UNIVERSITY OF CALGARY. The Influence of Tetrahydrofuran Hydrate Veins on Fine-Grained Soil Behaviour. William Edward Smith A THESIS

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1 UNIVERSITY OF CALGARY The Influence of Tetrahydrofuran Hydrate Veins on Fine-Grained Soil Behaviour by William Edward Smith A THESIS SUBMITTED TO THE FACULTY OF GRADUATE STUDIES IN PARTIAL FULFILMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF SCIENCE GRADUATE PROGRAM IN CIVIL ENGINEERING CALGARY, ALBERTA APRIL, 2016 William Edward Smith 2016

2 Abstract Gas hydrates are found in coarse-grained and fine-grained soil worldwide, within deepwater marine sediments and beneath permafrost. Natural gas hydrates can be formed within finegrained marine sediments as sub-vertical complex fibrous vein structures. A better understanding is required of the geomechanical behaviour of fine-grained hydrate-bearing soil that resemble fracture-hosted natural deposits, as they have the potential to pose a significant geohazard. This thesis presents a simple, repeatable laboratory procedure for the formation of simplified, vertical, cylindrical, synthetic tetrahydrofuran hydrate veins centred within fine-grained soil. The geomechanical impact of the different-sized tetrahydrofuran hydrate veins was then determined by carrying out consolidated and unconsolidated undrained compression tests on specimens. These results were then used to develop relationships between the hydrate vein size and the strength and stiffness of the fine-grained specimens. The application of these relationships to natural fine-grained sediments hosting gas hydrate veins is then discussed. ii

3 Acknowledgements I would like to express my deepest gratitude to my two co-supervisors, Dr. Jocelyn Grozic and Dr. Jeffrey Priest, who together formed a super-supervisory team that guided and supported me throughout my time at the University of Calgary. Dr. Grozic, thank you for all your enthusiastic support and for the chance to help you with tutorials for first year Statics it was the most personally rewarding experience of my time at U of C. Dr. Priest, thank you for your tireless commitment and for providing me with the once-in-a-lifetime opportunity to join you offshore India I will never forget it. I would like to acknowledge Drs. B. Jamieson, M. Maes, B. Moorman, R. Wan and R. Wong, who greatly enriched my post-graduate learning experience with their world-class courses. I would also like to thank the Civil Engineering technical staff, without whom I could not have surmounted the many interesting challenges and obstacles that presented themselves throughout my laboratory work. Special thanks to Mirsad Berbic for all his technical support. I would like to thank my fellow research-mates and friends in the department: Shmulik Pinkert who helped me begin my illustrious career in the gas hydrates laboratory, Umair Ashgar for our in-depth technical and philosophical discussions, Jithamala Caldera for all her guidance and optimism, Chee Wong for his vast technical knowledge and stimulating conversation, and Evan Wu for all his help in the lab as well as his inquisitive nature that led to innovative suggestions which proved invaluable to this research. iii

4 I am so grateful to have had supportive roommates over my short academic career who put up with my late study and work hours (Brendan, Danah, Jon 2, Dave, Pawel, Mike, Duncan and Paul). I would also like to thank all my friends from high school and Queen s based in Calgary, Ottawa and around the world, who have had to work around my student lifestyle while they pursue their successful careers. And of course, thank you so much to Rebecca for all her encouragement and patience, looking forward to our long-awaited and much-anticipated European adventure! Finally, I would like to acknowledge the two people who inspired me to attempt a brief foray into their much vaunted world of academia, Drs. Lorna J. Clark and Richard S. Smith. Without their constant love and support throughout my 24 years, this would not have been possible. And to my younger sisters Sarah, Jenny and Claire, you have all inspired me in your own way, and I can see only success in your bright futures, no matter what you choose to do. I never half-step cause I m not a half-stepper -Phife Dawg ( ) iv

5 Table of Contents Abstract... ii Acknowledgements... iii Table of Contents...v List of Tables... viii List of Figures and Illustrations...x List of Symbols, Abbreviations and Nomenclature... xvi CHAPTER ONE: INTRODUCTION Statement of Problem Research Question Objectives of Thesis Scope of Thesis Outline of Thesis...4 CHAPTER TWO: LITERATURE REVIEW Introduction to Gas Hydrates Formation and Stability Conditions Global Distribution Significance Natural Gas Hydrate Formation and Morphology Gas Availability and Migration Host Sediment and Hydrate Mode of Occurrence Laboratory Formation Techniques of Hydrate-Bearing Sediment Dissolved Gas Method Partial Water Saturation Method Hydrate Pre-mixing Method Analog Hydrate (Tetrahydrofuran) Previous Work on Geomechanical Behaviour of Hydrate-Bearing Sediment Strength Properties Consolidation Behaviour Dissociative Behaviour Summary...27 CHAPTER THREE: EXPERIMENTAL PROCEDURE Introduction Materials Fine-Grained Soil Synthetic Hydrate Specimen Preparation Soil Specimen Preparation Hydrate Vein Formation within Soil Vein Void Formation In Situ Formation Method Transfer Method Method Selection...51 v

6 3.4 Baseline Geomechanical Testing on Fine-Grained Soil Oedometer Consolidation Tests K0-Consolidation and Undrained (K0CU) Compression Tests Geomechanical Testing Apparatus Specimen Mounting and Cell Assembly K0-Consolidation Undrained Shear Geomechanical Testing on Hydrate-Bearing Soil Specimen Mounting and Cell Assembly Consolidated Undrained (CU) Triaxial Compression Testing Unconsolidated Undrained (UU) Triaxial Compression Testing...60 CHAPTER FOUR: LABORATORY RESULTS AND ANALYSIS Introduction Baseline Geomechanical Testing on Fine-Grained Soil Oedometer Consolidation Tests K0-Consolidation and Undrained (K0CU) Compression Tests Consolidated Undrained (CU) Compression Testing Isotropic Reconsolidation Results and Analysis Undrained Shear Compression Results and Analysis Issues Encountered Unconsolidated Undrained (UU) Triaxial Compression Testing Pressurization Results and Analysis Undrained Shear Compression Results and Analysis Summary...94 CHAPTER FIVE: DISCUSSION Introduction Quantifying the Geomechanical Impact of THF Hydrate Veins on Specimens Quantifying the Hydrate Veins Hydrate Vein Saturation Area Ratio Relationship between Hydrate Vein Saturation and Area Ratio Quantifying the Impact of Hydrate Veins on Sediment Strength Undrained Shear Strength Relationships Shear Strength Relationships from CU Test Results Quantifying the Impact of Hydrate Veins on Undrained Stiffness Predicting the Stiffness of a Material using Hookean Springs Undrained Stiffness versus Area Ratio Undrained Stiffness versus Hydrate Vein Saturation Discussion Theoretical Geomechanical Impact of Gas Hydrate Veins on Natural Sediment Theoretical In-Situ Strength Behaviour Theoretical In-Situ Consolidation Behaviour Theoretical In-Situ Dissociation Behaviour Summary vi

7 CHAPTER SIX: SUMMARY AND CONCLUSIONS Overview Summary of Laboratory Program Conclusions Limitations Significance and Contributions Recommendations and Future Work REFERENCES APPENDIX A: MATERIAL SPECIFICATION SHEETS APPENDIX B: OEDOMETER TEST RESULTS APPENDIX C: ANISOTROPIC CONSOLIDATION AND UNDRAINED SHEAR TEST RESULTS APPENDIX D: CONSOLIDATED UNDRAINED TRIAXIAL TEST RESULTS APPENDIX E: UNCONSOLIDATED UNDRAINED TRIAXIAL TEST RESULTS..194 vii

8 List of Tables Table 3.1: Characteristics of natural hydrate-bearing soils and prepared soil for this research Table 3.2: Data from plastic limit determination on prepared soil using ASTM D Table 3.3: Preliminary tests in the development of the THF hydrate formation procedure Table 3.4: Preliminary tests in the development of the in situ vein formation procedure Table 4.1: Summary of results from oedometer tests to 800 kpa vertical pressure on finegrained soil Table 4.2: Summary of results from undrained shear tests on anisotropically consolidated and isotropically reconsolidated fine-grained soil specimens Table 4.3: Summary of results from consolidated undrained tests on soil specimen and competent hydrate-vein-bearing specimens Table 4.4: Summary of results from consolidated undrained tests on non-competent hydratevein-bearing specimens Table 4.5: Summary of results from unconsolidated undrained tests on soil specimen and hydrate-vein-bearing specimens Table B1: Oedometer consolidation test on Preconsolidated Soil Table B2: Oedometer consolidation test on Preconsolidated Soil Table B3: Oedometer consolidation test on Preconsolidated Soil Table B4: Oedometer consolidation test on Slurried Soil Table C1: Data from anisotropic consolidation and undrained shear of specimen Table C2: Data from K0-consolidation and undrained shear of specimen Table D1: Data from CU test on specimen with no hydrate vein Table D2: Data from CU test on specimen with 0.75" diameter hydrate vein Table D3: Data from CU test on specimen with 1" diameter hydrate vein Table D4: Data from CU test on specimen with 0.25" diameter hydrate vein Table D5: Data from CU test on specimen with 0.50" diameter hydrate vein Table D6: Data from CU test on specimen with 0.50" diameter hydrate vein viii

9 Table D7: Data from CU Test on specimen with 0.75" diameter hydrate vein Table E1: Data from UU test on specimen with no hydrate vein Table E2: Data from UU test on specimen with 0.25" diameter hydrate vein Table E3: Data from UU test on specimen with 0.50" diameter hydrate vein Table E4: Data from UU test on specimen with 0.75" diameter hydrate vein Table E5: Data from UU test on specimen with 1" diameter hydrate vein Table E6: Data from UU test on specimen with 1" diameter hydrate vein ix

