displacements relative to toe bearing resistance. Therefore, determination of the shaft resistance has been a leading concern and many research studies have been carried out on this subject. The shaft resistance can be estimated by performing an effective stress or total stress analysis (Coduto 21), based on the generation and/or persistence of excess pore water pressure. If the time for dissipation of excessive pore water pressure that was generated during the installation procedure is comparable with some portion of service life of pile, then a total stress analysis may be considered. On the other hand, if dissipation is rather rapid, effective stress analysis is more appropriate. In case of the construction of drilled shafts in granular soils, dissipation is usually rapid, so the predominant condition during the service life is the steady-state condition and calculations will be more appropriate using an effective stress analysis. The relationship for the ultimate unit shaft resistance,, can be described as (Coduto 21): where equals the horizontal effective stress, and is the effective soil-foundation friction angle. The value of is usually developed through field investigations and laboratory tests. An example of such efforts may be found in Kulhawy et al. (1983) and Kulhawy (1991) and is presented in Table 2.1. Also, since the construction and installation of the pile introduces disturbance in the surrounding soil and the lateral earth
13 pressure is not certain after installation, relationships for the lateral earth pressure coefficient after installation, K, are reported as a function of their values prior to the installation, K, in a same manner as for the soil-foundation friction angle. One such relationship is presented by Kulhawy et al. (1983) and Kulhawy (1991) and shown in Table 2.2. Table 2.1: Approximate φ' f φ' values for the interface between deep foundations and soil (after Kulhawy et al. 1983 and Kulhawy 1991). Foundation Type Rough concrete 1. Smooth concrete (i.e., precast pile).8-1. Rough steel (i.e., step-taper pile).7 -.9 Smooth steel (i.e., pipe pile or H-pile).5 -.7 Wood (i.e., timber pile).8 -.9 Drilled shaft built using dry method or with temporary casing and good construction techniques Drilled shaft built with slurry method (higher values correspond to more careful construction methods) 1..8-1.
14 Table 2.2: Approximate ratio of coefficient of lateral earth pressure after construction to that before construction (after Kulhawy et ai. 1983 and Kulhawy 1991). Foundation Type and Method of Construction Pile-jetted.5 -.7 Pile-small displacement, driven.7-1.2 Pile-large displacement, driven 1. 2. Drilled shaft-built using dry method with minimal sidewall.9-1. disturbance and prompt concreting Drilled shaft-slurry construction with good workmanship.9-1. Drilled shaft-slurry construction with poor workmanship.6 -.7 Drilled shaft-casing method below water table.6 -.9 In a similar approach, Tomlinson and Woodward (28) suggested the following approximation: where K represents the coefficient of lateral earth pressure which here obtained from, is the average vertical effective stress along the soil layer, and is the effective friction angle of the soil layer. To eliminate the individual evaluation of the parameters such as lateral earth pressure coefficient and pile-soil friction angle, Burland (1973) suggested the implementation of the -method. The -method is an alternative approach to Equation 2.7, with the values
15 of acquired from full-scale static load tests. The values of are then used for similar soil profiles and foundations. To benefit from this method, the original soil layer is divided into several sub-layers, and appropriate values of is used for each layer according to characteristics of soil in the layer considered, as: where all of the parameters have been previously defined. For drilled shafts installed in granular soil, especially sand, with N 6 15, O'Neill and Reese (1999) recommend: for calculation of with a maximum value of 19 kpa, where, z is the depth to midpoint of soil layer. Rollins et al. (25) suggested a modification to this equation for cases where gravel is present. 8 where z is the depth to midpoint of soil layer in meters.
16 2.2.2 Estimation of Bearing Capacity of Augered Cast-in-Place Piles 2.2.2.1 Ultimate Toe Bearing Capacity of ACIP piles Despite the use of augered cast-in-place piles for seven decades, there are fewer relationships available for them as compared to conventional drilled shafts. Neely (1991) studied 66 ACIP piles that were installed mostly in sand. Neely (1991) assumed that the ultimate pile resistance occurs at pile head movement of 1 percent of pile diameter; therefore, the values for ultimate loads were obtained either from the field measurements or stability plot method and corresponding toe bearing capacities were calculated. Consequently, an empirical relationship was developed for toe bearing capacity of augered cast-in-place piles in granular soil, specifically in sand: where is the ultimate unit toe bearing capacity in kpa, is the SPT blow count below toe and a depth of about D t. Figure 2.5 illustrates the experimental data used by Neely (1991) to generate Equation 2.13.
