Applied Science, Faculty of
Civil Engineering, Department of
DSpace
UBCV
Moffat, Ricardo
Doctor of Philosophy - PhD
University of British Columbia
Internal instability of widely-graded cohesionless soils results from an inability to prevent loss of the finer particles in the presence of forces induced by seepage flow and, on occasion, vibration. Instability is governed by a combination of geometrical and hydromechanical constraints, and therefore largely controlled by soil type and seepage regime. Performance monitoring of embankment dams suggests the phenomenon is an important factor in the overall stability of these structures. A large permeameter test device was designed and commissioned to perform unidirectional seepage flow tests on four widely-graded cohesionless soils, all of which are potentially unstable. The objective was to assess the main hydromechanical factors that influence the onset of instability. A modified slurry mixing technique, with discrete deposition, was found satisfactory for reconstitution of homogeneous saturated test specimens. Test variables examined include hydraulic gradient, rate of increase in hydraulic gradient, flow direction, and vertical effective stress. The onset of internal instability was found to be triggered either by an increase in hydraulic gradient, or by a decrease in vertical effective stress. The relation between effective stress and hydraulic gradient, where the onset of instability occurs, is used to characterise a hydromechanical constraint to internal stability. Each soil yielded a hydromechanical boundary, and the difference between them is attributed to the shape of the grain size distribution of the soil and to the variation of vertical effective stress in the direction of flow where suffosion initiated. A limit-equilibrium approach is used to explain the experimental findings on hydromechanical constraints, based on a geometrical characterization of the gradation curve and for an assumed distribution of effective stress between coarser and finer fraction particles of the soil. Analysis of the experimental data established that the hydromechanical boundary is not unique: it is governed by both the magnitude of vertical effective stress and the variation of that stress in the direction of flow.
EXPERIMENTS ON THE INTERNAL STABILITY OF WIDELY GRADED COHESIONLESS SOILS by RICARDO MOFFAT B.Sc, University de Chile, 1996 Civil Engineering, University of Chile, 1998 M.So, The University of British Columbia, 2002 A THESIS SUBMITTED IN PARTIAL FULFILMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY in THE FACULTY OF GRADUATE STUDIES Civil Engineering THE UNIVERSITY OF BRITISH COLUMBIA June 2005 ©Ricardo Moffat, 2005. ABSTRACT Internal instability of widely-graded cohesionless soils results from an inability to prevent loss of the finer particles in the presence of forces induced by seepage flow and, on occasion, vibration. Instability is governed by a combination of geometrical and hydromechanical constraints, and therefore largely controlled by soil type and seepage regime. Performance monitoring of embankment dams suggests the phenomenon is an important factor in the overall stability of these structures. A large permeameter test device was designed and commissioned to perform unidirectional seepage flow tests on four widely-graded cohesionless soils, all of which are potentially unstable. The objective was to assess the main hydromechanical factors that influence the onset of instability. A modified slurry mixing technique, with discrete deposition, was found satisfactory for reconstitution of homogeneous saturated test specimens. Test variables examined include hydraulic gradient, rate of increase in hydraulic gradient, flow direction, and vertical effective stress. The onset of internal instability was found to be triggered either by an increase in hydraulic gradient, or by a decrease in vertical effective stress. The relation between effective stress and hydraulic gradient, where the onset of instability occurs, is used to characterise a hydromechanical constraint to internal stability. Each soil yielded a hydromechanical boundary, and the difference between them is attributed to the shape of the grain size distribution of the soil and to the variation of vertical effective stress in the direction of flow where suffosion initiated. A limit-equilibrium approach is used to explain the experimental findings on hydromechanical constraints, based on a geometrical ii characterization of the gradation curve and for an assumed distribution of effective stress between coarser and finer fraction particles of the soil. Analysis of the experimental data established that the hydromechanical boundary is not unique: it is governed by both the magnitude of vertical effective stress and the variation of that stress in the direction of flow. iii Table of Contents Page Abstract ii List of Tables viii List of Figures ix List of Symbols xiv Acknowledgments xv 1.0 Introduction 1 1.1 Seepage-induced erosion 1 1.2 On the phenomenon of internal erosion 2 1.2.1 Terminology 2 1.2.2 Description of the phenomena 5 1.2.3 Empirical rules 6 1.3 Purpose and scope of the study 7 1.4 Organization of the thesis 9 2.0 Literature Review 13 2.1 Experimental studies 13 2.1.1 Compatibility of base soil and filter 14 2.1.2 Internal instability studies 20 2.2 Analytical studies 24 2.2.1 Empirical criteria: internal instability 24 2.2.1.1 Kezdi criterion 24 2.2.1.2 Kenney and Lau criterion 2 5 2.2.2 Analytical models 26 2.2.2.1 Limit equilibrium analysis of particle migration 26 2.2.2.2 Porous media theory 27 2.3 Effects of non-plastic f i n e s 32 2.4 Implications for professional practice 34 iv 3.0 Test Device 49 3.1 Introduction 49 3.2 Permeameter test device 50 3.2.1 Permeameter cell 51 3.2.2 Top boundary and axial loading system 51 3.2.3 Bottom boundary and reaction system 53 3.2.4 Water supply system 54 3.2.5 Seepage control system 55 3.3 Instrumentation 56 3.3.1 Axial load 56 3.3.2 Axial displacement 57 3.3.3 Variation of water pressure 57 3.3.4 Weight of inflow/outflow water tanks 5 8 3.3.5 Water temperature and dissolved oxygen content 59 3.3.6 Data acquisition and control system 59 3.3.7 Photo and video equipment 61 3.4 Influence of the equipment 62 3.4.1 Diameter and length of the specimen 62 3.4.2 Filter screen opening size 63 3.5 Summary 65 4.0 Experimental Investigation 75 4.1 Material properties 75 4.1.1 Glass beads 75 4.1.2 Soils 76 4.2 Specimen reconstitution 78 4.2.1 Modified slurry preparation technique 78 4.2.2 Placement technique 79 4.2.3 Pre-test assembly of the permeameter 80 4.3 Test procedure 80 4.3.1 Consolidation 80 4.3.2 Seepage flow 81 v 4.4 Post-test observations 4.5 Programme of testing 4.6 Conclusions 83 84 86 5.0 Results 94 5.1 Preliminary tests to commission the apparatus 94 5.2 Main test programme 96 5.2.1 Axial displacement 97 5.2.2 Hydraulic conductivity 98 5.2.3 Pore volume exchange 100 5.2.4 Transition material (T-0) 101 5.2.4.1 TestT-0-25-D 101 5.2.4.2 Test T-0-100-D 104 5.2.4.3 Test T-0-175-D 106 5.2.5 Transition material (T-5) 108 5.2.5.1 T-5-25-D 108 5.2.5.2 T-5-175-U 109 5.2.5.3 T-5-50-D 112 5.2.5.4 T-5-50-D(R) 113 5.2.5.5 T-5-25-U 115 5.2.5.6 T-5-30-U 116 5.2.6 Core material (C-20) 119 5.2.6.1 C-20-50-U 119 5.2.6.2 C-20-85-U 121 5.2.7 Core material (C-30) 123 5.2.7.1 C-30-25-U 123 5.2.7.2 C-30-50-U 125 5.2.7.3 C-30-80-U 127 5.2.7.4 C-30-100-U 128 5.3 Summary 130 vi 6.0 Analysis and Discussion 178 6.1 Geometrical analysis of the tested soils 178 6.2 Generalized nature of particle migration 180 6.2.1 Soils with no fines content (T-0) 181 6.2.2 Soils with some fines content (T-5) 182 6.2.3 Soils with significant fines content (C-20, C-30) 183 6.3 Factors influencing the onset of instability 185 6.3.1 Equipment 185 6.3.2 Ffydromechanical influences on internal stability 185 6.3.2.1 Rate of change of i a v and pore volume exchange 186 6.3.2.2 Effective stress 188 6.3.2.3 Local hydraulic gradients 191 6.3.2.4 Ffydromechanical boundaries 192 6.3.2.5 Flow direction and stress variations 195 6.3.2.6 Seepage velocity 197 6.4 Proposed model for internal erosion 198 6.4.1 Model formulation 199 6.4.2 Model application 204 6.5 Summary 206 7.0 Conclusions and Recommendations 223 Bibliography 229 Appendix A: Preliminary tests 237 Appendix B: Variation of local hydraulic gradient with time 248 Appendix C: Soil specimen before and after failure 257 Appendix D: Effective stress model using FLAC 271 vii L I S T O F T A B L E S Table 2.1 Summary of selected studies Table 4.1 Characteristic properties of the commissioning tests Table 4.2 Characteristic properties of the soil Table 4.3 Geometrical stability analysis of the soils Table 4.4 Summary of test programme Table 5.1 Summary of preliminary tests Table 5.2 Vertical displacement Table 5.3 Summary of laboratory testing Table 6.1 Effects due to the variation of grain size distribution Table 6.2 Input parameters of the model viii LIST O F F I G U R E S Figure 1.1. Hydrological view of soil piping 10 Figure 1.2. Sadd el-Kafara dam (after Smith 1971) 11 Figure 1.3. Cases of soil migration 11 Figure 1.4 Example of widely graded soils present in nature 12 Figure 2.1. Selected studies on particle migration 37 Figure 2.2. Retention criteria (Terzaghi, 1939) 37 Figure 2.3. Functional requirements of granular filters 38 Figure 2.4. Permeameter (after Bertram, 1940) 39 Figure 2.5. Permeameters (after Pare, 1982) 40 Figure 2.6. Permeameter test assembly (after Lafleur, 1984) 41 Figure 2.7. Permeameter device (after Sherard et al., 1984) 42 Figure 2.8. Permeameter setup (after Tomlinson and Vaid, 2000) 43 Figure 2.9. Test permeameters (after Kenney and Lau, 1985) 43 Figure 2.10. Test permeameter (after Skempton and Brogan, 1994) 44 Figure 2.11. Permeameter setup (after Honjo et al., 1996) 44 Figure 2.12. Permeameter configuration (after Moffat, 2002) 45 Figure 2.13. Stability criterion of Kezdi (1979) 45 Figure 2.14. Stability criterion of Kenney and Lau (1985) 46 Figure 2.15. Limit equilibrium for a single particle (based on Indraratna and Radampola) 46 Figure 2.16. Critical hydraulic gradient (after Indraratna and Radampola, 2002) 47 Figure 2.17 Changes in porosity in the coarse-fine granular mixture (after Vallejo, 2001) 47 Figure 2.18 Coarser-finer fraction mixtures (based upon Vallejo, 2001) 48 Figure 2.19 Peak strength versus percentage of fine fraction (after Vallejo, 2001) 48 Figure 3.1. Components of the large permeameter 67 Figure 3.2. Rigid-wall permeameter 68 Figure 3.3. Permeameter cell (downward seepage) 69 Figure 3.4. Loading plate and rod 70 Figure 3.5. Base frame a) openings, b) visual observation 70 Figure 3.6. Two-stage filter system 71 ix Figure 3.7. Arrangement of the water pressure transducers and load cell 72 Figure 3.8. Load cells a) top and b) bottom 73 Figure 3.9. Weigh scales (flow measurement) 73 Figure 3.10. Dissolved oxygen meter, thermocouples, and LVDT 74 Figure 3.11. Labview software 74 Figure 4.1. Glass beads 89 Figure 4.2. Grain size distribution of glass beads specimens 89 Figure 4.3. Grain size distributions of soil specimens 90 Figure 4.4. Particle shape: gravel fraction 90 Figure 4.5. Shape curves for the soils (after Kenney and Lau, 1986) 91 Figure 4.6. Variation of D'^/d'gs with percent passing (after Kezdi, 1979) 91 Figure 4.7. Mixing of soil particles 92 Figure 4.8. Placement technique 92 Figure 4.9. Overview of the test programme 93 Figure 5.1. Unit displacement material T-0 135 Figure 5.2. Unit displacement material T-5. 135 Figure 5.3. Unit displacement material C-20 and C-30 136 Figure 5.4. Hydraulic conductivity of the test specimen (material T-0) 136 Figure 5.5. Hydraulic conductivity of the test specimen (material T-5 downward flow) 137 Figure 5.6. Hydraulic conductivity of the test specimen (material T-5 upward flow) 137 Figure 5.7. Hydraulic conductivity of the test specimen (material C-20) 138 Figure 5.8. Hydraulic conductivity of the test specimen (material C-30) 138 Figure 5.9. Volumetric flux through transition specimens 139 Figure 5.10. Volumetric flux through core specimens 139 Figure 5.11. Vertical displacement on test T-0-25-D 140 Figure 5.12. Local hydraulic gradient at i a v = 5 (test T-0-25-D) 140 Figure 5.13. Water head distribution (test T-0-25-D) 141 Figure 5.14. Hydraulic conductivity (based on DPT) along the specimen length 141 Figure 5.15. Local hydraulic conductivity (T-0-25-D) 142 Figure 5.16. Size distribution after testing (T-0-25-D) 142 Figure 5.17. Vertical displacement in test T-0-100-D 143 x 3? Figure 5.18. Local hydraulic gradients at failure using DPTs (T-0-100-D, i a v = 8) 143 Figure 5.19. Water head distribution (test T-0-100-D) 144 Figure 5.20. Hydraulic conductivity (based on DPT) along the specimen length 144 Figure 5.21. Local hydraulic conductivity during the test (T-0-100-D) 145 Figure 5.22. Size distribution after testing (T-0-100-D) 145 Figure 5.23. Vertical displacement in test T-0-175-D 146 * Figure 5.24. Local hydraulic gradient when attempting to reach in = 30 (T-0-175-D) 146 Figure 5.25. Water head distribution (test T-0-175-D) 147 Figure 5.26. Hydraulic conductivity (based on DPT) along the specimen length 147 Figure 5.27. Local hydraulic conductivity during the test (T-0-175-D) 148 Figure 5.28. Grain size distribution after testing (T-0-175-D) 148 Figure 5.29. Vertical displacement in test T-5-25-D 149 Figure 5.30. Water head distribution (test T-5-25-D) 149 ^ Figure 5.31. Hydraulic conductivity (based on DPT) along the specimen length 150 Figure 5.32. Local hydraulic conductivity (T-5-25-D) 150 Figure 5.33. Vertical displacement in test T-5-175-U 151 Figure 5.34. Water head distribution (test T-5-175-U) 151 Figure 5.35. Hydraulic conductivity (based on DPT) along the specimen length 152 Figure 5.36. Local hydraulic conductivity (T-5- 175-U) 152 Figure 5.37. Grain size distribution after testing (T-5-175-U) 153 Figure 5.38. Vertical displacement in test T-5-50-D 153 ,„., Figure 5.39. Water head distribution (test T-5-50-D) 154 Figure 5.40. Hydraulic conductivity along the specimen length (T-5-50-D) 154 Figure 5.41. Local hydraulic gradients during the test (T-5-50-D) 155 Figure 5.42. Vertical displacement in test T-5-50-D(R) 155 Figure 5.43. Water head distribution (test T-5-50-D(R)) 156 Figure 5.44. Hydraulic conductivity (based on DPT) along the specimen length 156 Figure 5.45. Local hydraulic conductivity (T-5-50-D(R)) 157 Figure 5.46. Grain size distribution after testing (T-5-50-D(R)) 157 * Figure 5.47. Vertical displacement in test T-5-25-U 158 Figure 5.48. Water head distribution (test T-5-25-U) 158 xi Figure 5.49. Hydraulic conductivity (based on DPT) along the specimen length 159 Figure 5.50. Local hydraulic conductivity (test T-5-25-U) 159 Figure 5.51. Vertical displacement in test T-5-30-U 160 Figure 5.52. Water head distribution (test T-5-30-U) 160 Figure 5.53. Hydraulic conductivity along the specimen length (T-5-30-U) 161 Figure 5.54. Local hydraulic conductivity (T-5-30-U) 161 Figure 5.55. Grain size distribution after testing (T-5-30-U) 162 Figure 5.56. Vertical displacement in test C-20-50-U 162 Figure 5.57. Water head distribution (test C-20-50-U) 163 Figure 5.58. Hydraulic conductivity (based on DPT) along the specimen length 163 Figure 5.59. Local hydraulic gradients during the test (C-20-50-U) 164 Figure 5.60. Grain size distribution after testing (C-20-50-U) 164 Figure 5.61. Vertical displacement in test C-20-85-U 165 Figure 5.62. Water head distribution (test C-20-85-U) 165 Figure 5.63. Hydraulic conductivity (based on DPT) along the specimen length 166 Figure 5.64. Local hydraulic gradients during the test (C-20-85-U) 166 Figure 5.65. Grain size distribution after testing (C-20-85-U) 167 Figure 5.66. Vertical displacement in test C-30-25-U 167 Figure 5.67. Water head distribution (test C-30-25-U) 168 Figure 5.68. Hydraulic conductivity along the specimen length (C-30-25-U) 168 Figure 5.69. Local hydraulic gradients during the test (C-30-25-U) 169 Figure 5.70. Grain size distribution after testing (C-30-25-U) 169 Figure 5.71. Vertical displacement in test C-30-50-U 170 Figure 5.72. Water head distribution (test C-30-50-U) 170 Figure 5.73. Hydraulic conductivity along the specimen length (C-30-50-U) 171 Figure 5.74. Local hydraulic gradients during the test (C-30-50-U) 171 Figure 5.75. Grain size distribution after testing (C-30-50-U) 172 Figure 5.76. Vertical displacement in test C-30-80-U 172 Figure 5.77. Water head distribution (test C-30-80-U) 173 Figure 5.78. Local hydraulic gradients during the test (C-30-80-U) 173 Figure 5.79. Hydraulic conductivity along the specimen length (C-30-80-U) 174 xii Figure 5.80. Grain size distribution after testing (C-30-80-U) 174 Figure 5.81. Vertical displacement in test C-30-100-U 175 Figure 5.82. Water head distribution (test C-3 0-100-U) 175 Figure 5.83. Hydraulic conductivity along the specimen length (C-30-100-U) 176 Figure 5.84. Local hydraulic gradients during the test (C-30-100-U) 176 Figure 5.85. Grain size distribution after testing (C-30-100-U) 177 Figure 6.1. Unified stability chart 209 Figure 6.2. Grain size distribution variation 209 Figure 6.3. Kenney and Lau (1986) analysis of material C-3 0 210 Figure 6.4. Kezdi (1979) analysis of material C-30 210 Figure 6.5. Local hydraulic gradient (i a v = 24, test C-30-80-U) 211 Figure 6.6. Factors that influence the hydromechanical response of the specimen 212 Figure 6.7. Idealized response of vertical effective stress under water seepage 212 Figure 6.8. Vertical effective stress in the test specimen (T-0) 213 Figure 6.9. Vertical effective stress in the test specimen (T-5) 213 Figure 6.10. Vertical effective stress in the test specimen (C-20) 214 Figure 6.11. Vertical effective stress in the test specimen (C-30) 214 Figure 6.12. Influence of hydraulic gradient on internal instability (T-0) 215 Figure 6.13. Influence of hydraulic gradient on internal instability (T-5) 216 Figure 6.14. Influence of hydraulic gradient on internal instability (C-20 and C-30) 216 Figure 6.15. Critical hydraulic gradient for a™' ~ 0 (after Skempton and Brogan, 1994) 218 Figure 6.16. Influence of effective stress (C-30-100-U) 218 Figure 6.17. Seepage velocity at the onset of instability 219 Figure 6.18. Proposed model for internal instability 219 Figure 6.19. Representative unit volume 219 Figure 6.20. Effective stress distribution factor, G* or G** 219 Figure 6.21. Comparison of i c r for transition materials using the model 220 Figure 6.22. Comparison of icr for core materials using the model 221 Figure 6.23. Values of icr for test T-5-50-D 221 Figure 6.24. Effects of variation of Ovm'on hydromechanical boundaries 222 xiii LIST OF SYMBOLS A Cross section of the permeameter Cu Coefficient of uniformity D n Grain size corresponding to n % finer (used for the coarser fraction or filter layer) d n Grain size corresponding to n % finer (used for the finer fraction or base layer) e Void ratio F Mass fraction smaller than D GR Gradient Ratio Gs Specific gravity H Mass fraction measured between particles sizes D and 4D (Kenney & Lau, 1985) h Water head hy Head difference between ports i and j i a v Average hydraulic gradient iy Hydraulic gradient between ports i and j k Hydraulic conductivity ky Hydraulic conductivity between ports i and j L Soil specimen length LVDT Linear variable differential transformer Q volumetric flow rate of water Re Reynolds number yw Unit weight of water p Density xiv A C K N O W L E D G M E N T S Thanks to my wife, Claudia, for her support and love during this study. I also wish to acknowledge the understanding of my two sons, Benjamin and Martin, on the occasions that I worked long evenings and weekends at UBC. I am also grateful of my Research Supervisor, Dr. Jonathan Fannin, P. Eng. for his support, advice and engineering judgement through the study. Without his advice and understanding ^ this research would not have been possible. This research was funded by British Columbia Hydro and Power Authority (BCHydro) to wish I want to express my thanks for the financial support provided. In particular, I would like to express my gratitude to Stephen Garner, P. Eng., Specialist Engineer at BCHydro Engineering, for his keen interest in implications of the laboratory investigation for , performance monitoring of earthfill dams. I also want to thank the members of my Supervisory Committee: Dr. Roger Beckie, P. Eng. Dr. John Howie, P. Eng., Dr. Dawn Shuttle, C. Eng., and Dr. Dharma Wijewickreme, P. Eng., for their input and constructive comments, on the thesis during its preparation, I thank all of the support obtained at the University of British Columbia, in particular that from the staff of the Civil Engineering workshop. xv Finally, I wish to recognize the Faculty of Applied Science, at the University of Chile, for the support given during the period of my studies in Canada. xvi 1.0 INTRODUCTION 1.1 Seepage-induced erosion Seepage-induced erosion is observed both in natural soil deposits and in engineered fill structures. A comprehensive review of factors governing such erosion in natural deposits is provided by Jones (1981). Emphasis is placed on the nature of soil piping from a hydrological perspective (Figure 1.1). Flow of water concentrates in "pipes" or conduits called macropores in the hydrology literature, formed by internal erosion, where the hydraulic conductivity is substantially higher than the adjacent soil. The initiation and * development of soil pipes is primarily attributed to climatic factors, human activity, soil chemistry, soil morphology, composition, and physico-chemical properties of the soil. The relative contribution of each factor determines the rate at which erosion progresses at a specific site. Engineered fill structures, like some natural soil deposits, may also be susceptible to internal v erosion as a result of seepage induced particle migration. One of the oldest engineered fill structures is that of Sadd el-Kafara or "Dam of the pagans", in Egypt, which was likely built some time between 2950 and 2750 B.C (Smith, 1971). The upstream face was covered with limestone blocks to provide for seepage control and resistance to erosion and wave action (see Figure 1.2). The configuration of the blocks was not watertight, yielding conditions that were conducive to internal erosion. This case could be the first known failure due to internal erosion in earth dams. l Interestingly, in 1800 B.C. the king of Babylon carried out extensive irrigation works and had a legal code with rules and regulations pertaining to the operation of dams and canals, with severe penalties applied for inadequate preservation of dams (Smith, 1971). Recent concerns with regard to sinkholes in earth dams, attributed to internal erosion, have been brought to the attention of the geotechnical community. A summary of the statistics of cases of piping were presented by BCHydro and Hydro Quebec at the 53rd Canadian geotechnical conference (2000). Piping through the embankment fill or through the foundation soils is identified as one of the most important causes of failure. A growing awareness of the topic has generated a companion interest on research in support of a more confident understanding of the phenomenon. 1.2 On the phenomenon of internal erosion The grain size distribution of a widely graded soil may be taken to comprise, respectively, a "finer fraction" and a "coarser fraction" (Kezdi, 1979). With respect to the phenomenon of internal erosion, the kind of particle migration considered in this thesis is that involving a considerable portion of the finer fraction of a widely-graded cohesionless soil. 1.2.1 Terminology Different terms have been proposed to distinguish between types of soil migration involving the finer fraction of a soil. From observations of upward seepage flow in sands, Terzaghi (1939) associated "piping" with a behaviour in which "the discharge increases... indicating an increase of the average permeability of the sand... and finally a mixture of sand and water breaks though" at the ground surface. The retrogressive phenomenon of piping in embankment dams is defined somewhat differently as "the progressive development of internal erosion by seepage, appearing downstream as a hole discharging water" (ICOLD, 1994). Piping has been used in the same sense as internal erosion and usually relates to erosion that presents a specific zone for its development that looks like a "pipe". This type of failure has , been reported close to locations such as tubes, cables, instruments or other matrix features that promote arching, or leaves free spaces within the surrounding soil. The erosion develops rapidly and moves new fines until a continuous conduit is formed. Suffusion refers to the "phenomenon where water, while seeping through the pores, carries along fine particles without destroying the soil structure" (Kezdi, 1979). Flow then carries * the smallest particles between the larger ones of the soil matrix, structural particles, or soil skeleton. Kovacs (1981) defined suffusion as the "redistribution of fine grains within the layer... when the solid volume of the layer is not changed, only the local permeability is altered". Hence, internal suffusion has been used to describe a movement of fines within the soil that affects only the local permeability. This takes place when the hydrodynamic forces are large enough to move fine particles inside the skeleton of the soil. In contrast, external% suffusion occurs at the surface where fine particles are carried from one layer (e.g. base soil) into an adjacent layer (e.g. filter). Kovacs (1981) defined external suffusion as "when the 3 volume of the solid matrix is reduced, accompanied by an increase in permeability, but the stability of the skeleton composed by the coarse grains is unaffected". In general suffusion is considered a slow process that shows a slight and slow change of permeability without disturbance to the skeleton of the soil. The term "erosion" has been widely used to describe soil migration in different geological areas. Erosion may consist of migration of fine particles as well as migration of larger particles from the soil matrix. Internal erosion may lead to the formation of an open conduit through the soil (piping). This may also occur when cavities already exist in the skeleton, leading to high seepage velocities. External erosion may occur at the ground surface of an earth structure, due to high exit gradients and the resultant seepage force. Contact erosion is described as erosion due to parallel flow between two materials. Kenney and Lau (1985) referred to suffosion as "the transport of small particles from a soil". The phenomenon suffosion has been defined as a transport of the finer fraction of the soil but that affects the coarser fraction as well. Therefore, it could be defined as migration of the finer fraction that occurs simultaneously with a rearrangement of the coarser fraction ("skeleton"), yielding an overall change in volume. This definition of suffosion is used in the thesis. 4 1.2.2. Description of the phenomena Seepage of groundwater exerts a force on soil particles, which acts in the direction of flow. The seepage force per unit volume of soil is theoretically given by "iy w ", where " i " is the hydraulic gradient (Terzaghi, 1939). If the grain size distribution comprises a finer and coarser fraction, this force has the potential to dislodge finer fraction particles from the matrix of the coarser fraction, thereby triggering the onset of internal instability. Thereafter, the velocity of flow may be sufficient to transport those particles, causing migration of the finer fraction through voids between the coarser grains. Therefore, there are two concerns: (i) can particle migration initiate (internal instability), and (ii) can the migrating particles pass through the soil. The onset of internal instability is deemed to be governed by two constraints, a geometrical and a hydromechanical constraint. Kovacs (1981) defined the geometrical constraint as that for which "the size of the pores are smaller than the smallest particle", and the hydromechanical constraint as "the critical velocity or hydraulic gradient, above which the fine grains start to move". In potentially unstable soils, the finer grains are able to move inside the matrix of the soil "skeleton". In addition to the geometrical constraint, there are different conditions, or factors, influencing the onset of internal instability. These factors include: - Hydraulic conditions o Seepage velocity o Hydraulic gradient 5 o Flow direction, reversing flow Geometrical conditions o Porosity o Constriction size o Shape of the distribution size curve - Other conditions o Effective stress o Vibration (earthquake induced or other movements) o Air content o Temperature o Viscosity of the fluid Resistance to soil migration is derived from the size of the finer fraction and the impediment to movement resulting from the structure of pore size constrictions. A soil may be considered "inherently stable" when these geometrical constraints restrict the movement of the finer fraction. 1.2.3. Empirical rules When dealing with seepage flow of water through a zoned earth fill structure and its foundations, it is important to ensure that each individual layer is internally stable under the expected field conditions (see Figure 1.3). Additionally, it is necessary to verify the filtration compatibility of adjacent layers with respect to soil retention and relative permeability or 6 hydraulic conductivity. It is for these reasons that empirical criteria have been developed, both to assess the internal stability of a soil, as well as to verify the filter requirements of adjacent soils. There is a considerable body of experience with the use of granular soils as a filter material, and that with the use of geotextiles is steadily growing. The filtration process depends on the formation of a stable interface between the base soil and the filter material, termed "self-filtering layer". To ensure this stable interface, it is necessary to satisfy appropriate design criteria. Such design criteria are empirical and typically derived from experimental studies on reconstituted laboratory test specimens. 1.3 Purpose and scope of the study Widely-graded and gap-graded soils are present in nature and consequently encountered in construction practice, for example at the Portage Mountain Dam (see Figure 1.4) in British Columbia, as mentioned by Morgan and Harris (1967). A sinkhole incident at the WAC Bennett dam (formerly known as Portage Mountain Dam) was reported by Steward and Watts (2000): widely-graded materials of the Core and Transition zones may have played an important role in the potential for a seepage-induced instability. However, although internal erosion is considered to be the second most common cause of failure in earth dams (Foster et al. 2000), its mechanisms and the conditions to trigger failure remain poorly understood. Accordingly, the influence of both geometrical and hydromechanical constraints must be identified with equal confidence in order to properly account for internal erosion in 7 potentially unstable soils. In this regard, hydromechanical constraints are largely unknown, and constitute the more uncertain of the two concerns. The primary purpose of this study is to establish the conditions associated with the onset of particle migration in widely-graded cohesionless soils. The importance of particle migration in cohesive soils is also recognized as a companion phenomenon, for which a specific body of literature exists, however it falls outside the scope oh this study. An experimental investigation is reported that addresses the hydromechanical conditions that trigger internal erosion with specific reference to the influence of effective stress and hydraulic gradients. A large permeameter test device was designed and commissioned to perform a series of tests, involving unidirectional seepage on soils representative of selected zones in the WAC Bennett Dam, British Columbia. A multi-stage test procedure was developed to assess conditions necessary to trigger the onset of instability. The general nature of internal erosion is described, and behaviour of Transition and Core materials under seepage flow is then contrasted. Results of this study are compared with the findings of others, and a conceptual model of failure is established that determines the susceptibility to instability with reference to hydromechanical constraints. In addition, a unified approach to the assessment of geometrical constraints is proposed based on the work of Kezdi (1979), Kenney & Lau (1985, 1986) and Moffat (2002), to evaluate the potential for internal stability of soils. 8 1.4 Organization of the thesis In Chapter 2, a review is given of selected filtration studies reported in the literature. The differences between various laboratory tests are examined, and factors that influence seepage-induced instability in soils described. A summary is provided of the more commonly-used methods to assess constraints to instability, together with an evaluation of aspects that could influence the hydromechanical stability of a potentially unstable soil. Chapter 3 describes the design and instrumentation of a large permeameter test device used in this laboratory study. Properties of the materials tested, specimen reconstitution technique, and multi-stage test procedure are reported in Chapter 4. Chapter 5 presents results of a preliminary series of tests to commission the test device, and the main test programme which addresses hydromechanical constraints governing the onset of instability. It is followed by an analysis and interpretation of the experimental data in Chapter 6. A series of conclusions and recommendations are presented in Chapter 7. 9 Subsurface flow from matrix into upstream end of pipes Pipes beneath surface Overland flow from pipe outlet Inflow through pipe "blow - hole" Pipes formed at change in • soil properties Throughflow seepage from saturated part of soil -Turbulent flow of water -Laminar flow of water Figure 1.1. Hydrological view of soil piping (after Atkinson, 1978, shown in J. Jones, 1981) 10 Figure 1.2. Sadd el-Kafara dam (after Smith 1971) 100 2 80 < x »-e so z UJ o rr LU a 40 20 Grovel 3/4" Sand Silt #4 ^10 **40 100 10 I GRAIN SIZE, mm. ^200 V \ X i 1 I ... .,.„.,, i i i ! 1 \ \ IN. 1 N ? \ > I > \ I 1 \ ! i \! j 1 1 \ 1 1 \ . 1 1 \ . \ N i \ 1 1 1 1 1 1 i N . 1 \ ^ >^ ! 1 1 1 1 1 1 1 1 1 i i i i i i i Typical Sandy Grovels Typical Sand Figure 1.4 Example of gap and widely-graded soils in nature (Morgan and Harris, 1967). 12 2.0 L I T E R A T U R E R E V I E W Cohesionless soils may exhibit, under the influence of seepage flow, finer fraction migration through the interstices of the matrix formed by the coarser fraction. This may lead to an increase in the permeability of the soil and to stability concerns. This phenomenon has been named internal instability, and it was first observed in studies of filter/base compatibility. Empirical methods were proposed to recognize grain size distribution of internally unstable materials, where vibration was commonly used to trigger internal erosion of the specimen. Limited research has been performed on the hydromechanical component of internal erosion. A few physical models have been proposed to identify the hydromechanical influence in the onset of internal stability of soils. The current understanding of the phenomenon is reviewed with reference to experimental studies, analytical studies, related models in porous media theory, and a consideration of non-plastic fines (see Figure 2.1). 