10 List of Figures and Illustrations Figure 2.1: Hydrate stability envelopes for onshore (a) and offshore (b) deposits, showing zones of hydrate stability based on the geothermal gradient (after Collett, 2002) Figure 2.2: Locations of sampled (purple) and inferred (red) gas hydrate occurrences in marine sediments and permafrost, with the location of some of the projects discussed in this thesis highlighted in red (after Collett et al., 2009) Figure 2.3: Schematic cross-section showing the five distinct geographic locations in which gas hydrate deposits can form, with the two most likely locations of dissociation in the near future highlighted in red (after Ruppel, 2011) Figure 2.4: Model of mass movement by slip along a dissociating hydrate glide plane, posing a potential mechanism for seafloor instability (after McIver, 1982) Figure 2.5: (a) Thin, high angle gas hydrate lenses from the Krishna-Godavari (KG) Basin; (b) Partially dissociated core from the KG Basin; (c) Massive gas hydrate nodule from the KG Basin; (d) Gas hydrate layer and nodule from the Gulf of Mexico; (e) Hydratebearing sandstone from Mount Elbert; (f) Gas hydrate in gravel from Mallik, Canada permafrost-hosted deposits (after Winters, 2011) Figure 2.6: Schematic illustration of potential fracture mechanisms: (a) Shear failure along pre-existing features due pore pressure increase, (b) Hydraulic fracturing due to increase in pore pressures, leading to zero effective stress in the horizontal stress direction and tensile failure, (c) Hydrate heave due to volume increase as hydrate forms (after Daigle & Dugan, 2010) Figure 2.7: X-ray CT images of samples from the Krishna-Godavari Basin showing pervasive hydrate veins forking and branching (white) and ice (blue) (after Rees et al., 2011) Figure 2.8: Descriptions of hydrate distribution habit using different formation techniques. The physical properties of the hydrate-bearing sediment depend on the saturation and distribution of the hydrate (black) within soil grains (gray) (Waite et al., 2009) Figure 2.9: Stress (solid) and volumetric strain (dashed) versus axial strain for four methane hydrate-bearing sands at different hydrate saturation values (indicated on the diagram in percentage) and the same effective confining stress, showing an increase in peak strength, stiffness and dilation with increasing hydrate saturation (after Masui et al., 2006) Figure 2.10: (a) Cohesion (triangles) increasing and friction angle (circles) constant with increasing hydrate saturation in natural and laboratory-formed coarse-grained hydratebearing sediment. (b) Dilation angle increase with increasing saturation (after Masui et al., 2006; Soga et al., 2006) x

11 Figure 2.11: (a) Peak strength and (b) Young's modulus at 50% of failure stress versus methane hydrate saturation for cementing and pore-filling hydrate (after Ebinuma et al., 2005; Masui et al., 2005). The offset in peak strength is due to a difference in the effective confining pressure (1 MPa versus 3 MPa) Figure 2.12: Stiffness plotted against effective confining pressure for precipitated silt and kaolinite with increasing hydrate saturation. Trends show a non-linear increase in stiffness with increasing hydrate saturation, but virtually no increase in stiffness with increasing confining stress at hydrate saturations of 50% and 100% (after Yun et al., 2007) Figure 2.13: Shear strength plotted versus initial effective stress for kaolinite (A) and precipitated silt (B) at different hydrate saturations, showing a non-linear increase in shear strength with hydrate saturation, but little increase in strength with increasing confining stress at hydrate saturations of 50% and 100% (after Yun et al., 2007) Figure 2.14: Overconsolidation (OCR) ratio versus depth for samples from the Krishna- Godavari Basin, Mahanadi Basin and Andaman Islands (NGHP-01 project), the Blake Ridge (ODP Leg 164 project) and the Cascadia Margin (IODP X311 project), indicating that results vary significantly, but that samples taken from cores in which a portion of the sediment was formerly hydrate-bearing (NGHP-01 and ODP Leg 164) exhibit a decreasing OCR with depth (after Winters, 2011) Figure 2.15: Consolidation results on samples recovered from the Ulleung Basin, including sediments taken above (2B-3H, 6B-14H, 6B-16H) and below (6C-9H) the hydrate occurrence zone, and formerly-hydrate-bearing sediments (6B-17H) compared with expected in situ effective stresses (red) calculated from results presented by the authors (after Lee et al., 2013) Figure 3.1: Flowchart summarizing the testing procedure adopted for this research program including specimen preparation, baseline testing and geomechanical testing program Figure 3.2: Grain size distribution curve of the prepared fine-grained soil compared to formerly gas-hydrate-bearing soil recovered from the KG Basin (after Clayton et al., 2008) and the Gulf of Mexico (after Winters, 2011), as well as basin averages from the KG Basin (after Winters, 2011) and Ulleung Basin (after Lee et al., 2011) Figure 3.3: Liquid limit determined from fall cone penetrometer results. The liquid limit of the soil (~34%) is defined as the water content when penetration depth is equal to 20 mm Figure 3.4: THF hydrate cylindrical vein before dissociation (a) and during dissociation (b, c, d) with veins breaking into distinct segments along planes of weakness Figure 3.5: The specially constructed consolidation cell mounted in a load frame, with the aluminium top plate connected by ram to the load cell and porous discs fitted to the top and base plate allowing for the drainage of excess pore water during consolidation xi

12 Figure 3.6: Hydraulic jack used to extrude cylindrical consolidated soil specimens from 70 mm internal diameter sampling tube (left) Figure 3.7: Vein void installation in specimen using 0.50" wood auger hooked up to drill press. Excessive specimen deformation was prevented by confining the specimen within a latex rubber membrane, stainless steel split mold and steel dummy pedestal Figure 3.8: Specimen temperature as measured throughout the vein drilling procedure, showing the initial cooling after extrusion, warming during the vein drilling process, and specimen re-cooling before hydrate formation Figure 3.9: In situ hydrate vein formation method with (a) the THF-water mixture poured into the vein void and (b) the specimen after overnight storage within the hydrate stability field Figure 3.10: Preliminary Test 6 described in Table 3.4 showing (a) ice lenses, (b) full hydrate vein formation, (c) de-structured soil after melting of ice lenses Figure 3.11: Aluminium foil mold containing a 0.25" hydrate cylinder, which proved impossible to unwrap without fracturing into segments Figure 3.12: Triaxial system showing (a) upper and lower cooling systems, (b) with double wall cells and (c) with insulation, hooked up to refrigerated circulators Figure 3.13: Schematic illustration of triaxial system showing modifications made to maintain specimen at 2⁰C, including refrigerated circulators pumping coolant through copper piping within cell fluid and below the base plate, and water reservoir containing water cooled to 1⁰C Figure 4.1: (a) Consolidation data from one oedometer test on slurry and three tests on preconsolidated soil. (b) Data from Preconsolidated Soil 1 test used to verify the preconsolidation pressure (~100 kpa) using the Casagrande Method (Casagrande, 1936).. 99 Figure 4.2: Determination of compression and recompression indices from oedometer tests on slurried soil (a) and preconsolidated soil samples (b, c and d) Figure 4.3: Effective stress paths followed during anisotropic consolidation tests showing the stress increments applied for K=0.38 and K=0.75 anisotropic consolidations, along with stress levels at which the specimen returned to its original diameter, indicating a K0 value of approximately 0.38 for the soil Figure 4.4: Void ratio versus logarithm of vertical effective stress for oedometer and K0 consolidation tests. The recompression slope during isotropic reconsolidation is greater than seen in oedometer test results, however the soil appears to be less compressible once virgin compression is initiated Figure 4.5: (a) Plot of deviatoric stress versus strain for the anisotropically consolidated and isotropically reconsolidated specimens. (b) Similar A f values are observed for the xii

13 isotropically reconsolidated (to 100 kpa) and K 0.75 specimens, with a lower value for the K 0.38 specimen Figure 4.6: (a) Effective stress paths from undrained shear tests on the isotropically reconsolidated specimen and two anisotropically consolidated specimens at the same effective confining pressure (800 kpa), along with derived critical state line. (b) Effective stress paths for undrained shear tests on similar clayey silt (75% Sil-Co-Sil silt and 25% kaolin) on isotropically reconsolidated (T5 and T8) and overconsolidated (T6 and T7) specimens, showing similar dilatant behaviour (Dayarathne and Hawlader, 2015) Figure 4.7: Plot of volumetric strain versus square root of time during isotropic reconsolidation of specimens to 100 kpa effective stress. Greater volumetric strain is observed in vein-bearing specimens, which is counterintuitive as these specimens contain less compressible soil, implying the change in volume is due to the dissolution of the THF hydrate vein in addition to soil consolidation Figure 4.8: Deviatoric stress versus axial strain for three soil specimens with two different hydrate vein diameters (0.75" and 1"). The maximum deviatoric strength is chosen as the failure criteria. Specimens display an increase in peak strength and stiffness with increasing hydrate vein diameter Figure 4.9: (a) Excess pore pressure and (b) pore pressure coefficient versus axial strain. A decrease in A f is seen with increasing vein diameter. The soil exhibits a dilatant tendency with decreasing pore pressure coefficient after peak, but since the coefficient is never negative the specimen volume does not increase from its original volume Figure 4.10: Deviatoric stress versus mean effective stress, showing the presence of hydrate veins enhances the strength and allows the soil to exceed its critical state Figure 4.11: Images of 1" (a & b) and 0.75" (c & d) diameter hydrate-vein-bearing specimens post-shear (before and after being cut open) illustrating the differences in their failure modes (blue), the remaining THF hydrate (red) and the disappearance of THF hydrate at the base of the specimens Figure 4.12: Deviatoric stress versus axial strain for hydrate-vein-bearing specimens with diameters of 0.25", 0.50" and 0.75" showing similar stiffness and similar or lower peak deviatoric stress than non-hydrate-bearing soil Figure 4.13: Post-shear images of exposed hydrate veins for hydrate-vein-bearing specimens with diameters of 0.25" (a), 0.50" (b & c) and 0.75" (d) shown outlined with colours used in stress-strain plot in Figure Figure 4.14: Stress-strain plots from unconsolidated undrained compression tests on specimens containing hydrate veins of different diameters Figure 4.15: Images of specimens cut open after compression showing different failure modes. Hydrate veins of 0.25" (a), 0.50" (b), 0.75" (c) and 1" (d & e) diameter shown xiii