17 Figure 2.5: The relationship between SPT N values and the toe bearing resistance, after Neely 1991. Recently, a modification to the Neely (1991) relationship was made by Stuedlein et al. (212), in which the results from 11 test cases reported by Gurtowski (1997, 29) were added to data base created by Neely (1991). Stuedlein et al. (212) suggested a lower value for the upper bound of the ultimate toe bearing capacity as: based on the new load test data, where represents the SPT N value. This relationship described by mean bias, where bias defines as ratio of measured values to those predicted using proposed model, of 1.1 and COV in bias of 23 percent. Stuedlein et al. (212) proposed relationship provides more appropriate results than the expression recommended by O Neill and Reese (1999), as presented in Figure 2.6.
18 Figure 2.6: The relationship between SPT N values and the toe bearing resistance, with comparison between the Stuedlein et al. (212) expression with FHWA method, after Stuedlein et al. (212). 2.2.2.2 Ultimate Shaft Resistance of ACIP Piles For the calculation of the ultimate shaft resistance of ACIP piles, Neely (1991) used the same database of 66 ACIP test piles and found out that the coefficient varies with depth until 24 m, and remains constant for pile length greater than 24 meters. He also proposed a maximum value for ultimate shaft resistance based on filed observations as: where is the average unit shaft resistance, represents the average β coefficient that could be obtained from the values demonstrated in Figure 2.7, and is the average vertical effective stress along length of pile.
Length of Pile (m) 19 5 1 15 2 25 3 1 2 3 β-coefficient Figure 2.7: Average β-coefficient for use in Equation 2.15 (Neely 1991). Stuedlein et al. (212) suggested a relationship for -coefficient values based on the results reported by Gurtowski (1997, 29) in addition to database prepared by Neely (1991). To determine this expression, the ultimate shaft resistance was assumed to be mobilized at pile head movements of 5 to 7.5 percent of pile diameter. This relationship is: ( ( ) )
Length of Pile (m) 2 where the coefficient is defined as function of pile length and is pile length. This expression was characterized with a mean bias and COV in bias of 1.8 and 4 percent respectively, as shown in Figure 2.8. Since the expressions for toe bearing capacity and ultimate shaft resistance suggested by Stuedlein et al. (212) provide more adequate predictions than those developed for drilled shafts, Equations 2.14 and 2.16 will be used to calculate the ultimate toe bearing capacity and ultimate shaft resistance as described in chapter 5. Figure 2.8: The relationship between depth of embedment and β coefficient. The data points consist of experiments conducted in the Stuedlein et al. (212) and Neely (1991) experiment. after Stuedlein et al. (212).
21 2.3 Prediction Models of Pile Settlement in Granular Soils 2.3.1 Vesic (1977) Settlement Method One of the widely used tools to predict and calculate the settlement of piles was proposed by Vesic (1977) which can be included in the category of empirical relationships. This method can be used with either drilled shafts or driven piles. The Vesic (1977) method separates the total settlement of a pile into three components: settlements due to load transferred to the pile toe,, settlement due to load transferred along the pile shaft,, and the settlement due to the distortion of shaft,, as: In the Vesic (1977) method, load on the shaft is defined as the smaller of the applied load on the pile head and the ultimate shaft resistance. Load on the toe is chosen as the smaller of the difference between the applied load and the shaft resistance, which should be considered as zero if the shaft resistance were greater than applied load, and the toe resistance. The Vesic (1977) method requires that the ultimate shaft and toe bearing resistance by means of a selected capacity model, such as those described in Section 2.2. Then, the settlement due to the elastic deformation of the shaft can be calculated by:
22 where represents the load on the toe, is the load on the shaft, equals to the pile length, is the pile cross sectional area, equals elastic modulus of the pile material, and is the load distribution factor, which ranges from and generally assumed to be.6. Figure 2.9 illustrates the effects of load distribution factor along a pile. The settlements in pile toe; due to load transferred through pile toe and along the pile, can be calculated by: where is the toe settlement of the pile due to load transferred through the toe of the pile, equals the toe settlement of the pile due to load transferred along the shaft, is the net load on the toe, is the load on the shaft, equals the pile toe diameter, equals the ultimate unit toe resistance, and are empirical coefficients depending on the soil type and installation method. Table 2.3 presents some typical values for according to soil type and pile installation method. The values for can be calculated using: ( )
23 Table 2.3: Values for C p coefficient (after Vesic 1977). Soil Driven Piles Drilled Piles Sand (dense to loose) Clay (stiff to soft) Silt (dense to loose).2-.4.9-.18.2-.3.3-.6.3-.5.9-.12 Figure 2.9: The load distribution factor for different typical loading cases: (a) uniform distribution, α =.5, (b) extreme case of linear distribution, α =.67, (c) extreme case of linear distribution, α =.33. After Vesic (1977).