2.1 Experimental studies Studies of particle migration in soil mechanics began in an attempt to prove the soil retention criterion of Dis/dgs < 4 proposed by Terzaghi (1939), illustrated in Figure 2.2. Minor revisions to this criterion for soil retention have been proposed from interpretation of* laboratory studies. A schematic illustration of the causal relations between characteristics of a granular filter, and functional requirements against which performance is assessed, is given in Figure 2.3. Three of those functional requirements are: base soil retention, permeability, and internal stability. The characteristic size of the coarser fraction (for example, D15) 13 influences the size distribution of the pores, and hence the capacity for retention of a base soil, and the permeability of the filter itself. The distribution and size of the finer fraction particles also exerts an influence on the permeability. Lastly, the gradation shape of the filter and base material determines the potential for internal instability. Usually, retention and permeability requirements are given greatest importance in laboratory studies, but careful consideration of widely-graded and gap-graded soils is also recommended, due to the potential for internal stability in these materials. The maximum fines content (< 75 Lim, Unified Soil Classification System) is restricted to no more than 5% and the plasticity of the fines is also limited. There is no restriction on the width of the filter, or the number of filter layers if it is used in a multi-layer system of filter. These requirements are further considered in ICOLD (1994), wherein the internal stability of a soil is also recognized as an important subject when analyzing soil filters. 2.1.1 Compatibility of base soil and filter Bertram (1940) reported the results of tests to verify the soil retention criterion proposed by Terzaghi (1939). This first systematic study established many key considerations for laboratory investigations of filtration phenomena. Specifically, the importance of using distilled de-aired water was identified to prevent any air in the water supply coming out of solution in the soil specimen and to avoid any decrease in permeability due to suspended solids in laboratory tap water. Details of the constant head test device are shown in Figure 2.4. Tests were either of 2 ri (at i = 18 to 20) or 4 h (at i = 6 to 8) duration, with unidirectional 14 flow imposed in a downward direction (and for selected cases in an upward direction) through the base soil located above a filter layer. Any incompatibility between the base and filter soils was found to initiate very quickly, with the visible movement of soil ceasing after three to five minutes. The minimum critical ratio between the 15 per cent size of the filter (D15) and the 85 per cent size of the base (dgs) was found to be approximately 6. It implied a margin of safety against inadequate retention of the base soil when using the Terzaghi criterion. The U.S. Army Corps of Engineers (1941) studied filter requirements for underdrains using two permeameters. A small permeameter of 7.6 cm diameter was used to test a thick layer of filter material and thin layer of fine base material. A second permeameter, 21.3 cm in diameter, was used with perforated discs to simulate drainage pipe. In both cases the specimen length was 16.5 cm. The material passing the discs was collected and weighed in each test. The duration of the test was between 2.5 and 5 hours in the small permeameter, and from 15 to 30 minutes in the large permeameter. The base material was a very fine sand and coarse silt, which was believed to be most susceptible to soil migration. A mixture of concrete sands and gravels was used as the filter. The flow was typically downward, and the loss of water head was measured at various points along the permeameter. Regular tap water was used in these tests. The side of the permeameter was tapped with a rubber mallet. No surcharge was applied to the soil specimen. Relatively stable conditions prevailed when the ratio Dis/dgs was equal to 5 or less in the small permeameter tests. When this value exceeded 5, the fine base material washed through the filter material. Introduction of large size particles into the filter material made it easier for the base fine to move through the filter. 15 Filter material should be packed densely in order to reduce changes in the gradation due to soil migration. Karpoff (1955) evaluated the design criteria for protective filters based on test results published by the US Bureau of Reclamation (1947) and (1955) using a well-graded silt, a uniformly-graded fine sand and medium sand, and a well-graded gravelly sand as the base soils. The filter materials were uniform gradations of medium or coarse sand or fine to medium gravel. A few selected details of the test methods are reported for comparison in Table 2.1. From these tests, Karpoff made the following observations that have since been widely adopted in filtration criteria: - "The filter material should pass the 75 mm (3") screen for minimizing particle segregation and bridging during placement. Also filters must not have more than 5 per cent minus 0.075 mm (sieve N° 200) particles to prevent excessive movement of fines in the filter and into drainage pipes causing clogging." - "The gradation curves of the filter and the base material should be approximately parallel in the range of the finer sizes, because the stability and proper function of protective filters depend upon skewness of the gradation curve of the filter toward the fines, giving a support to the fines in the base." - "In the designing of filters for base materials containing particles larger than 4.75 mm (sieve N° 4) size the base material should be analyzed on the basis of the gradation of material smaller than N° 4 size." 16 Karpoff (1955) recommended that these requirements, and additional rules for the ratio Dso/dso and Dis/dis for uniform and broadly-graded soils, should be met in designing a protective filter. Pare et al. (1982) performed a series of tests to simulate various base/filter combinations encountered in the LG3 dam, Quebec. The soils included widely-graded and gap-graded particle size distributions. Filtered de-aired or tap water was used depending on the water flow obtained in each soil specimen. The permeameter was relatively large, yielding a specimen 91 cm in length and 61 cm in diameter. A constant level tank allowed for control of the differential water head across the soil specimen. Tests were run under downward or horizontal flow of water, without any surcharge pressure (see Figure 2.5). Generally, filter tests in both vertical and horizontal directions were performed under five hydraulic gradients, with successive values of 0.5, 1.5, 2.5, 3.5, and 5. The average duration of each hydraulic gradient was 48 hours. Pare et al. (1982) confirmed that de-aired filtered water is beneficial, particularly in the case of long duration tests carried on fine materials. Gap-graded tills that were considered initially as potentially unstable, remained stable under severe conditions simulated in the laboratory. It is important to consider that Pare et al. (1982) were concerned that unstable soils could fail in a typical filter/base configuration. Lafleur (1984) performed tests on base soils that were well-graded gravelly silt-sands, considered representative of tills used extensively for dam construction in the James Bay project, Quebec. The test setup is shown in Figure 2.6. A constant water head was applied using calibrated springs to support water tanks, thereby maintaining a constant differential 17 water level. The soil specimen had a diameter of 15 cm with a base soil length of 15 cm and 20 cm for the filter. An effective cell pressure of 100 kPa, and back-pressure of 800 kPa, were applied to the soil specimen. An hydraulic gradient up to 8 was applied, and the test duration was between 50 and 880 h. All of the base soil movement was found to take place in the first 50 h. Analysis of laboratory data indicated that the limiting criterion of Bertram was applicable when the size d 8 5 (Dgs of the finer fraction) is based on the fraction smaller than sieve No. 4, thereby confirming the general observations of Karpoff (1955) (and the design criterion for soil retention first proposed by Terzaghi, 1939) Sherard et al. (1984) carried out an investigation to gain an improved understanding of the fundamental properties and behaviour of granular filters. The permeameter used is shown in , Figure 2.7. The base soil was a uniform fine, medium or coarse sand. The filter material was a uniform coarse sand, uniform gravel, or well-graded sandy gravel. The test specimens were 10 cm in diameter, with a length between 5 and 10 cm (base soil) and 12.5 to 17 cm (filter soil). The test was run with downward flow, and without surcharge, for a duration of 5 to 10 min. Afterwards, the soil specimen was placed on a shake-table for 60 s if little or no base soil had migrated through the filter layer. Results suggested the existing filter criterion, Di5/d85 < 5, is conservative but should remain as the basis for judging filter acceptability. This rule applies to base soils that are internally stable. Criteria based on the ratios D50AI50 and Dis/dis, such as those used by Karpoff, were not believed to have an experimental or theoretical basis, and it was suggested that they should be abandoned. It was also argued that the particle size distribution curve of the filter is not required to be similar in shape to the base soil. 18 Lafleur et al. (1989), used the same equipment and test methodology (as described by Lafleur, 1984; and Lafleur et al., 1986), to further study the response of base soils. The base soils were very well-graded silty gravel-sands, with a trace to some clay, and gap-graded silty gravels. The filter materials were uniform gravels, and well-graded gravel-sands. There were two series of tests. The first series, called compatibility tests, involved filters of increasing coarseness. In the second series, called screen tests, the quantity of soil particles lost during the self-filtration process was measured using glass beads as base materials. In the first series, after consolidation an effective cell pressure of 100 kPa, hydraulic gradients were applied in steps up to a value of 8 during an average test duration of 77 h. These constant head tests were performed on specimens that were 150 mm in diameter, with a length of 150 mm (base soil) and 200 mm (filter soil). In the second series of tests, a variety of diameter of glass beads were mixed to obtain bases with different gradation curves and a conventional square mesh sieve was used as a filter. Downward flow was maintained using a constant head tank, with hydraulic gradients ranging from 2.5 to 6.5. Vibration was applied to impede the formation of arches at the filter interface. The test duration was 2.5 h, after which sieve analyses were performed on different layers. It was found that the self-filtration process in the base-filter interface is very important to control the stability of the base soil. It was also identified that the broadness coefficient (CB = ® FA , where O ' F is the actual opening size / o of the filter, and d 0 minimum particle size) had a significant influence in the self-filtration process. A key factor was the opening size of the voids on the filter material. This must be compared with an indicative size of the base soil to ensure minimal particle migration. 19 Tomlinson and Vaid (2000) studied the influence of confining pressure on the retention capabilities of fihWbase systems (see Figure 2.8). Different values of the ratio Dij/dgs were evaluated at various levels of confining stress, hydraulic gradient, rate of gradient increase, and filter thickness. Downward flow of tap water was imposed. Surcharge pressure was varied between 50 and 400 kPa. The applied hydraulic gradient was in the range of 0 to 25. No vibration was applied during the test. The base soil specimen had a diameter of 10 cm and a length of 4 cm. Analysis of the data revealed no failure for Dis/dgs < 8 and spontaneous failure at Dis/dgs > 12. In addition, the critical gradient at which failure is triggered was found to decrease with an increase in Dis/dgs. At a given Dis/dgs, the critical gradient decreased with increase in confining pressure. 2.1.2. Internal instability studies An important finding from filter/base studies was to recognize that certain materials, despite satisfying Terzaghi's retention criterion (D15 < 5 dgs), fail because of internal erosion of the finer fraction of the soil itself (base or filter layer). Therefore, it was necessary to define which materials are internally unstable. The following summary addresses the main laboratory studies carried out to establish grain size distributions that are geometrically potentially unstable. Kezdi (1979) reported an investigation carried out by Bernatzik (1947) where "poorly-graded" materials, mixed from two components, underwent significant internal erosion. No further details were given about these tests. Two further examples were 20 mentioned in Kezdi (1979) from a report (in Hungarian) by Kezdi and Konyary (1971) where layers of different materials were tested to verify the condition of self-filtering. The results of these tests were used to propose an empirical geometrical method to assess internal instability of soils. Kenney and Lau (1985) studied the effect of disturbing forces such as seepage and vibration on internal stability. Compacted cohesionless soils were investigated using a constant-head device (see Figure 2.9) where specimens of approximate diameter 245 mm or 580 mm were tested. The smaller specimen had a length of 450 mm (base soil), and the larger one a length of 860 mm (base soil). A surcharge of 10 kPa together with flow conditions maintaining an hydrodynamic number R' > 10 were used during testing, where R'= qDi/nv and q is the unit flux, D 5 is the 5 % particle size, n is porosity, and v is the kinematic viscosity of water. Test duration was in the range between 30 and 100 hrs. (see Table 2.1). A mild vibration was applied to the specimen throughout the test, and was found to have a profound influence on the response of some of the soils. The results obtained during this study leave open the possibility that many existing dams have zone layers of questionable potential for instability. A method to assess that potential was proposed: discussions of the paper (Sherard and Dunnigan, 1986; Milligan, 1986; and Ripley, 1986) subsequently led to changes in the proposal that are discussed later in Section 2.2.1.2. Skempton and Brogan (1994) performed piping tests on sandy gravels. The objective was to compare the theoretical value of the critical gradient to cause "segregation piping" under upward flow on stable and unstable materials. The tests apparatus comprised a permeameter 21 13.9 cm in diameter, with internal piezometers located at different positions. The base specimen was approximately 15.5 cm in length, see Figure 2.10, with a screen and filter material below it. The test procedure imposed an incremental increase of the upward flow of water until piping occurred. Internal erosion yielded a big increase in flow. An abrupt transition was noted from unstable to stable response. The response found was similar to experimental results of Kenney and Lau (1985, 1986). Further, the empirical method proposed by Kezdi (1979) was found to be similar to their findings. Honjo et al. (1996) examined aspects of self-filtration behaviour in widely and gap-graded cohesionless base soils, using permeameters of diameter 15 and 32 cm. The specimen length was 10 cm, and it was supported on a metal screen filter (see Figure 2.11). A light surcharge of 0.9 kPa was applied to the specimen. Downward flow of distilled water was imposed, either with head control (at a hydraulic gradient between 2.5 to 14) or using a pump (for hydraulic gradients up to 52). Continuous tapping by a rubber hammer while maintaining the downward flow of water was also applied to the specimen. All tests were run for a period of two hours. The mass of soil passing the screen was measured to define the potential for internal stability. For gap-graded soils, a gap ratio of 4 (defined as the large diameter / smaller diameter in the gap range) was considered as an upper limit for stability. Garner and Sobkowicz (2002) described the results of laboratory tests that demonstrate that seepage and gas exsolution can combine to trigger directly the onset of instability. The nature of piping within an internally unstable material was studied through CTscan images. Gap-graded materials were tested using a permeameter of 95 cm length and approximately 22 30 cm diameter under a seating load of 600 kPa. A filter soil was placed over a base soil and subjected to upward flow, and run for about 2 months increasing the overall hydraulic gradient in stages. A second phase consisted in the exsolution of gassy water where suffosion was observed for the first time after about 10 hours and continued for approximately 72 hrs until catastrophic suffosion was observed. Moffat (2002), used an ASTM Gradient Ratio test device (see Figure 2.12), to provide further confidence to the use of geometrical approaches in assessing the potential for internal instability. A surcharge of 25 kPa was applied to the top of a specimen of 10 cm. diameter and 10 cm length. Distilled de-aired water was used, with downward flow at hydraulic gradients between 0.1 and 15. Each test had an approximate duration of 10 hrs of seepage and 1 hour of seepage and vibration. The Kezdi (1979) and Kenney and Lau (1985, 1986) methods (see Section 2.2.1), based on shape of the gradation curve, both proved successful. Kenney and Lau (1985, 1986) may be unduly constrained by a limit on the value of the percentage finer (F, %) over which the H:F boundary of 1:1 is evaluated. It was therefore suggested to evaluate the (H/F) value for 0 < F < 30%, independent of whether the grain size distribution is widely graded (WG) or narrowly graded (NG) as originally proposed by Kenney and Lau (specific details of this method are given later in the chapter). Conditions such as the flow regime, energy imposed by tapping the specimen and geometry of the mesh filter screen were found important when evaluating the mass of soil migrating. This presents a challenge when attempting to characterize the response at the field scale from laboratory element tests. It was recommended that permeameter studies should have the 23 following characteristics: (i) measurement of axial load at top and bottom of the soil specimen in order to characterize effective stress along the specimen length; (ii) companion measurement of pore water on opposite sides of the permeameter, to better characterize spatial variations of the hydromechanical conditions. 2.2. Analytical studies 2.2.1 Empirical criteria: internal instability Interpretation of laboratory permeameter studies have led to criteria in the literature (Kezdi 1979; Kenney and Lau, 1985-86) that appear to define a geometrical threshold to the onset of internal instability in cohesionless soils. These methods are based on analysis of the grain size distribution curve. No hydromechanical constraints, for example hydraulic gradient and effective stress, are considered in these methods. They were validated by laboratory testing and are empirical: confidence in their applicability to other soils depends on the similarity of those soils with the ones used in development of the criteria. 2.2.1.1 Kezdi criterion Kezdi (1979) describes an empirical method, based on laboratory tests, where the original grain size curve is split into a coarser and finer fraction: Kovacs (1981) reported that Kezdi first proposed this method in 1969. Each component or fraction should satisfy the filter/base retention criterion of Terzaghi (1939), to be considered stable. Interesting, De Mello (1975) 24 also advocates the same method to check the acceptability of "skip-grading" materials with respect to internal erosion. The method makes use of the Terzaghi-type ratio D'is/d'gs where, in this case, D'15 is the diameter for 15 % mass passing in the coarser fraction and d\s is that corresponding to 85% mass passing in the finer fraction (see Figure 2.13). According to this method, D'15 characterizes the pore size constriction of the coarser fraction, and d'ss is the representative grain size of the finer fraction. Consequently, when the finer fraction particles are smaller than the pore size constriction then the soil will be potentially unstable under seepage. A value D'is/d'gs > 4 would indicate the soil is potentially unstable. The method is ideally suited to assess the potential for internal instability of bi-model or gap graded soils where the percentage of finer fraction (Sf) does not exceed 35% of the soil mass. This ensures that the skeleton particles are in contact with each other and not dispersed in a fine matrix. 2.2.1.2 Kenney and Lau criterion Kenney and Lau (1985) carried out permeameter tests on well graded sandy gravels. Results of the seepage tests were used to propose a method for evaluating the potential for grading instability based on the shape of the grain size curve (see Figure 2.14). In this method, the value of F corresponds to the "mass fraction smaller than" a particle diameter D. The value of H corresponds to the mass fraction between particle sizes D and 4D. The rationale for 25 selecting a ratio of 4D is the size of the predominant constrictions in the void network of a filter is approximately equal to one-quarter the size of the smallest particle making up the filter. The postulated boundary between stable and unstable grading curves was initially defined as H/F =1.3, over the lower portion of gradation to F<20% for a widely-graded primary fabric, and F<30% for narrowly-graded soils (1985). The boundary was amended to Ff/F = 1.0 (Kenney and Lau, 1986) upon discussion of the data by Milligan (1986). 2.2.2 Analytical models Various attempts have been made to develop models that could be use to predict the behaviour of internally unstable soils under seepage flow. Physical explanations of the process have been given using limit equilibrium analysis for individual particles (Indraratna and Radampola, 2002) in a pore channel, as well as using balance equations considering porous media theory (Wilhelm and Wilmanski, 2002). 2.2.2.1 Limit equilibrium analysis of particle migration An explicit solution for the critical hydraulic gradient to move a base particle within a pore channel was proposed by Indraratna and Radampola (2002) for cohesionless soils. Limit equilibrium in the direction of flow was implemented. A drag force, frictional force between the particle and the pore wall, weight of the particle, and uplift force are considered to act on 26 the base particle (see Figure 2.15). This mathematical formulation was verified with laboratory tests on a fine gravel filter and base sands. An important parameter that influenced the critical hydraulic gradient was the minimum equivalent diameter (d0). It depends on the porosity of the filter (n) the shape of the particle (a), and the equivalent diameter (Dh). The minimum equivalent diameter is given by: d0=2.67 h- Eq. 2.1 \-n a knowing: Dh =—l—- (Kovacs, 1981) Eq. 2.2 A where the distance between the maximum and minimum diameter is divided into n equidistant intervals. The average diameter in the i-th interval is D;, and the weight of grains in this interval compared with the total weight of the sample is represented by ASj. The relation between critical hydraulic gradient, found in the laboratory, and the pore diameter do, is shown in Figure 2.16. 2.2.2.2 Porous media theory Porous media theory has made significant contributions to the development of mathematical models in soil mechanics and other engineering fields. Fundamentals of the theory are 27 described in the following paragraphs. Porous media theory can be used in soil mechanics to explain the behaviour of unstable soils under seepage flow. Solids consist of different solid components with closed and open pores. The pores are filled with fluids and there is interaction between solid and fluid components. Exact description of the solids is almost impossible; therefore a volume fraction concept is used. Mixture theory, in combination with the volume fraction concept, is used to study the behaviour of the different components that define the behaviour of soils. Boer (2000) reports that, in 1936, Fillunger founded the mechanical theory of porous media, but his work was forgotten for years. Then mixture theory was finally developed at the end of the 1950s and in the 1960s. In the 1970s, the concept of volume fraction, together with elements of mixture theory, was first used to obtain a constitutive model for a fluid-saturated porous solid. Arguments of continuum mechanics were also used at this time. A balance equation of forces and inertia was defined. In the following paragraphs the key concepts behind this theory are explained. These general concepts are used later in the thesis to explain the influence of different conditions, such as local hydraulic gradient and effective stress, on the onset of internal stability. To account for materials that are composed of different fractions (for example, water, air, soil particles, etc.), it is necessary to define a notion of volume that describes the material as a continuum. With the volume fraction concept, there is a control space where it is assumed 28 that the pores are statistically distributed and that the actual arrangement of particles is composed of volume elements of the real constituents. Neither the geometry of the pore structure nor the exact location of the individual components is considered in this macroscopic model. The control space of porous solid Bs, with boundary 8Bs consists of constituents with volumes a 1 . Therefore the volume fraction can be defined as: »'(*,/) = — Eq. 2.3 da where x is the position vector of the actual placement and t the time. The volume fractions n1 must satisfy the following condition for the k constituents: k X w ' = 1 Eq. 2.4 1=1 The main kinematic assumptions in the porous media theory are: - each spatial point x is simultaneously occupied by materials points of all the constituents at the time t. The material points proceed from different reference positions Xi at time t=to; and, - each constituent is assigned an independent state of motion. The motion of the constituents is defined as a chronological change of placements (fr. The spatial position vector x of the material can be derived from this succession of placements 29 and a reference position vector Xi at time t = to, from which the following relation holds at time t: Eq. 2.5 Velocity and the acceleration of a material point of a constituent are defined as: Eq. 2.6 dt dt2 Eq. 2.7 In addition, balance equations have to be used for each constituent, and the interaction between the constituents and external agents must be considered. The balance equations are composed of: balance of mass, balance of momentum and moment of momentum, as well as balance of energy. The sum of all interaction effects should be zero. Considering a continuum with a multi-component material where each component itself is a continuum filling the entire space, the temporal rate of change of G a (any physical property, for example, mass) can be expressed using Reynold's transport theorem by: Eq. 2.7 where: 30 VGCX: velocity field Ia: temporal rate of production of property G a that takes place at U . Using this approach an equation can be formulated for each constituent mass balance equation. The superposition of each one of these mass balance equations must reflect the balance equation of the whole mixture. Therefore for each constituent the following equation can be derived: where p' is the mass supply per volume element. Another important equation is that derived from the balance of momentum. This equation is obtained when the material derivative of the momentum is equal to the sum of external forces. Then the equation of balance of momentum for a can be obtained from: where p a is the unit density, T a is the stress, b a is the external acceleration, a« is the acceleration, and v a is the velocity. This concept of momentum balance of each component has the potential to explain hydromechanical influences governing suffosion. Balance of momentum of each component must be considered and the balance momentum for the mixture is that given by superposition of the momentum of all the constituents. This is dt Eq . 2.8 ctivT" + paba +pa = paaa +pavt a Eq. 2.9 31 the general conservation principle. More details on the derivation of these equations are given in Bear (1972). 2.3 Effects of non-plastic fines This thesis examines the internal stability of soils with different fines contents. Variations in fines content may cause a change in the hydromechanical threshold to onset of internal instability. In the empirical methods described previously, the maximum amount of "mobile" finer fraction has been established to be between 20 to 30% of the total mass (Kenney and Lau, 1985). This suggested limit is now examined with reference to literature describing the effect of fines on the shear strength of coarse-grained soils. Shear strength of a soil containing a sand- silt fraction is an important issue in numerous geotechnical problems. Experimental data show that a change in fines content may either decrease or increase the shear strength of the soil. The behaviour seems to vary depending on the range of fines content. Essentially, the finer fraction content influences the behaviour of gap-graded granular mixes. Inter-granular void ratios have been used to characterize the behaviour of these soils. Thevanayagam (1998) performed an experimental study and obtained a correlation between void ratio and undrained shear strength. The basic idea was that the soil behaves as a skeleton of particles arranged in an organized fashion: In some cases, the coarser fraction would carry most of the load and the finer fraction would not participate significantly in load transfer. 32 This observation must be considered in any limit equilibrium analysis of the finer fraction particles, since the finer fraction has less strength in this case (see also Skempton and Brogan, 1994). Thevanayagam, proposed to use a modified void ratio (inter-granular and inter-fine void ratio) to explain the change in shear strength when introducing non-plastic fines. It was found that the large strain undrained shear strength (sus) depends on the inter-granular void ratio. Thevanayagam (2002) also defined a threshold limit for fines content (FCth) at which the strength of the soil is controlled mainly by the of fines. At low fines content, the intergranular friction (skeleton of the soil) plays an important role, and finer particles play only a secondary role. Vallejo (2001) studied the shear strength and stiffness of mixtures of sand-gravel sizes of glass beads and mixtures of sand and silt. The fines concentration was changed to determine its influence on the shear strength of the mixture. Porosity changes in the granular mixtures were used to explain changes in its shear strength. A change in porosity occurs with percentage by weight of fine particles (see Figure 2.17). After the minimum porosity is reached, there is no more space between the coarse grains for additional fine grains. Beyond this point, the coarse grains become separated and the fine matrix becomes the dominant material in the mixture (see Figure 2.18). The percentage of finer fraction that separates the behaviour of the material was shown to be in the range 25 to 30% in weight (see Figure 2.17). This percentage is equivalent to the maximum value adopted as "mobile" fines in studies of soil migration for material supported by the coarse fraction. It compares well with the change in peak shear strength found in binary mixtures (see Figure 2.19, after Vallejo, 2001), as a function of normal stress applied in direct shear tests. 33 Accordingly, finer fraction content of a soil may influence its behaviour in several aspects, including: - Strength of the soil mixture - Load transfer between finer and coarser fractions - Maximum amount of finer fraction (mobile fraction) inside the voids of the skeleton A l l of these aspects likely influence the onset of internal stability under seepage flow and the method used to assess whether a soil is potentially unstable or not and wil l be considered in this thesis. 2.4. Implications for professional practice As mentioned in Chapter 1, the W A C Bennett Dam experienced a sinkhole incident that led to a programme o f research to establish a likely explanation for its probable cause. Internal erosion by suffosion of earthfill materials at different zones of the dam was identified as a candidate mechanism. The potential for suffosion is governed by geometrical constraints. The onset of any instability is governed by hydromechanical constraints. Although the literature contains empirical methods for evaluating the geometrical potential for internal instability, the confidence with which they can be applied in practice is severely limited by uncertainty regarding their application to soils of varying gradations, and by a lack of experimental observations on conditions that trigger the onset of instability itself. A primary aim of this thesis is to examine factors contributing to the occurrence of suffosion; in widely-34 graded cohesionless soils, and postulate a hydromechanical explanation that will support the evaluation of soils that are deemed potentially unstable in geotechnical practice. 35 Table 2.1 Summary of selected studies Year Author Specimen Size Surcharge Water Quality Hydraulic gradient Flow direction Vibration Duration 1940 Bertram base h = 6 cm d = 5 cm unif. d = 10 cm broad. no surcharge distilled and de-aired 6-8; 18-20; 10-12 (i =50 preliminary test) downward or upward no 2-4 hrs; 161/2hrs 1941 U.S. Army Corps of Eng. h = 16.5 cm d = 7.6, 21.3 cm no surcharge tap water 1 to 2 downward some upward tapping 15 - 30 min 2.5 - 5 hrs 1955 Karpoff, K.P. h = 20 cm d = 20 cm no surcharge not mentioned 1 to 24 downward no not mentioned 1979-1984 Sherard, et al. h = 6.5 in. base d = 4 in. no surcharge fine soil: distilled coarse soil: tap water 4 kg/cm2 pressure through the slots downward or horizontal with and without vib. few minutes 1982 Pare, J. h = 95 cm d = 61 cm de-aired and tap water depending on flow 0.5 to5 downward or horizontal no aver: 48 hrs max. 250 hrs 1984 Lafleur, J h = 15 cm d = 15 cm back pressure not mentioned 8 downward no 50 - 880 hrs 1985 Kenney and Lau h =20-50 cm d =24.5-58 cm 10kPa recirculated water R'>10 downward mild vib. 30-100 hrs 1994 Skempton and Brogan h =15.5 cm d = 13.9 cm no surcharge not mentioned 0-1 upward no 1.5 hrs 1996 Honjo et al. h =10 cm d = 15 and 30cm 0.9 kPa tap water, no de-aired 36-52 pump pres. 2.5 -19 constant head downward tapping 2 hrs 2000 Tomlinson & Vaid h =4 cm d = 10 cm 50 - 400 kPa tap water 25 downward no 3 1/2 hrs Current UBC h = 10 cm d = 10 cm 25kPa distilled and de-aired 0.1-15 downward automatic vibration 11 hrs 36 Geometric Hydromechanical Terzaghi (1939) Bertram (1940) Karpoff (1955) Lafleur (1984) Sherard (1984) Tomlinson and Vaid (2000) Indraratna and Radampola (2002) Kezdi (1979) Kenney and Lau (1985) Skempton and Brogan (1994) Honjo et al. (1996) Moffat (2002) Skempton and Brogan (1994) Indraratna and Radampola (2002) Garner and Sobkowicz (2002) Current study Figure 2.1. Selected studies on particle migration Figure 2.2. Retention criteria (Terzaghi, 1939) 37 Base soil retention Uniformly graded « Widely-graded Gradation shape Gap-graded Finer fraction particle size (dss, Dm. or D30) 1 Coarser fraction particle size (D1S or Ded) Permeability: Minimum particle size (DS o r DO) Par*tele size : irJjSfBbutlWI : Maximum particle size (Dim) Granular filter 3 L~ Thickness Mineralogy Placement and durability I Figure 2.3. Functional requirements of granular filters 38 39 40 Filter paper Figure 2.6. Permeameter test assembly (after Lafleur, 1984) 41 Figure 2.7. Permeameter device (after Sherard et al., 1984) 42 Carmecttort to dHfefenliaf pressure transducer Top pS&tien 1.5 mm ffl»ab Bottom pedestal Ssmpte collector OO<J0O0O0O000OOOO0OO OOOOCOOOOOQOOOOOOOO ooooGooocsociOooaaaao ooooooooooooooooooo iX)pcyo .OOOOQQ T I SOU Hter Figure 2.8. Permeameter setup (after Tomlinson and Vaid, 2000) Uj"'»*:?' M'S-rt v.-5i> • it>;!?t • Tf5:S! :;;0f;>::.:r<:<jr': Sf...j s,?