14 outlined with colours used in stress-strain plot shown as Figure 4.14, and the shear band through the 1" vein (d) shown in blue Figure 5.1: Undrained shear strength from UU tests versus (a) area ratio and (b)hydrate vein saturation. The transition from soil controlled strength behaviour (red) to hydrate vein controlled behaviour (blue) is extrapolated (dashed lines) to predict a threshold value at which the two behaviours transition Figure 5.2: Vein stress (load on specimen divided by hydrate vein area) versus axial strain for horizontally fractured vein-bearing specimens. An approximately constant peak for the three different vein sizes suggests that the soil has little to no impact on the undrained shear strength in UU tests, and that their peaks represent the compressive strength of hydrate which controls the strength behaviour Figure 5.3: Deviatoric stress at failure versus (a) the area ratio and (b) hydrate vein saturation for CU and UU tests on specimens. The significant increase in deviatoric stress at failure for vein-bearing CU specimens indicates that the strength in CU tests may be influenced by the interaction between the soil and hydrate vein strength Figure 5.4: Deviatoric stress versus axial strain for different tests on specimens with ~1" diameter hydrate veins. Different hydrate vein failure modes for UU tests give rise to differences in peak strength. A much higher peak strength is measured in the CU test, which exceeds the estimated compressive strength of the THF hydrate, indicating that the isotropically reconsolidated soil provides additional strength to the specimen Figure 5.5: Mohr circles of effective stress and Mohr-Coulomb failure envelopes for a CU test on a specimen with no hydrate vein (green) and for a UU test on a specimen with a 1" diameter hydrate vein (purple), as well as a tentative failure envelope for a CU test on a specimen with 1" diameter hydrate vein (dotted red). The failure envelope for the 1" diameter hydrate vein is defined assuming no change in the friction angle but an increase in cohesion Figure 5.6: Comparison of undrained stiffness versus area ratio for (a) UU and (b) CU compression tests, showing that UU results follow the hydrate-controlled stiffness relationship after a predicted threshold ratio, while the CU results follow the parallel Hookean spring theory Figure 5.7: Comparison of undrained stiffness versus hydrate vein saturation for (a) UU and (b) CU compression tests, showing that UU results follow the hydrate-controlled stiffness relationship after a predicted threshold value while the CU results follow the parallel Hookean spring theory Figure 5.8: Schematic illustration of a layer of fine-grained marine soil containing continuous vertical gas hydrate vein networks of sufficient size to provide an increase in stiffness Figure 5.9: Theoretical consolidation behaviour of hydrate-bearing fine-grained soil before and after vein formation, resulting in the soil being at a higher metastable void ratio than would be expected at the same in situ effective stress state xiv

15 Figure 5.10: Potential void ratio change due to hydrate dissociation from its metastable state to its expected state given the effective stress conditions on the normal consolidation line (NCL), and potential further collapse to its critical state line (CSL) due to the transfer of overburden pressure from the hydrate vein network to the soil Figure A1: Specification Sheet for EPK Kaolin Figure A2: Specification Sheet for Sil Industrial Minerals Ground Silica Flour 325 Mesh Size171 xv

16 List of Symbols, Abbreviations and Nomenclature A A f A p A r(thresh) A r A soil A specimen A vein C C C 0 C c C r c u C4H8O CH4 CK0U CO2 CU E E 0.5% Pore pressure coefficient Pore pressure coefficient at failure Area of piston Threshold hydrate vein area ratio Hydrate vein area ratio Cross-sectional area of soil Cross-sectional area of specimen Cross-sectional area of hydrate vein Circumference Effective cohesion Initial circumference Compression index Recompression index Undrained shear strength Tetrahydrofuran Methane K0-consolidated undrained Carbon dioxide Consolidated undrained Young s modulus Secant Young s modulus to 0.5% Axial Strain xvi

17 E 50 E h E eq E i E sec E soil E u(soil) E u e e 0 e 1kPa e cs e f e soil e vein F F eq F max F specimen GIP H H 0 Young s modulus at 50% of failure stress Young s modulus of hydrate Equivalent Young s modulus Initial tangent Young s modulus Secant Young s modulus Young s modulus of soil Undrained stiffness of soil Undrained elastic modulus Void ratio Initial void ratio Void ratio at 1 kpa on critical state line Critical state void ratio Void ratio after each consolidation stage Void ratio of hydrate-bearing soil component Void ratio of vein void Force Equivalent force Maximum axial load on the specimen Axial load on the specimen Gas-in-place Height Initial height xvii

18 H2S JGS K K 0 K 0(NC) k k h k eq k soil KG L L eq L soil L specimen L vein LL LVDT M M g M gsl m Ma Hydrogen sulfide Japanese Geotechnical Society Stress ratio Coefficient of lateral earth pressure at rest Coefficient of lateral earth pressure at rest for normally consolidated soil Spring constant Spring constant of hydrate Equivalent spring constant Spring constant of soil Krishna-Godavari Length Equivalent length Length of soil Length of specimen Length of hydrate vein Liquid limit Linear voltage displacement transducer Slope of critical state line in q-p space Methane gas concentration Methane gas solubility limit Slope of best fit line Million years ago xviii

19 n 0 N2 r 0 OCR P p PI PL q r S h S u S u(soil) S vh(thresh) S vh THF u u 0 u a u b u c u f Initial porosity Nitrogen Initial radius Overconsolidation ratio Position Mean effective stress Plasticity index Plastic limit Deviatoric stress Radius Hydrate saturation Undrained shear strength Undrained shear strength of soil Threshold hydrate vein saturation Hydrate vein saturation Tetrahydrofuran Pore pressure Initial pore pressure Average pore pressure Pore pressure at top of sample Pore pressure at base of sample Pore pressure at failure xix

20 USCS UU V V 0 V h V T(soil) V s(soil) V s V v(soil) V v V vein ε 1 ε 3 ε a ε r ε rgauge ε v lens ε V ps ε V thaw ε V tot ε V π Unified Soil Classification System Unconsolidated undrained Volume Initial volume Hydrate volume Total soil volume of the specimen Volume of solids within host soil Volume of soil solids Volume of voids within host soil Volume of voids Vein volume Major principal strain Minor principal strain Axial strain Radial strain Radial strain measured using circumferential strain gauge Volumetric strain Volumetric strain due to hydrate structure collapse Volumetric strain due to effective stress changes involved with the depressurization production method Volumetric strain due to hydrate dissociation Total volumetric strain due to hydrate dissociation Pi xx

21 σ σ σ 1 σ 1 σ 3 σ 3 σ ch σ v σ vc σ vein σ vein(max) σ vo (σ 1 σ 3 ) max τ φ φ cs φ Total stress Effective stress Total major principal stress Effective major principal stress Total minor principal stress Effective minor principal stress Compressive strength of THF hydrate Vertical effective stress Past maximum vertical effective stress or preconsolidation pressure Vein stress Maximum vein stress Current vertical effective stress Maximum deviatoric stress Effective shear strength at failure Effective friction angle Critical state friction angle Angle of dilation xxi

22 Chapter One: Introduction 1.1 Statement of Problem Gas hydrates are naturally occurring, ice-like compounds that are stable under low-temperature and high-pressure conditions. Their molecular structure allows for the encasement of various gas molecules within a crystal lattice water structure. Methane gas hydrate deposits occur naturally in deepwater sediments along the world s outer continental margins and onshore beneath permafrost in Arctic regions. Global interest in methane gas hydrates has been generated due to its recognized potential as an unconventional natural gas resource, its potential role in climate change, and its impact as a geotechnical hazard. Marine hydrates have the potential to pose a geohazard when temperature and/or pressure conditions change, leading to hydrate dissociation. Hydrate dissociation involves the release of free gas and liquid water into the host sediment pore space at volumes several times larger than the solid hydrates. This can lead to the generation of excess pore pressure, and result in soil strength reduction and volumetric deformation. Hydrate dissociation during deep sea drilling or production can lead to hazards that include borehole instability, gas blowouts and large-scale reservoir subsidence (Nimblett et al., 2005). Gas hydrate dissociation has been suggested as a potential trigger for several historical and active submarine slope failures globally (Grozic, 2010; Vanneste et al., 2014). With expected increases in sea bottom temperatures due to climate change and increasing human activity on the seafloor, the likelihood of submarine landslides could increase, threatening offshore infrastructure (pipelines, seafloor equipment, etc.) and generating tsunami waves that threaten coastal regions (Locat and Lee, 2002). 1

23 Natural gas hydrates can be found in all sediment types, from clay to gravel. However, they are most common within fine-grained sediments, which may pose the greatest risk in terms of slope instability as their ability to dissipate excess pore fluid is low (Kayen and Lee, 1991). Methane gas hydrates form within fine-grained sediment either within the pore space in localized areas of higher pore size, or as discrete nodules, lenses and veins in areas of higher permeability caused by a local increase in grain size or faults (Waite et al., 2009). An example of an extensive finegrained hydrate-bearing deposit is within the Krishna-Godavari Basin, where hydrates are formed as grain-displacing, sub-vertical veins in complex fibrous structures (Rees et al., 2011). To date, the study of the geomechanical behaviour of hydrate-bearing coarse-grained sediments has been the emphasis within the research community due to the economic interest in this reservoir type coupled with the complexity of forming and testing gas hydrates within finegrained sediments. Studies on the geomechanical properties of natural gas hydrates within finegrained marine sediments have been attempted, however changes in temperature and pressure using conventional core recovery, storage and transfer techniques result in significant hydrate dissociation, leading to a degradation of in situ properties (Priest et al., 2014; Winters et al., 2008; Yoneda et al., 2015). More recently, pressurized transfer and triaxial systems have been developed that maintain samples at in-situ stresses and temperatures throughout the coring and testing process (Priest et al., 2015; Yoneda et al., 2013). However thus far, no geomechanical results on fine-grained hydrate-bearing samples have been published using these systems. Due to the difficulty and expense associated with testing natural samples, several experiments involving the formation and testing of laboratory analogues of natural hydrate-bearing fine- 2