24 2.3.2 Introduction to t-z and q-z Models A significant assumption that is invoked when calculating the bearing capacity of piles using the relationships described in Section 2.2 is that the ultimate shaft and toe resistance was mobilized completely and simultaneously. This assumption is not valid for serviceability limit states, which is prevalent over the service life of a pile. Load transfer models help to overcome this problem by introducing a relationship between mobilized resistance and its ultimate value as function of a particular parameter such as the head movement of the pile or relative movement of each point on the pile relative to the surrounding soil. These models consist of an expression for the mobilized shaft resistance (also known as shear stress along the shaft, t, which equals f s in the ultimate case), and mobilized toe resistance (which is known by q and q t in ultimate case) as function of the relative displacement between the pile and soil, z. The load transfer models for shaft and toe resistance are termed t-z and q-z models, respectively. The main idea behind t-z and q-z method is to model the behavior of the surrounding soil as a series of springs with nonlinear stiffness. Employment of nonlinear stiffness is one of the advantages of t-z and q-z models over conventional approaches, since they can estimate the expected nonlinear behavior of the soil during loading. A schematic demonstration of one such model is presented in Figure 2.1. There are two methods to produce t-z and q-z models, theoretical and empirical. In the theoretical approach, the load transfer mechanism is usually developed based on theoretical concepts such as
25 elasticity, whereas in empirical approaches the load transfer is estimated using actual measurements. Figure 2.1: Schematic concept used in t-z method modeling (after FHWA 213). 2.3.2.1 Theoretical t-z Curves for Linear Elastic Soils Kraft, et al. (1981) reasoned that geometric parameters of piles, such as diameter and length, and soil characteristics have a critical influence on pile stresses, empirical relationships that are established using a limited number of experiments cannot be used as a general model for all conditions. Kraft, et al. (1981) proposed that the load transfer mechanism is more significant than ultimate resistances where considering service loads. Therefore, they adapted existing approaches to obtain ultimate resistances and used
26 theoretical concepts to create a t-z model. For piles under service load, Kraft, et al. (1981) used a linear elastic soil model developed by Randolph and Wroth (1978). The assumptions used included that the deformation mode of the soil around the pile is considered to be the shearing of concentric cylinders, and radial deformations are neglected. A brief outline of the expressions, mostly from Randolph and Wroth (1978) are presented in the following pages. Figure 2.11: Concentric cylinder model for settlement analysis of axially loaded piles (modified from Randolph and Wroth 1978). Vertical equilibrium of the soil element of Figure 2.11 can be expressed as: where is the shear stress, and is the total vertical stress. Since after loading a pile, the rate of change in shear stress around the pile is much larger than the vertical total stress, which can be considered constant, the Equation 2.22 can be simplified to:
27 The shear strain can be calculated from in which is the radial displacement and is neglected, represents the vertical displacement. According to the relationship between shear strain and shear stress for linear elastic soils as, the settlement of the pile shaft can be determined from (Randolph and Wroth 1978): where equals the distance at which shear stresses in the soil become negligible, and equals the shear modulus of the soil at distance. There are several, mostly empirical, relationships for calculating, such as the relationship proposed by Randolph and Wroth (1978): where is the pile embedment depth, represents the factor of vertical homogeneity of soil stiffness which is the ratio of soil shear modulus at the middle of the pile to its amount at pile tip, and is the Poisson's ratio of the soil. This model results in an equivalent linear t-z curve, as shown in Figure 2.12.
28 Figure 2.12: Linear t-z curve obtained using Randolph and Wroth (1978) (after FHWA 213). 2.3.2.2 Empirical t-z Curves Equation 2.1 implies that the maximum shaft resistance and toe bearing are mobilized simultaneously, and is independent of pile movement. However, results from load tests on instrumented piles do not support this assumption. It has been observed that for small loading, pile movement occurs mostly near the top of the pile and is resisted largely by shaft resistance. Further increases in the applied load results in larger movements of the pile head due to the elastic deformation of the pile and downward movement of the pile toe, such that the total pile head movement is described by (Vijayvergiya 1977): where is the movement of the pile head, represents the movement of the pile toe, and is the elastic compression of the pile.