> h - X - { :>•.. >•?' J>,^ ' / • » * s !t t v ( » » • ' « i ' t ' f'ji ' -^ ' . •>, i - - t ' r ' r ,» < ^ i ,< <)' J 0 < J 5" *V &x rr i*s on \ r <V i >~>f*l< ' Figure 2.9. Test permeameters (after Kenney and Lau, 1985) 43 Figure 2.10. Test permeameter (after Skempton and Brogan, 1994) Figure 2.11. Permeameter setup (after Honjo et al., 1996) 44 Figure 2.12. Permeameter configuration (after Moffat, 2002) Figure 2.13. Stability criterion of Kezdi (1979) 45 S A N D GRAVEL Fine iMediuml Coarse Fine | Medium 1 Coarse 1 Gra in s ize D. m m 0.02 0.060.1 0.2 0.61.0 2 6 10 20 60 100 200 1.0f Figure 2.14. Stability criterion of Kenney and Lau (1985) Filter/base system • Limit Single particle —* equilibrium do: minimum pore diameter Weight of the particle Uplift force Hydrostatic foroe Frictional force Assumptions Assemblage of unconnected fluid conduits d: particle size Inertia forces neglected Porosity, n Dh: average diameter N: number of pores in the model tanC*) 6s: length of 6 s • filter pore ohannel thickness Ciitical hydraulio gradient, icr Figure 2.15. Limit equilibrium for a single particle (based on Indraratna and Radampola) 46 40f> Figure 2.16. Critical hydraulic gradient (after Indraratna and Radampola, 2002) $ i.a&«raN>ry Fefssrtssgs by Weight of Coarss (tse) and Pirn faf) Past^Am Figure 2.17 Changes in porosity in the coarse-fine granular mixture (after Vallejo, 2001) 47 Transition: finer fraction supported Finer fraction supported-coarser fraction floating Figure 2.18 Coarser-finer fraction mixtures (based upon Vallejo, 2001) so 101 100 80 * "n - 84.3 W>a i t - - "n-^sekPa 60 10 20 0 • 0 20 10 ' 60 80 .100 Percentage by Welghi ot Coarse (uc) and Fine foil Particles Figure 2.19 Peak strength versus percentage of fine fraction (after Vallejo, 2001) 48 3. TEST DEVICE 3.1 Introduction Review of the literature reveals a considerable body of experience on empirical rules to establish geometrical constraints governing the potential for internal instability in a soil. However, the companion understanding of hydromechanical influences is very limited. Accordingly, a permeameter test device was designed and commissioned with the specific intent of examining hydromechanical parameters that govern the onset of internal instability. The design procedure took into consideration features and experience reported in previous studies, most notably those of Bertram (1940), Lafleur (1984), Kenney and Lau (1985), and Garner and Sobkowicz (2002). A primary focus of this experimental study was the relation between effective stress and hydraulic gradient, in soils that experience a seepage-induced migration of particles. Investigation of this relation governed the design of the main features of the permeameter test assembly. Following reconstitution of the tests specimen in the rigid-wall permeameter, it was subjected to a stage of consolidation under axial load. Upon completion of consolidation, it was subject to seepage flow. The distribution of effective stress in the test specimen was determined from measurements of axial load, and therefore total stress, together with measurement of pore water pressure. Hydraulic gradient was calculated knowing the spatial variation in pore water pressure along the length of the specimen. Additional consideration was given to any axial deformation of the specimen as a result of consolidation, seepage 49 flow, and the onset of failure by particle migration. Other variables such as water flow, temperature were also recorded, and used in analysis of the specimen response to flow. 3.2 Permeameter test device The arrangement of the permeameter test device is illustrated in Figure 3.1. The assembly comprises a permeameter cell that mounts in a reaction frame. The test specimen is subjected to unidirectional seepage flow by means of a water supply system that provides flow from an inlet tank to an outlet tank. Flow was imposed in either a downward or an upward direction through the specimen. In design and fabrication of the test device, particular attention was paid to details governing: (1) axial loading of the top surface of the specimen; (2) a reaction system; and, (3) quality of the supply water. In using instrumentation to monitor the response of the specimen, particular attention was paid to: (1) variation of effective stress along the length of the test specimen; and, (2) the average hydraulic gradient applied across the specimen, and any local variation of that gradient along the length of the specimen. 50 3.2.1 Permeameter cell The rigid-wall permeameter comprises an approximately 1 meter long acrylic tube in which the test specimen is reconstituted (see Figure 3.2). The acrylic tube has an internal diameter of 279 mm and wall thickness of 13 mm. The permeameter accommodates a test specimen 279 mm in diameter (D), reconstituted to a length (L) that varied between 30 and 60 cm (see Fig 3.3). The acrylic cell sealed to a top and base plate through the use of O-rings mounted between the cell and plates: compression was applied using six tie rods between the two plates (see Figure 3.2). The test assembly was designed to impose unidirectional seepage in either the upward or downward direction. For the case of downward seepage (see Figure 3.3.) inflow was introduced through an inlet port (dia. = 15.7 mm) on the top plate. Outflow was collected in three concentric and interconnected channels in the base plate, and discharged through and outlet port (dia. = 15 mm). The configuration was designed to minimize any zone of preferential flow through the specimen. 3.2.2 Top boundary and axial loading system The axial loading system comprises a piston mounted on a reaction frame. The piston was a double-acting air cylinder, Nopak model CLP6-10x5-CJ-NN-l3/4 -4, with a stroke of 150 mm that could apply a maximum load of 2200 kgf. Axial load was transmitted to the test 51 specimen by a rod and loading plate (see Figures 3.1 and 3.3), to a maximum contact or vertical effective stress of 350 kPa. The loading rod passes through the top plate inside a "frictionless seal system", details of the frictionless sealing system are shown in Figure 3.3, Detail A. The design is modeled on that of a triaxial cell. The sealing action relies on injecting air into a cavity at the same pressure as the water below. Vertical motion of the loading plate is guided by a pair of stainless steel linear ball bushings. A measurement of friction-induced loss of axial force on the loading plate gave a maximum value of approximately 1 kgf. The loss is deemed negligible since it gives an error in vertical stress of 0.15 kPa, which is less than 1% of the minimum stress applied during the test series. The 12.4 mm thick loading plate is made of stainless steel and has a diameter of 276 mm, which ensures a narrow gap between the edge of the plate and the inside wall of permeameter. A total of 20 perforated holes, 12.8 mm diameter, are located symmetrically around this plate (see Figure 3.4). The gap between plate and wall of the acrylic tube was filled by a thick nonwoven geotextile that was bonded to the edge of the plate to prevent any passage of fine particles through the gap (see Figure 3.3, Detail B). A wire mesh screen, of diameter equal to that of the loading plate, is located between the loading plate and top surface of the soil specimen. The opening size of the screen was selected in accordance with the direction of the seepage flow and grain size distribution of the soil tested. 52 3.2.3 Bottom boundary and reaction system The configuration of the bottom boundary is very similar to that of the top boundary in as much as it comprises a wire mesh screen, and base frame that provides a reaction to the axial force of the loading rod, self weight of the specimen, and any seepage forces (see Figure 3.3. Detail C). Again, the bottom configuration is intended to be "frictionless", and includes a thick geotextile to seal against the acrylic tube. The test specimen is supported on a base frame made of anodized steel. This frame allowed the observation of particle migration at the bottom of the specimen (see Figure 3.5a). The frame comprises an upper perforated plate, with holes 8 to 13 mm in diameter, located symmetrically to ensure a uniform flow of water (see Figure 3.5b). The upper plate is fixed to a lower perforated plate, by four struts; the frame sits on a base load cell. A wire mesh screen was placed between the test specimen and upper perforated plate of the base frame. The opening size of the screen was selected in accordance with the direction of seepage flow and grain size distribution of the soil tested. The arrangement of acrylic tube and base frame allowed for visual observations of any particle migration in the test specimen and, importantly, its spatial and temporal variation across the cross-sectional area of the base, in tests with downward seepage flow. 53 3.2.4 Water supply system Seepage flow to the permeameter was from an inlet tank that is charged from a water reservoir (see Figure 3.1). A water supply system was used to produce clean de-aired water in the reservoir, from the tap water in the laboratory. The supply system comprises two elements: (1) a two-stage filter system; and, (2) a de-airing system (see Figure 3.1). In the two-stage filter system filter system water is first passed through a sand filter, type ZWMULTIMD, manufactured by Millipore (see Figure 3.6). It removes all particles larger than 10 um. Water is then passed through a carbon filter, model P72017, manufacture by Millipore. This second stage removes particles larger than 3 pm. The filter system requires periodic maintenance, wherein the filter sand is cleaned and the carbon units changed to ensure good quality of the water. Filtered water is held in a storage reservoir, where it is de-aired. The reservoir is a water tank manufactured by John Wood, with a capacity of 270 litres. The process of de-airing involves introducing the water into the reservoir, after first imposing a vacuum of approximately 80 kPa on the reservoir tank itself. Vacuum is pulled using the laboratory supply line, and controlled by a Fairchild regulator, model 16222. A filter prevents water from entering the vacuum line. Experience showed that, when the water is left under the vacuum for more than 6 hours, the dissolved oxygen content falls below 2.5 mg/L. 54 This system was used to prepare water in the storage reservoir for supply to the inlet tank of the permeameter. Depending on the hydraulic conductivity of the test specimen, and the test duration, the inlet tank would be re-supplied with clean de-aired water, as needed. 3.2.5 Seepage control system Seepage through the test specimen was controlled by means of the inlet and outlet tank. The tanks are Jacuzzi models AP-260/302, manufactured by Aqua P.O.D., with a maximum capacity of 320 litres. Each tank has an internal membrane that provides an impermeable air-water interface. It prevents air from dissolving into the de-aired water supply from the storage reservoir. The operating principle was one of head-control rather than flow-control. Head-control was achieved through control of the differential air pressure on the air/water interface of the inlet and outlet tank respectively and therefore the differential water head across the test specimen. Differential air pressure was controlled by an electro-pneumatic regulator, a Fairchild model TA600-505, with input voltage of 0 to 5 VDC and output pressure range of 0 to 415 KPa. A feedback system that comprises: a) the previously mentioned electro-pneumatic regulator, and b) computer software. This system controls changes in pressure between the outlet and inlet valves (measured with total pressure transducers) to keep a constant overall hydraulic gradient during each stage of the test. 55 3.3 Instrumentation An array of transducers was used, both to enable control of seepage flow as described above, and to monitor the response of the test specimen to seepage flows. Given the objectives of the experimental study, particular emphasis was placed on determining the variation of effective stress and hydraulic gradient. 3.3.1 Axial load Effective stress on the top surface of the test specimen is calculated knowing the axial force in the loading rod. It was recognized, at the time of design, that sidewall friction would cause a variation in effective stress along the length of the specimen. To account for sidewall friction, load is measured at both the top and bottom of the specimen. The top load cell is located on the loading rod, above the top plate and therefore outside of the permeameter cell. In contrast, the base load cell is located inside the permeameter, between the base frame and base plate (see Figure 3.7). The top load cell is an Interface model 1210AF-5K, with a capacity of 2250 kgf (see Figure 3.8a). The base load cell is a waterproof model 41/5700-02 manufactured by Sensotec, with a capacity of 2250 kg (see Figure 3.8b). Given the resolution of these transducers and attributes of the data acquisition system, the deduced values of vertical effective stress on the top and bottom boundaries of the specimen are believed accurate to ±0.5 kPa. 56 3.3.2 Axial displacement The length of the reconstituted test specimen (L) changes, first in response to consolidation under the influence of axial loading, and second in response to seepage flow. Additionally, potential exists for a change in specimen length as a result of loss of soil particles following the onset of internal erosion. In order to monitor these changes in specimen length, axial displacement of the top loading plate was measured throughout a test. Displacement was measured using a Linear Variable Differential Transformer (LVDT) mounted on the loading rod. The LVDT was custom manufactured, and measurements of displacement are believed accurate to ±0.1 mm. 3.3.3 Variation of water pressure A differential water head and therefore average hydraulic gradient, i a v , is imposed across the test specimen. Accordingly, water pressure varies along the length of the specimen throughout a test. The water head distribution was measured and used to calculate the variation of local hydraulic gradient across discrete intervals of the specimen length. Measurements were made with two arrays of transducers. One array comprises 7 total pressure transducers (TPT) that connect to ports on the sidewall of the permeameter. The other array comprises 6 differential pressure transducers (DPT) that connect across ports, at the same elevation, on the opposite side of the permeameter. A schematic layout of the port locations is shown in Figure 3.7. The distance between each successive port location is 125 57 mm. At the base of the specimen, between ports 6 and 7, the head loss occurs over a discrete interval of 30 mm. Length of a reconstituted specimen, which varied by soil type, determined the corresponding interval at the top boundary. The total pressure transducers are a Schaevitz model PS 10000, 0 to 5 VDC output, with a pressure range between 0 and 350 kPa. The differential pressure transducers are a model PDW/E972-05-01, manufactured by Sensotec. The range of these sensors is -70 to 70 kPa, with an output signal between -5 and 5 VDC. The transducers are believed accurate to ±0.5 kPa and ±0.1 kPa respectively. 3.3.4. Weight of inflow/outflow water tanks A measurement of volumetric flow (Q) is required for calculation of hydraulic conductivity of the specimen. In addition, seepage velocity is also deduced from a measurement of water flow through the specimen. Therefore, an automated system was designed to continuously measure seepage flow, by means of the change in mass of water held in the inlet and outlet tanks. These water tanks were each mounted on a weigh scale that measured the total mass, and hence volumes of seepage flow, over the duration of a test (see Figure 3.9). The weigh scales are Mars Scales MSG series, with a capacity of 900 kg and an accuracy of ±0.1 kg. Output of the load cell in each weigh scale is read directly by the data logger. The output was recorded every second, to check for any sudden change in water flow. Hydraulic conductivity is calculated with an accuracy that varies with elapsed time of measurement and 58 average hydraulic gradients. For example, with an elapsed time of 60 min., values of\ hydraulic conductivity are believed accurate to ±5xl0" 5 cm/sec for i a v =1.0, and ± lxl0" 6 cm/sec for i a v = 20. 3.3.5 Water temperature and dissolved oxygen content The permeameter is located in a temperature and humidity-controlled laboratory. Accordingly water in the storage reservoir and inlet/outlet tanks is held at a constant temperature of 20° C ± 1° C. A thermocouple is used to verify the temperature of the water flowing through the test specimen. The custom manufactured type "J" thermocouple mounts through connected to a port located below the soil specimen, to measured water temperature inside the cell. It is accurate to ±0.5°C. The dissolved oxygen content of the water was periodically measured at both the inlet and outlet, to verify it remained de-aired. The range of measurement is 0 to 19.9 mg/L, with an accuracy of + 0.3 mg/L (see Figure 3.10). 3.3.6 Data acquisition and control system Acquisition and control in each stage of testing was carried out by an automatic system. Data acquisition reads the output of the instrumentation (18 transducers and 1 regulator) and 59 writes the data to storage. A total of 24 input channels and 2 output channels were used to * assist 18 instruments. To control the imposed average hydraulic conductivity, an electronic regulator was used. To assist instruments and the electro-regulator, a power supply and A/D board system connected to a personal computer was used. Commercial software was used to acquire the data and to control the test. The acquisition system comprises: (1) power supply for the A/D board and another one for w the instruments; (2) Input and output modules; (3) ethernet network module; (4) a personal computer; (5) software. Power to the A/D board was provided by power supply model PS-4 manufactured by National Instruments. Power to the instruments was provided by a custom made power supply with an output voltage of approximately 15 Volts. The A/D board comprises different modules. The PP-AI-100 module manufactured by" National Instruments includes 8 single ended analog input channels. These are 16 bits resolution channels with a current input range between 0 to 20 mA or voltage between -10.4 to 10.4 volts, and with a maximum reading frequency of 6 Hz. There is an electronic filter on each channel, which can be configured for 50 Hz or 60 Hz. The FP-TC-120 module also manufactured by National Instruments includes 8 differential inputs in the milli-volts range or to be used with thermocouples. This module has the same resolution and filter capabilities as the previous module. Output component of the system consists in an analog output module includes two 0 to 5 V analog output channels, generated with 12-bit of resolution. One of these channels is used to control the electronic regulator used to apply pressure in the inlet water tank. 60 The Ethernet network interface connects a node of a module to a high-speed Ethernet network. It manages communications between the host PC and the inlet and outlet modules. A personal computer, Pentium 4 manufactured by Seanix, was used during testing to collect and save the data read through the A/D board. The software used to control and save data from the data acquisition system is the National Instrument Lab VIEW Full development system 6.1. Lab VIEW is a graphical development environment that helps you build data acquisition monitor and control system. Lab VIEW is used to create a graphical program called a virtual instrument (VI) where each one of the instruments and control systems are display (see Figure 3.11). 3.3.7 Photo and video equipment A digital camera, model DSC-P71 manufactured by Sony, and two digital camcorders, model DCR-TRV18 manufactured by Sony, were used to keep a visual record of the test specimen at various stages of a test. These digital images enable a comparison of any physical changes in.the specimen, resulting from seepage flow, as well as the response to the onset of internal instability in tests. 61 3.4 Influence of the equipment Attributes of laboratory test equipment can affect the behaviour of the soil specimen observed during testing. Knowing the main characteristics of the permeameter, and their likely effects on the behaviour of the soil, ensures they are properly addressed in the experimental programme. 3.4.1 Diameter and length of the specimen The difference between diameter of the test specimen and the maximum particle size of the soil may influence results obtained in soil mechanics testing. Boundary effects, such as arching, also exert an influence on the soil specimen. The standard method for permeability testing of granular soils requires a minimum ratio of 8 to 12 (ASTM D 2434) between the diameter of the permeameter cell and the maximum particle size of the tested specimen. The permeameter cell used in this research had an internal diameter of 279 mm. This diameter yields a ratio of 3.7 for the transition materials and 14.7 for the core materials. Consequently only core material fully satisfies the ASTM guidance. Yet, only approximately 15% in mass of the transition material yields a ratio smaller than 8 recommended by the ASTM. Therefore these are isolated particles that are not in contact. In addition, visual observations during failure confirmed that the behaviour of core and transition materials was very similar. 62 Consequently, the size of the test permeameter was considered satisfactory for the programme of testing. During a previous study, it was found that a ratio between length and diameter of the specimen equal to 1.0 or larger was adequate for internal erosion assessment (Moffat, 2002).1 The minimum ratio between length and diameter during this programme of testing was 1.2. Therefore, the length of the soil specimens was considered adequate. Horizontal deformation of the acrylic cell was limited by the use of metallic ring clamps placed at four locations on the permeameter cell. This technique increased the rigidity of the •-acrylic cell and reduced the horizontal deformation to less than 0.15 % of the diameter of the specimen (Test C-30-100-U). It is believed the magnitude of this deformation did not influence significantly the response of the test specimens. 3.4.2 Filter screen opening size In an internally unstable soil, soil particles move inside the skeleton of the soil. Therefore, to allow complete failure and release of the finer fraction the outlet boundary of the specimen has to allow the free passage to particles. In contrast, this exit boundary has to be able to retain the skeleton of soil in order to avoid losing the complete specimen even if the soil is 63 not internally unstable. Therefore, these two conditions define the minimum and maximum opening size of the wire mesh screen used at the outlet face of the soil specimen. Wire mesh with opening sizes of either 2.76 or 1.14 mm were used in the exit face of the soil specimen. Each mesh allowed movement of the finer fraction of the soil and maintained the coarser fraction in place. All transition materials had an exit wire mesh opening equal to 2.76 mm. This opening size is equivalent to D 4 0 of the transition material, and therefore is located in the flat zone of the grain size distribution curve (see Figure 4.3). This is a clear division between the coarser and finer fraction of the soil. The exit wire mesh used for core material had an opening size of 1.14 mm for all tests except one using 2.76 mm. The opening size of 1.14 mm is equivalent to approximately D70 of the core material, and the wire mesh of 2.76 mm is equivalent to approximately Dg2 of the grain size distribution. Both opening sizes are located in the coarser section of the grain size distribution curve (see Figure 4.3) and therefore they allow free passage of the finer portion through its openings. There was no clear difference between the behaviour of the core specimen using the 2.76 mm and the 1.14 mm opening wire mesh. Experience (Moffat, 2002) shows that differences in opening size of the exit mesh do not influence the onset of internal stability, however the amount of soil loss during specimen reconstitution could be larger when the wire mesh opening is bigger. 64 3.5 Summary A large permeameter device was designed and commissioned to perform unidirectional seepage tests, under either downward or upward flow, on soils with a potential for internal instability. Commercial software was used to control the apparatus and acquire the data. Attention was given to the quality of the inflow water, which was filtered, and then de-aired in two stages, before use in testing. This was done to ensure saturation of the soil specimen, and removal of all suspended solids greater than 3 pm. The magnitude of axial stress on the top and bottom surface of the specimen is believed of great importance to a systematic interpretation of the test behaviour. More specifically, it enables a more proper account of side wall friction effects, and the corresponding distribution of effective stress within the specimen during testing. Consequently, a "frictionless" seal is used where the loading rod passes through the top plate of the permeameter, and a top and bottom load cell are used to measure directly the corresponding loads on the specimen. Water pressure is measured through two independent arrays of transducers located on opposite sides of the permeameter wall. This allows for a calculation of local hydraulic between port locations prior to failure, and a monitoring of spatial and temporal changes in local hydraulic gradient after failure. 65 Given the objective of assessing the main hydromechanical factors that influence the response of internally unstable soils, the permeameter was designed to ensure a confident assessment of the relation between effective stress and hydraulic gradient at discrete intervals within the test specimen. 66 Sand Carbon filter filter Pressure regulatorr\-Air supply TS TJ Computer S 'r 1" 1! Air supply Figure 3 .1. Components of the large permeameter 67 68 26.4 mm Water supply T T Air pressure T T Water pressure Detail A ^ Loading plate Wire mesh screen Nonwoven Geotext i le Detail C Nonwoven Geotext i le Water supply Wire mesh screen Base plate Figure 3.3. Permeameter cell (downward seepage) 69 70 Figure 3.6. Two-stage filter system 71 TPT1 Top load r - i V1 | cell \ J - L o 1 Loading plate TPT2 I M d TPT3 l « M TPT 4 TPT5 1= TPT 6 mmm= Base_..._^ plate TPT 7 I fcd LVDT ->[ZZ3 DPT 1 , > d Z 3 DPT 2 >\ I DPT3 N I DPT4 > f Z Z ] DPT 5 DPT 6 996 mm Base load cell Figure 3.7. Arrangement of the water pressure transducers and load cell 72 a) b) Figure 3.8. Load cells a) top and b) bottom Figure 3.10. Dissolved oxygen meter, thermocouples, and LVDT. Figure 3.11. Labview software 74 4.0. EXPERIMENTAL INVESTIGATION The test programme involved: (i) a series of preliminary tests that were undertaken to commission the permeameter device, and (ii) a main series of tests intended to establish hydromechanical constraints governing internal instability. All test specimens were subjected to a stage of consolidation, followed by seepage flow at increasing values of average hydraulic gradient that, in most cases, resulted in the onset of internal instability! In total, 3 tests were performed to commission the device, on reconstituted specimens of glass beads. Thereafter, a total of 15 tests were performed on reconstituted specimens of widely graded cohesionless soil. Properties of the materials used in testing are reported, together with a description of method of specimen reconstitution and an outline of the test procedure. 4.1. Material properties 4.1.1. Glass Beads Glass beads were selected for the preliminary tests because they provide for: (1) easy mixing of different size ranges to obtain a target gradation; (2) greater visual insight to fabric of the reconstituted specimen and, (3) a better indication of water quality during testing. 75 The glass beads used were of the Series A and P manufactured by Potters Industries. They have a minimum round fraction of 70% (true spheres), and visual inspection confirms them to be very rounded (see Fig 4.1). The glass beads are made of cerium oxide, yielding a specific gravity in the range of 2.46 to 2.5. Test specimens were prepared by mixing different size ranges of beads to obtain one of three gradation curves (see Figure 4.2), equivalent to sand and gravely sand. Material properties are reported in Table 4.1. The first test was made on Material "F", a uniformly graded "fine sand" equivalent. The second test was a base gradation "F" against a filter gradation "C", the latter being equivalent to a coarse sand and fine gravel. The third test was on a gap-graded material, obtained from a mix proportion of 40% Material "F" and 60% Materials "C". Details of the procedure used in the preliminary tests, and the results obtained, are reported in Appendix A. 4.1.2. Soils Following the successful commissioning of the apparatus, the main programme of testing was performed on four widely-graded cohesionless soils (see Figure 4.3), deemed representative of materials used in construction of zoned earthfill dams. Specifically, two soils (T-0 and T-5) are representative of transition materials placed next to, and downstream from the core; and two soils (C-20 and C-30) are representative of core materials. Transition materials were obtained from a borrow source near the location of the W. A.C Bennett dam in British Columbia. Core material was obtained from a program of sonic drilling on the W.A.C 76 Bennett dam. All soils used in the main program of testing were supplied by the BC Hydroelectric and Power Authority. Upon receipt, soil was processed by washing and sieving to obtain the target gradation used for testing. The remaining processed original materials were placed in dry storage. Visually particles have a sub-angular shape (see Figure 4.4), and a specific gravity of 2.7 was measured for the fine sand fraction of the soil (the same value is assumed for the coarser fraction). Grain size distributions of the transition materials indicate a sandy gravel (T-0), and a sandy gravel with some silt (T-5). Each has the same D 5 0 of 7.4 mm, see Table 4.2, and both exhibit the same distinct gap in the gradation curve at 0.35 <F<0.4, where F is the percent passing by mass. The gap occurs in the 1 mm to 5 mm grain size range, namely that of the coarse sand to fine gravel. Grain size distributions of core materials indicate silty sands with some gravel. One soil has a fines content of 30 % (C-30), and the other a silt content of 20 % (C-20). Each has the same D50 of 0.42 mm, see Table 4.2, and both exhibit a subtle gap in the gradation curve. The gap occurs at 0.2 < F < 0.25, in soil C-20, and at 0.25 < F < 0.32 in soil C-30. It involves the 0.06 mm to 0.15 mm grain size range, namely that of the coarse silt to fine sand. All four soils are deemed internally unstable using the empirical methods both of Kenney and Lau (1986) and of Kezdi (1979), see Figure 4.5 and 4.6 respectively. A summary of the minimum values of H/F and maximum values of D'i5/d'85 is given in Table 4.3. It is important to note that these soils contain finer particles than those originally used in 77 development of the two empirical methods. Further, the two methods evaluate the potential for instability accordingly to geometrical constraints, but do not provide information concerning the hydromechanical conditions necessary to trigger the onset of instability. 4.2. Specimen reconstitution The objective of the reconstitution technique was to replicate a homogeneous and saturated , specimen. This was achieved through preparation of a saturated slurry that was placed manually in the permeameter cell. In construction, grain size distributions that exhibit a distinct gap are generally recognized as susceptible to segregation, and similar concern exists for reconstitution of laboratory test specimens. Therefore, particular care was taken to prevent this occurrence as described below. 4.2.1 Modified slurry preparation technique A slurry deposition technique proposed for reconstitution of triaxial specimens (Kuerbis and Vaid, 1988) was modified for the purposes of this study, wherein the permeameter specimen was reconstituted in a series of layers. Each layer was prepared using a batch of approximately 6 kg of dry soil, prepared to the target grain size distribution (see Figure 4.7). The batch was placed inside a vacuum container, to de-air the soil at a vacuum pressure of approximately 60 kPa for a minimum of 12 hrs. During this time, removal of any entrapped air was facilitated by frequent stirring of the mix. Depending on the length of the test 78 specimen each test typically required between 7 and 12 batches of soil, prepared over a period of 1 to 2 days prior to testing. 4.2.2 Placement technique Before placing the first batch of saturated slurry, the permeameter was prepared to receive it. This required filling the permeameter with de-aired water to yield a thin film of standing water over the surface of the base frame. A wire mesh screen was placed on top of the base frame. The opening size of this base wire mesh varied according to soil type and direction of flow (see Table 4.4). Then, the first layer of slurry was placed directly onto the wire mesh with the aid of a large spoon. Experience showed that, although some of the particles of the first batch of soil pass through the wire mesh, little to no particle loss was observed thereafter. Placement of successive slurry-mixed layers was done under the same thin film of a standing water, to maintain saturation and prevent segregation of the material (see Figure 4.8). Each batch yielded a reconstituted layer thickness of approximately 5 cm. As the advancing water level reached the elevation of a port location on the sidewall of the permeameter, the water line to the respective transducer (TPT or DPT) was saturated and connected. The placement technique of slurry mixing and discrete deposition yielded specimens with void ratios (ec) between 0.28 and 0.44 after completion of consolidation (see Table 4.4). Homogeneity of the specimen was indirectly assessed by visual observation, and by measurements of the pressure transducers at the beginning of each test (as will be shown for 79 each test in Chapter 5). In tests with downward flow, the base wire mesh constitutes the exit wire mesh; in the case of upward flow, the base wire mesh was considerably smaller in opening size that the exit wire mesh on the top of the specimen (see Table 4.4). 4.2.3. Pre-test assembly of the permeameter Upon completing deposition of the last batch of slurry mix, the top surface of specimen was levelled using a spoon and a wire mesh screen placed over it. The loading plate was then placed in contact with that wire mesh screen, and hence the specimen, prior to seating the top plate and sealing the permeameter. Thereafter, the permeameter was filled with water, and the LVDT and top load cell installed. The piston was advanced to apply a small seating force to the load cell and loading rod. At this point a measurement of hydrostatic pressure was checked using both arrays of pressure transducers (TPT and DPT). 4.3. Test procedure 4.3.1 Consolidation Consolidation occurred upon applying a vertical effective stress on the top of the specimen (dt). A condition of double-drainage exists in the permeameter, involving dissipation of excess pore water pressure in both the upward and downward direction. It was important to control the magnitude of these transient hydraulic gradients, in order to avoid a potential for internal migration during the consolidation stage. Therefore, the target value of oY (see Table 80 4.4) was applied slowly, in increments of 5 to 10 kPa. Each new increment was maintained until local hydraulic gradient in the specimen diminished to near-zero values. The required time varied from a few seconds, for transitions materials, to 2 hrs. for core materials. Successive increments of load were applied until the target effective stress at the top of the specimen was reached and consolidation completed. At this point, the specimen length (Lc) was measured and used to calculate the void ratio (ec). 