24 grained specimens have been carried out (H.-S. Kim et al., 2013; Yun et al., 2007). The hydrate distribution within the host soil has been shown to directly affect the sediment s macroscopic physical properties (Waite et al., 2009), and these laboratory-formed analogues may not have resembled the grain-displacing distribution habits observed in nature, thereby limiting the applicability of results to the modelling of natural systems. A better understanding is needed of the geomechanical behaviour of fine-grained hydrate-bearing sediments that resemble the fracture-hosted deposits found in nature before and after hydrate dissociation, which is integral in assessing the submarine slope instability and production response of hydrate-bearing sediments. 1.2 Research Question The overarching research question addressed in this dissertation is: How do discrete, segregated gas hydrate structures influence the geomechanical behaviour of fine-grained sediments? 1.3 Objectives of Thesis To address the research question posed, a number of more focused objectives can be identified: Establish a simple, repeatable procedure to enable the formation of simplified hydrate vein structures within fine-grained soil that mimic naturally-occurring structures. Determine the impact of various hydrate vein sizes on the geomechanical behaviour of a specimen under different effective stress conditions. Establish relationships between the hydrate vein size and the geomechanical behaviour of the fine-grained soil in which they are hosted. 3

25 1.4 Scope of Thesis The research question will be addressed by carrying out a laboratory investigation of artificiallyformed specimens. The specimen formation procedures focus on creating simplified vertical, cylindrical, synthetic hydrate veins within a fine-grained soil matrix, to mimic hydrate structures seen in nature. This was achieved by drilling a cylinder of soil out of consolidated clayey silt specimens and forming tetrahydrofuran (THF) hydrate within this void. Given successful hydrate formation using this method, unconsolidated undrained shear tests were carried out on specimens with differing hydrate vein diameters. Consolidated undrained shear tests were also carried out to determine the effect of differing vein sizes and effective confining pressure on the geomechanical behaviour of the specimen. 1.5 Outline of Thesis It is essential to review the current knowledge of hydrate-bearing sediment in order to contextualize this research and its contributions to the understanding of the behaviour of hydratebearing fine-grained sediments. Chapter Two provides an introduction to gas hydrates in nature and discusses previous work on laboratory formation techniques and geomechanical studies, which form the basis of our current understanding. Chapter Three addresses the first objective of this thesis by presenting a simple, repeatable laboratory procedure for the formation of simplified hydrate veins within fine-grained soil, and describes the experimental methodology undertaken to investigate the geomechanical effect of hydrate veins. 4

26 Chapter Four presents the results and analysis of the baseline testing program carried out on the experimental soil. The impact of the hydrate veins on soil behaviour is then investigated by presenting and analyzing consolidated and unconsolidated undrained compression test results on hydrate-vein-bearing sediment. By determining the geomechanical response of hydrate-bearing specimens under different stress conditions, the second objective of this thesis is addressed. Chapter Five presents relationships that quantify the geomechanical impact of THF hydrate veins on sediments based on the experimental results analyzed in Chapter Four, which addresses the third objective of this thesis. The application of these relationships to natural gas hydrate systems is then discussed, focussing on fine-grained sediments. Chapter Six summarizes the thesis, provides conclusions regarding the effect of hydrate veins within fine-grained sediment, and includes recommendations for future studies based on the limitations of this experimental study. 5

27 Chapter Two: Literature Review 2.1 Introduction to Gas Hydrates Gas hydrates are crystalline compounds, in which hydrogen-bonded water molecules form a rigid open lattice that encages gas molecules of low molecular weight. Methane (CH4) is the most commonly hosted gas, with over 99.9% of natural hydrates containing methane. Other gases contained within hydrate include ethane, propane, isobutene, and non-hydrocarbons such as CO2, N2 and H2S (Kvenvolden, 1988) Formation and Stability Conditions Several factors affect the formation and stability of gas hydrates in natural and laboratory environments, including pressure, temperature, gas composition, free water volume, salinity, sediment type and the presence of catalysts/inhibitors. Methane gas hydrates form in water when the pressure and temperature conditions are conducive to stability and when the methane gas concentration in the pore fluid (M g ) exceeds the solubility limit (M gsl ), itself a function of pressure, temperature and salinity. If stability conditions are no longer met then hydrate dissociation takes place, leading to free gas and water production, resulting in a significant volume increase (Kwon et al., 2008). Dissociation occurs when pressure decreases or temperature increases, which can occur due to human activity on the seafloor (e.g. drilling, production) and environmental changes (e.g. long-term sea level and temperature changes). If the gas concentration in the pore water falls below the solubility limit (M g < M gsl ) then dissolution of the hydrate crystal occurs, involving only a small volume increase (Lu et al., 2008; Sultan et al., 2004). 6

28 The typical locations within sediment where gas hydrates are stable are shown in Figure 2.1. The base of the hydrate stability zone depends on the geothermal gradient, but is also influenced by the methane available in the pore water and its solubility. In order for hydrate to form, the gas concentration within the pore water must exceed the solubility limit (M g > M gsl ). If M g falls below M gsl, hydrate dissolution will occur until the gas concentration increases such that M g = M gsl, at which point dissolution and formation occur at the same rate (Waite et al., 2009). As a result, this phenomenon can determine the base of the hydrate occurrence. While pressure-temperature conditions can be suitable for hydrate growth at the seafloor, hydrates are typically not found on the seafloor except at active methane vents. This may be due to rising gas being consumed by hydrate formation at depth (Xu and Ruppel, 1999), chemical processes consuming the methane in shallow sediment (Egorov et al., 1999), or low methane concentrations in seawater promoting rapid hydrate dissolution (Rehder et al., 2004) Global Distribution Gas hydrates can be found globally where pressure and temperature conditions are conducive to hydrate stability, in marine sediments from the seafloor to depths more than 3 km below sea level, and within and below onshore permafrost deposits beginning at depths of approximately 130 m (Boswell and Collett, 2011). Offshore hydrates can be inferred from seismic reflectors coincident with the base of the hydrate stability zone known as bottom-simulating reflectors, which represent the boundary between hydrate-bearing and free-gas-bearing sediments (Kvenvolden, 1993). In addition, offshore drilling expeditions have led to the direct observation of gas hydrates through sampling in marine basins worldwide, including the Gulf of Mexico, the 7

29 Cascadia margin of North America, the Black Sea, the Caspian Sea, and offshore Peru, India, China, South Korea and Japan (Collett et al., 2009). Gas hydrate deposits have been found within and below Arctic permafrost in Canada and Alaska, and alpine permafrost on the Qinghai-Tibet Plateau in China (Collett et al., 2009). Figure 2.2 illustrates the known and inferred locations of gas hydrates, demonstrating their ubiquitous distribution on offshore continental slopes and within and below permafrost Significance Potential Energy Resource Since the late 1960s, methane gas hydrates have been identified as a potential energy resource due to the significant volume of gas contained in hydrate within the geosphere. Current estimates of the total gas contained within hydrate deposits vary over several orders of magnitude, depending on assumptions made about the global volume of hydrate-bearing sediment, the average degree of hydrate saturation of the pore space, and the cage occupancy of gas molecules within the hydrate lattice. Global gas-in-place (GIP) estimates range from to m 3 (Boswell and Collett, 2011). As low-saturation, regional-scale accumulations are used to calculate the total GIP, concentrated local deposits are not accounted for in this calculation despite their relative importance for resource evaluation. The most likely hydrate deposits to be commercially produced using current technologies are concentrated gas hydrates within sand reservoirs (Moridis et al., 2008). Proof-of-concept of gas production from hydrate reservoirs was actualized by field-scale production tests onshore at the Mallik site in Canada (Yamamoto and Dallimore, 2008) and offshore in the eastern Nankai 8

30 Trough (Yamamoto, 2013). Hydrate-bearing sands are amenable to production due to high reservoir permeability, which also leads to hydrate accumulations of high saturation, the ability to transmit pressure/temperature perturbations from the wellbore to induce hydrate dissociation, and the ability to allow gas flow back to the wellbore for extraction to surface (Boswell and Collett, 2011). Gas hydrates can be found in petroleum provinces that are currently being exploited, thus hydrates present an intriguing late-stage field development opportunity using existing infrastructure to improve production methods and rates (Boswell and Collett, 2011). Potential Agent in Global Climate Change Gas hydrates have the potential to supply vast quantities of methane to the atmosphere. For example, if just 0.1% of methane using a conservative estimate of the GIP was to be liberated, atmospheric methane concentrations would increase from 1774 ppb (IPCC, 2007) to around 2900 ppb (Ruppel, 2011). Methane is a more potent greenhouse gas than CO2, but it oxidizes to CO2 after a decade in the atmosphere. Models indicate that following large-scale hydrate dissociation, the long-lived CO2 oxidation product presents a greater warming potential than methane (Archer et al., 2009). Climate warming events throughout the geological record have been attributed to hydrate dissociation, such as the 600 Ma Neoproterozoic flooding of continental shelves after glaciation (Jiang et al., 2003), the 183 Ma Jurassic anoxic event (Hesselbo et al., 2000) and the Ma Paleocene-Eocene thermal maximum (Dickens et al., 1995). Shallow gas hydrate dissociation may occur in the next few hundred years based on projected warming rates of 0.2⁰C/decade (IPCC, 2007), however most of the deepwater hydrates present in large volumes (~95.5% of global volume) are expected to remain stable over the next 1000 years 9