29 The downward movement of any desired depth of the pile can be calculated by knowing the load distribution as a function of the pile depth. where is the downward movement of pile at any desire depth, represents the movement of the pile head, is the area of the pile, defined as pile modulus, and is the load at the considered depth. Coyle and Reese (1966) proposed a criterion for the determination of load transfer in clay, presented in Figure 2.13(a). Later, a similar criterion by Coyle and Sulaiman (1967) was proposed for sand as demonstrated in Figure 2.13(b). Figure 2.13: (a) Normalized load transfer curves for clay proposed by Coyle and Reese 1966 and (b) Normalized skin friction curves for sand proposed by Coyle and Solaiman 1967. According to Vijayverjia (1977) the load transfer mobilized at any depth can be represented using: ( )
3 where is the unit shaft resistance mobilized along a pile segment at movement, represents the ultimate unit shaft resistance, and defines as the critical movement of the pile segment at which is mobilized. According to experiments by Vijayverjia (1977), can considered equal to about.3 inches or 8 mm for sands. Figure 2.14: Normalized skin friction curve for clay and sand proposed by Vijayverjia (1977). O Neill and Reese (1999) developed charts from full scale experiments of drilled shafts to relate mobilized shaft resistance normalized by the ultimate shaft resistance, with pile settlement normalized by the pile toe diameter in percent, shown in Figure 2.15. As indicated in the graph, O Neil and Reese (1999) predicted that the ultimate shaft resistance would be reached in a settlement of about 1 percent of the pile diameter. They also included minor decreasing trend in the shaft resistance values after reaching the maximum value, for use with deflection softening soils.
31 1.2 1.8 f s A s Mobilized f s A s Ultimate.6.4.2 Range for Deflection Softening Response Range for Deflection Hardening Response Trend.5 1 1.5 2 δ/b (%) Figure 2.15: Normalized curves of load transfer in side friction vs. settlement for drilled shafts in sand (O'Neill and Reese 1999). 2.3.2.3 q-z Curves for Toe Bearing Capacity Results from load tests show that ultimate toe resistance, if at all possible to be achieved, is mobilized at very large toe movements. Therefore,, the critical displacement, can be defined as the point at which the maximum unit bearing capacity of the pile toe is mobilized. The critical displacement also is typically defined as a function
32 of the pile toe diameter and ranges from.4d t to.6d t where D t is pile toe diameter. The mobilized unit toe bearing pressure at any movement can be obtained from (Vijayverjia 1977): 3 where is the ultimate toe bearing, equals the critical displacement corresponding to the, and refers to the toe bearing mobilized at movement of. Values for at the pile toe can be obtained from Equation 2.3 or other proposed relationships. Figure 2.16 shows normalized q-z curve described by (Vijayverjia 1977) for clay and sand, where the ratio of gradually increases to its maximum value and remains constant thereafter. This behavior is not typical of most piles bearing in sand. Figure 2.16: Normalized q-z curve for clay and sand (Vijayverjia 1977). Later, O Neill and Reese (1999) developed charts from the full scale experiments to define a relationship between the mobilized toe bearing resistance normalized by the ultimate toe resistance, and the pile settlement normalized by the pile toe diameter in percent. The resultant graph is provided in Figure 2.17. This graph does not indicate an
33 ultimate resistance, as there typically is not an ultimate toe bearing resistance in granular soils. However, in practice, the pile settlement of about 1 percent of the pile diameter is considered as the point where the toe bearing capacity would be mobilized. 1.8 1.6 1.4 q t A t Mobilized q t A t Ultimate 1.2 1.8.6.4 Range Trend.2 2 4 6 8 1 12 δ/b (%) Figure 2.17: Normalized curves of load transfer in toe bearing vs. settlement for drilled shafts in sand (after O'Neill and Reese 1999).
34 2.4 Summary In this chapter, a brief overview of drilled foundations and their applications were presented. A detailed description of augered cast-in-place (ACIP) piles, including installation method, was provided. The existing relationships for calculation of ultimate shaft and toe resistances were discussed and relevant load transfer models were summarized. It has been noted that most of the existing relationships and load transfer models are generated from the drilled shaft specific experimental results; therefore, application of such expressions for ACIP piles can reduce the accuracy of the loaddisplacement model. In the following chapters the database and procedures to generate a ACIP pile specific load-displacement model will be presented to address the apparent gap in ACIP pile design.