4.3.2 Seepage flow The purpose of the main test programme was to establish the influence of hydromechanical constraints on the onset of internal instability. In total, 9 tests were performed on transition materials and 6 tests on core materials (see Table 4.4 and Figure 4.9). Upon completion of consolidation, unidirectional seepage flow was imposed in a multi-stage test procedure. All tests commenced with an average hydraulic gradient, i a v = 10 across the soil specimen. Thereafter, the hydraulic gradient was increased in increments of 1, until failure of the specimen was observed or the capacity of the apparatus was exceeded. The duration of each stage was typically 90 min. However, in core materials, the duration was increased up to 12 hours in order to have sufficient volumetric flow to deduce the hydraulic conductivity of the specimen. Failure occurred in all but one test: test T-5-25-D did not exhibit instability at i a v ^ 65. All tests were subjected to unidirectional seepage. No tests were subjected to vibration. 81 The initial value of i a v = 1.0 represents the lowest that can be applied, given the configuration of the feedback control system and accuracy of the instrumentation. It required special attention: relative water levels in the inlet and outlet tanks were adjusted to ensure that the differential water head did not initiate a hydraulic gradient larger than 1. Thereafter, for i a v > 1, the electronic regulator was used to apply air pressure to the inlet tank and increase the hydraulic gradient across the specimen. The total duration of a test depend primarily on soil type (T-0, T-5, C-20, C-30) and the magnitude of effective stress imposed on the specimen. Longer tests, and tests on more permeable soils, required a more frequent re-supply of the inlet tank with clean de-aired water. This process of recharging the inlet tank from the storage reservoir, and empting the outlet tank, usually took a couple of hours. On occasion, seepage flow was stopped overnight (unless a long stage was run, in the case of core materials) and during a weekend. In these cases, flow of water was interrupted; care was taken, when next initiating flow, to gradually increase the average hydraulic gradient up to the previous maximum value. No failure was encountered during this process, which is attributed to the slow rate of increase of i a v . The duration of a test, including specimen reconstitution, consolidation, multi-stage seepage, and post-test measurements, was generally about 1 month. The only difference between tests involving downward and upward flow was that, for upward flow, the inlet tank was connected to the bottom valve of the permeameter. All steps in the test procedure were otherwise identical. 82 During the multi-stage of seepage flow, measurements of vertical load on top and bottom boundaries, distribution of water pressure using the TPTs and DPTs, water temperature, and axial displacement, were made at a frequency of 1 H z . The data were stored for later analysis. Video and photographic records were also saved, and notes were taken of any significant event that occurred during the test. Following the onset of any internal instability, the hydraulic gradient across the specimen was immediately reduced to observe the first evidence of failure by seepage-induced erosion. 4.4. Post-test observations When either failure or the maximum capacity of the apparatus was reached, the inlet and outlet valves were closed. Thereafter, the top plate and loading plate were removed, with care taken not to disturb the upper portion o f the specimen. The top wire mesh screen was then removed, and water drained from the specimen through the base plate. Upon exposure, a forensic examination of the specimen was carried out. It was extracted, by spoon, in three sections (top, middle, and bottom) and a sieve analysis made on each of the sections. A photographic record was made of visual observations. The initial location and development of soil migration within the specimen was the primary objective o f the photographic record. After recovery of the specimen, all the components of the apparatus were cleaned and prepared for the next test. 83 4.5 Programme of testing An overview of the test programme is given in Figure 4.9. The three tests on glass beads were integral to the design and commissioning of the test apparatus. As noted earlier, the main test program involved tests on four widely-graded cohesionless soils. The test code defines the type of material, fines content, effective stress on the top of the specimen (25 to 175 kPa), and the direction of seepage flow (D, downward; U, upward). Two test specimens deserve a special mention. Specimen T-5-25-D did not fail during downward seepage flow, leading to a decision to subject it to upward seepage flow following the same procedure as if the specimen had just been reconstituted. The new test (T-5-175-U) subsequently exhibited instability. Post test observation on specimen T-5-50-D indicated a failure of limited extent, evidence of a composition that was largely intact, with a hydraulic conductivity comparable to that of the initial. Therefore it was decided to run a new test using the same specimen (T-5-50-D-R). Transition material T-5 was run either in downward or upward direction to examine the change of hydromechanical conditions that trigger internal erosion. As mentioned, during specimen reconstitution soil losses created a high hydraulic conductivity layer at the bottom of the specimen. This layer always showed low local hydraulic gradients. Therefore, to obtain high hydraulic gradients close to the exit face of the specimen, all core materials were run using an upward-flow direction. 84 The summary of the test programme shown in Table 4.4 indicates that the specimen length varied between 55.6 cm and 32.5 cm, yielding a ratio of length to diameter between 2.0 and 1.2, after initial completion of consolidation. Generally the specimen length was shortened for tests on soils with higher fines content. This enabled the application of a larger hydraulic gradient across the specimen. Reconstitution and consolidation of transition materials yields values of void ratio between 0.28 and 0.38, and corresponding dry unit weights between 20.8 and 19.3 kN/m3. For core materials, the void ratio was between 0.36 and 0.44, with a dry unit weight in the range 19!5 to 18.4 kN/m3 (see Table 4.4). Calculation of these void ratios is predicated on an assumed percentage loss of fine particles during specimen reconstitution. The assumed loss was based on the opening size of the base wire mesh, and the fraction of the grain size distribution smaller than that opening. More specifically, the amount of soil passing during reconstitution was estimated as 10% of the fraction of soil smaller than the opening of the wire mesh screen: this method was independently verified using data from Moffat (2002), from tests on a small permeameter and different wire mesh openings. Therefore, values of soil loss during reconstitution are estimated to be 4.0% for T-0 and T-5 materials, 3.3% for C-30 material, and 2.5% for C-20 material. The difference in void ratio is believed to be due to the grain size distribution of each material (T-0, T-5, C-20, and C-30) and magnitude of a t o ' . All specimens were reconstituted on top of the base wire mesh which varies in size as shown in Table 4.4. As noted earlier this screen represents the exit wire mesh opening in tests with downward flow. With upward flow, the exit wire mesh was the top wire mesh. Two different 85 wire-mesh openings were selected for the exit boundary (see Table 4.4). An opening size of 2.76 mm was used for transition materials, which corresponds to F « 0.4 on the grain size distribution curve. For core materials, an opening size of 1.1 mm was used in all but one test, which corresponds to F » 0.7 or 70% passing. 4.6 Summary A specimen reconstitution technique and multi-stage testing procedure were developed for testing four soils of different gradations (T-0, T-5, C-20, and C-30). The soil used was non-plastic, with sub-angular particles. A total of 18 tests were performed (3 tests on glass beads to commission the apparatus, and 15 tests on the four soils) under different conditions of top effective stress (a™') and seepage (iav). A modified slurry mixing and discrete deposition technique was used in these widely-graded materials to prevent segregation and obtain a uniform specimen for testing. A minor soil loss was observed to occur through the lower wire mesh screen during specimen reconstitution. It yielded a zone of approximately 4 to 7 cm adjacent to the wire mesh screen where the soil was fines deficient. Otherwise, the technique yielded reproducible homogeneous specimens. A routine test procedure was developed to assess the effect of increasing hydraulic gradient (iav) on potentially unstable soils. Initial consolidation of the soil specimen was followed by multi-stage increments of i a v , with either upward or downward seepage flow, until the onset of failure in the test specimen or the maximum capacity of the equipment was reached. 86 Table 4.1. Characteristic properties of the commissioning tests Material Type D 5 0 dso Dis/dss Gap (H/F)m l„ Specimen (mm) (mm) Dis'/d 8 5 ' ratio (kPa) length (cm) F Uniform 0.14 — — — > 1 25 43 C/F Filter/Base 1.95 0.14 7.5 — — 25 43.5 G3 Gap graded 1.40 — 7.5 5.6 0.95 25 46 Table 4.2. Characteristic properties of the soil (main programme) Fines Soil type Material content D 8 5 Dso D 6 0 D10 CU (% <74 um) (mm) (mm) (mm) (mm) T-0 Transition 0 38.5 7.4 11 0.15 73 T-5 Transition 5 38.5 7.4 11 0.13 85 C-20 Core 20 3.4 0.42 0.6 0.021 29 C-30 Core 30 3.0 0.42 0.6 0.018 33 Table 4.3. Geometrical stability analysis of the soils Soil type (D-is/d'ssW (H/F)m l n (1) (H/F)m,„ F<0.2 F<0.3 T-0 13.7 @ 30% 0.95 0.29 T-5 14.3 @ 30% 0.9 0.29 C-20 7.7 @ 5% 0.67 0.67 C-30 10 @ 5.6% 0.67 0.42 (1) Kenney and Lau method (1986) (2) Modified Kenney and Lau 87 Table 4.4. Summary of test programme Length, Top Void Base Test code L c effective ratio Dry unit Exit wire wire weight mesh mesh (cm) stress (kPa) e c» (kN/m3) (mm) (mm) T-0-25-D 55.2 25 0.38 19.3 2.76 2.76 T-0-10OD 55.5 100 0.33 20.0 2.76 2.76 T-0-175-D 55.6 175 0.33 20.0 2.76 2.76 T-5-25-D 44.2 25 0.28 20.8 2.76 2.76 T-5-175-U 44.0 175 N.A. N.A. 2.76 2.76 T-5-50-D 45.0 50 0.29 20.5 2.76 2.76 T-5-50-D-R 42.4 50 N.A. N.A. 2.76 2.76 T-5-25-U 45.5 25 0.30 20.3 2.76 2.76 T-5-30-U 45.0 30 0.29 20.5 2.76 2.76 C-20-50-U 34.0 50 0.37 19.3 1.14 0.15 C-20-85-U 34.5 85 0.39 19.1 1.14 0.15 C-30-25-U 35.5 25 0.44 18.4 2.76 0.15 C-30-50-U 35.4 50 0.44 18.4 1.14 0.15 C-30-80-U 33.5 80 0.36 19.5 1.14 0.15 C-30-100-U 32.5 100 0.39 19.1 1.14 0.15 e c*: void ratio after consolidation 88 GRAVEL SAND Coarse Medium Fine #4 #10 #20 #40 #6 #140 10.0 1.0 Grain size; mm Figure 4 . 2 . Grain size distribution of glass beads specimens 0.1 89 GRAVEL SAND FINES Coarse Medium Fine Silt Clay Grain Size; mm Figure 4.3. Grain size distributions of soil specimens Figure 4.4. Particle shape: gravel fraction 90 70 • T-0 •T-5 —*--C-20 — A - -C-30 •g 50 -ro Q 0 10 20 30 40 50 60 70 80 90 F, mass fraction smaller than; % Figure 4.5. Shape curves for the soils (after Kenney and Lau, 1986) r 18 I— , , , , , 1 . r 1- 0 0 5 10 15 20 25 30 35 40 45 F, mass fraction smaller than; % Figure 4.6. Variation of D'is/d'gs with percent passing (after Kezdi, 1979) 91 Figure 4.7. Mixing of soil particles Figure 4.8. Placement technique 92 Test Programme Commissioning tests Main programme at —^• as F/C c CD 0) c 03 o 73 O C-20 50-U 85-U I. Core materials C-30 25-U \ J 50-U 80-U T-0 25-D 100-D 175-D Transition materials T-5 25-D 50-D 50-D-R 25-U 30-U 175-U 25 100-U Figure 4.9. Overview of the test programme 93 5.0 RESULTS Results obtained using the apparatus described in Chapter 3 are reported, with emphasis placed on understanding hydromechanical influences that govern the behaviour of potentially unstable soils under steady seepage flow. Specifically, the performance of the apparatus is described with reference to three preliminary tests used to commission its operation. Thereafter, results are given for nine tests on Transition materials and six tests on Core materials. These results of the main test programme are described with reference to hydraulic conductivity, pore volume exchange, seepage velocity, effective stress, and hydraulic gradient. 5.1 Preliminary tests to commission the apparatus Commissioning tests were performed on three different configurations of test specimen, all of which were reconstituted from glass beads. The specimens were prepared by mixing of different sizes of spherical beads (Gs = 2.5) to obtain the intended gradation curves (see Figure 4.2). Specimen F comprised a uniform "base gradation": inspection of the initial water head distribution indicates that the reconstitution technique yielded a homogeneous specimen (see Appendix A Figure A.1). Specimen F/C comprised a two layer "base soil: filter" system, where each water head distribution appears uniform suggesting that each component (F and C) are homogeneous (see Appendix A Figure A.2). In contrast, the gap-graded specimen, G3, suffered particle loss through the bottom wire mesh during specimen reconstitution, yielding a homogeneous specimen only at the middle and bottom portion of 94 the specimen (see Appendix A, Figure A.3). Experience indicates this to be a typical feature in reconstitution of gap-graded soils. Particle loss causes an increase in the hydraulic conductivity close to the bottom of the specimen, over a length estimated at 3 to 5 cm above the wire mesh screen. Experience also showed that, notwithstanding the value of D 1 5 ' /dgs' was lower than that for Transition and Core materials tested, the spherical shape of the glass beads yields a very challenging material with which to work. Specimen F exhibited a water head distribution that was linear throughout testing. In contrast, specimen F/C experienced a change in the local hydraulic gradient that is attributed to failure of the "filter/base" interface (see Figure A. 4), associated with a migration of the finer "base" particles into the interstices of the coarser "filter" below, that was confirmed by visual inspection of the specimen. Similarly, with specimen G3, a decrease in local hydraulic gradient O 2 3 ) is associated with particles of the finer fraction of the gradation migrating from the layer between ports 2 and 3 (see Figure A.5). They moved downstream, resulting in higher values of i34 and '145 because of a corresponding increase of hydraulic conductivity in those zones. Approximately 200 sec later complete failure is obtain in the specimen. A piping zone is then observed from top to bottom of the test specimen (see Figure A.6). This type of change in local hydraulic gradient would be used to define failure by internal erosion. A comparison between the local hydraulic gradients calculated from differential pressure transducers (DPT) and the total pressure transducers (TPT) revealed that before failure both sets of transducers gave the same values of ijk. However, after failure, these values were quite different (see Figure A.7). The explanation is that the piping zone was observed on only one side of the permeameter near the TPT array, and far from the corresponding DTP array. 95 When comparing local hydraulic gradients at failure it is possible to observe that the values obtained in tests F/C and G3 are reasonably close (see Table 5.1). Additional details of the three commissioning tests performed are shown in Table 5.1 and in Appendix A. The performance of the tests apparatus was considered adequate to continue with the main series of tests. Instrument measurements showed that it was possible to identify internal erosion events, if the event either within the specimen or at one side of the permeameter wall. All instruments performed as expected and accuracy was in the range anticipated in design. 5.2 Main test programme The main objective of these tests was to establish and quantify the influence of hydromechanical parameters governing the onset of internal instability. Accordingly, it was also important to gain a qualitative understanding of the phenomenon of soil migration itself. Results are reported, in sequence, for the two Transition materials (T-0 and T-5) and two Core materials (C-20 and C-30). The variation of effective stress and local hydraulic gradient within each test specimen was determined as explained in Appendix D. The local hydraulic gradient between two ports locations (ijk) was measured using two sets of independent pressure transducers (TPT and DPT). The results obtained were used to identify trends in the hydromechanical conditions that are important for the analysis of internally unstable soil and in defining the onset of instability. 96 5.2.1. Axial displacement Vertical displacement of the loading plate during testing had several causes. In downward flow, vertical displacement was measured due to seepage-induced consolidation as well as re-arrangement of the soil fabric after internal migration of the finer fraction. Seepage-induced consolidation was deduced from the measurement of axial displacement. The seepage-induced strain was less than 1.5% during downward tests and less than 0.7% in upward tests. In addition, most of the upward tests showed almost no vertical displacement due to seepage-induced consolidation. A summary of seepage induced displacements is given in Table 5.2. After failure, measured axial displacements were considerably higher than the seepage-induced displacements (see Table 5.2). The test was stopped immediately after failure was observed in the specimen. The intent was to observe the initial characteristics of failure, such as the location and extent of disturbance. For this reason, vertical displacement measured immediately after failure was not the maximum displacement that it is possible to obtain due to failure by internal stability. Comparison between tests can be done only to show the general trend in the response of the soil. Specimens of transition material T-0 yielded vertical strain after failure in the range of 2 to 7% (in addition to initial consolidation) of the initial height of the specimen after consolidation. The average vertical displacement in T-0 material was 4.5% (see Figure 5.1). 97 Specimens of transition materials T-5 yielded a vertical displacement after failure in the order of 0 to 8% displacement, with an average vertical displacement of 4.2% (see Figure 5.2). In material T-5, upward flow tests had an average vertical displacement after failure equal to 2.4% and downward tests had an average vertical displacement equal to 6.8%. Therefore, both materials showed similar displacement after failure of the tests. However, when comparing only the downward tests in T-5 with tests using T-0, an increase in the displacement after failure can be observed. In core material C-20 vertical displacement was in the range 1 to 4% of the initial height. Smaller axial displacement was measured in material C-30, in the range between 0.2 to 1% of the total height (see Figure 5.3). Therefore, core material showed smaller axial displacement than transition material tested in upward flow. 5.2.2. Hydraulic conductivity Hydraulic conductivity of the soil in the zone between two ports of measurement was deduced knowing the local hydraulic gradient and the volumetric flow rate. Variation of hydraulic conductivity along the length of the specimen is reported in Figures 5.4 to 5.8, with location of the mid-point between the port locations expressed as a ratio of distance above the base of the specimen (z) to its length at the initiation of seepage flow (Lc). Two curves are presented for each test: they describe the initial values obtained in the first stage of,, seepage at im « 1 , and the final values obtained immediately prior to failure in the last stage of the test. 98 The three tests with material T-0 had an initial uniform hydraulic conductivity of lx l 0"2 cm/sec along all but the lowest portion of the specimen (see Figure 5.4). The lower 3 to 5 cm consistently showed a larger hydraulic conductivity, which is attributed to loss during specimen reconstitution of the finer fraction of the gradation through the bottom wire mesh. Comparison of the data (T-0) suggests a slight but consistent decrease in hydraulic conductivity prior to failure, which is attributed to seepage-induced consolidation during testing. Transition materials T-5 had an initial hydraulic conductivity between 3X10"4 and 6xl0"3 cm/s (see Figures 5.5 and 5.6). Once again, the lowest portion of the specimen yielded a higher local hydraulic conductivity due to losses during specimen reconstitution. The wider range in initial values implies the specimens were less homogeneous than T-0 specimens, which is attributed to the 5% fines content and greater propensity for segregation during reconstitution. Generally, pre-failure values again show a small reduction in hydraulic conductivity as does the final value for test T-5-25-D that did not fail. Upward test T-5-175-U shows a small increase of hydraulic conductivity on the top layer as fines leave the specimen. Specimens of material C-20 yielded a uniform hydraulic conductivity in the range 1 to 4xl0"5 cm/sec. Losses during specimen reconstitution are again confirmed through the higher values of hydraulic conductivity in this zone (see Figure 5.7). No significant difference between the hydraulic conductivity of both specimens suggests that reproducibility of the 99 specimens is good, and there was almost no difference between the initial and pre-failure values of hydraulic conductivity. Core materials C-30 likewise yielded a very uniform hydraulic conductivity in the range 1 to 3x10"5 cm/sec (see Figure 5.8), except test C-30-100-U that yielded slightly smaller hydraulic conductivity of 4xl0"6 cm/sec. The general trend is that of hydraulic conductivity decreasing with a t 0 ' . Again, evidence of soil loss in the bottom portion of the specimen was observed. Reproducibility appears to be excellent between the different specimens considering the difference in a t 0 ' . There was no clear trend in seepage-induced changes in hydraulic conductivity between the initial and pre-failure states. 5.2.3. Pore volume exchange Pore volume exchange represents the amount of water that passed through the voids of the soil specimen during a test. It is calculated with reference to the cumulative volume of seepage flow and the initial average void ratio of the specimen after consolidation (ec), and permits a comparison among the four materials examined in testing (see Figures 5.9 and 5.10). In addition, the total elapsed time of seepage flow during each test is reported separately in Table 5.4. Tests on wash-out transition materials (T-0) yield a common response, which is attributed to the similar values of hydraulic conductivity and the similar duration of each stage of seepage flow during testing. Likewise, the companion transition material (T-5) also yields a common 100 response, but one associated with a considerably lower pore volume exchange as a result of the relatively lower hydraulic conductivity (see Figure 5.9). As expected, core materials showed much smaller pore volume exchange compared with transition materials. There is a little difference in the flow response between the C-20 and C-30 materials (see Figure 5.10), which is consistent with the similar values of hydraulic conductivity. 5.2.4 Transition material (T-0) The transition material T-0 is a sandy gravel with essentially no fines (see Figure 4.3). Three tests were performed, all with downward flow, for a vertical effective stress at the top of the specimen in the range 25 to 175 kPa. All exhibit instability, at an average hydraulic gradient between 8 and 45. A summary of results for each test is presented in the following paragraphs. 5.2.4.1 Test T-0-25-D Seepage flow was imposed in downward direction, commencing at i a v = 2, and sequentially increased to failure (at i a v = 11) over a period of flow of nearly 14 hours (see Table 5.3). Displacement of the loading plate increased gradually with increasing i a v up to 10, whereupon a marked increase occurred leading to the onset of failure (see Figure 5.11). The 101 test was stopped at maximum axial displacement of the loading plate of 3.8 cm. The observed displacement is attributed to two different processes: (1) consolidation induced by the imposed seepage force, and (2) re-arrangement of structure due to particle migration and loss of particles from the specimen. Initially, at 1 < i a v < 4, visual observations indicated a minor loss of soil, typically at the beginning of each stage. It was not associated with any preferential zone, either within the specimen or at the rigid wall of the permeameter. On occasion, for example in this test at i a v = 5, the local hydraulic gradients revealed minor changes with time (see Figure 5.12), which are attributed to a spatial re-adjustment of the finer-fraction within the test specimen. Visual observations confirmed the episodic passage of finer fraction particles from the specimen. In this case, the duration of the event was about 1 minute, and it did not propagate within the specimen. Similar events occurred at i a v = 7 and 10. Thereafter nothing was observed until the onset of failure itself. The distribution of water head in the specimen is almost bi-linear, with a negligible loss over the lowermost 5 cm and a nearly linear variation above this zone (see Figure 5.13). This can be attributed to the homogeneity of the reconstituted specimen, above a lower portion that experienced loss of its finer fraction during placement. The generalized shape of the relation is unchanged with increasing i a v , which implies no significant spatial change in grain size distribution prior to the onset of failure. Initial hydraulic conductivity of the reconstituted specimen was approximately lxlO"2 cm/sec (see Fig 5.14), constant over much of the specimen length. It increased to lxl0 ' 1 cm/sec in the zone immediately adjacent to the lower wire mesh screen. The conductivity showed very 102 little variation with increasing i a v , prior to the onset of failure at i a v =11. The post-failure variation indicates a significant increase of K, to an average value of 2.4x10"' cm/sec, which is attributed to extensive loss of the finer fraction that affected the entire specimen. The variation of local hydraulic conductivity throughout the test (see Figure 5.15), which is believed to be a sensitive index of physical change, suggests the minor episodic loss of soil at iav = 5, 7 and 10 had little influence in the overall integrity of the test specimen. Uniformity of the test specimen is further illustrated by similar values of k through ports 1 to 6. The onset of instability is evident from the increase of hydraulic conductivity that is common throughout the specimen. The variation of local hydraulic gradient in the final stage of testing is shown in Figure B . l . Related data for the onset of instability are reported in Table 5.3. Average hydraulic gradient was increased (to an i a v = 11) over a period of approximately 250 sec, whereupon a stable behaviour was observed to an elapsed time of 750 sec. At this point, a significant local change in the fabric of the specimen yielded a change in hydraulic conductivity of some layers, inducing a companion change in local hydraulic gradient (ijk). More specifically, i45 and is6 show a decrease that is accompanied by an increase in '123 and i34. At the same time visual observations confirmed the onset of a significant particle migration and loss, and independently, a relatively large displacement of the top loading plate. Inspection of the specimen before and after the instability event revealed an extensive zone without fines and with holes (piping zone) (see Figures C.l and C.2). The quantity lost from 103 the specimen is shown in Figure C.3. The affected area includes only a small segment of the permeameter wall (see Figure C.4a), implying this boundary is not a preferential zone for soil migration. Post-test observations of grain size distribution within the test specimen (see Figure 5.16) reveal the nature of the finer fraction loss resulting from failure. In contrast to the original gradation, that at the base of the specimen shows the greatest loss of finer fraction particles. The measured grain size distribution of soil lost from the specimen represents only the finer fraction of the original gradation. Size of the base wire mesh screen does not appear to have controlled the size of soil passing, since its maximum particle (2 mm) is significantly smaller than the 2.76 mm opening size (see Table 4.4). 5.2.4.2 Test T-0-100-D Seepage flow was imposed in the downward direction, commencing at i a v = 1, and sequentially increased to failure (at i a v = 8) over a period of flow of nearly 7 hours and 30 minutes (see Table 5.3). Displacement of the loading plate was negligible for i a v < 2, and increased gradually thereafter to i a v = 7. A marked increased occurred at i a v= 8 leading to the onset of failure (see Figure 5.17). The test was stopped at maximum axial displacement of 10.3 mm (see Table 5.3). Initially, at 1 < i a v < 7, an episodic and minor loss of soil was observed during stages of seepage flow that is attributed to a re-adjustment of finer fraction particles that again did not 104 lead to failure (see Figure 5.18). Distribution of water head was again bi-linear as a result of loss of finer fraction in the lower 3 to 5 cm of the specimen (see Figure 5.19). The hydraulic conductivity of the specimen was l.OxlO"2 cm/sec at the beginning of the test (see Figure 5.20), excepting a zone at the bottom of the specimen. Given a comparison with T-0-25-D (see Figure 5.14), this indicates excellent reproducibility of the reconstituted specimen, and no influence of Ot 0 ' on the hydraulic conductivity. Post-failure measurements did not show a significant change, but exhibit a relatively higher value at the middle of the specimen where soil migration is believed to have started. The variation in local hydraulic conductivity (see Figure 5.21) again confirms no significant change prior to failure. In the final stage of the test (i a v = 8) the soil specimen exhibited a stable response through to an elapsed flow time of approximately 16 minutes (see Figure B.2), when a considerable quantity of soil migration and loss was observed. The migration occurred close to the array of TPTs, therefore hydromechanical conditions at failure are reported with respect to those data (similarly, it was observed at commissioning tests, see Appendix A). The test was continued until a large amount of finer fraction was removed from the specimen. Conditions at the onset of internal stability are summarized in Table 5.3. The overall condition of the soil specimen before and after failure (see Figure C.5) suggests development of an internal connection between these zones where finer fraction is deficient. In Figure C.6, two layers of the specimen after failure revealed holes on their surface at different positions, suggesting migration developed in the whole area. It is expected the holes 105 would be larger if the test was continued for a longer period. Grain size distribution curves (see Figure 5.22) indicate a loss of the finer fraction that is greater at the bottom of the specimen. 5.2.4.3 Test T-0-175-D Seepage flow was imposed in downward direction commencing at i a v = 1, and sequentially increased to failure at i a v = 45 over a period of flow time of nearly 29 hours (see Table 5.3). < Axial displacement of the loading plate was negligible up to i a v = 7, increased linearly for 8 < i a v < 30, and yielded a marked increase at i a v = 31 and 32 that was accompanied with visual observations of soil migration. Following this episodic loss, and subsequent episode at onset of failure at i a v = 45 (see Figure 5.23). The maximum displacement of the loading plate measured during this test was 23.1 mm. Initial seepage flows, at 1 < i a v < 8, produced soil migration at the beginning of each stage. For stages with 8 < i a v < 14, migration was observed both at the beginning of a stage and at other periods during each stage, from the base of the specimen. During these stages, two small voids formed at the wall of the permeameter, on opposite sides. Minor soil migration continued to i a v = 30, whereupon a considerable migration was observed at the bottom-center of the specimen, accompanied by displacement in the loading plate. Changes in the calculated hydraulic gradients were also observed at this stage (see Figure 5.24). However, 106 the local hydraulic gradients returned to original values, and the event is considered one of particle re-arrangement rather than onset of failure, which occurred later (see Figure B.3) Water head distribution curves (see Figure 5.25) indicate a linear variation at the beginning of the test, above the lowermost 4 cm of the specimen, and suggest a subtle change in shape at i a v = 14. The initial hydraulic conductivity was approximately lxlO'2 cm/sec (see Figure 5.26), and it had a tendency to decrease with seepage-induced consolidation of the soil (see Figure 5.27). The large increase at the top of the specimen is attributed to axial displacement of the loading plate that made it difficult to obtain an accurate value of hydraulic conductivity over what became a very short length of soil between ports 1 and 2 at the top of the specimen. The variation of local hydraulic gradient in the final stage is shown in Figure B.3. After a period of approximately 400 sec a change in fabric of the soil induced a decrease in the local hydraulic gradient values. The whole specimen showed a decrease in hydraulic gradient. At the same time visual observation confirmed particle migration and displacement of the loading cell. This episode was considered as a representative failure and the values obtained at this stage characterize the hydro-mechanical values at the onset of internal instability. In Table 5.3 key values at the onset of internal stability are shown. Inspection of the soil after failure (see Figure C.7) revealed zones with depleted finer-fraction that suggest preferential flows within the specimen. Layers of the specimen after failure (see Figure C.8), suggest zones with low finer-fraction at the top layer and depleted finer fraction 107 at the bottom layer. Grain size distributions of selected layers after testing (see Figure 5.28) suggest that the soil lost is that of the finer fraction of the original gradation. 5.2.5 Transition material (T-5) The transition material T-5 is a sandy gravel with 5% silt content (see Figure 4.3). Six tests were performed three with downward flow and three with upward flow, on 4 soil specimens (see Table 5.3); tests T-5-175-U and T-5-50-D-R were performed on the same specimens as T-5-25-D and T-5-50-D respectively. The range of vertical effective stress applied to the top of the specimens was between 25 and 175 kPa. All tests, except T-5-25-D, exhibit instability at an average hydraulic gradient between 6 and 36. A summary of results for each test is presented in the following paragraphs. 