31 due to the time expected for the warming front to reach them (Ruppel, 2011). The most sensitive sediments to warming are those located at the feather edge of the gas hydrate stability zone on upper continental slopes shown as Sector 2 in Figure 2.3 (Ruppel, 2011), and on the Arctic continental shelves shown as Sector 3 in Figure 2.3, where hydrate dissociation and subsea permafrost thawing may be occuring due to warming and inundation (Lachenbruch, 1994; MacDonald, 1990; Maslin et al., 2010). While there is a greater volume of methane within the world s upper continental slopes (~3.5%), the methane from dissociating Arctic Ocean shelf sediments (0.25%) may be more likely to enter the atmosphere rapidly as methane rather than CO2 (Ruppel, 2011). However, direct evidence that hydrate dissociation is currently contributing to elevated methane concentrations in seawater in these locations is lacking and there remains considerable uncertainty regarding the role of gas hydrate dissociation in relation to atmospheric methane concentrations (Ruppel, 2011). Potential Geohazard Gas hydrate dissociation is an endothermic process that results in the release of free gas and water into the pore space of the sediment in which it is hosted. Therefore, given sufficient heat transport to drive hydrate dissociation more rapidly than pore pressure dissipation, excess pore pressures may be generated leading to an effective stress reduction, which is a function of the sediment s hydrate saturation and permeability (Grozic and Kvalstad, 2001). Since soil strength is directly related to the effective stress, dissociation can lead to sediment instability. Gas hydrate dissociation has long been proposed as a potential mechanism for triggering and propagating submarine failures (McIver, 1982). A number of historic and active slope failures 10

32 have been suggested to be initiated or propagated in part by hydrate dissociation, for example on the continental slope of the west coast of Africa, in the fjords of British Columbia and on the Beaufort Sea continental margin (Kvenvolden, 1999). The effective stress loss due to hydrate dissociation can lead to the development of weak zones followed by slope failure, as shown in Figure 2.4. Submarine landslides can pose a risk to offshore infrastructure (e.g. seafloor equipment) and generate tsunami waves that threaten coastal regions (Locat and Lee, 2002). Several parameters affect the susceptibility of a slope to instability through hydrate dissociation. Low permeability clayey sediments may experience greater instability compared to high permeability sandy sediments, due to greater excess pore pressure development during dissociation (Kayen and Lee, 1991). Similarly, the presence of a low permeability cap layer over hydrate-rich sand layers can reduce dissipation of excess pore pressure and lead to instability at the base of the hydrate stability zone where most hydrate dissociation may take place (Xu and Germanovich, 2006). Modelling of submarine slope failures due to hydrate dissociation indicate that a pore space hydrate saturation of 5% in shallow water can lead to a sufficient reduction in effective stress to cause sediment failure (Nixon and Grozic, 2006). As previously discussed, the thickness and location of the hydrate stability zone are important in determining the slope stability of the system, for example hydrates in Arctic Ocean continental shelves and along upper continental slopes at the feather edge limit of the stability zone are most likely to dissociate in the near future, and therefore are most likely to experience slope instability. Gas leakage and blowouts, well-site subsidence and borehole collapse are other hydrate dissociation-induced geohazards (Collett and Dallimore, 2002) that can occur during oil and gas 11

33 exploration and production. These events can occur when drilling through or producing from gas hydrate deposits, which can thermally and mechanically disturb the hydrates, leading to uncontrolled gas flow or increases in formation pressure that can overcome confining stresses and lead to failure (Rutqvist and Moridis, 2010). 2.2 Natural Gas Hydrate Formation and Morphology Gas Availability and Migration Methane gas can be derived from microbial and thermogenic sources, and its availability is an important control on the location of hydrate formation (Collett et al., 2009; Kvenvolden, 1988). Microbial (biogenic) gas can be generated from the seafloor to several hundred metres below the seabed (Parkes et al., 1990), while thermogenic methane is produced under high pressure and temperature conditions more than 1 km below the seabed (Floodgate and Judd, 1992). While the majority of gas hydrate deposits are formed from biogenic gas sources, thermogenic gas sources have been proposed offshore in the Black Sea and onshore in the Mackenzie Delta and Northern Alaska (Collett, 2002). Sites with a thermogenic gas source are typically characterized by faults, seeps, diapirs and mud volcanoes (Booth et al., 1996). A combination of the two sources have been suggested in the Gulf of Mexico and Nigeria (Booth et al., 1996). The volume of biogenic gas generated locally within the sediment pore space is generally insufficient to account for the high saturations observed in hydrate deposits (Kvenvolden, 1993). Therefore, several models have been proposed for the migration of gas through the sedimentary column into the hydrate stability field: (1) Diffusion, (2) Dissolved gas in migrating water or (3) As a bubble/continuous gas phase. The diffusion of gas is a relatively slow process, and may not 12

34 result in concentrated hydrate accumulations (Xu and Ruppel, 1999). The last two models require permeable pathways through which the fluid can migrate, such as along fault systems or permeable sediment layers (Collett et al., 2013). Jain and Juanes (2009) relate grain size to gas transportation mechanism using a coupled fluid flow and geomechanics model, concluding that capillary invasion is favoured in coarse-grained sediments and fracturing dominates fine-grained sediments. Therefore the migration mechanisms imply that hydrates tend to form veins within fracture networks in fine-grained sediment and pore-filling deposits in coarse-grained sediment Host Sediment and Hydrate Mode of Occurrence Methane hydrates have been observed within both coarse and fine-grained sediment. The morphology of gas hydrates observed in field studies suggest that a correlation exists between grain size and mode of hydrate occurrence (Booth et al., 1996). Methane hydrate can be observed in core samples disseminated relatively homogeneously within the pore space of coarse-grained sediments, and inhomogeneously distributed within fine-grained sediment as nodules, sheets, lenses, and veins (Waite et al., 2009) as shown in Figure 2.5. Coarse-Grained Sediment Methane hydrates in coarse-grained sediments have been identified on the North Slope of Alaska, the Nankai Trough offshore Japan, and the Mallik permafrost site in Canada, disseminated within the sediment pore space (Dallimore and Collett, 2005; Fujii et al., 2009; Park et al., 2008; Yamamoto, 2013). This morphology arises as the higher sediment permeability allows for gas migration while the lower capillary pressures due to the larger pore sizes allow for hydrate nucleation (Torres et al., 2008). Hydrate forms within the pore space in three distribution 13

35 habits: (1) At low saturation the hydrate is pore filling, in which hydrates nucleate within the pore space without bonding particles (Helgerud et al., 1999); (2) At 25-40% hydrate saturation this becomes load bearing, where hydrates strengthen the soil skeleton by becoming part of the load-bearing framework (Yun et al., 2007); (3) Cementation occurs at low hydrate saturations when a small amount of hydrate forms at particle contacts, bridging particles together thereby dramatically increasing the strength (Dvorkin et al., 1999). It has been shown that hydrates exhibit a cementing behaviour when formed in the presence of excess gas, while pore-filling and load-bearing habits occur when hydrates precipitate from dissolved aqueous gas. Most hydrate within coarse-grained sediment is likely characterized by pore-filling/load-bearing models (Buffett and Zatsepina, 2000). Fine-Grained Sediment Fine-grained hydrate-bearing sediments have been observed in the Blake Ridge offshore the western U.S., the Gulf of Mexico, offshore Taiwan, Hydrate Ridge offshore western Canada, the Krishna-Godavari Basin offshore India and the Ulleung Basin offshore Korea (Winters, 2011). Hydrates within fine-grained sediment such as clays commonly exhibit a grain displacing morphology, and can exist as discrete nodules, planar fracture-filling, layered deposits or complex vein structures (Cook et al., 2008; G. Y. Kim et al., 2013; Rees et al., 2011; Tréhu et al., 2004; Winters, 2011). This morphology occurs due to the high capillary pressures within clays, inhibiting hydrate nucleation in the interstitial pore space between particles (Torres et al., 2008). While hydrate saturations at a local scale (e.g. segregated within veins) can be up to 100% 14

36 (Winters et al., 2008) and as high as 85% at the sample scale (e.g. as vein structures) (Winters, 2011), hydrate saturations on a broader regional scale are typically much lower. The fractures in which hydrates precipitate can be generated by three mechanisms illustrated in Figure 2.6: (1) Hydraulic fracturing due to high overpressures generated by free gas or pore fluid below the hydrate stability zone (Flemings et al., 2003; Jain and Juanes, 2009; Weinberger and Brown, 2006); (2) Shear failure along pre-existing soil features driven by pore fluid pressure (Hornbach et al., 2004); (3) Heave due to the volume increase associated with the formation of hydrate crystals, forcing sediment grains apart (Daigle and Dugan, 2010). Hydrate formation is theorized to occur during or after sediment fracturing, but this is poorly understood. Hydrates have also been found in the pore-filling habit within fine-grained sediments in localized areas of comparatively higher permeability and pore size, for example within layers of silt, silty sand or diatoms (Bahk et al., 2013; Ginsburg et al., 2000). Historically, numerous scientific expeditions have identified disseminated hydrates in finegrained sediments, generally in samples with hydrate saturations of less than 10% (Waite et al., 2009). However, the descriptor disseminated is used for core description where the hydrate is invisible to the naked eye, and so is not necessarily equivalent to pore-filling as the hydrate may have already dissociated before core inspection (Holland et al., 2008). Hydrate dissociation within fine-grained samples results in the destruction of the soil fabric and obscures the original morphology, so the concept of disseminated hydrates within fine-grained sediment may be a result of the difficulties in recovering and observing intact sediment. It has been suggested that 15

37 once dissociated, thin veins of hydrate in fine-grained sediments might be classified as disseminated in the absence of pressure core imaging (Holland et al., 2008). Until recently, samples recovered from the Krishna-Godavari Basin have provided the bestdocumented example of a fine-grained reservoir in which hydrates are fracture-hosted. The hydrate deposits in this basin form discrete, grain-displacing, sub-vertical veins with no evidence of disseminated pore-filling hydrate (Rees et al., 2011). X-ray CT scanning of samples revealed heterogeneous sub-vertical veins dipping at 50-80⁰ from the horizontal in a complex structure that forks and branches as shown in Figure 2.7, with an average hydrate saturation of 20-30% and some portions as high as 60% (Rees et al., 2011). The fibrous nature of hydrate distribution was suggested to be due to the infill of hydraulic fractures, leading to the development of up to centimetre-thick veins over time (Rees et al., 2011). The hydrate-bearing sediments are typically high plasticity clays that exhibited lower shear strength than would be expected given the in-situ vertical effective stress (Priest et al., 2014; Winters, 2011). 2.3 Laboratory Formation Techniques of Hydrate-Bearing Sediment Forming methane hydrate is time-consuming due to the low solubility of methane in water. Several laboratory methods were developed that balance ease of formation with creating a hydrate distribution resembling natural specimens. These methods were typically perfected by forming hydrate within coarse-grained sediment, and then extended to fine-grained soil. Laboratory studies that resulted in the successful formation of hydrates within fine-grained soil are highlighted. 16