35 3 Research Objectives The global objective of this study is to develop an ACIP pile specific t-z model to use with the granular soils of Western Washington. The proposed t-z model will be incorporated with previously established q-z models, such as those by O Neill and Reese (1999), to create a load-displacement and load transfer prediction model that can be used to estimate the behavior of augered cast-in-place piles. The following specific objectives comprise the goals of this study: 1. Develop the load transfer behavior for each of six instrumented piles based on strain gage measurements; 2. Determine empirical t-z curves derived from field measurements of each of six instrumented test piles; 3. Determination of a suitable approach for determining the uncertainty in the empirical t-z model; 4. Generate a general t-z model and quantify the associated uncertainty; and, 5. Simulate the observed load test data and evaluate the accuracy and variability of the load-displacement behavior. The research program performed to achieve the objectives outlined above includes: 1. Development of the soil profiles for each test case and the determination of appropriate geotechnical parameters; 2. Generation of the empirical t-z curves and normalizing the curves to generalize the observed behavior;
36 3. Selecting and calibrating existing q-z curves for use with ACIP piles; 4. Predicting the observed load-displacement curves of the instrumented and independent un-instrumented ACIP piles in West Washington granular soils; 5. Using the statistics of variability in the t-z and q-z model to simulate numerous load-displacement curves using Monte Carlo Simulation; and, 6. Comparing the actual and estimated loads at specific displacements to estimate the variability of the new procedure.
37 4 Database of Loading Test Cases This chapter provides the details of the full-scale experimental load tests previously conducted on ACIP piles that form the basis for this study. The experimental data are used in this study to develop a new t-z and q-z model based load-displacement prediction tool. The measurements in this database are adopted from two sets of test series, termed the 1997 and 29 Test Series. 4.1 Static Pile Load Test Data for 1997 Test Series 4.1.1 General A series of static load tests were performed in the state of Washington and reported by Gurtowski (1997). The loading tests were performed in accordance with the Quick Load Test Method as described in ASTM D-1143-74. Static load tests were performed for six ACIP piles, the length and diameter of which varied from 9 to 22 m and 46 to 46 mm, respectively. Table 4.1 summarizes the geometric details and location of the test piles. Table 4.1: Geometric details and locations of test piles for the 1997 Test Series, after Gurtowski (1997). Load Test Name Diameter (mm) Length (m) Location 1997-1 46.4 9.45 Keyport, WA 1997-2 46.4 9.45 Keyport, WA 1997-3 46.4 14.94 Bremerton, WA 1997-4 46.4 21.34 Bremerton, WA 1997-5 46.4 12.2 Richland, WA 1997-6 457.2 1.36 Seattle, WA
38 The pile loading tests were conducted in the Western Washington region, for which the geological strata consists of dense sand to silty sand and very dense gravel to gravelly sand. Occasionally, a relatively thin weak silt or clay layer was observed in some boring logs. None of the test cases in the 1997 Test Series reached an ultimate resistance; therefore, Gurtowski (1997) applied the stability plot procedure to estimate the value of the ultimate loads. The stability plot procedure is described by Neely (1991) and does not account for soil-pile interaction effects. To improve the results from stability plot method, Gurtowski (1997) applied the t-z model proposed by Kraft (1981) and a q-z model by Vijayvergiya (1977) to estimate loads at large displacements. The simulated loads resulted in a 12 to 3 mm deflection of the pile heads, which corresponded to a deflection of 3 to 7.5 percent of pile diameter. The ultimate values proposed by Gurtowski (1997) will be utilized as references for comparison purposes. Figure 4.1 presents the loading test results along with stability plot method prediction and proposed ultimate loads reported by Gurtowski (1997). Test piles in the 1997 Test Series were subjected to loads up to 3336 kn, achieved for pile 1997-6 and which resulted in a maximum head deflection of 28 mm. On the other hand, the smallest maximum load corresponded to pile 1997-2, equal to 884 kn and 3.6 mm of pile head movement. The ultimate loads and corresponding pile head displacements proposed by Gurtowski (1997) showed a wide range of values, with the minimum of 1446 kn and pile head displacement of 12.7 mm for pile 1997-1 and the maximum head deflection of 31 mm with 3114 kn load for pile 1997-4.