5.2.5.1 Test T-5-25-D Downward seepage flow was, at i a v = 1 and sequentially increased to failure at i a v = 62 over a cumulative flow time of nearly 67 hours (see Table 5.3). Negligible displacement was measured up to i a v = 5, whereupon a linear increase commence through to the end of the test (see Figure 5.29), which was stopped at an axial displacement of 5.4 mm due to suffosion of the specimen (see Table 5.3). A minor loss of fines (< #200) was observed, at the center of the specimen base, with each increment of seepage flow. In addition, the finer fraction was observed to pass in very small 108 quantities at the beginning of each stage. A couple of small voids formed near the base of the specimen, likely in zones filled only with fine grains during specimen reconstitution. Some soil migration was detected at the walls of the permeameter close to these voids. Thereafter the specimen was stable all through the end of the test. The distribution of water head reveals the development of an extensive zone, in the lower-third of the specimen, which experienced loss of finer fraction (see Figure 5.30). Examination of the curves suggests it likely developed at i a v greater than 9, given the greater linearity to the relation at lower gradients. Initial hydraulic conductivity of the T-5 material was one order of magnitude smaller than the T-0 material (see Figure 5.31). The initial values ( i a v = 1) suggest reconstitution gave a uniform specimen, and that seepage flow induced considerable change. An increase occurs adjacent to the lower wire mesh, attributed to particle losses, and a decrease occurs in the remainder of the specimen that was found common to this material. Variation of local hydraulic conductivity through the test is consistent with the absence of soil migration (see Figure 5.32), in this test that did not exhibit internal instability. The test was stopped at i a v = 65, upon reaching the maximum capacity of the permeameter. Local hydraulic gradients in this final stage are shown in Figure B.4. 5.2.5.2 Test T-5-175-U Upward seepage flow was imposed at i a v = 1, and increased sequentially to failure at i a v = 32 over a period of nearly 32 hours (see Table 5.3). Negligible displacement of the loading plate was measured up to i a v = 4; a minor increment at i a v = 5 was again followed by negligible 109 movement prior to failure (see Figure 5.33). The test was stopped at maximum axial displacement of 23 mm (see Table 5.2). Evidence of particle movement was detected during stages with 1 < i a v < 30. This included "boiling" in small voids on the wall of the permeameter, and a companion minor boiling at the openings of the top loading plate. Suffusion of fines (< #200) was also observed during these stages. The distribution of water head in the test specimen shows that the first 5 cm of the specimen experienced loss of its finer fraction during specimen reconstitution (see Figure 5.34). Water head distribution was nearly linear above this zone for earlier stages. The shape of the curves is maintained as the average hydraulic gradient is increased. This suggests that particles loss did not have a preferential zone and it was experienced in the whole specimen. Initial hydraulic conductivity of the test specimen was approximately lxlO"3 cm/sec (see Figure 5.35). It diminished with seepage flow at all locations except the top, where it increased for i a v ^ 6, and remained unchanged thereafter. Likewise, it remained unchanged at the base. A post-failure analysis (at the same oY) indicates a significant increase, to an average value of 2xl0"2, cm/sec which is attributed to an extensive loss of finer fraction that affected the entire specimen. Variation of local hydraulic conductivity shows a general decrease, with exception of the top layer that showed an increase at i a v > 5 (see Figure 5.36). The decrease is attributed to a seepage-induced consolidation that changed the void ratio of 110 the specimen. The increase in the top layer is associated with some loss of fines on the top layer. Variation of local hydraulic conductivity at the onset of internal stability is shown in Figure B.5: the negative value of ijk signifies upward flow. The onset is characterised by a sudden decrease in i34 and companion increase of gradient in the adjacent layers, attributed to loss of the finer fraction in layer 34 and compensation in the other layers as a results of the head control system. It occurred during the increment of seepage from i a v = 31 to i a v = 32, and led rapidly to a complete failure of the entire specimen. Data for the onset of internal instability are reported in Table 5.3. Comparison of the specimen before and after failure indicates that a considerable amount of the finer-fraction of the gradation washed out (see Figure C.9). Visual observations of layers of the specimen after failure show that the center of the top layers had voids without finer-fraction material (see Figure CIO). At a depth of 27 cm. the specimen the finer-fraction content was homogeneous and smaller than the original gradation in the complete area of the permeameter (see Figure C. 1 Ob). Post-test observations of grain size distribution (see Figure 5.37) indicates the finer loss was about 8 to 12 % of the total mass of the specimen or approximately 25 % of the original finer fraction content. Grain size distribution of the soil passing was not plotted as there was not recovered enough material. I l l 5.2.5.3 Test T-5-50-D Downward seepage flow was imposed at i a v = 1, and increased sequentially to failure at i a v = l over a cumulative flow period of nearly 10 hours (see Table 5.3) There was a negligible displacement of the loading plate prior to i a v = 10 (see Figure 5.38), attributed to a seepage-induced consolidation. The specimen was stable at 1 < i a v < 10, with only minor fines migration (< #200) and loss of the finer fraction at the base of the specimen. Of note was the fact that migration did not occur at the walls of the permeameter and that, at the end of these stages, the appearance of the specimen was consistent with the initial state. The distribution of water head again shows a linear curve with changes in the slope at bottom and top of the specimen (see Figure 5,39). This is attributed to a uniform specimen that experience loss of finer fraction during specimen reconstitution. No significant change in shape of these curves is noticed until the onset of instability when a change in fabric of the middle zone is noticed. The initial value of hydraulic conductivity was approximately lxl0" 3 cm/sec (see Figure 5.40), the specimen experienced only a relatively minor decrease to approximately 4xl0"3 cm/sec before failure (at i a v =11), after which the hydraulic conductivity was approximately lxl0" 2 cm/sec. As the test was stopped after internal erosion most of its particles stayed in the soil. This was demonstrated when the specimen was tested under seepage again and the average hydraulic gradient was very close to its original value before failure. 112 The trend in local hydraulic conductivity of each layer (see Figure 5.41) is one of little change with increasing flow, prior to the episode of instability that yields an increase that is common to all locations, followed by a return to values similar to those preceding the event. The response is attributed to re-arrangement of the skeleton of the soil, after vertical displacement of the loading plate, yielding a very uniform specimen in the upper 15 to 40 cm portion (see Figure 5.39). The instability was triggered by increasing to 11, and resulted in a migration of fines and finer fraction from the base of the specimen. It initiated at an elapsed time of 150 sec. in the test stage, with a significant decrease in i45 (see Figure B.6). The entire specimen responded, and then exhibited a recovery or return to stability within 100 sec. that was maintained over the remaining nearly 10 min. of the test. Visual observations (see Figure C . l l ) suggest that the central portion of the specimen was essentially "intact", with no evidence of voids or significant loss of finer-fraction. 5.2.5.4 Test T-5-50-D(R) Given the apparent localised and episodic instability of test T-5-50-D, the "intact" central portion, and the values of hydraulic conductivity that were similar to pre-failure values, it was decided to perform an additional test on the same specimen. Seepage flow was imposed in a downward direction, commencing at i a v = 1, and sequentially increased to failure at i a v = 6 over a commutative flow period of nearly 15 hours (see Table 5.3). 113 Negligible axial displacement was observed at i a v ^ 11 (see Figure 5.42). Thereafter a modest linear increase was observed, prior to failure. The maximum displacement measured in this test was 6.8 mm. The specimen exhibited a generally stable behaviour for i a v = 1 and 35, with only minor and episodic fines migration observed at the base and at the wall in a zone with a small void. Finer fraction migration was observed for short periods in the latter stages of the test. These episodes were followed by stable periods, and did not progress beyond localized internal stability. Initial water head distribution suggests that finer fraction moved retrogressively from the bottom to the top of the specimen on the previous test (T-5-50-D), leaving the bottom portion of the test with the larger amount of finer fraction. The variation of water head distribution reveals a progressive change in shape from concave upwards (at i a v < 11) to concave downward (at i a v > 21) indicative of a developing of particle migration prior to the onset of internal instability (see Figure 5.43). Initial overall hydraulic conductivity was approximately 2xl0"3 cm/sec, which is similar to the other T-5 specimen (see Figure 5.44). No significant variation of local hydraulic gradient was observed during the test (see Figure 5.45), with the exception of the upper most portion (K23) of the specimen. The onset of internal stability yields a common change that was similar at all locations, which implies no preferential movement of fines. Variation of local hydraulic gradient, at the onset of instability, is shown in Figure B.7. A transient re-arrangement of particles at approximately 75 sec of elapsed time in this final 114 stage preceded a significant instability that initiated with a marked decrease in is6 at about 270 sec. Visual observations confirmed the existence of voids that would refill with finer-fraction and create new voids. Complete failure was observed after approximately 5 minutes, characterize by a "pipe" that formed from top to bottom of the specimen (see Figures C.12 and C.13). Grain size distribution curves (see Figure 5.46) suggest that the passing finer fraction to be at least 13 % of the original T-5-50-D specimen. Post observations of grain size distribution of the test specimen after testing are shown in Figure 5.46. In contrast of the original size, size distribution of different layers of the specimen after failure showed approximately 13 % loss in the finer fraction component. The loss affected the complete specimen in a very similar manner. 5.2.5.5 Test T-5-25-U Upward seepage flow was imposed at i a v = 1, and sequentially increased to failure at i a v = 6, over a flow period of nearly 6 hours (see Table 5.3). Negligible axial displacement occurred up to i a v = 5 (see Figure 5.47). Instability was associated with a movement of 4.2 mm, attributed to seepage forces against the stress-controlled top loading plate. Some fines migration (< #200) was observed from the top of the soil specimen in the early stages of seepage at 1 < i a v < 4. Minor localized boiling was observed at i a v = 5 on the lower portion of the specimen. The test was continued until the onset of a more generalized internal instability. 115 Although the distribution of water head at i a v = 1 indicates a relatively uniform specimen at the beginning of the test (see Figure 5.48), increasing flow suggest the lower portion of the specimen to be relatively more permeable (see also Figure 5.49). The Initial hydraulic conductivity was approximately 1 x 10"3 cm/sec, and decreased moderately with increasing i a v as a result of seepage forces. A post failure measurement indicates the top layers of the specimen were most affected by soil migration associated with internal erosion. The spatial variation in hydraulic conductivity, but little temporal variation, is well-illustrated in Figure 5.50. After failure hydraulic conductivity of the top layer, k 2 3, showed an increase in its value due to the loss of finer fraction by internal erosion. The variation of local hydraulic gradient during the onset of internal stability at i a v = 6 (see Figure B.8) shows that, after approximately 1000 sec. of stable flow, the top layer of soil experienced a sudden loss of water head (see i 2 3) attributed to a release of finer fraction particles. Visual observations confirm "boiling" at different locations, on the top surface. The behaviour stopped for a few minutes, was subsequently followed by a catastrophic failure associated with release of most of the uppermost finer fraction at 1800 sec. of elapsed time in this final stage. 5.2.5.5 Test T-5-30-U Since the T-5-25-U specimen remained intact in its centre and bottom portions and the hydraulic conductivity of these layers was practically unchanged (see Figure 5.50), it was 116 decided to run another test using the same specimen. No changes were made to the specimen but a portion of the soil lost returned filling the voids left on test T-5-25-U. Upward seepage flow was imposed at i a v = 1, and sequentially increase to failure at i a v = 13, over a cumulative flow period of nearly 13 hours (see Table 5.3). Negligible displacement of the loading plate occurred for i a v < 12 (see Figure 5.51). For 1 < i a v < 8, the specimen was stable and exhibited only minor fines migration (< #200). Between i a v = 9 and 11, some boiling was detected at the permeameter wall in the lower portion of the specimen. Small openings or cracks were observed in the same area at i a v = 12. Minor boiling of particles was observed in some of these openings. The test was continued until the onset of internal stability, at i a v =13. The initial distribution of water head showed that the top and bottom portions of the specimen lost were finer fraction deficient (see Figure 5.52), due to failure of the uppermost layer in the previous test and specimen reconstitution loss respectively. The shape of the water head distribution remains unchanged until the onset of internal stability, indicating no changes of the fabric of the specimen. The after failure distribution suggest loss of finer fraction particles between ports 3 and 5. The initial overall hydraulic conductivity of 4 x 10"4 cm/sec decreases stage by stage until the onset of instability ( i a v = 13) when there was a considerable increase (see Figure 5.53). Before the onset of internal instability there was a layer with small hydraulic conductivity close to port 4. After failure, the overall hydraulic conductivity was approximately 117 5 x 10"1 cm/sec, and a new zone of relatively low hydraulic conductivity was close to port 5. The change in hydraulic conductivity suggests that considerable amount o f finer fraction was loss during failure. This is confirmed by local hydraulic conductivity plot (see Figure 5.54). Local hydraulic gradients at the onset of instability (see Figure B.9) reveal a sudden decrease in Us that occurred in advance o f a decrease in i34. A t the same time, visual observations confirmed significant internal erosion at the bottom and centre of the specimen. At an elapsed time o f 150 to 200 sec, a common loss of local hydraulic gradient commenced at all locations that suggests the development of a conduit across the complete specimen. Observations o f the specimen before and after failure (see Fig. C.14) indicate a preferential flow zone with practically all the finer fraction eroded. Inspection o f internal horizons (see Figure C.15a) revealed a large zone with a relatively low finer fraction (right side), however the configuration o f testing caused the finer fraction to refill the specimen after flow of water was shut down at the end of the test. Accordingly, grain size analysis after failure, which showed a loss o f about 12 % in the finer fraction (see Figure 5.55) is interpreted as a generalised indication o f conditions resulting from seepage flow Again it was not possible to plot the grain size distribution of the soil lost. 5.2.6 Core material (C-20) 118 The core material C-20 is a very silty sand with some gravel (see Figure 4.3). Two tests were performed with upward flow, at a top vertical effective stress of 50 and 85 KPa. The average hydraulic gradient at failure was found to be 27 and 37 respectively. A summary of the tests results for each test is presented in the following paragraphs. 5.2.6.1 Test C-20-50-U Upward seepage flow was imposed at i a v = 1 and increased sequentially to failure at i a v = 27 over a period of nearly 112 hours (see Table 5.3). Axial displacement of the loading plate (see Figure 5.56) was negligible up to i a v = 25. The onset of internal instability yielded a maximum downward displacement of 3.53 mm before the test was stopped. Visual observations showed the specimen to be stable for 1 < i a v < 26, with only minor fines migration observed at different locations. No movement was observed at the permeameter wall. The water head distribution confirmed the specimen was homogeneous except the bottom portion where part of the finer fraction was lost during specimen reconstitution (see Figure 5.57). After failure, water head distribution is almost vertical between ports 5 and 6. This is attributed to suffosion of the finer fraction at the onset of instability under upward flow. Initial overall hydraulic conductivity of this specimen was approximately 2 x 10'5 cm/sec (see Figure 5.58). Hydraulic conductivity did not vary significantly for i a v > 5 (see Figure 119 5.59). The onset of instability produced an increase in local hydraulic conductivity that was largest for k56, suggesting a significant particle migration in that layer. The variation of hydraulic gradient leading to the onset of instability (see Figure B.10) confirms is6 exhibit a sudden decrease after 80 sec of flow in this final stage. This event was confirmed by visual observation of upward soil erosion at the base of the specimen that progressed with time and to reach the top of the specimen (see Figure B.10). The companion increase in hydraulic gradient at the top between 200 and 500 sec is believed a consequence of the finer fraction moving within the specimen. Finally, the top layer released its finer fraction, yielding a substantive and generalised (rather than localised) instability failure of the specimen (see Figure C.16). Importantly, no preferential erosion was observed at the walls of the permeameter (see Figure C.17). Grain size distribution curves (see Figure 5.60) reported at different depths do not reflect the damage of specific zones where suffosion occurred. However, the grain size distribution of the soil passing through the exit wire mesh corresponds to the finer fraction of the soil and is smaller than the opening of the exit wire mesh (1.14 mm) for this test. 5.2.6.2 Test C-20-85-U Upward seepage flow was imposed at i a v = 1 and increased sequentially to failure at i a v = 37 over a period of nearly 127 hours (see Table 5.3). Again, there was negligible axial 120 displacement prior the onset of instability (see Figure 5.61), which yielded a maximum displacement of 13.3 mm. No movement or loss of the finer fraction was observed during seepage with 1 < i a v < 34. Indeed, only a minor fines migration occurred and its rate of loss decreased with time. A short horizontal crack appeared in the specimen approximately 10 mm above the wire mesh, and within the zone between ports 6 and 7, at i a v = 35. It increased in length during the stage with i a v = 36 but it did not propagate further prior the onset of instability in zone 5-6 at i a v = 37. The water head distribution shows the lower 5 cm of the specimen experienced loss of its finer fraction material during specimen reconstitution (see Figure 5.62). Water head distribution was bilinear above this zone. The shape of the distribution remained largely unchanged with increasing i a v , which implies no significant change in the soil specimen prior to failure. A post-failure curve suggests the finer fraction was lost in the portion of the specimen at 5 to 15 cm above the wire mesh, between ports 5 and 6. Initial hydraulic conductivity is approximately 4xl0' 5 cm/sec (see Figure 5.63) and uniform in most of the specimen except the lowermost zone where loss of the finer fraction yields a higher value of k56«5xl0' 4 . It remained constant prior to the onset of failure at i a v = 37: post-failure values indicate a significant increase (see Figure 5.63), attributed to an extensive loss of the finer-fraction that affected the entire specimen. The value of i a v decreases with failure 121 because the control system is unable to maintain the target gradient of 37 as the hydraulic conductivity increases with loss of finer fraction. With the onset of instability, at i a v = 37 (see Table 5.3), local hydraulic gradients i34 and is6 showed a sudden decrease after about 300 sec of elapsed flow; it was later followed by a decrease of 145 after approximately 600 sec (see Figure J3.ll). Visual observations confirmed the instability in a similar manner as other tests on C-20 material. Then a short horizontal crack in zone 5-6 was observed about 8 cm from the top of the specimen, followed shortly thereafter by a second crack in zone 6-7 about 1 or 2 cm from the bottom wire mesh screen. A sudden increase on the opening size of the bottom cracks was followed by horizontal flow and fines filling the cracks again. The failure progressed with vertical soil migration and apparent "pulses" of water pressure (detected by instruments) leading and abrupt instability in the top layer and a release of finer fraction particles at a very high rate. Comparison of the specimen before and after failure (see Figure C.18) shows that finer fraction left a large area along the specimen length. Clear piping can be visually observed on the specimen as it is shown in a bottom layer (see Figure C. 19) Grain size distributions at different depths did not reveal the magnitude of the failure, because the unstable zone was localized (piping) through the complete specimen (see Figures 5.65 and C.19). However, grain size analysis of the zone of piping showed the soil lost approximately 18 % of its original particles (see piping zone curve on Figure 5.65). The size distribution of soil that was lost from the specimen and recovered was a silt/sand with a 122 maximum particle size of 1 mm that was smaller than the opening size of the exit wire mesh screen. 5.2.7 Core material (C-30) The core material C-30 is a very silty sand with some gravel and had a fines content of 30 % (see Figure 4.3). Four tests were performed with upward flow, at a top vertical effective stress in the range between 25 to 100 kPa. The average hydraulic gradient at failure was found to be in the range 13 to 29 respectively. A summary of the tests results for each test is presented in the following paragraphs. 5.2.7.1 C-30-25-U Upward seepage flow was imposed at i a v = 1, and increased sequentially to failure at i a v = 13, over a very short period of only 3 minutes (see Table 5.3). This test was an exception to the method of multi-stage testing, in which hydraulic gradient was continually increased to failure in a single stage, with the intent of examining rate effects. Practically no axial displacement was measured up to i a v = 8 (see Figure 5.66). The onset of failure resulted in a 13 mm upward displacement that was followed by a downward movement of 6.6 mm as the seepage force diminished with decrease in i a v . After the onset of internal stability under upward flow a maximum axial displacement of 6.6 mm was 123 measured. This reveal a loss of finer fraction particles and a rearrangement of the soil matrix yielding a decrease of the void ration of the soil matrix. Water head distribution indicates a uniform specimen, with the same exception for all tests of a finer fraction deficiency in the reconstituted soil against the base wire mesh screen (see Figure 5.67). The shape of the water head distribution changed slightly, when i a v increased, in the zone between ports 4 and 5. This change indicates minor migration of the finer fraction in the uppermost portion of the specimen. The post-failure measurement indicates the loss of finer fraction was through the whole specimen. The initial hydraulic conductivity of the reconstituted specimen was approximately 2xl0' 5 cm/sec (see Figure 5.68), and increased by 2 orders of magnitude after failure. The very fast rate of increment in flow, however, means values of hydraulic conductivity obtained during this test were not as accurate as those in the companion tests. The negligible variation of local hydraulic conductivity through the test suggests no significant loss of fines or finer fraction before failure (see Fig 5.69). Failure began at the middle of the specimen and propagated rapidly upwards to the top causing fast release of finer fraction. This was confirmed by measurements of water pressure that yield a sudden decrease in the calculated value of 145 after approximately 115 sec from the beginning of the test (see Figure B. 12). The subsequent instability of the top layer is evident in a rapid decrease of \M, at which point the flow of water increased substantially. 124 1 A large amount of the finer fraction was removed in the zones of piping (see Figures C.20 and C.21). These zones had a relatively small cross-sectional area, and it is believed they would have been more extensive if the flow had been maintained longer. The grain size distribution of the bottom, middle and top layers is very similar to the original gradation, which is attributed to the small area of the piping zone (see Figure 5.70). However, the soil passing through the wire mesh filter from was a finer-rich gradation of the finer fraction. 5.2.7.2 C-30-50-U Upward seepage flow was imposed at i a v =1.4 and increased sequentially to failure at i a v = 14 over a period of nearly 139 hours to i a v = 14 (see Table 5.3). N o displacement was measured to i a v = 13 (see Figure 5.71). Then, after the event considered as failure at i a v = 14, an axial displacement of approximately 1 mm was measured. The test was continued thereafter to a total displacement of 4.2 mm (at i a v =19). Water head distribution during this test again revealed an initial uniform specimen, except for the lowermost material adjacent to the base mesh screen (see Figure 5.72). Thereafter, the shape of the curves is constant suggesting no significant change in the soil specimen prior to the onset of internal instability, and a soil migration that starts at the base and progress upwards. The initial overall hydraulic conductivity was approximately 1 x 10' 5 cm/sec (see Figure 5.73). Upon failure, it increased by approximately 1 order of magnitude in the lower half of 125 the specimen, with no change in the upper half. Thereafter the test was continued and the average hydraulic gradient was increased, causing a further increase in hydraulic conductivity over a progressive larger portion of the specimen. The variation of local hydraulic conductivity again shows the uniformity of the test specimen, and suggests the first layers to lose finer fraction were 5-6 and 6-7. Subsequently layers 4-5 and then 3-4 exhibit a similar increase in hydraulic conductivity (see Figure 5.74). During the onset of internal instability, at i a v = 14, the local gradient is6showed a marked decrease after approximately 2000 sec (see Figure B.13). Internal erosion was not observed at the walls of the permeameter, suggesting the change occurred inside the specimen. With continuation of the test, and increased of i a v the development of small cracks were observed together with boiling at the top of the specimen and minor finer fraction migration. Some cracks would open (empty) and close (fill) after a few minutes, as a result of particle movement. Diagonal cracks were observed, the episodic movement of finer fraction become more energetic, and finally led to "eruption boiling" events at the top of the specimen until the test was stopped. Instability yielded a piping zone that extended from bottom to top of the specimen (see Figure C.22). Visual observation of different horizons of the specimen confirms the existence of a piping zone (see Figure C.23), with a zone without finer fraction at the front of the specimen. Similarly to test C-30-25-U, grain size distributions of bottom, middle and top layers are very similar to the original gradation (see Figure 5.75). This is attributed to the small area of the 126 piping. However, most of the soil passing had particles smaller than 1 mm suggesting that most of the finer fraction of the piping zone was lost during instability. 5.2.7.3 C - 3 0 - 8 0 - U Upward seepage flow was imposed at i a v = 1, and increased sequentially to failure at i a v = 29 over a period of nearly 193 hours (see Table 5.3). Negligible axial displacement was measured up to i a v = 28, and the onset of internal instability (see Figure 5.76). The test was stopped at maximum axial displacement of 8.2 mm (see Table 5.2). Only minor fines migration was observed at the top of the specimen with 1 < i a v < 28, and there was no indication of impending failure prior to its onset. The distribution of water head indicates a uniform specimen that was unchanged by seepage flow to i a v = 21 (see Figure 5.77). Local hydraulic conductivity also confirmed the uniformity of the soil and shows a significant change of the specimen fabric after failure (see Figure 5.78). The subtle change at 24 < i a v < 29 is attributed to a minor finer fraction movement within the 5-6 zone of the specimen and slight increase from the initial overall hydraulic conductivity of 1 x 10"5 cm/sec (see Figure 5.79). The final stage at i a v = 29 witnessed a relatively long period of stability, of approximately 6500 sec, prior to the onset of a failure in which the local hydraulic gradients i45 and is6 showed a significant decrease in its value, followed thereafter by i34 (see Figure B .M). Visual observations confirmed the particle migration. A horizontal crack was visually observed 127 between ports 4 and 5 at i a v = 29. After a period of approximately 6500 sec of this final stage there was a release of finer fraction at top of the specimen. Outbursts of finer fraction particles were observed at the centre of the specimen and no soil1 migration at the walls of the permeameter. Subsequently, soil boiling and a zone with no finer fraction were observed between ports 5 and 7, thereafter small pipes develop at two zones, increasing water flow and transporting finer material upward. Catastrophic failure was reached soon after, with soil migration at different locations of the specimen. Although instability affected a considerable area (see Figure C.24), once the flow of water was stopped some of the finer fraction dropped down to voids in the specimen. Nonetheless, Figure C.25 shows a top layer (Z/L c = 1.0) and mid-point layer (Z/L c = 0.66), after failure, which reveal a location with no finer fraction (right side) and one with a significant deficiency. Like the specimens of C-30 core material, there is no significant difference in grain size distribution of the post-test and pre-test specimen (see Figure 5.80) yet the soil passing is again very fine rich, suggesting a very localized washing-out of the finer fraction in a zone of piping. 5.2.7.4 C-30-100-U This test was run with a slightly different procedure, where upward seepage flow was imposed at i a v = 1 and increased sequentially to failure at i a v = 28 over a period of nearly 198 hours (see Table 5.3). Then a t ' was reduced in stages from 100 kPa to 65 kPa. The objective was to examine the response of the specimen and compare it with previous tests. Negligible 128 displacement was measured up to i a v = 27 (see Figure 5.81). A maximum upward displacement of 12 mm was measured at the onset of internal stability followed by a downward maximum displacement of 18.4 mm, attributed to loss of finer fraction particles and a rearrangement of the soil matrix. No visual evidence of finer fraction migration accompanied by a negligible displacement was observed at 1 < i a v < 28. Unloading of o t ' is continued until the onset of internal instability was reached. Initial water head distribution reveals a uniform soil specimen, except for the lowermost material adjacent to the base mesh screen that lost finer fraction during specimen reconstitution (see Figure 5.82). No change in the water head distribution was observed until the onset of internal instability, attributed a constant soil fabric before failure. A significant loss of finer fraction between ports 5 and 7 is revealed by the water head distribution after instability. Uniformity of the specimen is again confirmed by the hydraulic conductivity along the specimen (se Figure 5.83), where the initial value was approximately 6xl0"6 cm/sec, that remained practically constant through the test until the onset of instability was reached. The final hydraulic conductivity was approximately 2xl0"3 cm/sec making evident the release of finer fraction of the specimen at the onset of instability. Similarly, local hydraulic conductivity shows practically no variation during the test, with an increase in all layers after instability (see Figure 5.84). 129 During the last stage after approximately 150 sec 156 showed a sudden decrease, followed by a decrease of 'i6i and 134 and an increase of i45 attributed to the gain of finer fraction or to maintain i a v in a head control system (see Figure B.l5). Visual observations confirmed this failure in addition to noise attributed particle movement and displacement of the loading plate, followed by soil migration observed on the permeameter wall. Internal instability was visually observed from the lowermost portion of the specimen to approximately 6 cm below the surface of the specimen, however, the uppermost portion of the specimen did not experience soil migration. It can be observed that the uppermost portion remained almost intact and evidence of horizontal-vertical particle flow paths in the rest of the specimen (see Figure C.26). Visually particle migration was observed going upward, horizontal and even downward direction. Once again as in other core specimens there is no evidence of change in the grain size distribution before and after the onset of instability (see Figure 5.85), attributed to the localized failure at the middle and lower most portions of the soil specimen that did not propagate to the upper portion of the specimen. 5.3 Summary A total of 15 tests are reported for the main test programme. Three downward flow tests were performed on the T-0 material, three downward and three upward flow tests were performed on the T-5 material, two upward flow tests were performed on the C-20 material, and four upward tests were carried out on the C-30 material. 130 The general nature of internal erosion is described from visual observations and interpretation of instrument measurements. The behaviour under unidirectional seepage of core and transition materials is compared and contrasted in order to distinguish the phenomena of suffusion and suffosion. Suffusion took the form of episodic events of finer migration and a time-dependent process of fines migration (<75 um). Suffosion was associated with the onset of internal instability, and was the main focus of the investigation. Piping often occurred as a progression of internal erosion from a localised zone of instability to a preferential zone or conduit along the specimen length. No influence of the rigid wall was observed in any of the soil gradations tested. Homogeneity of each specimen is confirmed from visual observations and the deduced profiles of hydraulic conductivity along the specimen. These profiles confirm that the specimens are uniform, but with a zone of higher hydraulic conductivity at the lower 4 to 7 cm as a result of finer fraction losses during specimen reconstitution. Material T-5 was the first gradation tested with fines: experience showed it was less homogeneous than the T-0 specimens, based on the deduced hydraulic conductivity profile. The specimen reconstitution technique was therefore improved through discrete deposition of the slurry mix in a thinner film of standing water (5 mm instead of 10 mm). The onset of internal instability is defined from visual observations and the variation of the deduced values of local hydraulic gradient (ijk). At failure, loss of the finer fraction particles induces a companion increase in local void ratio, yielding an increase in local hydraulic 131 conductivity and decrease of ijk. This sudden decrease of ijk was therefore used to define the onset of internal instability. Other parameters may also be used to characterise failure by internal erosion, such as axial displacement of the top loading plate (8) and increase of volumetric flow rate (Q). A post-test analysis of grain size distribution gives an indication of the amount of finer fraction lost in each specimen. It was used to complement visual observations of the area affected by internal erosion, layer by layer, along the specimen length. 