38 2.3.1 Dissolved Gas Method The dissolved gas method involves circulating water containing hydrate-forming gas through a specimen held within the hydrate stability field. Since the gas solubility limits the hydrate saturation and affects the formation time, CO2 is often used due to its higher solubility (Katsuki et al., 2006). The dissolved gas method is typically limited to forming hydrate saturations below 60-70% (Waite et al., 2009), and when formed within coarse-grained soil results in heterogeneous hydrate nucleation on soil grains and growth into the pore space as shown in Figure 2.8. The dissolved gas method is therefore effective in mimicking the pore-filling and load-bearing distribution habit of natural hydrate-bearing coarse-grained soil, but the long formation time and low maximum hydrate saturation are significant drawbacks. Grozic and Kvalstad (2007) formed hydrate using the dissolved gas method within kaolin clay, which is of low sensitivity and relatively high permeability. A maximum estimated hydrate saturation of 7.7% of the pore space was attained after keeping the specimen within the stability range for 39 days Partial Water Saturation Method The partial water saturation method involves mixing soil with water to form a partially saturated specimen, pressurizing the system with methane gas, and then cooling the sample into the hydrate stability field to form hydrate. A variant of this method involves forming a saturated specimen and then introducing methane gas as a bubble phase before cooling into the stability field (Winters et al., 2002). Within coarse-grained soil, the partial water saturation method leads to a cementing habit due to hydrate formation at grain contacts (shown in Figure 2.8), bridging 17

39 sand grains at relatively low hydrate saturations and leading to a stiffer sediment skeleton than the pore filling habit (Priest et al., 2005). This hydrate distribution habit is limited to deposits formed in high gas flux areas (Bohrmann et al., 1998) or where gas is recycled into the hydrate stability zone (Guerin et al., 1999). An experimental study was successful in forming CO2 hydrates in partially saturated, remoulded clayey silt sediments from the Ulleung Basin at hydrate saturations of 28%, 47% and 63% (H.-S. Kim et al., 2013). The specimens exhibited what the authors termed weak cementation, a behaviour transitional between load-bearing and grain-cementing models, postulated to be due to weak bonding between hydrate crystals and clay mineral grains due to the presence of water film on mineral surfaces. However, the hydrate morphology within the sediment was not determined Hydrate Pre-mixing Method Hydrate can be formed as granules by spraying misted water in a pure methane gas atmosphere (Hyodo et al., 2005), or melting ice in the presence of methane under hydrate stability conditions (Stern et al., 1998). Gas hydrate granules can then be combined with soil at low temperature and consolidated to the target effective stress. In coarse-grained soils, the load-bearing contribution depends on the relative size of the soil grains and hydrate granules, shown in Figure 2.8. Li et al. (2011) created a methane hydrate-ice mixture with a hydrate-to-ice ratio of 3:7, and then mixed it with kaolin at atmospheric pressure and at -10⁰C, before compacting the mixture into a cylindrical specimen at 10MPa for geomechanical testing. 18

40 2.3.4 Analog Hydrate (Tetrahydrofuran) Tetrahydrofuran (THF) (C4H8O) is a hydrate former that is completely miscible in water, allowing for rapid hydrate synthesis and precise control of hydrate saturation within sediments (Lee et al., 2007). THF hydrate is formed by mixing THF with water at atmospheric pressure and temperatures below 4⁰C, greatly simplifying hydrate formation. THF molecules are polar while methane is non-polar, possibly altering the hydrate behaviour in the presence of polar water molecules; however the large THF molecule may weaken the ionic interaction between THF and water molecules, such that despite their chemical differences they are mechanically similar (Lee et al., 2007). THF hydrate nucleates on mineral grains and grows into the pore space similar to the dissolved gas method (Waite et al., 2009). As THF hydrate does not dissociate into free gas, the volume change due to dissociation is much less significant than for gas hydrates. Yun et al. (2007) formed THF hydrates within silt and kaolin clay at hydrate saturations of 50% and 100%. This was achieved by mixing dry soil with a THF-water solution to form a saturated paste, consolidating the specimen to the target effective stress, and then freezing. Several other studies following this procedure have also been carried out (e.g. Lee et al., 2010; Santamarina and Ruppel, 2010). However, hydrate distribution habits were not determined during these investigations, and it is therefore not known whether they resemble natural, segregated hydrate deposits within fine-grained sediment. 19

41 2.4 Previous Work on Geomechanical Behaviour of Hydrate-Bearing Sediment Strength Properties Introduction and General Trends The strength of hydrate-bearing sediments is commonly evaluated within the Mohr-Coulomb framework, where the effective shear strength at failure τ is: τ = C + σ tan φ (2.1) Where σ is the normal effective stress and C and φ are the effective cohesion and friction angle respectively. The cohesion is the cohesive resistance, and the friction angle includes resistance to interparticle sliding, rearrangement and crushing. These parameters can be determined using results from undrained and/or drained triaxial compression tests under different effective stress conditions to define the Mohr-Coulomb failure envelope. Triaxial tests can also be used to determine the soil stiffness (approximated by the Young s modulus, E), and the volume change during shear deformation, defined as the dilatancy and characterized by the angle of dilation (φ). Gas hydrates are stronger and stiffer than the soil in which they form, and their presence has been shown to increase sediment stiffness, enhance prefailure dilation and lead to higher shear strength. The strength of hydrate-bearing sediments has been found to be a function of the strain rate (Winters et al., 2004), confining pressure (Miyazaki et al., 2011b; Yun et al., 2007), temperature (Hyodo et al., 2005, 2002; Li et al., 2012), nature of pore fluid (Hyodo et al., 2013a), density and grain size of soil particles (Yun et al., 2007), hydrate formation habit (Priest et al., 2009) and degree of hydrate saturation (Ghiassian and Grozic, 2013; Hyodo et al., 2013a; Miyazaki et al., 2011a; Yun et al., 2007). 20

42 Previous Work on Hydrate-Bearing Coarse-Grained Sediments Numerous studies have been carried out on laboratory-formed and naturally-occurring hydratebearing coarse-grained sediments, including drained and undrained triaxial testing. The geomechanical impact of hydrate on coarse-grained sediments is better understood as they have been studied more thoroughly than their fine-grained counterparts. Generally, it has been shown that peak strength, stiffness, dilation, and strain softening after peak strength increase with increasing hydrate saturation as shown in Figure 2.9 (Masui et al., 2006; Miyazaki et al., 2011a). The strength increase is related to an increase in apparent cohesion, while the friction angle remains relatively constant, as shown in Figure 2.10 (Masui et al., 2006; Soga et al., 2006). However, the peak strength has been shown to increase with increasing effective confining stress, suggesting that there is a frictional contribution from the hydrate rather than being solely due to particle cementation (Hyodo et al., 2013a). At 100% hydrate saturation, the specimen strength and stiffness are dominated by the hydrate properties rather than the effective stress (Yun et al., 2007). The relationship between the hydrate saturation and coarse-grained soil behaviour is also affected by the hydrate distribution within the pore space. When the hydrate is located at grain contacts (cementing), a low hydrate saturation can lead to a pronounced increase in strength, stiffness and dilation angle, while pore-filling hydrates only affect the response when the saturation exceeds 30% and the hydrate becomes load-bearing (Ebinuma et al., 2005; Masui et al., 2005), shown in Figure However, the effect of the distribution habit decreases with increasing hydrate saturation. 21

43 Previous Work on Hydrate-Bearing Fine-Grained Sediments The influence of hydrate on the mechanical behaviour of fine-grained sediment is not as well understood due to the difficulty of forming hydrate in these sediments. Strength data are limited to a few laboratory studies on THF hydrate in kaolin clay (Li et al., 2012, 2011; Yun et al., 2007; Zhang et al., 2015), testing on recovered natural samples (Yun et al., 2006), and in-situ tests on sediment (Sultan et al., 2007). Undrained triaxial testing was carried out on THF-hydrate-bearing kaolin clay and silt formed by mixing dry soil with a THF-water solution (previously outlined in Section 2.3.4), the results of which are shown in Figure 2.12 and Figure 2.13 (Yun et al., 2007). Peak strength and stiffness were seen to increase non-linearly with increasing hydrate saturation. The stress dependency of the strength and stiffness reduced as hydrate saturation increased, and at 100% saturation the sediment behaviour was stress-independent. A more gradual stiffness degradation occurred with increasing strain for hydrate-bearing kaolinite specimens when compared to coarser grained samples, following the hyperbolic-type stress-strain model put forward by Duncan and Chang (1970). It was postulated that this was due to weak bonding between the clay and hydrate. Zhang et al. (2015) formed THF hydrates within silty clay using the same method at typical in-situ saturation values (5%, 10%, 15%, 25%, 35%, and 45%), noting an increase in cohesion and a linear increase in peak strength with increasing hydrate saturation. However, the hydrate distribution habit was not determined in these studies, and may not have greatly resembled heterogeneous fracture-dominated morphologies observed in nature, potentially limiting the applicability of the geomechanical relationships. 22

44 Undrained shear strength testing using a cone penetrometer under zero effective stress was carried out on fine-grained natural sediments recovered using pressure cores from the Gulf of Mexico, showing higher undrained strength in hydrate-bearing sediments than non-hydratebearing sediments from a similar depth (Yun et al., 2006). The undrained shear strength of hydrate-bearing clayey-sand specimens with less than 30% hydrate saturation from the Nankai Trough were found to be controlled primarily by effective stress (Santamarina et al., 2015). Cone resistance measurements using piezocones on in-situ shallow hydrate-bearing sediments offshore Nigeria showed increased strength (Sultan et al., 2007). As of yet, triaxial tests on natural hydrate-bearing sediments have not been performed due to the difficulty in preserving samples. For example, Yoneda et al. (2015) attempted triaxial compression tests of hydrate-bearing clayey-silty samples (5-30% estimated in situ hydrate saturation) after rapid depressurization from pressure cores, but by the time the samples were tested the hydrate had disappeared Consolidation Behaviour Introduction and Expected Trends The deposition of marine sediment on the seafloor leads to an increase in the vertical effective stress (σ v) on the previously-deposited sediment, which is generally assumed to result in onedimensional soil consolidation as lateral strains are prevented by the surrounding soil. The volume change during consolidation due to pore water expulsion and soil particle rearrangement can be expressed by a change in the soil s void ratio ( e), and the soil s volumetric response to effective stress is defined by the soil s compression index (C c ) which is the slope of the normal compression line on a void ratio versus log vertical effective stress (σ v) plot. If the vertical effective stress on the sediment is reduced by erosion, the soil s volume will increase, allowing 23