Head Displacement (mm) Head Displacement (mm) Head Displacement (mm) Head Displacement (mm) Head Displacement (mm) Head Displacement (mm) 39 5 Load (kn) 5 1 15 2 (a) 5 Load (kn) 5 1 15 2 (b) 1 1 15 15 2 25 3 Actual measurements Stability plot results Proposed ultimate by Gurtowski (1997) 2 25 3 Actual measurements Stability plot results Proposed ultimate by Gurtowski (1997) 5 Load (kn) 1 2 3 4 (c) 1 Load (kn) 1 2 3 4 (d) 1 2 15 2 25 3 Actual measurements Stability plot results Proposed ultimate by Gurtowski (1997) 3 4 5 Actual measurements Stability plot results Proposed ultimate by Gurtowski (1997) 1 Load (kn) 5 1 15 2 25 3 (e) 1 Load (kn) 1 2 3 4 (f) 2 2 3 3 4 5 Actual measurements Stability plot results Proposed ultimate by Gurtowski (1997) 4 5 Actual measurements Stability plot results Proposed ultimate by Gurtowski (1997) Figure 4.1: Comparison between actual load-displacement and predications from the stability plot method, along with the ultimate loads proposed by Gurtowski (1997): (a) Test 1997-1, (b) Test 1997-2, (c) Test 1997-3, (d) Test 1997-4, (e) Test 1997-5, (f) Test 1997-6.
4 4.1.2 Development of Soil Profile and Soil Parameters for the 1997 Test Series In order to evaluate the load-displacement response of the test piles, the soil profile for each pile must be developed from the vicinity of the pile. In the 1997 Test Series, at least one boring with the Standard Penetration Test (SPT) was performed near each test pile. Boring logs indicated the SPT N-values, soil stratigraphy, and ground water table elevations. These SPT N-values were used with empirical relationships to estimate soil parameters such as the unit weight and the friction angle. To determine the friction angle, the relationship proposed by Wolff (1989) was used. where is the friction angle in degrees and represents the SPT N-value. The unit weight of the soil layers was estimated using the range of values recommended in the ODOT Geotechnical Design Manual (29), which is modified from Meyerhof (1956), and outlined in Table 4.2. Soil profiles and parameters for all the piles were generated from boring data by means of Equation 4.1 and Table 4.2. A linear regression was assumed for computing the unit weights from Table 4.2. Figures 4.2 through 4.7 illustrate the SPT N-values, unit weight, friction angle, and a simplified soil profile for each test case. The observed ground water elevation is shown with a dashed line on the soil profiles.
41 Table 4.2: Criteria for the estimation of unit weight, adopted from ODOT Geotechnical Design Manual (29). SPT N-value Unit Weight (kn/m 3 ) to 4 11. - 15.7 4 to 1 14.1-18.1 1 to 3 17.3-2.4 3 to 5 18.9-22. greater than 5 2.4-23.6 (a) (b) (c) (d) Figure 4.2: Subsurface information for test pile 1997-1: (a) SPT N-value, (b) unit weight, (c) friction angle, (d) soil profile.
42 (a) (b) (c) (d) Figure 4.3: Subsurface information for test pile 1997-2: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile. (a) (b) (c) (d) Figure 4.4: Subsurface information for test pile 1997-3: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.
43 (a) (b) (c) (d) Figure 4.5: Subsurface information for test pile 1997-4: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile. (a) (b) (c) (d) Figure 4.6: Subsurface information for test pile 1997-5: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile.
44 (a) (b) (c) (d) Figure 4.7: Subsurface information for test pile 1997-6: (a) SPT N-value, (b) Unit weight, (c) friction angle, (d) soil profile. Generally, SPT N-values in granular soils, and hence the unit weight and friction angles, usually increase with depth. Despite some minor variations, this behavior is noted in the 1997 Test Series. Except for pile 1997-2 and 1997-6, which contains layers of cohesive soil in their profiles, the 1997 Test Series are embedded in mostly granular strata. Furthermore, the values of unit weight and friction angle extend from 14 to 23 kn/m 3 and 28 to 5 degrees. The unit weights are used to calculate the vertical effective stress along each pile. For cases where a layer of cohesive soil was presented in the soil profile, such as piles 1997-2, 1997-4 and 1997-6, no values of the friction angle are reported for cohesive layers.