132 Table 5.1. Summary of preliminary tests Test F F/C G3 Specimen length 43 43.5 46 Max. i a v 15 13.5 8.8 Max. local i 15.2 16.7 12.6 ijk at failure — 16.7 10.6 Max. displacement 2.5 cm 2.8 cm Table 5.2. Vertical displacement Test code Total axial strain (%) Seepage induced strain (%) T-0-25-D 7.00 0.851 T-0-100-D 2.16 1.441 T-0-175-D 3.96 1.457 T-5-25-D 1.22 1.22 T-5-175-U 5.23 0.705 T-5-50-D 5.78 0.400 T-5-50-D(R) 7.78 0.590 T-5-25-U 0.00 0.000 T-5-30-U 2.00 0.022 C-20-50-U 1.00 0.103 C-20-85-U 4.00 0.009 C-30-25-U 1.00 0.000 C-30-50-U 0.20 0.198 C-30-80-U 0.20 0.060 C-30-100-U 0.57 0.040 133 Table 5.3. Summary of laboratory testing Test code lav Failure icr CT vm Axial disp. Duration of zone (kPa) (5/Lc) seepage flow T-0-25-D 11 56 7.0 38.1 0.070 13 h 47 m T-0-100-D 8 45 12.2 62.1 0.022 7 h 2 6 m T-0-175-D 45 56 25.4 182.6 0.042 28 h 59 m T-5-25-D No failure • 0.011 67 h 10 m T-5-175-U 31 34 58.4 101.1 0.052 32 h 5 m T-5-50-D 11 45 18.2 57.7 0.058 10 h 7 m T-5-50-D(R) 36 56 57.0 105.6 0.020 15 h 1 m T-5-25-U 6 23 13.3 22.9 0.000 5 h 45 m T-5-30-U 13 35 23.0 41.2 0.020 12 h 36 m C-20-50-U 27 56 35.2 24.6 0.010 112 h 22 m C-20-85-U 37 56 25.4 21.3 0.040 126 h 52 m C-30-25-U 13 45 14.7 15.9 0.010 0 h 3 m C-30-50-U 14 56 15.8 20.9 0.002 139 h 0 m C-30-80-U 29 45 39.2 41.4 0.002 192 h 55 m C-30-100-U 28 56 18.2 19.3 0.057 198 h 20 m 134 0.08 0.07 0.06 o —i 0.05 j c 4) E 8 0.04 jo o. </> — 2 0.03 0.02 -0.01 -0 • - AT -• T-0-25-D T-0-100-D •T-0-175-D • — i — 0 5 10 15 20 25 30 Average hydraulic gradient, i a v Figure 5.1. Unit displacement material T-0. 35 40 45 u _j c a> E a> u « Q. (A C 3 0.07 0.06 0.05 0.04 0.03 0.02 0.01 0 -0.01 -0.02 • • • - T-5-25-D T-5-50-D - - m • •T-5-50-D(R) • • o - •T-5-175-U • A - •T-5-25-U - - D • •T-5-30-U OOOOOOOOOi •••••• 10 20 30 40 50 60 A A Average hydraulic gradient, i a v Figure 5.2. Unit displacement material T-5. 7b 135 0.08 0.02 i •C-20-50-U •C-20-85-U - - A -•C-30-25-U - - D - •C-30-50-U . - o - •C-30-80-U •C-30-100-U • 5 *•-. TQ 15 20 25. 30 35 4P -0.06 J Average hydraulic gradient, i, Figure 5.3. Unit displacement material C-20 and C-30 o.o -I—1— i 1 — i — 1 — i •— i ' — 1 — i •— i 1.E-06 1.E-05 1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 Hydraulic conductivity; cm/sec Figure 5.4. Hydraulic conductivity of the test specimen (material T-0) 136 1.0 -0.9 -0.8 - 6 T A \ .0' 0.7 -0.6 -=f 0 .5- '0 A \ f \ P 0.4 - * -T -5 -25 -D : initial --0--T-5-25-D: final 0.3 - -*r-T-5-50-D: initial 0.2 - --&--T-5-50-D: pre-failure -«-T-5-50-D(R): initial 0.1 - -D--T-5-50-D(R): pre-failure 0.0 - 1 — . — i 1 — i — i ,—, , , , , , , , 'A -i •— I "^ ~A "" 'O — 1 ' 1 — 1.E-06 1.E-05 1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 Hydraulic conductivity; cm/sec Figure 5.5. Hydraulic conductivity of the test specimen (material T-5 downward flow) 1.0 1.E-06 1.E-05 1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 Hydraulic conductivity; cm/sec Figure 5.6. Hydraulic conductivity of the test specimen (material T-5 upward flow) 137 1.0 1 — I—1 — — — I 1.E-04 1.E-03 1.E-02 1.E-01 Hydraulic conductivity; cm/sec Figure 5.7. Hydraulic conductivity of the test specimen (material C-20) 1.0 0.9 0.8 0.7 0.6 -j 0.5 0.4 -I 0.3 0.2 0.1 -I 0.0 1.E-06 —A-C-30-25-U: initial • - A - - C-30-25-U: pre-failure -•—C-30-50-U: initial - • • - C-30-50-U: pre-failure - •— C-30-80-U: initial • o • • C-30-80-U: pre-failure -•—C-30-100-U: initial •••»•• C-30-100-U: pre-failure 1.E-05 1E-04 1.E-03 1.E-02 Hydraulic conductivity; cm/sec Figure 5.8. Hydraulic conductivity of the test specimen (material C-30) 1.E-01 1.E+00 138 700 600 o O) c ro J= U X 0) 4) E 3 o > 0) o a. 500 300 200 A o T-0-25-D - - T J - T-0-100-D • A T-0-175-D T-5-25-D T-5-50-D T-5-50-D(R) T-5-25-U T-5-30-U + T-5-175-U 10 20 30 40 Average hydraulic gradient, ia 60 70 Figure 5.9. Volumetric flux through transition specimens 35 30 25 O) x *v o o E 1 15 > o CL 10 H 5] C-20-50-U C-20-85-U o C-30-25-U C-30-50-U A C-30-80-U C-30-100-U ..•••-'••jpa^ 0 ig^rfUt r B l • • • _ . . . . A A A A A A A . A A A A 10 15 20 25 Average hydraulic gradient, i a v Figure 5.10. Volumetric flux through core specimens 30 35 40 139 4.5 -r 4.0 -3.5 -E 3.0 -o ent 2.5 E 0) o re 2.0 Q. 5 1.5 -1.0 -0.5 -0.0 1 0 2 4 6 8 Average hydraulic gradient, ia Figure 5.11. Vertical displacement on test T-0-25-D 8.0 10 7.0 ..ft 6.0 0) TJ 5.0 2 u = 4.0 s TJ 3.0 re u o 2.0 1.0 0.0 i34 b f « * « i i23 i45 M2 i56 l""fl * "II If" 12 -i12 -i23 -i34 -i45 -i56 -i67 0 500 1000 1500 2000 2500 3000 3500 4000 Time; sec Figure 5.12. Local hydraulic gradient at i a v = 5 (test T-0-25-D) 140 60 0 4 , . . 1 ' 1 0 100 200 300 400 500 600 Water head; cm H 20 Figure 5.13. Water head distribution (test T-0-25-D) — • — iav = 1.0 —o—iav = 5.0 —*—iav = 11.0 • - X • • After failure 1.E-03 1.E-02 1.E-01 1.E+00 Hydraulic conductivity; cm/sec Figure 5.14. Hydraulic conductivity (based on DPT) along the specimen length 141 Average hydraulic gradient, i a v 0 2 4 6 8 10 12 1.E+00 H ' ' 1 1 1 | 1.E-03 Figure 5.15. Local hydraulic conductivity (T-0-25-D) GRAVEL SAND FINES Ccerse Medium Fine ait 100 10 1 0.1 0.01 Grain size; mm Figure 5.16. Size distribution after testing (T-0-25-D) 142 4 6 8 Average hydraulic gradient, i a v Figure 5.17. Vertical displacement in test T-0-100-D 10 12 Si. I56 167 -i12 -i23 -i34 -i45 -i56 -i67 500 1000 1500 2000 2500 3000 Time; sec Figure 5.18. Local hydraulic gradients at failure using DPTs (T-0-100-D, i a v = 8) 3500 4000 143 60 600 Water head; cm H20 Figure 5.19. Water head distribution (test T-0-100-D) 1.0E-03 1.0E-02 1.0E-01 1.0E+00 Hydraulic conducivity cm/sec Figure 5.20. Hydraulic conductivity (based on DPT) along the specimen length 144 1.E+00 o in I" 1.E-01 •o c o o u § 1.E-02 o o 1.E-03 Average hydraulic gradient, i a v 4 6 8 10 12 -o- k12 -a—k23 k34 -*—k45 -*— k56 • • • k67 Figure 5.21. Local hydraulic conductivity during the test (T-0-100-D) 100 90 liT 80 | 70 t> 60 B E 50 in o 40 ts 2 30 2 20 E 10 100 GRAVEL SAND FINES Coarsa Modum | Fine Sill 3" 1 1/2" 3/4" na" m e— #10 o-. #20 #40 #60 #100 #200 10 Grain size; mm Figure 5.22. Size distribution after testing (T-0-100-D) -^—Original —•— Bottom X Middle —Top ••0-- • Soil passing 0.01 145 2.5 2.0 i 0) E 1.5 o « Q. W z 1.0 n u r > 0.5 0.0 10 15 20 25 30 Average hydraulic gradient, i 35 Figure 5.23. Vertical displacement in test T-0-175-D 40 45 50 500 2000 1000 1500 Time; sec Figure 5.24. Local hydraulic gradient when attempting to reach i a v = 30 (T-0-175-D) 2500 146 3000 Water head; cm H 20 Figure 5.25. Water head distribution (test T-0-175-D) 147 Average hydraulic gradient, i a v 0 5 10 15 20 25 30 1.E+01 I 1.E-03 J . Figure 5.27. Local hydraulic conductivity during the test (T-0-175-D) GRAVEL SAND FINES Coarse j Medium J Fine Silt 100 10 1 Grain size; mm Figure 5.28. Grain size distribution after testing (T-0-175-D) 0.1 148 2.5 10 20 30 40 Average hydraulic gradient, ia 50 60 70 Figure 5.29. Vertical displacement in test T-5-25-D 149 1.0 0.9 £ 0.8 VI I 0.6 | 0.5 i d04 -f . 0.3 0.2 i 0.1 0.0 iav - 1.0 —Hr-iav = 9.0 —»—iav = 19 — • — iav = 28 —•—iav = 53 ... i a v = 62 1.E-05 1.E-04 1.E-02 1.E-03 Hydraulic conducivity; cm/sec Figure 5.31. Hydraulic conductivity (based on DPT) along the specimen length 1.E-01 1 .E+00 i/sec i/sec fc o 1.E-01 -j f 2f :onductiv 1.E-02 -u O 1.E-03 • 3 hydi 1.E-04 • n o _l 1.E-05 10 Average hydraulic gradient, i a v 20 30 40 50 •• 60 70 =i • a • • k23 -*—k34 -X—k45 —9t— k56 » < k67 Figure 5.32. Local hydraulic conductivity (T-5-25-D) 150 2.5 2.0 ] E u c E 1.5 o ra Q. « !° LO n u t > 0.5 0.0 • • • • « i 10 15 20 25 Average hydraulic gradient, i a v 30 35 . 40 Figure 5.33. Vertical displacement in test T-5-175-U 4 5 r 0 200 400 600 800 1000 1200 1400 1600 1800 Water head; cm H 20 Figure 5.34. Water head distribution (test T-5-175-U) 151 1.0 1-E-05 1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 Hydraulic conducivity; cm/sec Figure 5.35. Hydraulic conductivity (based on DPT) along the specimen length 1.E+00 u 01 1 1.E-01 o > I 1.E-02 •a c o I 1.E-03 E •a Ss •C 75 1.E-04 u o 1.E-05 • B • • k23 —A—k34 - X — k45 -e—k56 • c> - • k67 Average hydraulic gradient, ia, 10 15 20 25 30 35 Figure 5.36. Local hydraulic conductivity (T-5-175-U) 152 GRAVEL SAND FINES Coarse Medium Fine Silt 3" 11/2" 3/4" 3/8" #4 #10 #20 #4 #80 #10 #2 00 Grain size; mm Figure 5.37. Grain size distribution after testing (T-5-175-U) ti Original • o - Middle - H Top O Bottom 0.01 6 8 Average hydraulic gradient, ia 10 12 Figure 5.38. Vertical displacement in test T-5-50-D 153 154 1.E+01 S 1.E-04 o 1.E-05 Average hydraulic gradient, ia, 4 6 8 10 12 14 Figure 5.41. Local hydraulic gradients during the test (T-5-50-D) 10 15 20 25 Average hydraulic gradient, ia> 30 35 Figure 5.42. Vertical displacement in test T-5-50-D(R) 155 45 0 d 1 1 1 1 1 1 1 1 1 0 200 400 600 800 1000 1200 1400 1600 1800 Water head; cm H20 Figure 5.43. Water head distribution (test T-5-50-D(R)) 156 1.E+00 o oi in 1 1.E-01 ts 3 § 1E-02 o u £ 1.E-03 "5 o o 1.E-04 Average hydraulic gradient, i a v 10 15 20 25 30 35 40 Figure 5.45. Local hydraulic conductivity (T-5-50-D(R)) GRAVEL SAND FINES Coarse Medium Fine Silt 100 r 90 ; oN 80 ; U. C 70 ; ro £ a> 60 : "ro E 50 : in c o 40 ; o n k_ «^ 30 ; (0 in ro 20 j S 10 : 100 #20 #4 #6 #10 #20 0 — —t-1-10 Grain size; mm Figure 5.46. Grain size distribution after testing (T-5-50-D(R)) 0.1 • ^ r — Original ~e—Bottom © - Middle -B—Top X - soil passing 0.01 157 2.5 2.0 I 15 H d> § 1.0 o a I 0.5 H Average hydraulic gradient, ia, Figure 5.47. Vertical displacement in test T-5-25-U • ^ • • iav = 1.0 -a—iav = 2.0 - A — i a v = 3.0 —x—iav = 4.0 - * — iav * 5.0 — i a v = 6.0 A - After failure 200 300 400 500 Water head; cm H 20 600 700 800 Figure 5.48. Water head distribution (test T-5-25-U) 158 1.0 0.9 I 0.6 > o H 0.5 OA 0.3 0.2 0.1 0.0 1.E-04 • • - - lav » 1.0 iav = 2.0 - • — i a v = 3.0 - a — i a v = 4.0 - o — l a v = 5.0 -a—iav = 6.0 * ' After failure 1.E-03 1.E-02 1.E-01 Hydraulic conducivity; cm/sec Figure 5.49. Hydraulic conductivity (based on DPT) along the specimen length 1.E+00 1.E+00 o 41 m 1 1.E-01 o >< I 1.E-02 •o c o o = 1.E-03 3 2 •a >< 75 1.E-04 u o 1.E-05 Average hydraulic gradient, i a v 2 3 4 5 P -3 • D • k23 - A — k 3 4 -e—k45 - e — k 5 6 -k67 Figure 5.50. Local hydraulic conductivity (test T-5-25-U) 159 2.5 2 6 8 Average hydraulic gradient, i a v 10 12 14 Figure 5.51. Vertical displacement in test T-5-30-U 100 200 300 400 500 Water head; cm H20 600 Figure 5.52. Water head distribution (test T-5-30-U) 700 800 160 1.0 0.9 fi 0 8 * oi * 07 fi 1 0.6 > 1 0.5 R 0.4 0.3 0.2 0.1 4 0.0 . - o • * • - iav = 1.0 —A—iav = 4.0 —o—iav = 6.0 —a—iav = 8.0 - « — l a v = 10 —K—iav= 12 —e—iav = 13 • O - - After failure 1.E-05 1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 1.E+01 Hydraulic conduclvity; cm/sec Figure 5.53. Hydraulic conductivity along the specimen length (T-5-30-U) Average hydraulic gradient, i a v 0 2 4 6 8 10 12 14 16 1.E+02 1.E+02 18 8j 1.E+01 i 1 u >; 1.E+00 1.E-01 i o o o 1.E-02-); s o 1.E-04 4 1.E-05 -k23 -k34 -k45 -k56 -k67 Figure 5.54. Local hydraulic conductivity (T-5-30-U) 161 GRAVEL SAND FINES Coarse Medium Fine Silt 0.01 Grain size; mm Figure 5.55. Grain size distribution after testing (T-5-30-U) 2.0 E 1.5 o E 41 O 1 1.0 VI > 0.5 0.0 10 15 20 Average hydraulic gradient, i a v 25 30 Figure 5.56. Vertical displacement in test C-20-50-U 162 163 Average hydraulic gradient, i a v 0 5 10 15 20 25 30 1.E+00 -u | 1.E-01 -; <-> f 1.E-02 -; 1.E-06 J Figure 5.59. Local hydraulic gradients during the test (C-20-50-U) - A - k 3 4 - « - k 4 5 - A - k 5 6 - » - k 6 7 164 2.5 2.0 1.5 o a 2 1.0 u r > 0.5 1 0.0 • •••»• * i * » • ••• - - » • » , - • > • • • • • • - • • • - « • ? • » • 0 5 10 15 20 25 Average hydraulic gradient, l, v 30 -« • 7 4 35 40 Figure 5.61. Vertical displacement in test C-20-85-U 165 1.E-05 1.E-04 - - • - -iav* 2.0 iav = 4.0 — - T B — iav = 8.0 —X—iav= 16 —»—iav = 24 — . — iav = 32 —•—iav = 36 - • C- - - After failure 1.E-01 1.E-03 1.E-02 Hydraulic conducivity; cm/sec Figure 5.63. Hydraulic conductivity (based on DPT) along the specimen length 1.E+00 TS 1.E-04 o o 1.E-05 Average hydraulic gradient, i a v 10 15 20 25 30 35 40 Figure 5.64. Local hydraulic gradients during the test (C-20-85-U) 166 100 r 90 -ll' 80 • c IS 70 : 0) 60 ; "5 E 50 • UI c o 40 • tj a 3 0 : Ul UI re 20 -5 10 • 0 -GRAVEL SAND FINES Coarse Medium Fine Sill 3" 1 1/2" 3/4" 3/8" #4 #10 #20 #4 #60 #10 " 1 •«u-«*u-*' ~ — ' ''' —"—' — .. 200 1 0 0 - A — Original —*— Bottom O Middle Top — - soil passing B Piping zone 10 0.1 0 .01 Grain size; mm Figure 5.65. Grain size distribution after testing (C-20-85-U) Average hydraulic gradient, ia Figure 5.66. Vertical displacement in test C-30-25-U 167 1.0 T 0.9 -0.8 • in ! 0.7 -5 0.6 0) 0 £ o.5 _r o.4 •a o.3 0.2 o.i -; o.o L 1.E-05 • • • • - iav = 5.8 — © — i a v = 9.7 —9—iav= 12.9 - O - After failure 1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 Hydraulic conducivity; cm/sec Figure 5.68. Hydraulic conductivity along the specimen length (C-30-25-U) 1.E+01 168 Average hydraulic gradient, i. 1.E+00 1.E-05 Figure 5.69. Local hydraulic gradients during the test (C-30-25-U) GRAVEL SAND FINES Coarse Medium Fine Silt 20 t 10 : o f — • — • — • 1 ' • • .' ' —•— • 1 — ' — • 11 * * ' •'—•—•—^ 100 10 1 0.1 0.01 Grain size; mm Figure 5.70. Grain size distribution after testing (C-30-25-U) 169 2.5 8 10 12 Average hydraul ic gradient, i a v 14 16 18 20 Figure 5.71. Vertical displacement in test C-30-50-U 40 0 100 200 300 400 500 600 700 Water head; cm H 20 Figure 5.72. Water head distribution (test C-30-50-U) 170 1.0 0.9 i 0.8 0.7 -F 0.6 - : 01 > I 0.5 R 0.4 0.3 i 0.2 0.1 0.0 • - iav = 2.1 iav = 4.1 -»—iav = 8.2 —4—iav = 14 B After failure —6—iav = 18.5 1.E-01 1.E-06 1.E-05 1.E-04 1.E-03 1.E-02 Hydraulic conducivity; cm/sec Figure 5.73. Hydraulic conductivity along the specimen length (C-30-50-U) 1.E+00 1.E+01 8 1.E+00 ] J . 1.E-04 o 2 1.E-05 1.E-06 Average hydraulic gradient, i a v 6 8 10 12 14 16 18 20 Figure 5.74. Local hydraulic gradients during the test (C-30-50-U) 171 GRAVEL SAND FINES Coarse Medium Fine Silt 100 90 uT 80 | 70 | 60 ro E 50 VI VI to 2 40 30 20 10 0 i» Original —*— Bottom O - Middle Top A ~ soil passing 100 10 1 Grain size; mm Figure 5.75. Grain size distribution after testing (C-30-50-U) 0.1 0.01 0.0 15 20 Average hydraulic gradient, i a , 25 30 35 Figure 5.76. Vertical displacement in test C-30-80-U 172 1.E+01 S 1.E+00 =? 1.E-01 > % 1.E-02-I c o £ 1.E-03 o 1.E-05 Average hydraulic gradient, I 10 15 20 25 30 1.E-06 Figure 5.78. Local hydraulic gradients during the test (C-30-80-U) 173 174 Average hydraulic gradient, ia, Figure 5.81. Vertical displacement in test C-30-100-U 0 100 200 300 400 500 600 Water head; cm H20 Figure 5.82. Water head distribution (test C-30-100-U) 700 800 900 1000 175 1.0 0.9 0.8 0.7 0.6 > | 0.5 i \4 0-4 0.3 -E 0.2 0.1 0.0 -iav = 2.0 -iav = 4.0 -iav = 12 - iav = 20 - iav = 24 -iav = 28(70kPa) -iav = 28(65kPa) • After failure 1.E-06 1.E-05 1.E-04 1.E-03 1.E-02 1.E-01 Hydraulic conducivity; cm/sec Figure 5.83. Hydraulic conductivity along the specimen length (C-30-100-U) 1.E+00 1.E+00 Average hydraulic gradient, ia 10 15 20 25 30 1.E-06 Figure 5.84. Local hydraulic conductivity during the test (C-30-100-U) 176 GRAVEL SAND FINES Coarse Medium Fine Silt 3 " 1 1 / 2 " 3 / 4 " 3 / 8 " # 4 # 1 0 # 2 0 # 4 # 6 0 # 1 0 0 # 2 0 0 10 o r — . — • — . — , 1 1 0 0 1 0 1 0 . 1 0 . 0 1 Grain size; mm Figure 5.85. Grain size distribution after testing (C-30-100-U) 177 6.0 ANALYSIS AND DISCUSSION Test data, and companion visual observations, were examined to better understand the influence of geometrical and hydromechanical factors on potentially unstable soils. A unified method of analysis is proposed to assess the potential for internal instability, predicated on the two more commonly-cited methods in the literature. It is evaluated against the findings of selected experimental studies reported in the literature, and the findings of the current study. Thereafter, hydromechanical constraints to the onset of internal instability are analysed for each of the four soils used in testing. More specifically, the combined influence of effective stress and hydraulic gradient is examined, together with test duration and seepage velocity. A model is then proposed that reconciles an apparent discrepancy, in the four soils tested, between a ranking of their potential for instability based on geometrical influences and a ranking for the onset of that instability based on hydromechanical influences. 6.1. Geometrical analysis of tested soils Empirical rules have been extensively used in professional practice, to establish the potential for internal erosion. The four soils used in this research are evaluated using these rules, and compared with the findings of other studies on similar potentially unstable gradations. Relative plotting position on a unified stability chart, with axes defined by the Kezdi (1979) and Kenney and Lau (1985) indices, is used to establish the susceptibility to internal instability. 178 Grain size distribution curves for the transition (T-0 and T-5) and core (C-20 and C-30) materials are deemed potentially internally unstable by Kezdi and Kenney & Lau methods: a summary of the (H/F)min and (D'is/d'gsVax is given in Table 4.3. As noted previously, the location (F) of maximum values of D'is/d'gs and minimum values of Ff/F match very closely the location (F) of the gap in the corresponding grain size distribution (see Figures 4.3, 4.5, and 4.6). The two existing methods may be combined to yield a unified procedure to assess internal instability based on (Dis'/dgs'Vax and (F£/F)mjn. Data for stable and unstable gradations, reported for 17 soils by Kenney and Lau (1985, 1986), and another 17 gap-graded soils by Honjo et al. (1996) for which the finer fraction is less than or equal to 50%, are summarized in Figure 6.1. Additionally, results are included for 4 soils, for which data are taken from Skempton and Brogan (1994). Two distinct clusters appear in the data, with all unstable gradations defined by the quadrant (D'i5/d'85)max> 4 and a modified (H/F)mi„ < 1 as described in Chapter 2. All four gradations of core and transition materials tested in the current laboratory investigation plot in the unstable quadrant. Of further interest to the utility of the unified plot is the observation that the two core materials are located closer to the intersect (H/F)min =1 and (Dis'/dgs'Vax = 4 than the two transition materials. By implication, the core materials would likely exhibit a greater resistance to onset of failure by internal stability. Specifically, the geometrical ranking of susceptibility to internal instability is given by: T-5 > T-0 > C-30 > C-20. However, the 179 hydromechanical ranking obtained during testing is given by: T-0 > T-5 > C-30 > C-20 as will be shown below. Minor variations may occur in the target grain size distribution during reconstitution of the laboratory test specimen; for this reason, the sensitivity of the gradation curve analysis was examined for one of the materials, to gauge the general influence of such variations on findings implied by Kenney and Lau and the Kezdi methods of evaluation. Accordingly, the grain size distribution of the C-30 core material is compared with a curve that exhibits a minor variation (see Figure 6.2). The variation has a fines content of 28.0 % rather than 29.3 %, and the slope of curve is slightly different in the sand size. Although the grain size distribution is not changed significantly, the values of (Ff/F)min at F<0.3 change from 0.42 to 0.51 (see Table 6.1 and Figure 6.3), which represents a change of approximately 20 %. Similarly, values obtained using the Kezdi method also change (see Figure 6.4). Although the values of Dis'/dss' are very close to those of the original curve, the (Dis'/dgs^max at F « 5 % changes from 10 to 8.4 (see Table 6.1). Therefore, this suggests that the geometrical constraints to instability likely vary spatially within a specimen, even if there is only a minor variation in the grain size distribution. Instability will tend to exploit a localized path where the geometrical constraints are more susceptible to instability, and it is believed this might explain experimental observations of both migration in varying directions and different spatial locations of the onset of instability. 6.2. Generalized nature of particle migration 180 When the geometrical criterion of (H/F)m i n < 1 or (Dis'/dgsOmax = 4 is exceeded, there is a potential for seepage-induced internal erosion. If the seepage forces are large enough, the finer fraction may start to migrate through the interstices of the coarser fraction. Two different types of internal erosion are recognized; namely suffusion and suffosion (see Chapter 1). Suffusion is believed to be a retrogressive phenomenon as fines were observed to leave first the zone closer to the exit face of the specimen. Suffosion is either retrogressive or progressive depending on the layer where failure initiates. In some cases failure initiated at the exit face of the specimen, and the internal erosion continued thereafter in a retrogressive direction. In other cases failure initiated inside the specimen and internal erosion advanced progressively to the exit face. In discussing the generalized nature of particle migration, three groups of materials are distinguished: T-0 that is without fines; T-5 that has relatively low fines content; and the C-20 and C-30 materials for which the finer fraction has a relatively large fines content. 6.2.1. Soils with no fines content (T-0) Minor suffusion of finer particles was observed visually at the base of the specimen, when i a v was increased from one test stage to another. Most of particles came from the centre base area of the specimen, and very few from the periphery of the permeameter. The onset of internal instability (suffosion) is relatively fast (less than 3 minutes) and, from inspection of local hydraulic gradients (see Appendix B, Figures B . l to B.3), is characterized by an intensity that varies between tests. After the onset, changes in the structure of the 181 specimen occur but the specimen tends to recover for some time after most of the finer fraction particles are released. Visual observations suggest that finer fraction particles seem to migrate downstream at a similar velocity to the seepage velocity, accompanied by displacement of the loading plate, and noise of particle contact from the soil specimen. "Pipes" connected layers of the specimen approximately 1 to 5 minutes after the onset of instability. These pipes seemed to concentrate most of the seepage flow, and increased in size gradually until the test was stopped. Generally, for this material, internal erosion initiated in either zones 4-5 or 5-6 (see Table 5.3, and Figure 3.7), and moved retrogressively upstream through the specimen. Spatial analysis of finer fraction content after failure was made in Chapter 5, Section 5.2.4, and related photographs are presented in Figures CI to C.8. 6.2.2. Soils with some fines content (T-5) Qualitative evidence of suffusion of fines (< 75 p.m), was obtained through the change in colour of the water at the exit of all specimens, but no quantitative evidence was obtained. In contrast, suffosion was observed in 5 of 6 of the tests on this material. Failure was very similar to that observed in material T-0. Particle migration manifested itself through particle/water flow in channels that, after a couple of minutes, affected most of the specimen. These initial flow channels were typically vertical but also included short portions that were diagonal and sometimes horizontal. "Boiling" of particles was observed, during the test with upward flow, where migration of the finer fraction yielded voids or empty pockets at the wall of the permeameter. 182 A considerable proportion of the soil migrated from the specimen but, in tests with upward flow, subsequently returned into the specimen after the test was stopped. Post-failure observations indicates that soil migration typically led to the formation of "pipes" that grew to affect the complete area of the soil specimen in 4 of 6 tests after failure (see for example Figure C. 14 and C. 15). 6.2.3. Soil with significant fines content (C-20, C-30) Changes in the specimen before the onset of internal instability were easier to detect in these soils due to the larger fines content. Visual observations gave the impression that failure was relatively more severe in core materials than in transition materials. Fines migration by suffusion was marked by a change in local hydraulic gradient, for example that in the C-30-80-U test specimen during a stage of approximately 24 hours (see Figure 6.5). During this stage, while '134 and Us were slowly decreasing, is6 and 'i6i were increasing. The average hydraulic gradient was kept constant, however, fines were leaving the uppermost portion of the specimen, causing local hydraulic gradients to decrease. Cloudy water was visually observed at the exit face of the specimen, confirming the fines migration. This process was time dependent, as the decrease in hydraulic gradient was slow and continuous (see Figure 6.5). The same pattern was observed in other long stages of core materials. 183 Suffosion at the onset of internal instability was again observed to include combinations of vertical and horizontal flow. On occasion failure yielded "eruptions" of finer fraction on the exit face of the specimen. Pipes in core material were characterized by a considerably small diameter than those in transition materials, and did not appear to.grow by entraining finer fraction from adjacent soils in the same aggressive manner. Accordingly, the initial extent of the zone affected by instability appeared smaller (see Figure C.21 to C.23). Typically, grain size analysis of specimens after failure showed a size distribution that was very similar to the initial. Hence the size of the pipes that were observed usually was insufficient to provide for detection by change in average grain size. In general, tests may exhibit: (i) localized failure followed by re-establishment of a stable equilibrium, or (ii) localized failure followed by a period of unstable equilibrium within the specimen, other tests exhibit (iii) localized failure that triggers an immediate generalized failure throughout the specimen. Three of four tests on material C-30 provide good examples of this latter failure mechanism and its propagation through the specimen. Test C-30-50-U is an example of the failure type (i). All tests on material C-20 provide a good example of failure type (iii). 184 6.3. Factors influencing the onset of instability The onset of internal stability is affected by internal conditions in the soil, and may also be governed by aspects related to the permeameter and exit boundary of the specimen. These characteristics are now examined in an attempt to define the most important factors that influence the onset of internal instability (see Figure 6.6). 6.3.1. Equipment As mentioned on Chapter 3, it was believed the equipment would not exert an influence on the response of the test specimen to seepage flow. The rigid wall of the permeameter did not cause preferential flow along the wall of the permeameter. Clear examples are given by test T-0-25-D (see Figure C.4), test T-5-50-D(R) (see Figure C.13), and C-20-85-U (see Figure C.19) where particle migration was predominantly in the interior of the specimen. Further, when pipes were observed to contact the permeameter wall, the path would often return into the specimen as it progressed, as was observed during test T-5-175-U (see Figure C . l l ) and test C-30-50-U (see Figure C.22). 185 Specimen dimensions are deemed adequate, although it was recognized the maximum particle size for transition materials is slightly larger than the maximum diameter recommended by ASTM. Wire mesh screen openings at the exit face of the specimen were also expected to hold the coarser fraction of the soil yet allow the finer fraction to pass through (see discussion in Chapter 3). Test C-30-25-U was performed with an exit wire mesh 2.76 mm opening size compared to 1.14 mm for all other tests with C-30 material. The response of this test followed the same pattern of the other C-30 specimens, and the soil passing through after failure was very similar in size to the other specimens (see, for example, Figures 5.61 and 5.66). 6.3.2. Hydromechanical influences on internal stability Hydromechanical parameters believed to influence the onset of internal stability include direction of flow, seepage velocity, hydraulic gradient, and effective stress (see Figure 6.6). Testing of core and transition materials, enables a more extensive analysis of these variables. 6.3.2.1 Rate of change of i a v and pore volume exchange The amount of water flowing through the specimen, and the rate of change of i a v , were studied to establish if they might influence the onset of internal instability. 186 The response of materials with fines (T-5, C-20, and C-30) suggests that fines migration (< 75 pm) depends on the pore volume exchange or duration of seepage flow, as was described previously. This time dependent process was evident during long stages on core materials (see Figure 6.5). On the other hand, finer fraction migration at the onset of internal instability or suffosion leads rapidly to a catastrophic failure of the soil specimen. No relation was found between the duration of seepage flow in each test and the onset of suffosive action (see Figures 5.9 and 5.10). Rate of increase in i a v describes the temporal change of i a v across the test specimen. The laboratory investigation included one test on core material, C-30-25-U, that was taken to failure in 3 min. In contrast to the companion C-30-50-U and C-30-80-U tests, in which failure occurred after 139 h and 193 h respectively, test C-30-25-U represents a rate of increase in i a v that is extremely fast (see Table 5.3). A comparison of the three responses, prior to the onset of instability, indicates good agreement between mean vertical effective stress and local hydraulic gradient in the zone of failure as is shown in Section 6.3.2.4. Accordingly, it would appear likely that rate effects do not exert a significant control on the critical hydraulic gradient governing onset of instability. 187 6.3.2.2 Effective stress Soil specimens were consolidated under different values of effective stress on top of the specimen (a t0'), each yielding different values of effective stress at the bottom of the specimen (Obo')- After consolidation, aDo' was smaller than the applied value of ato', as a result of friction between the wall of the permeameter and the soil specimen. The idealized profile of vertical effective stress with depth, after consolidation, is given in Figure 6.7a. It describes the initial condition prior to seepage flow. In testing, oto' was maintained constant in all tests except C-30-100-U (see Chapter 5.2.7.4). Ideally, downward flow would increase the vertical effective stress due to seepage forces in the soil specimen; in contrast, upward flow would decrease the vertical effective stress in the bottom portion of the soil specimen (see Figure 6.7b, which is predicated on a homogeneous test specimen). Values of o t 0 ' and oD ' are deduced from top and bottom load cell measurements, respectively, and used to determine vertical effective stress at the port locations along the specimen. An analysis of vertical effective stress in the specimen was carried out using the computational code FLAC. The main results obtained are reported in Appendix D. Total vertical stress can be reasonably assumed to vary linearly along the length of the specimen. Accordingly, total stress on the top and bottom are calculated based on measured water pressures, and the known top and bottom vertical effective stress. Thereafter, the measured 188 variation of water pressure and assumed linear distribution of total stress enable a calculation of effective stress at each port location along the soil specimen. The deduced initial distribution of vertical effective stress after consolidation (bold line), and immediately prior to the onset of instability (dashed line) are shown for the T-0 test specimens in Figure 6.8. In these tests with downward flow, the vertical effective stress is seen to increase with seepage flow. The curves compare well with the idealized response shown in Figure 6.7b. Upon completion of consolidation, the initial value immediately prior to seepage flow yielded a ratio of Obo ' /c t o' in the range 0.32 to 0.6 for the T-0 material. Tests on material T-5 were conducted with downward and upward flow, yielding an increase and decrease of vertical effective stress respectively (see Figures 6.9a and 6:9b). Downward tests had an average initial value of the ratio Oho ' /oV in the range between 0.52 and 0.74. A small decrease is apparent in the pre-failure effective stress profile of test T-5-50-D(R), which was a consequence of the previous local failure in this same specimen (test T-5-50-D). Upward tests had an average initial value of abo ' / o t o ' in the range between 0.41 and 0.71. A small increase in the pre-failure effective stress profile of test T-5-30-U at the top of the specimen, is attributed to a low permeability layer between ports 4 and 5 that suggest that most of the seepage force was acting on that zone increasing the vertical effective stress above that zone. 189 Core material C-20, subjected to upward flow, exhibit values of 0.38 < abo'/oto' < 0.48 (see Figure 6.10). The companion tests on C-30 material similarly exhibit 0.36 < abo'/oV < 0-52 after consolidation (see Figure 6.11). A manual adjustment of the top load was found necessary during tests with upward flow to maintain a constant a t ' on the specimen, and the result was typically an increase in effective stress on the top of the specimen before failure (see Figures 6.10 and 6.11). Initial vertical effective stress was higher at top of the specimen, than the bottom, as a consequence of the friction on the wall of the permeameter cell. The magnitude of abo'/oV is governed by two factors: the length of the specimen; and the relative influence of self-weight compared with oV. All tests experienced a frictional loss, and this loss increased with o t0'. Inspection of the initial curves suggests the frictional loss is comparable for all soils, because the slope of the relation between of vertical effective stress and height above base is very similar for a similar value of o t 0 ' (see for example, the initial curves of T-0-25-D and T-5-25-U; and those of T-5-50-D, C-20-50-U and C-30-50-U). Figures 6.8 to 6.11 also show the deduced value of vertical effective stress (Ovm') immediately prior to the onset of internal stability, termed pre-failure. Test T-5-30-U experienced a vertical effective stress nearly zero close to port 5, and visual observations tend to confirm this low effective stress as some horizontal cracks were found to develop in the lower portion of the specimen (as described in Chapter 5.2.5.6). However, the onset of 190 internal instability was located above this point in zone 3-5 (see Figure B.9). The curves show a profile of vertical effective stress prior to failure that is variable and the variation exerts an influence in the onset of instability, as will be shown later in this chapter (see section 6.3.2.5). 