45 vertical strain to be recovered. The slope of the swelling line in e versus logσ v space is termed the recompression index (C r ), representing the elastic response of the soil (as the strain was recovered). If deposition recommences then a soil s volume change with effective stress will follow the recompression slope until it reaches its past maximum vertical effective stress (or preconsolidation pressure) (σ vc), after which it continuous along the normal compression line. The degree of consolidation of a soil is defined by the overconsolidation ratio (OCR), which is the ratio of the current effective stress (σ vo) to the preconsolidation stress. Previous Work on Formerly Hydrate-Bearing Sediment At present, no known studies have been conducted to investigate the consolidation behaviour of hydrate-bearing soils. However, results from studies on natural formerly hydrate-bearing sediment can be used to draw conclusions about the consolidation behaviour. Consolidation tests on fine-grained soil samples taken from the Krisha-Godavari Basin, Mahanadi Basin, and the Blake Ridge indicated that the sediment was overconsolidated (OCR>1) near the seafloor and the OCR decreased with depth, with deeper, often formerly hydrate-bearing fine-grained sediments found to be underconsolidated (OCR<1) as shown in Figure 2.14 (Winters, 2011, 2000). This was postulated to be due to either rapid sedimentation preventing dissipation of excess pore pressures, or inherent physical sediment characteristics (Winters, 2011). Formerly hydrate-bearing sediments from the Ulleung Basin were found to have a high compressibility and initial void ratio in spite of high in-situ effective overburden pressures. This 24

46 was suggested by some to be due to hydrate dissociation of the tested samples, leading to volume expansion and structural sediment change due to interactions involving the clay s electronic double layer and the pore fluid s changing ionic concentration (Kwon et al., 2011). However, consolidation test results shown in Figure 2.15 demonstrate that both formerly hydrate-bearing soil, and hydrate-free samples taken directly above and below hydrate occurrence zones all displayed high compressibility. Therefore, it has been suggested more recently that this is not due to hydrate dissociation but rather because of physical sediment characteristics such as a welldeveloped pore structure and the presence of montmorillonite (Lee et al., 2013). However, using the results reported by Lee et al. (2013), an approximation of the in situ vertical effective stress was calculated for each of the samples and compared to the preconsolidation pressure determined from consolidation test results. As seen in Figure 2.15, calculated in situ effective vertical stress are slightly higher than the preconsolidation pressures, indicating that sediments may have been underconsolidated for the depth at which they were found. It has been suggested that the underconsolidation observed in reservoirs with fracture-hosted, hydrate-bearing fine-grained sediments may be partially due to the enhancement of the sediment stiffness by hydrate vein networks, preventing full consolidation of the host sediment (Priest et al., 2014). However, this hypothesis has not yet been confirmed by experimental studies. 25

47 2.4.3 Dissociative Behaviour During Dissociation Hydrate dissociation results immediately in a volume expansion due to the release of free gas and water, and depending on the drainage conditions can cause an increase in the pore pressure resulting in a decrease in effective stress. Estimating the excess pore pressure development is highly complex and time-dependent, depending on the both the sediment permeability and the rate of gas hydrate dissociation, which is a function of hydrate saturation and temperature/pressure transfer rates through sediments (Nixon and Grozic, 2007). After Dissociation Subsequent to the dissipation of excess pore pressures, hydrate dissociation will decrease the strength and stiffness of the soil, due to the disappearance of the hydrate (Lee et al., 2013). Hydrate dissociation has also been shown to lead to a loss of volume within the sediment under drained, zero lateral strain conditions regardless of sediment type, effective stress level and hydrate saturation (Lee et al., 2010). Volume loss mechanisms include: bulk hydrate dissociation, sediment skeleton alteration and consolidation (Lee et al., 2010). The volume change may cause borehole settlement and large-scale subsidence of reservoirs post-production. The magnitude of contraction may depend on hydrate formation history, the in situ stress, hydrate distribution and saturation, sediment porosity, and sediment grain size (Lee et al., 2010). Lee et al. (2010) proposed equations for total volumetric strain (ε V tot ) when small shear strains are expected (e.g. around production wells on level ground), dependent on the volumetric strain 26

48 due to hydrate dissociation (ε V thaw ), hydrate structure collapse (ε V lens ) and effective stress changes due to the depressurization production method (ε V ps ): ε V tot = ε V thaw + ε V lens + ε V ps (2.2) If large shear strains are anticipated after dissociation (e.g. slope instability), the authors suggested employing the critical state model to determine the total volumetric strain: ε tot V = e cs e 0 = (e 1kPa λ log 1kPa ) e 0 (2.3) 1 + e e 0 Where e cs is the critical state void ratio after dissociation, e 0 is the initial void ratio, e 1kPa and λ are critical state parameters and p is the final mean effective stress after dissociation. p Dissociation both by depressurization and heating was carried out on methane hydrate-bearing sand specimens under some initial shear stress in a triaxial cell, leading to significant axial deformation (Hyodo et al., 2013b). However, when sufficient axial load was applied such that the samples were consolidated to the metastable zone between the failure envelopes for pure and methane-hydrate bearing sand, the sediment experienced collapse to the critical state line. 2.5 Summary Gas hydrates are solid ice-like compounds that form at low temperature and high pressure conditions in the presence of excess gas (most commonly methane). Methane gas hydrates are distributed ubiquitously worldwide where stability conditions are met, along offshore continental slopes and within and below high latitude and alpine permafrost. Interest has been generated in gas hydrates for three reasons: its potential as an energy source, its potential role in global climate change and as a potential geohazard. 27

49 Methane hydrate can be formed in both fine-grained and coarse-grained sediments under hydrate stability conditions when gas and free water are available. Hydrate is commonly observed disseminated relatively homogenously within the pore space of coarse-grained sediments, and heterogeneously distributed within fine-grained sediment in segregated deposits such as nodules, sheets, lenses, and complex interconnected sub-vertical vein structures. The distribution habit of the hydrate is important, as it affects the physical properties of the sediment. Gas hydrates are stronger and stiffer than the soil in which they form, and their presence has been shown to increase sediment stiffness, enhance prefailure dilation and lead to higher shear strength. Significant geomechanical investigations have been carried out on coarse-grained sediments, where it has been found that the strength of hydrate-bearing coarse-grained sediments is a function of the strain rate, confining pressure, temperature, nature of pore fluid, grain size, degree of hydrate saturation and hydrate formation habit. Different laboratory hydrate formation techniques have been adopted for coarse-grained soils, which led to different hydrate distributions (pore-filling, load-bearing, cementing) resulting in differing geomechanical behaviour. Large strain shear testing of hydrate-bearing fine-grained sediments has only been conducted on specimens formed using the THF hydrate formation method, where results indicated a non-linear increase in peak strength and stiffness due to a reduction in the stress dependency with increasing hydrate saturation (0%, 50%, 100%) (Yun et al., 2007). However, Zhang et al. (2015) used the same formation method and noted an increase in cohesion and linear increase in peak strength with increasing hydrate saturation at typical in situ values (5%, 10%, 15%, 25%, 35%, 28

50 and 45%). Researchers have compared geomechanical results on hydrate-bearing fine-grained soils to coarse-grained soils, suggesting that weaker bonding may exist between hydrate and clay minerals than granular material. However, the distribution habit of the THF hydrate within the fine-grained sediments was not determined in these studies and the applicability of results to the behaviour of natural hydrate-bearing fine-grained sediment is unclear. No published study has directly observed and characterized the geomechanical effects of gas hydrate distributed within fine-grained sediment as it is commonly observed in nature. Therefore, a greater understanding is needed of the mechanical effect that segregated gas hydrate structures (i.e. in sub-vertical veins) may have on the sediment in which they are hosted. The objectives of this thesis, to study the geomechanical response of fine-grained soils containing vertical hydrate veins, aim to begin to fill this gap in our current understanding of this important deposit. 29

51 A B Figure 2.1: Hydrate stability envelopes for onshore (a) and offshore (b) deposits, showing zones of hydrate stability based on the geothermal gradient (after Collett, 2002). 30

52 Figure 2.2: Locations of sampled (purple) and inferred (red) gas hydrate occurrences in marine sediments and permafrost, with the location of some of the projects discussed in this thesis highlighted in red (after Collett et al., 2009). 31

53 Figure 2.3: Schematic cross-section showing the five distinct geographic locations in which gas hydrate deposits can form, with the two most likely locations of dissociation in the near future highlighted in red (after Ruppel, 2011). 32

54 Figure 2.4: Model of mass movement by slip along a dissociating hydrate glide plane, posing a potential mechanism for seafloor instability (after McIver, 1982). 33

55 Figure 2.5: (a) Thin, high angle gas hydrate lenses from the Krishna-Godavari (KG) Basin; (b) Partially dissociated core from the KG Basin; (c) Massive gas hydrate nodule from the KG Basin; (d) Gas hydrate layer and nodule from the Gulf of Mexico; (e) Hydrate-bearing sandstone from Mount Elbert; (f) Gas hydrate in gravel from Mallik, Canada permafrost-hosted deposits (after Winters, 2011). 34