6.3.2.3. Local hydraulic gradients The magnitude of seepage force in the zone of failure depends on the local hydraulic gradient (ijk) across that layer: the higher the ijk, the larger is the seepage force acting on the finer fraction particles of the specimen. Internal instability occurs where grains of the finer fraction migrate through interstices of the matrix formed by the coarser fraction. Accordingly, failure is associated with an increase in void ratio (e) and, as a result of the loss of finer grains, an increase in the characteristic grain size (Ds). For the case of a saturated soil, the Kozeny-Carman equation establishes hydraulic conductivity as proportional to e3/(l+e) and to D s 2 and, therefore, a localised increase in e or D s leads to a localised increase in conductivity of the soil. Since changes in local hydraulic gradient (ijk) correlate inversely with changes in hydraulic conductivity, the onset of internal instability is defined, for purposes of analysis in this study, by a significant decrease in ijk over a relatively short period of time. The variation of ijk with time, in the failure stage of each test, is reported in Appendix B. The exception is test T-5-25-D, which did not fail. In general, local hydraulic gradient is seen to 191 change significantly during the failure stage, and the change is attributed to initiation of a seepage-induced migration of the finer fraction of the soil. Interpretation of the curves associates loss of the finer fraction with a decrease in ijk, and attributes a companion increase in any downstream zone to particle migration generating a net gain of finer fraction. Interpretation of these relations between ijk and time yields a value for the local hydraulic gradient just prior to the initiation of instability, termed the critical hydraulic gradient (iCT). Values of i c r for each test, established from inspection of the curves, were reported in Table 5.3 and reproduced in Table 6.2. Location of failure is defined by the zone, within the test specimen, where the decrease of ijk is first observed to occur, using measurements from either the array of differential or total pressure transducers. 6.3.2.4. Hydromechanical boundaries The variation of mean vertical effective stress (a'vm) in the zone where failure initiates, with local hydraulic gradient (ijk) across that zone, is shown in Figures 6.12 to 6.14. The relation defines a hydromechanical path to failure in 14 of the 15 tests, concluding with the value of i C T . The trend is typically one of increasing o'vm with downward flow, and nearly constant a'vm with upward flow. The curves for each material appear bounded by an envelope at which failure initiates. The envelope is defined generally by a linear relation between o'vm and ijk, and a small offset in hydraulic gradient at zero effective stress. The offset was introduced based in previous 192 studies (Skempton and Brogan, 1994) and is common to all envelopes. The boundary is strongly evident in the transition materials, T-0 and T-5, and in the C-30 core material. It is less apparent in the C-20 core material, for which only two tests were performed. Critical hydraulic gradient is seen to increase with increasing effective stress. The envelope does not appear unique to any soil, which implies that, at any value of effective stress, initiation of instability requires a larger hydraulic gradient in the core materials than the transition materials. More specifically, the T-5 transition material requires a larger gradient than the T-0 material, and the C-20 core material requires a larger gradient than the C-30 material. The resulting hydromechanical ranking obtained is that of T-0 < T-5 < C-30 < C-20. It is similar, but not identical to, the geometrical ranking of susceptibility to internal instability given by T-5 < T-0 < C-30 < C-20 (see Section 6.1). The linear failure boundaries are defined by two parameters: namely, the slope of the curve, and the interception on the "ijk" axis. All boundary curves yield an intercept with the "ijk" axis (icr), or zero vertical effective stress, smaller than 1.0. The classic study of Terzaghi on piping associated with heave establishes a theoretical critical hydraulic gradient of approximately 1 to cause the effective stress to be zero (upward flow towards a free surface). For the case of unstable sandy gravels at extremely low vertical effective stress, Skempton and Brogan (1994) observed "... piping of the sand grains can occur at hydraulic gradients one third to one fifth of the theoretical critical gradient...", (see Figure 6.15). The tentative relation appears insensitive to vertical or horizontal flow. Hence, based on this reasoning the intercept in Figures 6.12 to 6.14 was established at approximately 0.5. Thereafter, the slope 193 of the hydromechanical boundary is determined from the position of the best fit line, which was established manually with reference to the last data point of each path. The slope of the hydromechanical boundary is believed to depend partially on the geometric characterization of each material. Indeed, there is a difference between the postulated hydromechanical boundary of transition and core materials (see Figure 6.12 to 6.14) which appears consistent with the transition materials plotting farther from the stable quadrant on the unified chart than core materials (see Figure 6.1). The plotting positions explain the relatively greater resistance of the core materials, and the hydromechanical ranking of the two core materials. They do not, however explain the hydromechanical ranking of the two transition materials. Yet, the experimental results indicate a different slope for these materials, where T-5 specimens exhibit a higher i C T than T-0 specimens. The variation of effective stress in the zone of the onset of instability is proposed below to explain this apparent inconsistency in ranking of T-0 and T-5 materials (see Section 6.3.2.5). The hydromechanical paths obtained for T-0 specimens appear to define a hydromechanical boundary (see Figure 6.12). Two paths are shown for test T-0-25-D, as failure is believed to occur at the same time in layers 45 and 56 (see Figure B. 1). The hydromechanical paths obtained for T-5 specimens typically exhibit a decrease in vertical effective stress with upward seepage flow, except test T-5-30-U where local vertical effective stress increased slightly due to the seepage force and load control on top of the 194 specimen (see Figure 6.13). Five of the six tests exhibited failure by internal instability: test T-5-25-D did not fail at hydraulic gradients within the capacity of the apparatus, and a representative path is included for illustrative purposes. Inspection of curves shows test T-5-50-D to yield failure before reaching the postulated boundary, and the response is attributed to the variation of effective stress in the zone of failure itself. Interesting, the hydromechanical boundary seems to be independent of the direction of seepage flow, given failure in test T-5-50-D(R) is governed by the same hydromechanical boundary as three with upward flows. All the hydromechanical paths for C-20 and C-30 materials show a decrease in vertical effective stress (see Figure 6.14). As mentioned in Chapter 5 (see Section 5.2.7.4), one test was performed to study the effect of a reduction of vertical effective stress (ot') while maintaining i a v constant (C-30-100-U). The results suggest that' if rjvm' decreases, instability still occurs at the same hydromechanical boundary (see Figure 6.16). Failure was observed in the bottom layers but not observed in the top layer 3-4 (see Figures B.15 and C.26). This is entirely consistent with the hydromechanical path of zone 3-4 O 3 4 ) , which did not reach the hydromechanical boundary (see Figure 6.16). 6.3.2.5. Flow direction and stress variations The influence of upward versus downward flow, on the onset of instability, deserves careful analysis. It can be most easily considered through reference to the response of T-5 materials 195 that were subjected to flow in both directions. Test T-5-50-D yields a relatively lower critical gradient than other tests performed with this material. In contrast, test T-5-50-D(R) yields a critical hydraulic gradient that is consistent with the hydromechanical boundary defined by the upward tests (see Figure 6.13). Therefore, two specimens with the same flow direction and material, yield different hydromechanical boundaries. The difference is not explained by geometrical constraints. It is believed that the direction of the seepage flow relative to any variation in vertical effective stress also exerts an important role. In tests T-5-50-D and T-5-50-D(R) the vertical effective stress was higher at the bottom of the zone where failure initiated, zones 4-5 and 56 respectively (see Figure 6.9). Therefore, the direction of seepage flow is into a zone of increasing effective stress. The magnitude of the increase across the layer (Aavm') is significantly greater for test T-5-50-D(R) than T-5-50-D (see Table 6.2). This consequence for the hydromechanical boundary of flow direction and magnitude of stress variation is quantified below with reference to a limit equilibrium analysis. Further experimental evidence on the apparent role of vertical effective stress distribution along the specimen length prior to failure is also evident in tests C-30-80-U and C-30-100-U. In this case, the two tests exhibit a nearly identical profile of Ovm', and therefore values of Aovm', on a layer by layer basis (see Figure 6.11). Notwithstanding the very different hydromechanical paths imposed in the two tests, the same boundary was obtained in testing (see Figure 6.16), Again, these results confirm that not only Ovm', but Aovm', influence the onset of internal instability. Therefore, it is important to account also for the direction of flow relative to the change of vertical effective stress, since both seem to influence the 196 hydromechanical failure boundary of a material. This accounting for flow direction and variations of vertical effective stress was one of the main objectives of the model development reported below in section 6.4. 6.3.2.6. Seepage velocity Actual average seepage velocity though the soil specimen was calculated given the cross sectional area (A) of the permeameter cell, using measurements of specific discharge (q), and average hydraulic gradient (ia v), and knowing the porosity of the specimen. The determined seepage velocity represents an average on the length of the soil specimen. Drag forces on particles are linked to seepage velocities on the soil. This drag forces could trigger failure by internal stability. On the other hand, according to standard seepage theory the hydraulic gradient is the critical issue. Data from this laboratory testing programme are used to check the assumption about seepage velocities. Considering the 14 tests that exhibited failure, the relation between mean vertical effective stress (c'vm) at the onset of instability, in the failure zone, and actual average seepage velocity (va) just prior to instability, is given in Figure 6.17. The onset occurs at lower velocities in the core, than in the transition materials. More specifically, the instability in C-30 material occurs at a lower velocity than the C-20 material; likewise, the instability in T-5 material occurs at lower velocity than the T-0 material. The general trend is one of onset at a higher 197 velocity in materials with lower finer fraction. However, the results suggest that seepage velocity does not yield boundaries as were found for the local hydraulic gradient (ijk). Core materials C-20 and C-30 do not show any correlation between the different tests. Transition materials T-0 and T-5 indicate the critical seepage velocity increases with vertical effective stress. The reason why a trend between effective stress and seepage velocity was not found is explained by the fact that seepage velocities are controlled by the average hydraulic gradient, and not the local hydraulic gradient. 6.4. Proposed model for internal erosion A model based on porous media theory is applied to support the concepts regarding the hydromechanical boundaries of internal instability and to reconcile the apparent disconnect between geometric ranking and hydromechanical ranking as described above (Section 6.3.2.4). As reported in Chapter 2, a few studies have proposed physical models to determine the onset of internal particle migration in soils (eg. Indraratna and Radampola, 2002). A new model (see Figure 6.18) that considers the equilibrium of a three component system that comprises water, soil skeleton, and finer fraction is proposed (mixture theory). The features of this model include; (i) the ability to include geometrical analysis based on the grain size distribution shape; (ii) the capability of including the value of vertical effective stress and variation of vertical effective stress in the zone of failure; (iii) it can be used to model downward or upward flow, considering the effect of gravity force; (iv) strength parameters of the soil being tested such as friction angle; (v) simplicity and efficiency. The model can 198 distinguishing the general trend of hydromechanical boundaries for different materials and can easily be incorporated to a numerical code (such as FLAC). As shown in Figure 6.18, the main assumptions of this model are: (1) Inertia forces are neglected; (2) Seepage forces act only on the finer fraction particles; and, (3) Effective stress on the finer fraction is a proportion of the effective stress on the specimen. 6.4.1 Model formulation During this testing programme different parameters have been found to influence the onset of internal instability. Therefore, the formulation of a physical model based on fundamental principles that reflects the same characteristics observed during testing, is extremely difficult. A representative unit volume of soil comprises three components that can move independently but that influence each other through interaction forces (see Figure 6.19). Stress is the sum of three components defined by the fluid, coarser, and finer fractions: T = Tw+r=Tw+Tc+Tf=nTw+ncTc+nfTf Eq. 6.1 199 T c and T f are the equivalent "effective stress" on the coarse and finer fraction particles. Volume effects are considered through the use of n1 values that represent the volume fraction of each component. Accordingly, the volume fraction of these components is (see Chapter 2, Eq. 2.3): n w = n: fluid (water) fraction, or porosity nc: coarser fraction (= 1- n - nf) nf: finer fraction, where n is calculated from nf = — * 100 (Thevanayagam, 1998), knowing: Eq. 6.2 Sf e: void ratio, and Sf. finer fraction content by mass (%) Balance of momentum can be used to describe limit equilibrium of the secondary component (finer fraction). More specifically, the approach relates the three components with reference to the interaction forces between them. Momentum balance of each component, in the vertical direction, is given by (see also Chapter 2, Eq. 2.9): 8TW dV • + pwg = Rl-R2 + p w ^ : water ^ Eq. 6.3a dz. ° dt dT° dV + pcg = R3-Rl + pc — - : coarser fraction Eq. 6.3b dz dt 200 8Tf + pJg = -R3 + R2 + p dV, f dz dt : finer fraction Eq 6.3c where (see Figure 6.19): T1: stress in the " i " component p': unit density of the " i " component (pc=pf=2.7 Mg/m3; pw=1.0 Mg/m3) g: gravity Vj velocity of the " i " component RI: interaction force water/coarser fraction R2: interaction force water/finer fraction R3: interaction force finer/coarser fraction Consider the rate of change of velocity (^ -j: a small increase of i a v was applied to the soil specimen during testing, therefore the rate of change in seepage velocity is very low ((^L)=0). Additionally the coarser and finer fractions are not considered to move prior to the onset of instability. Therefore inertia forces are neglected in this model. Consequently, the last term on the balance of momentum equations is equal to zero. 201 Consider now the interaction forces RI, R2, and R3 between the three components. RI and R2 are function of the seepage force per unit volume given by iJlc • yw, where ijk is the local hydraulic gradient. It is assumed that due to the difference in size of the finer and coarser grains, the head loss occurs predominantly in the finer fraction. For purposes of modeling this concept is further simplified and seepage forces are assumed to act entirely on the finer fraction, and therefore RI = 0 and R2 = ijk •yv/. R3 is the interaction force per unit volume between the finer fraction and the coarser fraction of the soil. It is calculated directly from Coulomb theory, for an assumed value of the true friction angle ^ = 30° between both fractions, in all of the materials tested. While it is difficult to obtain a reliable evaluation of ty^. (Mitchell, 1993), the assumed value of 30° is reasonable for bulky minerals such as a mix of quartz and feldspar. Therefore: G*avm ••tamV) R3 = ^ Eq. 6.4. n1 Az where: Az: unit length (see Figure 6.4) G*: geometric-based factor (see Equation 6.5) Consider now the finer fraction transmits only a portion of the effective stress (c>vm') on the specimen. Similarly, Skempton and Brogan (1994) believed that in sandy-gravel mixtures "the greater part of the overburden load is carried on a framework of primary fabric of gravel 202 particles, leaving most of the sand under relatively small pressures". Therefore, the following assumption is postulated: TF=G*-oJ or TF =G~-CTJ Eq. 6.5 G* or G** is a geometric-based factor that determines the proportion of stress transferred to the finer fraction particles. Then the critical hydraulic gradient i c r can be derived from equations 6.3, 6.4, and 6.5 as follows: G* TI^ p g (cr 'tan(r^)-Arrv') — Downward flow Eq. 6.6a yw-nf-Az v yw or / = —. (cTv'-tan(^)+Acrv') + W P f 8 Upwardflow Eq. 6.6b yw-nf -Az yw Equation 6.6 suggests that the critical hydraulic gradient is a function of the effective stress on the soil specimen: as the effective stress increases the value of icr also increases. For this model (and with reference to Kezdi, 1979) it is proposed that G* depends on the difference in size of the coarser fraction, represented by D 1 5 ' , and the finer fraction, represented by dss'. When Di5'= dgV, the effective stress in the finer fraction is assumed half of that in the specimen (G* =0.5). On the other hand, if the finer fraction is very significantly smaller than the coarser fraction, say Dis'/dgs' = 20, it is assumed to take a very small proportion of stress (G* =0.1), as depicted by the limit values in Figure 6.20a. In defining the relation of G* 203 versus Dis'/dgs' between these arbitrary limits, the test data from one material (T-5) were used to establish a value of G* = 0.14 that yielded, by trial-and-error, a good fit to the data (see Figure 6.21). The T-5 material was selected because the specimens were tested with both upward and downward flows. A combination of these points (see Figure 6.20a), and an assumed logarithmic fit, allows a value of G* to be assigned to the three other materials (T-0, C-20, and C-30) based on their Dis'/dss' values. Alternatively, and with reference to the proposed unified plot (see Figure 6.1), G** can be obtained from the relative plotting that is a function of D 1 5 7d 8 5 ' and (H/F)m i n, where: G " = f ( ^ ' % t i - / l 6 + (// / Fmm-if). For arbitrary limit values and a similar fit to the T-5 material, the proposed logarithmic fit of G** is shown in Figure 6.20b. 6.4.2. Model application Values of critical hydraulic gradient, i C T , can be calculated from equation 6.6. and compared with the values of icr measured experimentally. Input parameters to the equation were deduced from laboratory measurements on T-0, C-20 and C-30 specimens just prior to failure (see Table 6.2). Calculated values of i c r are plotted together with measured values in Figures 6.21 and 6.22. The calculated values reveal the same trend in relation between a v m ' and i C T , in each of the three materials. 204 The model yields a different slope for the hydromechanical boundaries of core and transition materials, which is consistent with results obtained experimentally. This ranking in the general trend of material behaviour is attributed to the fact that core materials are closer to the stable zone on Figure 6.1. This gives a Dis'/dgs' value that is smaller in core materials, a corresponding geometric-based factor (G*) that is larger, and results in greater values of i c r . In addition, the value of n f is smaller in core materials, yielding a particle configuration that is closer to the minimum porosity (see Figure 2.19), which also results in greater values of i C T . In contrast the Dis'/dgs' of T-0 and T-5 materials are very similar (see Table 4.3) and, therefore, the value of G* derived from Figure 6.20a does not vary significantly. In addition, values of n f are very similar for the two transition materials (see Table 6.2). Consequently, the difference in the slope of the hydromechanical boundary cannot be explained through geometric considerations alone. Rather, the variation in effective stress Aavm' must be considered. Figure 6.9 shows the variation of vertical effective stress on test specimens T-5. The variation of effective stress in the zone of failure in test T-5-50-D is very small, and the corresponding influence of the Aavm' is believed insignificant in this test (see Table 6.2). Accordingly, while the hydromechanical path of test T-5-50-D appears to result in the onset of instability at a combination of CK™' and i C T that does not confirm to the boundary defined by other T-5 test (see Figure 6.23), the model calculation of i c r confirms the path is indeed consistent with an analysis that accounts for the influence of flow direction and variation of effective stress. Specifically, results from laboratory testing and modeling of test T-5-50-D suggest the variation of effective stress in the zone of failure may be an important parameter 205 when evaluating hydromechanical conditions necessary for the onset of internal instability. Consequently, it is believed that any material exhibits a threshold relation between Ovm' and icr (hydromechanical boundary) that is non-unique; it depends on the variation of vertical effective stress in the direction of seepage flow (See Figure 6.24). 6.5 Summary Geometrical analysis using methods reported by Kezdi (1979) and Kenney and Lau (1985 and 1986) indicates the four soils tested are potentially unstable. Experimental observations confirmed the internal stability of these widely-graded cohesionless soils, thereby extending the utility of these methods originally derived for sands and gravel materials to soils with significant silt content. It is proposed that a greater confidence may be gained in the assessment of these soils by combining the two methods in a unified plot. Effective stress and hydraulic gradient are deemed the most important parameters to define the onset of internal instability. Combinations of these parameters yielded hydromechanical boundaries for the onset of internal instability. These hydromechanical boundaries depend on the material (geometrical constraints) and variation of vertical effective stress in the zone of failure. The onset of instability is triggered either by an increase of hydraulic gradient, or by a decrease of vertical effective stress. Each yields the same response, and the onset is located 206 on the same hydromechanical boundary. Therefore, hydromechanical boundaries seem independent of the hydromechanical path (ijk and Ovm') followed prior to the onset of failure. A physical model is used to explain an apparent inconsistency in geometrical ranking of susceptibility to internal instability of the four soils (C-20, C-30, T-0, and T-5, with the latter being most susceptible), given the ranking of hydromechanical susceptibility based on the experimental results (C-20, C-30, T-5 and T-0, with the latter being most susceptible). The model is predicated on three assumptions: inertia forces are neglected, effective stress on the finer fraction is a portion of the effective stress on the specimen, and seepage force acts only on the finer fraction particles. A comparison of measured (laboratory) and calculated (model) icr yields the same ranking of hydromechanical susceptibility. Consequently, these assumptions influence the magnitude of i C T but not the relative ranking for each soil. This model both identifies and explains the importance of variation of vertical effective stress (Aavm') in the direction of flow, in addition to the governing influence of mean vertical effective stress (avm') and hydraulic gradient (ijk) across the zone of failure. Importantly, the contribution of Aavm' to the hydromechanical susceptibility is independent of the three assumptions. 207 Table 6.1. Effects due to the variation of grain size distribution C-30 Fines (H/F)min (OVd'asW content Original 29.3 0.42 10 Variation 28.0 0.51 8.4 Table 6.2. Input parameters of the model AfJvm' O'vm Test code G* nf (kPa) A Z (m) (kPa) 4>u i c r lab i c r model T-0-25-D 0.145 0.269 2.5 0.125 38.1 30 7.0 7.9 T-0-100-D 0.145 0.279 12.6 0.125 62.1 30 12.2 9.1 T-0175-D 0.145 0.279 31.6 0.125 182.6 30 25.4 30.6 T-5-25-D 0.140 0.290 No failure 30 T-5-175-U 0.140 0.290 64.5 0.125 101.1 30 58.4 49.2 T-5-50-D 0.140 0.287 -4.9 0.125 57.7 30 18.2 14.5 T-5-50-D (R) 0.140 0.289 -45.9 0.125 105.6 30 57.0 41.5 T-5-25-U 0.140 0.284 4.1 0.05 22.87 30 13.3 18.1 T-5-30-U 0.140 0.285 17.3 0.125 41.16 30 23.0 17.3 CT20-50-U 0.200 0.153 24.4 0.125 24.6 30 35.2 41.5 C-20-85-U 0.200 0.151 10.7 0.125 21.3 30 25.4 25.3 C-30-25-U 0.170 0.180 11.7 0.125 15.9 30 14.7 16.7 C-30-50-U 0.170 0.181 16.0 0.125 20.9 30 15.8 22.0 C-30-80-U 0.170 0.191 37.2 0.125 41.4 30 39.2 44.9 C-30-100-U 0.170 0.187 8.7 0.125 19.3 30 18.2 15.2 C-30-100-U 0.170 0.187 34.6 0.125 40.9 30 39.3 43.6 208 16 14 — 12 . - \ H A x 1 0 CO E "LO I? 8 "lO Q 6 H X X O o 8' o 0 o o o o $ - r -0.5 o Kenney & Lau: Unstable • • • Kenney & Lau: Stable o o O Honjo et al.: Unstable • • # Honjo et al.: Stable A A A Skempton & Brogan: Unstable A A ' A Skempton & Brogan: Stable X X X Current stud/ 2.5 • • 1.5 (H/F) mm F<0.3 Figure 6.1. Unified stability chart 3.5 GRAVEL SAND FINES Coarse Medium Fine Silt Clay #10 #200 1 0.1 Grain size; mm Figure 6.2. Grain size distribution variation 209 70 0 V- , — , 1 r 1 , , 1 0 10 20 30 40 50 60 70 80 90 F, mass fraction smaller than; % Figure 6.3. Kenney and Lau (1986) analysis of material C-30 210 •35 -30 ' . " 2 5 c .a 2 -20 o cs •5 -15 cs u 5 -10 - 5 H •56 -i34 -i45 •I56 -i67 "67 0 10000 20000 30000 40000 50000 60000 70000 80000 90000 100000 Time; sec Figure 6.5. Local hydraulic gradient ( i a v = 24, test.C-30-80-U) 211 T i m e rate of change of iai M Inertia forces Test duration — Fines migration (< #200 sieve) w S e e p a g e ve loc i t y Rate of particle transport Geometric constraints D1S7d& •j (H/F)tili HN Unified chart Finer fraction content {8<>. nf ErTectrve stre$s Hydromechanical boundaries; & vs. ler , Flow direction Upward HN Downward Experimental results Critical hydraulic gradient, icr Hydromechanical aspects of Internal instability Figure 6.6. Factors that influence the hydromechanical response of the specimen. to \l/ \l/ 4/ \[/ \t~ 'to sj/ si/ \L \l/ \l/ \|/ /|\ /|s /j\ /|S /|\ /|s bo a) time; t 0 . J T T T T T T j * b 0 | , b upward downward flow flow b) time; t Figure 6.7. Idealized response of vertical effective stress under unidirectional flow. 212 - * — T-0-25-D: initial • - A • -T-0-25-D: pre-failure —•—T-0-100-D: initial - - o - - T-0-100-D: pre-failure - • -T -0 -175 -D : initial -T-0-175-D: pre-failure Port 1 100 150 200 Vertical effective stress, a'^; kPa 250 17 300 Figure 6.8. Vertical effective stress in the test specimen (T-0) 50 45 40 E 35 o m 2 30 2 > 25 re .c O) 20 15 H 10 5 -*—T-5-50-D: initial - A • • T-5-50-D: pre-failure -»-T-5-50-D(R): initial •o--T-5-50-D(R): pre-failure Port 2 3 100 150 Vertical effective stress, a'™; kPa 200 250 Figure 6.9a. Vertical effective stress in the test specimen (T-5, downward flow) 213 40 0 10 20 30 40 50 60 70 80 90 100 Vertical effective stress, a '™; kPa Figure 6.10. Vertical effective stress in the test specimen (C-20) 214 40 60 80 Vertical effective stress, a'^; kPa Figure 6.11. Vertical effective stress in the test specimen (C-30) 200 100 120 •T-0-25-D: ,45 •T-0-25-D: i56 •T-0-100-D: i45 -T-0-175-D: i56 10 20 30 70 40 50 60 Hydraulic gradient, i j k Figure 6.12. Influence of hydraulic gradient on internal instability (T-0) 80 90 100 215 200 • * -T -5 -25-D: i56 - • -T -5 -50-D: i45 - B - T - 5 - 5 0 - D ( R ) : i56 - • - T - 5 - 2 5 - U : i23 - X - T - 5 - 3 0 - U : i35 —i—T-5-175-U: i34 10 20 30 70 40 50 60 Hydraulic gradient, i j k Figure 6.13. Influence of hydraulic gradient on internal instability (T-5) 200 80 90 70 80 40 50 60 Hydraulic gradient, i J k Figure 6.14. Influence of hydraulic gradient on internal instability (C-20&C-30) 90 100 100 216 1-2 08 0-4 0 Unstable t —x-• Stable x Upward vertical flow + Horizontal flow 0 (H/F)r Figure 6.15. Critical hydraulic gradient for ovm' ~ 0 (after Skempton and Brogan, 1994) 200 ra 180 f 160 I • i 140 j % 120 4> « 100 > 0) ra o t 01 > I : onset of rjt' reduction -e— C-30-25-U: i45 -B—C-30-50-U: i56 -A—C-30-80-U: i45 ••••C-30-100-U-i34 -•— C-30-100-U-i45 -o—C-30-100-U-i56 ..-"£'-30 10 20 30 40 50 60 Hydraulic gradient, ijk Figure 6.16. Influence of effective stress (C-30-100-U) 70 80 90 100 217 200 rs 0. 180 160 • | 140 </> > I 01 "fS o 0) > 100 80 60 40 20 f A C-20 A C-30 • T-0 • T-5 0.0001 0.001 0.01 0.1 1 Actual average seepage velocity, va; cm/sec Figure 6.17. Seepage velocity at the onset of instability -^flwd«e<iti«>"| -H * 1 Internal instability Volumetric based appiosch Limit tquilibrium Mixture theory 3 components Finer fraction Coarser fraction Assumptions Finer fraction content, Sf Weigth of particles Interaction forces Coarser vs. finer fraction —I R3 H Watefvs.soil Seepage force acts on finerfraction R1=0 R2 Inertia forces neglected Effective stress Ion finer fraction! Mi Figure 6.18. Proposed model for internal instability 218 219 0.6 0.5 -I 0.4 0.3 0.2 0.1 Verification zone y = -0.06Ln(x) + 0.20 T-5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 Figure 6.20b. Effective stress distribution factor, G** 4.0 4.5 5.0 200 180 - 160 to Q. • | 140 H > ! 1 2 0 « 100 > 4 3 I 80 f .8 60 %m <D > 40 -J 20 A / * 1 ,-"fj * 10 20 j—t—L 70 A T-0: Lab A T-0: Model • T-5: Lab. B T-5: Model 80 30 40 50 60 Local hydraulic gradient, i c r Figure 6.21. Comparison of i c r on transition materials using the model 90 100 220 200 180 - 160 f | 140 j i120 « 100 I 80 S 60 I— <u > 40 H 20 - • •""" l i • C-20: Lab. H C-20: Model OC-30: Lab. « C-30: Model 10 20 30 40 50 60 70 Local hydraulic gradient, i c r Figure 6.22. Comparison of i c r on core materials using the model 80 90 100 200 180 - 160 ro a - i 140 H > "b 8" 120 in 2 100 I I 80 8 60 5 > 40 -20 -0 A / A T-5-50-D 10 20 30 40 50 60 70 Local hydraulic gradient, i c r Figure 6.23. Values of i c r for test T-5-50-D A T-0: Lab A T-0: Model • T-5: Lab. B T-5: Model 80 90 100 221 200 180 A 160 H £ 140 •e> 120 </> vt £ 100 v> > u % 60 « u a> > 80 40 20 A 20 A C T *>0 A C T ' < ( ) 30 40 50 60 J - t - t 70 0 10 80 Hydraulic gradient, l j k Figure 6.24. Effects of variation of r j vm' on hydromechanical boundaries 90 100 222 7. CONCLUSIONS AND RECOMMENDATIONS 7.1 Conclusions The topic of this research is an experimental analysis of internal instability of widely graded cohesionless soils. The intent is to better understand the interpretation of permeameter data, from a series of unidirectional flow tests on soils representative of the filter and transition layers in a zoned embankment dam, and thereby evaluate hydromechanical influences governing the onset o f instability. Specifically the objectives of this study were as follows: • Design and commission a large permeameter test device, and associated controls, to perform internal instability tests under unidirectional seepage flow in either downward or upward flow directions; • Develop a routine test procedure to assess the effects of increasing seepage flow on potentially unstable specimens; • Describe the general nature of internal erosion from visual observations of the test specimen and its measured response; • Compare and contrast the behaviour o f Transition and Core materials under seepage flow; • Contrast the results of this study with a limited experimental database for internal instability testing; 223 • Assess the main hydromechanical factors that influence the response of internally unstable soils; and, • Establish a physical model to explain and link the experimental findings on hydromechanical influences with the potential for instability based on a geometrical characterization of the soil gradation curve. A programme of experimental research was undertaken to meet these objectives. A summary of the findings is presented below that addresses the apparatus and instrumentation, the test procedure and the interpretation of results. A large rigid-walled permeameter was designed and fabricated to accommodate a soil specimen of 279 mm diameter and variable length (between 325 and 556 mm). Important features of the apparatus are summarized below: • The measurement of axial load on the top and bottom surface of the specimen, thereby enabling more confident account of side-friction on the distribution of stress; • The use of a "frictionless" seal on the loading rod, which enables an unqualified measurement of applied load to the specimen; • A measurement of water pressures along the specimen length, yielding values of local hydraulic gradients in the soil to capture the onset of localised failure by internal erosion; • Analysis of the data indicates the apparatus provides for consistency of specimen reconstitution, and reproducibility of results during testing. 224 The specimen reconstitution technique and multi-stage testing procedure were developed for testing of 4 soil gradations (T-0, T-5, C-20, and C-30). The following observations are made from experience gained in the test programme: • A modified slurry mixing and discrete deposition technique was found satisfactory for reconstitution of a saturated specimen of widely graded soil; • Tests were performed with either upward or downward seepage flow. It transpires • that both flow directions were important to the identification of key parameters governing internal instability; and • The onset of instability does not appear to be governed by the rate at which the hydraulic gradient increases The test results yield a greater understanding of the response, to seepage flow, of soils that are potentially unstable. The most important insights are summarized below: • Suffusion and suffosion were both observed during testing. Suffusion took the form of episodic events of finer fraction migration associated with increments of i a v , and fines migration (< 75 Lim) associated with a time-dependent process. Suffosion was associated with the onset of internal instability, and was the prime focus of this study. Piping, when it occurred, took the form of a progression of that internal erosion from a localised zone of instability to a spatially-preferential failure along the specimen length. None of the visual observations and post-test analyses suggest the rigid-wall of the permeameter controlled the localization of internal instability. 225 Methods reported by Kezdi (1979) and Kenney and Lau (1985 and 1986) indicate that the four soils examined in testing are potentially unstable, and experimental observations verify this assessment. These methods were empirically developed for sands/gravels, but they appear to work for silts/sands/gravels, thereby extending their utility in engineering practice. Greater confidence in the assessment of a soil gradation may result from unifying the two methods (with a small modification to Kenney and Lau), and a comparison of limited data from other published studies supports this proposed unified approach. Effective stress and local hydraulic gradient are deemed the most important parameters that define the onset of internal instability. Combinations of these two parameters yield a hydromechanical boundary or threshold conditions to the onset of internal instability. Average vertical effective stress on the specimen and critical hydraulic gradient exhibit a linear relation, where the resistance to internal erosion increases with vertical effective stress. The hydromechanical boundary defined by civm' versus i C T is different for each of the four soils tested. The difference is attributed partly to the grain size distribution, represented by values of D 1 5 ' / dgs' or (H/F)min, given the relative difference between these values for transition (T-0 and T-5) and core (C-20 and C-30) materials. However, the difference cannot be fully explained by geometric constraints alone, given the T-0 and T-5 gradations are so similar, yet the hydromechanical boundary is markedly different. 226 Further analysis has shown the variation of vertical effective stress (Aavm') in the zone of failure also acts to influence the threshold conditions of avm' versus i C T . The onset of internal instability is triggered either by an increase in hydraulic gradient, or by a decrease in effective stress. Each phenomenon yields the same response, and the same combination of avm' and i C T , when the distribution of vertical effective stress along the specimen length is comparable. Rate of increase i a v was not found to exert a significant control on the critical hydraulic gradient (icr) governing the onset of instability. Projections of the hydromechanical boundaries suggest that when the effective stress is zero, the critical hydraulic gradient is less than 1.0, implying good agreement between these experimental data and data on "segregation piping" reported by Skempton and Brogan (1994). A three-component model is described. This model is used to reconcile an apparent inconsistency in ranking of potential for internal instability based on geometrical considerations, and ranking of hydromechanical susceptibility based on the experimental results. More specifically, given a geometrical ranking of decreasing potential for instability (T-5, T-0, C-30, C-20) and a hydromechanical ranking of decreasing susceptibility to onset of instability (T-0, T-5, C-30, C-20) that is different, the model explains the difference by means of ijk, avm', and Aavm'-The model establishes that the boundary to the onset of internal instability is not unique, rather it is governed by both avm' and its variation Aavm' in the direction of seepage flow through the zone in which failure initiates. Accordingly, for a given direction of flow, a decrease in avm' would yield a lower icr than an increase in a™'-227 This finding has profound implications for flows into zones of low stress in earthfill dams. 7.2 Recommendations for future studies In further studies of the internal stability of soils, it is recommended that consideration be given to: • Testing of additional materials with different gradation curves (yielding a methodical assessment of Dis'/dgs' and (H/F)mjn on the proposed unified chart of geometrical potential); • Examining the influence of specimen reconstitution technique, and its implications for fabric of the soil tested in the permeameter. • Defining a hydromechanical boundary for fines migration (< 75 pm) • Developing the model as a predictive tool for evaluation of earthfill dams, and .through detailed analysis of performance monitoring data, better define the spatial and temporal variations of hydraulic gradient and effective stress in those large earth structures. 228 BIBLIOGRAPHY Bear, J. 1972, "Dynamic of fluids in porous media", America Elsevier company, Inc. New York, London, Amsterdam, 764 p. Bertram, G.E. 1940, "An experimental investigation of protective filters", Harvard Soil Mechanics Series No. 7, Harvard University, No. 267. 21 p. Boer, R. 2000, "Contemporary progress in porous media theory", Appl. Mech. Rev, No. 53-12. pp. 323-370. De Mello, F. B. 1975. "Some lessons learned from unsuspected, real and fictitious problems in earth dam engineering in Brazil. Proceedings, 6 t h Regional Conf. for Africa on soil mechanics and foundation engineering, Durban, S. Africa, Sept. 1975, pp. 285-304. Foster, M. , Fell, R., and Spannagle, M. 2000. "A method for assessing the relative likelihood of failure of embankments dams by piping". Canadian Geotechnical Journal Vol. 37. No. 5. pp. 1025-1061. Garner, S.J and Sobkowicz, J.C. 2002, "Internal instability in gap-graded cores and filters". Annual conference of the Canadian Dam Association, Victoria 2002. 10 p. Honjo, Y., Haque, M.A. and Tsai, K.A. 1996 "Self-filtration behavior of broadly and gap graded cohesionless soils". In proceedings, Geofilters 96, Ecole Polytechnique de Montreal, Montreal, Quebec, Canada, p. 227-236. ICOLD. 1994, " Embankment dams: Granular filters and drains". Bulletin No. 95. International Commission on Large Dams. Paris, France. 256 p. 229 Indraratna, B. and Radampola, S. (2002). "Analysis of critical hydraulic gradient for particle movement in filtration". ASCE Journal of Geotechnical and Geoenvironmental Engineering, Vol. 128, pp. 347-350. Jones, J..A.A. (1981). "The nature of soil piping a review of research". Geo Books, Norwich. 301 p. Karpoff, K.P. 1955. "The use of laboratory tests to develop design criteria for protective filters". Proceedings, American Society for Testing and Materials 5. pp. 1183-1198. Kenney, T.C., and Lau, D. 1985, "Internal stability of granular filters", Canadian Geotechnical Journal, Vol. 22, No. 2. pp. 215-222. Kenney, T.C., and Lau, D. 1986, "Reply: Internal stability of granular filters", Canadian Geotechnical Journal, Vol. 23, No. 3. pp. 420-423 Kezdi, A. 1979. Soil Physics. Amsterdam: Elsevier. 160 p. Kezdi, A. and Konyary, Z. 1971, "Laboratory permeability tests" (in Hungarian). Diploma work, T. U. Budapest. Kovacs, G. 1981. Seepage hydraulics. Elsevier Scientific Publishing Company, Amsterdam, 730 p. Kuerbis, R.H., and Vaid, Y.P. 1988, "Sand sample preparation-the slurry deposition method", Soils and Foundations, Vol. 28. No.4. pp. 107-118. Lafleur, J. 1984, "Filter testing of broadly graded cohesionless tills", Canadian Geotechnical Journal, Vol. 21, No. 4. pp. 634-643. Lafleur J. Mlynarek J., and Rollin A. L 1989, "Filtration of broadly graded cohesionless soils", ASCE Journal of Geotechnical Engineering, Vol. 115, No. 2. pp. 1747-1767. Milligan, V. 1986. Internal stability of granular filters: Discussion. Can. Geotech. J. 23. 230 pp. 414-418. Mitchell J. K. 1993. "Fundamentals of Soil Behavior". University of California, Berkeley. 2 n d ed. John Wiley&Sons, Inc. 437 p. Moffat, R. 2002. "A laboratory study of particle migration in cohesionless soils". M.A.Sc Thesis, University of British Columbia, Vancouver, Canada. 119 p. Morgan, G.C and Harris, M.C. 1967. "Portage mountain dam. II. Materials". XIX t h Canadian Soil Mechanics Conference, II, Vol. IV, No.2. pp. 142-166. Pare, J., Ares, R., Cabot, L., and Garzon, M. 1982. "Large scale permeability and filter tests at LG 3". Proceedings, 14th International conference of Large Dams. Rio de Janeiro. Brazil, pp. 103-122. Ripley, CF. 1986. Internal stability of granular filters: Discussion. Can. Geotech. J. 23. pp. 255-258. Sherard, J.L. 1984 "Basic properties of sand and gravel filters" ASCE, Journal of Geotechnical Engineering, Vol 110, No 6. pp. 379-395 Sherard, J.L. and Dunnigan, L.P. 1984. "Filters for silts and clays". ASCE, Journal of the Geotechnical Enginnering Vol. 110, No. 6. pp. 397-414. Sherard, J.L. and Dunnigan, L.P. 1986. "Internal stability of granular filters: Discussion. Canadian Geotechnical Journal. 23. pp. 418-420. Skempton, A. W and Brogan, J.M. 1994, "Experiments on piping in sandy gravels", Geotechnique 44, No.3. pp. 449-460. Smith, N., 1971, "A history of dams", Peter Davies, London. 279 p. 231 Steward, R.A. and Watts, B D . 2000, "The WAC Bennett Dam sinkhole incident", Proceeding of the 53 rd Canadian Geotechnical Conference, Montreal, Canada, pp. 39-46. Terzaghi, K. 1939. Soil mechanics: a new chapter in engineering practice. 45 th James Forrest lecture. Journal Instn. CivilEngineers 12. Pp. 106-142. Thevanayagam, S. 1998. "Effects of fines and confining stress on undrained shear strength of silty sands". J.of Geo. and Geoenv. Eng. Vol.123, No. 6. pp. 479-490. Thevanayagam, S. Shenthan, T., Mohan, S., and Liang, J. 2002. "Undrained fragility of clean sands, silty sands, and sandy silts". J.of Geo. and Geoenv. Eng. Vol. 128, No. 10. pp. 849-859. Tomlinson, S.S. and Vaid, Y.P. 2000. "Seepage forces and confining pressure effects on piping erosion". Can. Geotech J. 37. pp. 1-13. U.S. Waterways Experiment Station. 1941. "Investigation of filter requirements for underdrains". U.S. Army Corps of Engineers. Technical memorandum No. 183-1. Vallejo, L.E. 2001. "Interpretation of the limits in shear strength in binary granular mixtures". Canadian Geotechnical Journal, No. 38. pp. 1097-1104. Wilhelm, T. and Wilmanski, K. 2002, On the onset of flow instabilities in granular media due to porosity inhomogeneities", Int. Journal of Multiphase Flow, No. 28. pp. 1929-1944. 232 Additional bibliography Cedergren, H.R. 1989. "Seepage, drainage and flow nets", John Wiley and Sons, Inc. 3 r d edition. 466 p. Cedergren, H.R. 1972. "Seepage control in earth dams". Embankment-Dam Engineering, Hirschfiled, R. and Poulos, S. eds. Wiley, New York, pp. 21-45. Cubrinovski, M. , and Ishihara, K. 2002. "Maximum and minimum void ratio characteristics of sands". Soils and Foundations, Vol. 42, No. 6, pp. 65-78. Fannin, R.J., Vaid, Y.P., and Shi, Y.C. 1994, "Filtration behaviour of nonwoven geotextiles", Canadian Geotechnical Journal, Vol. 31, No 4, pp. '555-563. Fannin, R.J., Vaid, Y.P., and Shi, Y.C. 1994, "A critical evaluation of the Gradient Ratio test", Geotechnical Testing Journal, Vol. 17, No. 1. pp.35-42 Fannin, R.J., Vaid, Y.P., Palmeira, E M . and Shi, Y.C. 1996, "A modified Gradient Ratio device", Recent developments in geotextiles filters and prefabricated drainage geocomposites, ASTM STP 128, Editors Shobha K. Bhatia and L. David Suit. pp. 100-112. Fannin, R.J. and Pishe, R. 2001, "Testing and specifications for geotextile filters in cyclic flow applications". Geosynthetics conference 2001. pp 423-435. Fisher, G.R., Mare, A.D. and Holtz, R.D, 1999, "Influence of procedural variables on the Gradient Ratio test". Geotechnical testing J. Vol.22, pp. 22-31. Foster M. and Fell, R., " Assessing embankment dam filters that do not satisfy design criteria". J. of Geotechnical and Geoenvironmental Engineering, 2001. pp. 398-407. 233 Foster M , Robin Fell, and Matt Spannagle, 2000. "A method for assessing the relative likelihood of failure of embankment dams by piping", Canadian Geotechnical ' Journal, Vol. 37 No. 5. pp. 1025-1061. Frost, J.D. 2003, "A critical assessment of the moist tamping technique". Geotechnical Testing Journal. Vol. 26No.l. pp. 1-14. Hameiri, A. and Fannin, R.J. 2001, "On the design and commissioning of a cyclic Gradient Ratio test device", ASTM Geotechnical Testing Journal Vol. 25, No. 3. pp 1-11. Ffawley, R. 2001, "Filtration performance of geotextiles in cyclic flow conditions: A laboratory study" MASc thesis, The University of British Columbia, Vancouver, Canada. 141 p. Honjo, Y. and Venesiano D. 1988, "Dam filters: Physical behavior, probability of malfunctioning and design criterion". Takenaka Technical research report, No. 40. Honjo, Y. and Venesiano D. 1989, "Improved filter criterion for cohesionless soils". Journal of geotechnical engineering, Vol. 115, No. 1, pp. 75-94. Indraratna, B. and Vafai, F. (1997). "Analytical model for particle migration within base soi-filter system". ASCE Journal of Geotechnical and Geoenvironmental Engineering, Vol. 123, pp. 100-109. Kenney, T.C., Chacal, R., Chiu, E., Ofeogbu, G.I., Omanga, G.N, and Ume, C A 1985, "Controlling constriction sizes of granular filters", Canadian Geotechnical Journal, Vol. 22, No. 1. pp. 32-43. 234 Kohler, H.J. 1993, "The influence of hydraulic head and hydraulic gradient on the filtration process". Proceeding, Filters in Geotechnical and Hydraulic Engineering, Brauns, Heibaum & Schuler (eds). Germany, pp. 225-240. Locke , M. , Indraratna, B. and Adikari, G., 2001. "Time-dependent particle transport through granular filters". Journal of Geotechnical and Geoenvironmental Engineering, Vol. 127, No.6. pp. 521-529. Loebotsjkov, E.A. 1969. "The calculation of suffusion properties of noncohesive soils when using the non-suffosion analog" (in Russian). International Conf. on Hydraulic Research, Brno, Czechoslovakia, pp. 135-148. Nishigata, T., Fannin, R.J. and Vaid, Y.P. 2000, "Blinding and clogging of a nonwoven geotextile". Soils and Foundations, Vol. 40, No. 4. pp. 121-127. Salgado, R., Bandini, P., and Karim, A. 2000. "Shear strength and stiffness of silty sand". Journal of Geotechnical and Geoenvironmental. Engineering, Vol. 126, No.5. pp. 451-462 Shi, T. 1993. "Filtration behaviour of non-woven geotextiles in the gradient ratio test", MASc thesis, University of British Columbia, Vancouver, Canada. 127 p. Singh, B and Varshbey, R. 1995. "Engineering for embankment dams" Rotterdam, Brookfield, VT. A. A. Balkema. 732 p. Stewart, R.A. and Garner, S.J. 2000, "Performance and safety of WAC Bennett dam; a seven year update". Proceeding of the 53rd Canadian Geotechnical Conference, Montreal, Canada, pp. 97-105. Terzaghi, K. 1922. "Soil failure at barrages and its prevention" (in German). From theory tp practice in soil mechanics. Selection of writings of Karl Terzaghi. pp. 445-449. 235 Terzaghi, K., Peck, R.B and Mesri, G. 1996. Soil mechanics in engineering practice, 3 r d ed. New York, NY: John Wiley and Sons. 592 p. Thevanayagam, S. and Mohan, S (2000). "Intergranular state variables and stress-strain behaviour of silty sands".Geotechnique. Vol.50, No. 1. pp. 1-23. U.S. Department of the Interior, Bureau of Reclamation, 1987, "Design standards No. 13 - Embankment Dams", Chapter 5 - Protective Filters. Vaid, Y.P and Negussey, D. 1986, "Preparation of reconstituted sand specimens", Soil Mechanics series No. 98, Department of Civil Engineering, The University of British Columbia. Also In advanced triaxial testing in soils and rock. American Society for testing and Materials, Special Technical Publication No. 97. pp. 119-131. Vardoulakis, I. and Papanastasiou, P. 2001, "Sand erosion in axial flow conditions", Transport in Porous media, No. 45. pp. 267-281. Zaslavsky, D. and Kassiff, G. 1965. "Theoretical formulation of piping mechanism in cohesive soils". Geotechnique, Vol. 15, No. 3. pp. 305-316. 236 Appendix A: Commissioning tests Three commissioning tests were carried out to confirm the good performance of the test apparatus and of the instruments. In addition, the specimen reconstitution technique was implemented to observe the homogeneity of the specimens obtained. A.l Procedure used during commissioning tests The following steps were carried out during preliminary testing using the large permeameter: The acrylic cell is placed over the bottom plate The bottom load cell and bottom frame are placed. A wire mesh screen is placed over the frame. The permeameter is filled with water to 1 or 2 cm above the wire mesh screen Layers of glass beads are placed using the modified slurry with discrete deposition technique. When the target specimen length is achieved, the cell is filled with de-aired water and the top load plate is placed over the specimen. Top load cell is attached and top load is applied in steps to the determined target vertical stress (0-175 kPa) Total and differential piezometers are connected to the cell avoiding air bubbles inside the hoses. 237 Flow of water is started through the permeameter in a downward direction (reverse flow is also possible). Stages are run starting with an overall hydraulic gradient of 1.0 and increased in steps ofi=1.0. The test is run until considerable migration of material is observed. After completion of testing, the specimen is dismantled. A Photographic record is taken at different layers, and the data are analyzed. A.2. Test F Specimen F was a uniform stable material (see Figure A.l) . The specimen was saturated during reconstitution and had a porosity of 43% indicating a loose packing. Particle migration was not expected because the grain size distribution was uniform. Top load was applied in steps up to an effective stress of 25 kPa. Bottom load cell readings were acquired during this process. A maximum average hydraulic gradient (iav) equal to 15 was applied to the specimen. Pressure transducers readings show uniform hydraulic gradients from top to bottom (see Figure A.2). It is important to mention that the reconstitution technique yielded a homogeneous specimen. In fact, uniform distribution of total head reflects that the specimen was uniform after specimen reconstitution. During the test the minimum i a v that was adequate to obtain accurate measurements was equal to 1. 238 As expected, internal erosion was not observed. Side wall friction affected the increment of effective stress at the bottom of the specimen after an application of top load. Measured values of seepage force were also affected by the side wall friction. A.3. Test on sample F/C Specimen F/C corresponds to a filter/base configuration where C is the filter gradation and F is the base gradation (see Figure A . l ) . Gradation C was a uniform material located below the base material. Downward flow was applied and observations made regarding material moving from the base to the filter specimen. Total head distributions were initially uniform in the top portion (base material), but as expected, there was a change in the slope of the head distribution at the filter zone (see Figure A.3). No soil migration was observed for 1 < i a v < 4. At i a v = 5 soil movement was noticed at the contact between filter and base materials. At i a v = 7 an increase of fine migration was noticed. Fine-fraction particles were observed passing the wire mesh filter. At i a v = 9 small "pipes" were noticed on the wall of the permeameter. At i a v = 10 there was a catastrophic failure of the specimen and the test was stopped after exhausting the supply of water in the inlet tank. After failure of the specimen (filter incompatibility), a change in head distribution was noticed indicating that soil migration occurred at different locations on the specimen. The local hydraulic gradient at failure reached a maximum value equal to i34 = 17.8 in the zone 239 close to the interface between base and filter material (see Figure A.4). This local hydraulic gradient was the critical amount necessary to move base particles through the skeleton of the filter material. A.4 Test on sample G3 Test G3 was prepared by mixing gradations F and C. The proportion of the mixing was 40% of material F and a 60% of material C. The result was a potentially unstable gradation with a Dis'/dgs' equal to 7.5. This gradation was very similar to the gradation G4-C used by Honjo et al. (1996). Loss of material during specimen reconstitution increased the hydraulic conductivity in the first 5 cm of the specimen (see Figure A.5). Soil migration was observed from the beginning of the test with downward flow. However, this migration decreased rapidly and no migration was observed after a few minutes. Similar observations on rate of loss were made by Bertram (1940). A zone with low finer fraction content was observed at the top of the specimen at i a v = 3. The size of this zone increased slightly from i a v = 3 to 6. At i a v = 7 this zone seems to refill with fines from the adjacent areas. Catastrophic failure occurred at i a v = 8 and a local hydraulic gradient of is6 = 12.8 at the zone of failure (see Figure A. 6). The critical zone was found close to the bottom of the specimen, above 5 cm from the wire mesh screen. At this hydraulic gradient the top zone with low fine 240 content connects with a zone with low fine content on the bottom forming a "pipe" (see Figure A.7). Additionally, it was found that a zone with low fine-fraction (material A) was found at the top of the specimen indicating that migration reached up to the top of the specimen. A.5. Effective stress The effect of friction between the permeameter wall and the glass beads reduced the effective stress at the bottom of the specimen during consolidation and during the seepage stages of the tests. In Figure A.8, are shown the calculated values of effective stress at bottom versus the applied values of top effective stress for tests F and test G3. The ratio between change in bottom effective stress and the applied top stress were in between 15 to 30%. This suggests that the effective stress changes with location on the specimen and that the friction on the walls must be considered when trying to predict the value of effective stress in a specific zone of the specimen. Wall friction also influenced the values of seepage force calculated at the bottom of the specimen. In Figure A. 9 the effective stress values obtained for tests F, F/C, and G3, are shown. The values of seepage induced stresses at the bottom of the specimen were in the range between 58 to 98% of the theoretical values. The filter/ base test (F/C) gave the lowest values of measured seepage stress. This was believed to be due to the fact that most of the head losses, that created seepage force, were on top of the specimen (soil F had the lowest hydraulic conductivity value). These results proved that the local distribution of hydraulic 241 gradients was important to determinate the values of effective stress in different zone of the specimen. 242 0 100 200 300 400 500 Hydraulic head; cm HzO Figure A. 1. Water head distribution (test F) 600 700 Figure 0 100 200 300 400 500 Hydraulic head; cm HzO A.2. Water head distribution (Test F/C) -o-Final -s—Initial 600 700 243 50 45 E 40 u £ 3 5 g> 30 I <o 25 > o \% 20 <u o S 15 (A 5 10 5 0 o .O' -o--Final -a—Initial Material loss during specimen reconstitution 100 Figure 200 300 400 500 Hydraulic head; cm H 2 0 A. 3. Water head distribution (test G3) 600 700 •f 500 1000 1500 2000 Time; sec Figure A.4. Local hydraulic gradient at failure (test F/C) 2500 3000 3500 244 18 Time; sec Figure A . 5. Local hydraulic gradient at failure (test G3) Figure A . 6. Piping zone after failure (test G3) 245 ? 2 0 0. 10 <A 0) a 15 w a) > o a) 10 a> E o ts o CO 5 On • n 0 5 10 15 20 25 30 35 Top effective stress (KPa) Figure A. 8. Load induced stress at the bottom of the specimen. 40 • TestF A Test G3 45 50 246 50 45 40 a Q. £ - 3 5 m w a> £ 30 in <u 5? 25 Co a OJ a> or) in z v •a o 15 •o c — 10 a ' = 25 KPa - • - A • . ' A --n O • • • o ° A A A-;:-- o ® • Test F ® o . o O Test F/C -O o ° ATest G3 eiEr— 1 1 1 1 1 : 4 6 8 10 Overall hydraulic gradient, i 1 7 12 14 Figure A.9. Seepage induced stress calculated during commissioning tests. 247 Appendix B Variation of local hydraulic gradient with time 248 3000 Figure B . l . Onset of instability (test T-0-25-D; i a v = 11) — i12 — i23 — i34 — ,45 — i56 500 1000 2000 1500 Time; sec Figure B.2. Onset of instability (test T-0-100-D; i a v = 8); using TPT's 2500 3000 249 200 400 600 800 Time; sec Figure B.3. Onset of instability (test T-0-175-D; i a v = 45) 1000 1200 1400 140 0 500 1000 1500 2000 2500 3000 3500 Time; sec 4000 Figure B.4. Test T-5-25-D (no instability observed during testing) 250 -80 300 Time; sec Figure B.5. Onset of instability (Test T-5-175-U; i a v = 31) 80 800 Time; sec Figure B.6. Onset of instability (Test T-5-50-D; i a v = 11) 251 70 0 100 200 300 400 500 600 700 Time; sec Figure B.7. Onset of instability (Test T-5-50-D(R); i a v = 36) -20 0 500 1000 1500 2000 2500 3000 3500 4000 Time; sec Figure B.8. Onset of instability (Test T-5-25-U; i a v = 6) 252 — i23 — i34 — i45 — i57 100 200 300 Time; sec 400 500 600 Figure B.9. Onset of instability (Test T-5 30-U; i a v = 13) -70 100 200 300 Time; sec 400 500 600 Figure B.10. Onset of instability (Test C-20-50-U; i a v = 27) 253 -140 0 100 200 300 400 500 600 700 800 900 1000 Time; sec Figure B. 11. Onset of instability (Test C-20-85-U; i a v = 37) -50 -r— -45 --40 --35 1] 0 50 100 150 200 250 300 350 Time; sec Figure B.12. Onset of instability (Test C-30-25-U; i a v = 13) 254 -25 -20 g -15 O 1 -io I —i34 — i45 — i 5 6 — i67 1000 i56 145 i34 i67 2000 3000 Time; sec 4000 Figure B.13. Onset of instability (Test C-30-50-U; i a v = 14) 5000 6000 -140 -120 -100 c 03 TD tn cn o 3 ro *_ T 3 >< X 1000 2000 3000 4000 5000 6000 Time; sec 7000 8000 9000 Figure B.14. Onset of instability (Test C-30-80-U; i a v = 29) 255 Time; sec -100 Figure B.15. Onset of instability (Test C-30-100-U; i a v = 28) 256 Appendix C Soil specimen before and after failure 257 Figure C l . Specimen T-0-25-D before failure 258 Figure C.3. Soil passing at bottom of the specimen. 259 a) b) Figure C.5. Soil specimen T-0-100-D a) before failure, b) after failure Figure C.6. Specimen layers after failure a) top; b) at 33 cm. depth 260 a) b) Figure C.7. Soil specimen T-0-175-D; a) before failure, b) after failure Figure C.8. Soil layers after failure a) top b) 13 cm depth (T-0-175-D) 261 a) b) Figure C.9. Soil specimen T-5-175-U; a) before failure, b) after failure Figure CIO. Soil layers after failure a) top b) 27 cm depth (T-5-175-U) 262 a) b) Figure C . l l . Soil specimen T-5-175-U; a) before failure, b) after failure. Figure C.13. Soil layers after failure a) top b) 25 cm depth (T-5-50-D(R)) 264 a) b) Figure C.15. Soil layers after failure a) top b) 30 cm depth (T-5-30-U) Figure C.16. Soil specimen C-20-50-U; a) before failure, b) after failure. 265 Figure C. 17. Soil layers after failure a) top b) 26 cm depth (C-20-50-U) a) b) Figure C.18. Soil specimen C-20-85-U; a) before failure, b) after failure. 266 Figure C.20. Soil specimen C-30-25-U after failure. 267 Figure C.22. Soil specimen C-30-50-U; a) before failure, b) after failure. 268 Figure C.24. Soil specimen C-30-80-U; a) before failure, b) after failure. 269 a) b) Figure C.25. Soil layers after failure a) top b) 22 cm depth (C-30-80-U) 270 Appendix D: Effective stress model using FLAC A numerical model was used to understand how vertical effective stress change when the soil specimen is subjected to seepage flow. The results obtained were used to propose a simpler model that allows a reasonable and efficient calculation of vertical effective stress at any elevation within the soil specimen. Fast Lagrangian Analysis of Continua (FLAC) is an explicit, finite difference program. A set of governing equations (differential equations) is solved using variables defined in discrete points in space. The explicit method used by FLAC can be used for elements of any shape and therefore we are able to specify boundaries of any shape. This explicit method is applied to nonlinear, large-strain, physical instability problems. FLAC can handle constitutive models directly without adjustment to the solution technique. Flow modeling using FLAC is computed in parallel with the mechanical modelling. Through this method it is possible to capture the effects of fluid/solid interaction. Changes in pore water pressure cause changes in effective stress and mechanical volume change induce a change in pore water pressure. An interface was used to model the boundary between the soil and permeameter wall. This interface plane is characterized by Coulomb sliding with properties of friction, cohesion, normal and shear stiffness, and tensile strength. This interface is described through the following values. Ks = 8el0 Pa/m; Kn = 8el0 Pa/m; 8 = 25. These values are an estimation of the actual values made by trial and error, and small variation does not change the main findings from the modeling. Values of hydraulic conductivity measured during testing were used in the FLAC model. Flow values, bottom stress, effective stress at the critical hydraulic gradient are obtained and compared with linear analysis and measured values. 271 An axi-symmetrical configuration was used to model the cylindrical shape of the permeameter. A grid of 20 elements of width and 40 elements height was used. Fixed vertical and horizontal displacement were applied at the bottom of the model. Fixed horizontal displacement was applied to the center of the axis-symmetric model. The load was applied at the top of the model only in the mesh corresponding to soil. A Mohr coulomb model was used to characterize the soil. Material properties estimated by trial-and-error are: - density = 2000 kg/m3 - Bulk modulus = 4 x 108 Pa - Shear modulus = 2 x 108 Pa - Cohesion = 0 - Friction angle = 28° - Dilation = 0 - Hydraulic conductivity = variable, depending on test - Poisson ratio, p. = 0.3: The initial state was obtained through given initial distribution of total stress followed by consolidation under its own self-weight yielding the initial stresses after consolidation. Displacements obtained during this first stage are shown in Figure D . l . Earth pressure coefficients (k) obtained under this initial state are analyzed. The value of k at rest condition under linear elasticity should be equal to 0.4. This theoretical value is obtained through the following equations: 3(1-2/*) and K = 272 -E:ODE-O<5 •7.00E-M •e.ooE-06 _ -5 OOE-D(S -5 OGE-Od •-2.om.a6 •t.ooE-oe O.OOEi-00 (MBWr 5MCT74J- I f;Sr*.r-6 i I 11 ' i Figure D. 1. Initial displacement during self-weight application Figure D.2 shows values of k for three configurations. The first one only considers self-weight and no wall/permeameter friction. The second curve corresponds to a model with no wall friction and an applied top effective stress of 10 kPa. The third curve is a model with a top effective stress of 25 kPa and a wall friction angle equal to 25°. 35 30 25 cu .O E 20 3 I 15 5 10 5 0 -• B A • m A 0-2 0.25 0.3 0.35 • self-weight d=0 Est = 10 kPa d=0 Ast = 25 kPa d=25 0.4 0.45 Ko 0.5 0.55 0.6 Figure D.2. k values obtained through FLAC model 273 Comparison with laboratory testing Test T-5-50-D Loading stage: The consolidation stage was modeled with FLAC and the load measured at the bottom was compared with values obtained from numerical analysis after complete dissipation of pore water pressure. In Figure D.3 values obtained with the model are compared with values measured during test T-5-50-D. It can be observed that values obtained with FLAC are adequate to represent what happens in the laboratory. This confirms that the parameters used to model the material are adequate. 80 j 70 -60 • ra Q. 50 • i/> </> O 40 • Ul E 0 0 30 -ffi 20 • 10 -<> 0 -0 20 40 60 80 100 120 140 160 180 200 Top stress; kPa Figure D.3. Measured and calculated boundary stress values (T-5-50-D) Seepage stage: After consolidation, downward seepage flow was applied to the specimen. The hydraulic gradient was increased in stages to observe the behaviour, and the resultant force from the friction between water and the soil particles (seepage force) yielded a change in the measured load at the bottom of the specimen. These measured values are compared with the change in effective stress generated in the model when the average hydraulic gradient is increased on the soil specimen. In Figure D.4 are shown the measured and modeled values of deduced seepage force. 274 1600 1400 •ft, 1200 a 1000 a ° 800 CD D) g. 600 co cu W 400 200 0 4-A-0 4 A • • Measured E3 Theoretical A F L A C 10 20 30 Hydraulic gradient, i a v 40 50 Figure D.4. Seepage forces There is a very good agreement between the FLAC and the measured values. The theoretical curve does not consider friction at the permeameter wall, and therefore is higher than the modeled and measured curves. Theoretical values are calculated using the equation: Seepage • force = V •yw-i where: i: hydraulic gradient V: volume of the specimen yw: unit weight of water Test: T-5-25-D Loading stage: This shorter soil specimen (44.2 cm) yielded a higher ratio between the measured load at the bottom and the applied load at the top. This was consistent with the 275 observed values and takes into account the fact that a shorter specimen had less friction because the smaller perimeter area of the specimen. Using FLAC the value of the ratio between bottom and top load was approximately 51 % as is shown in Figure D.5. There was a small deviation from a straight line due to the horizontal displacement on the permeameter walls. 10 15 20 Top stress; kPa 25 30 Figure D.5. Bottom stress at bottom using FLAC Seepage stage: In Figure D.6 are shown the values of seepage force obtained through FLAC. These values are in agreement with the values measured during the test but lower than the theoretical values where friction is not considered. 276 1800 1600 _ 1400 -j ra f. 1200 H Cu 1000 S, 800 -I ra g- 600 cu W 400 200 -| 0 A-A • • A • Measured H Theoretical A F L A C 10 20 30 40 Hydraulic gradient, i a v 50 60 70 Figure D.6. Seepage forces calculated using FLAC, measured, and theoretical values. Linear approximation A simple method to estimate the distribution of vertical effective stress along the specimen during testing was used in interpretation of test data. A linear change in total vertical stress is assumed to calculate the effective stress state of the specimen (see Figure D.7). Knowing the top and bottom effective stress (using load cells) and the pore water pressure in both locations, the total vertical stress at the top and bottom was calculated. Assuming a linear variation of total stress, and knowing the pore water pressure at each level, it is possible to deduce the effective stress for each layer. This approximation does not directly consider the soil/wall friction. However, the measured value of effective stress at top and bottom (used to obtain the total stress) directly consider that friction on the walls. A comparison with FLAC model is discussed in the following paragraphs. 277 Top stress: Bottom stress Figure D.7. Simplified method to calculate effective stress Values of effective stress calculated at failure Values of effective stress were obtained using the linear approximation described above. In test T-0-175-D the effective stress at failure under a hydraulic gradient equal to i a v =45 occurred at an effective stress of 182.6 kPa at the point of failure (estimated at 9 cm from the bottom of the specimen). Using FLAC the estimated value is equal to 170 kPa, which is very close to the value obtained through this assumption (see Figure D.8). This, together with the linear variation of total stresses observed during modeling, validates the method used to calculate the vertical effective stress at different layers of the soil. Model values of vertical effective stress were slightly higher than the calculated values using a linear approximation of total vertical stress. However, the difference in between FLAC modeling and linear approximation is less than 10 %. Therefore, this approximation can be considered adequate to represent the value of the effective stress at the zone of failure. 278 Discussion and Conclusions To acquire the fundamental relation between critical hydraulic gradient and the effective stress of the layer of soil that fails under internal erosion, if is important to evaluate the effective stress in the soil specimen. It has been established that the following parameters are important to estimate the value of the vertical effective stress in the failure zone: Specimen length - Friction at the walls of the permeameter - Hydraulic gradients - Hydraulic conductivity - Top effective stress (oV) The linear approximation of total stresses together with the measured values of hydraulic gradient can be used to obtain a good approximation to the actual values of vertical effective stress. Considering this, and recognising the approximation is both reasonable and more efficient that the numerical analysis, the simplified liner approximation is used for further analysis to deduce the vertical effective stress along the specimen. These values are used the determination of a relation between critical hydraulic gradient and effective stress at failure. C*10".lbo:W Figure D.8. Effective vertical stress (FLAC) 279
Thesis/Dissertation
eng
Civil Engineering
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Experiments on the internal stability of widely graded cohesionless soils
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