56 A B C Figure 2.6: Schematic illustration of potential fracture mechanisms: (a) Shear failure along preexisting features due pore pressure increase, (b) Hydraulic fracturing due to increase in pore pressures, leading to zero effective stress in the horizontal stress direction and tensile failure, (c) Hydrate heave due to volume increase as hydrate forms (after Daigle & Dugan, 2010). A B C D Figure 2.7: X-ray CT images of samples from the Krishna-Godavari Basin showing pervasive hydrate veins forking and branching (white) and ice (blue) (after Rees et al., 2011). 35

57 A B C Figure 2.8: Descriptions of hydrate distribution habit using different formation techniques. The physical properties of the hydrate-bearing sediment depend on the saturation and distribution of the hydrate (black) within soil grains (gray) (Waite et al., 2009). Figure 2.9: Stress (solid) and volumetric strain (dashed) versus axial strain for four methane hydrate-bearing sands at different hydrate saturation values (indicated on the diagram in percentage) and the same effective confining stress, showing an increase in peak strength, stiffness and dilation with increasing hydrate saturation (after Masui et al., 2006). 36

58 A B Figure 2.10: (a) Cohesion (triangles) increasing and friction angle (circles) constant with increasing hydrate saturation in natural and laboratory-formed coarse-grained hydrate-bearing sediment. (b) Dilation angle increase with increasing saturation (after Masui et al., 2006; Soga et al., 2006). 37

59 A B Figure 2.11: (a) Peak strength and (b) Young's modulus at 50% of failure stress versus methane hydrate saturation for cementing and pore-filling hydrate (after Ebinuma et al., 2005; Masui et al., 2005). The offset in peak strength is due to a difference in the effective confining pressure (1 MPa versus 3 MPa). 38

60 A B Figure 2.12: Stiffness plotted against effective confining pressure for precipitated silt and kaolinite with increasing hydrate saturation. Trends show a non-linear increase in stiffness with increasing hydrate saturation, but virtually no increase in stiffness with increasing confining stress at hydrate saturations of 50% and 100% (after Yun et al., 2007). A B Figure 2.13: Shear strength plotted versus initial effective stress for kaolinite (A) and precipitated silt (B) at different hydrate saturations, showing a non-linear increase in shear strength with hydrate saturation, but little increase in strength with increasing confining stress at hydrate saturations of 50% and 100% (after Yun et al., 2007). 39

61 Figure 2.14: Overconsolidation (OCR) ratio versus depth for samples from the Krishna-Godavari Basin, Mahanadi Basin and Andaman Islands (NGHP-01 project), the Blake Ridge (ODP Leg 164 project) and the Cascadia Margin (IODP X311 project), indicating that results vary significantly, but that samples taken from cores in which a portion of the sediment was formerly hydrate-bearing (NGHP-01 and ODP Leg 164) exhibit a decreasing OCR with depth (after Winters, 2011). 40

62 A B C Figure 2.15: Consolidation results on samples recovered from the Ulleung Basin, including sediments taken above (2B-3H, 6B-14H, 6B-16H) and below (6C-9H) the hydrate occurrence zone, and formerly-hydrate-bearing sediments (6B-17H) compared with expected in situ effective stresses (red) calculated from results presented by the authors (after Lee et al., 2013). 41

63 Chapter Three: Experimental Procedure 3.1 Introduction As highlighted in the literature review, gas hydrates form naturally as fracture-filling veins within fine-grained soil and may increase the strength and stiffness of the host sediment. To further investigate this potential behaviour, a series of laboratory tests were carried out. This chapter highlights the laboratory procedures designed to test this hypothesis and answer the thesis objectives presented in Section 1.3. Natural hydrate veins exhibit complex geometries that make them difficult to replicate. Therefore, a simplified formation process was developed to mimic natural veins by forming vertical cylinders of synthetic hydrate centred within pre-consolidated clayey silt specimens. The veins were aligned with the principal stress direction in a triaxial test to replicate natural nearvertical structures. Cylindrical veins were chosen due to the difficulty of creating thin, planar veins typically observed in nature, while also creating a specimen that responds to radial stress (σ 3 ) axisymmetrically (ε 2 = ε 3 = ε r ), simplifying mechanical analysis. Clayey silt was consolidated to a vertical stress (100 kpa) to mimic stress conditions on near-seafloor sediment that are likely to pose the greatest risk to slope instability. A triaxial test apparatus was used to investigate the geomechanical behaviour of the hydrate-bearing specimens, due to its versatility and simplicity. Prepared specimens were subjected to different effective stress conditions prior to shearing specimens in undrained compression. This chapter details the characteristics of the experimental soil and hydrate, the hydrate vein formation procedure, and descriptions of the apparatus and testing methods used to investigate the influence of hydrate veins on soil strength. A flow chart summarizing the testing program outlined in this chapter is shown in Figure

64 3.2 Materials Fine-Grained Soil The soil used in this laboratory investigation is intended to resemble typical natural fine-grained marine soils within which hydrates are hosted, because as discussed previously, soil properties have a significant impact upon the distribution habit of natural gas hydrates. Table 3.1 highlights characteristics of natural fine-grained soils recovered from marine drilling expeditions in several locations worldwide where hydrate was observed. In the Krishna-Godavari (KG) Basin, the host soil was of high plasticity, and found to range from silty clay to clayey silt based on grain size. Soil in the Ulleung Basin was found to be silt to clayey silt of medium to high plasticity, while samples from the Northern Gulf of Mexico contained a higher clay fraction. Given the variability of silt and clay content in natural soils, laboratory prepared soil was chosen to be a mixture by weight of 35% EPK Kaolin (Appendix A: Figure A1) and 65% Sil Industrial Minerals Flour 325 mesh ground silica (Appendix A: Figure A2). The grain size distribution of the prepared soil (determined from particle size distributions for the clay and silt from the manufacturers material specifications) closely resembles natural hydratebearing fine-grained soil, as illustrated in Figure 3.2. The liquid limit (LL) of the soil was determined to be 34% using a fall cone penetrometer, and its plastic limit (PL) was determined to be 18% using ASTM Standard D4318; the data from these tests are shown in Table 3.2 and Figure 3.3 respectively. Results indicate the soil had a plasticity index (PI) of 16 and an activity of 0.46, making the soil an inactive, low to medium plasticity, clayey silt classified as ML using the Unified Soil Classification System (USCS). The soil is of lower plasticity than natural KG and Ulleung Basin sediments due to the use of kaolinite rather than more plastic clay minerals 43

65 (e.g. illite, montomorillinite) that may be present in natural samples. The specific gravity of the experimental soil is 2.64, calculated as a weighted average of the manufacturer-provided specific gravities for silica silt and kaolin clay Synthetic Hydrate Methane gas is the predominant hydrate former in natural systems, but given the low permeability of the soil, low solubility of methane in water and high pressures required for hydrate formation, tetrahydrofuran (THF) was determined to be more suitable for this study. THF allows for rapid and homogeneous synthesis when mixed with water at temperatures below 4.15 C and atmospheric pressures. As THF has a higher vapour pressure than water, preferential vaporisation of THF can lead to incomplete hydrate formation. Carrying out differential scanning calorimetry measurements, Zeng et al. (2006) determined that a combination of ice and THF hydrate was obtained using a molar ratio of 1:17, and that using a ratio of 1:15 ensured complete THF hydrate formation when cooled to below 2.35 C. The re-formation of THF hydrate is accelerated after dissociation, suggesting a memory effect (Zeng et al., 2006). Extensive preliminary testing was carried out to determine the formation, dissociation and dissolution characteristics of THF hydrate. A summary of the most significant tests is shown in Table 3.3. From these trials (numbers in brackets), several important conclusions were reached: 1) THF volatilizes at a rate of 1 ml per 18 hours (1) making it important to minimize the formation time so that sufficient THF is present to ensure complete hydrate formation. 44

66 2) The formation time was reduced from 48 hours (2, 3) to 23 hours by pre-cooling THF and water and mixing regularly (4). Further reduction in the formation time to 16 hours was achieved by including a small amount of clay to increase the nucleation sites (5, 6), while cooling to -20⁰C reduced the formation time to 1.5 hours (7). 3) Hydrate formed at a molar ratio of 1:15 was found to be stiffer and more competent than when formed at 1:16 and 1:17, and contained less macroscopic structural defects (14). 4) Once THF hydrate was formed, there were no observable changes in the structure or volume with increasing storage time when kept within hydrate stability conditions (8). 5) Hydrate formed at a molar ratio of 1:15 took 30 minutes to completely dissociate at room temperature, compared to 20 minutes at a molar ratio of 1:17 (9, 10). 6) The onset of THF hydrate dissociation began after 5 minutes at room temperature and initiated along sub-horizontal fracture planes within the hydrate and spaced evenly along the column as shown in Figure ) The degree of dissolution of THF hydrate depends on the amount of free water in contact with the hydrate (11, 13). If the water volume in contact with THF hydrate is sufficiently small, the dissolved THF can reach the required molar ratio for hydrate formation (12). From tests conducted, a rapid and repeatable formation process was adopted to form THF hydrate. This included mixing THF and water at a molar ratio of 1 THF: 15 H2O and cooling the well-stirred mixture to below -20⁰C. It was assumed from the preliminary tests that hydrate dissolution into the pore water within the soil would not be significant over the time span required for geomechanical testing. 45

67 3.3 Specimen Preparation The procedure used to prepare specimens was designed to balance the practicalities of forming synthetic hydrate within saturated fine-grained soil with attempting to mimic the natural mechanisms that govern the in-situ formation of gas hydrates. The procedures developed within this thesis were built on previous work on THF hydrate in fine-grained soil (Yun et al., 2007), and extended using techniques for forming vertical sand columns in cylindrical clay specimens (Sivakumar et al., 2004) Soil Specimen Preparation Cylindrical soil specimens were prepared by consolidating a soil slurry under a vertical stress of 100 kpa, and subsequently extruding 70 mm diameter by 140 mm high consolidated specimens from the soil. An effective vertical consolidation stress of 100 kpa was chosen to form a specimen of sufficient strength such that excessive soil deformation was prevented during soil extrusion and void creation, while weak enough such that the impact of the hydrate vein on the specimen s geomechanical behaviour could be observed. This effective stress value is typical of that experienced by fine-grained sediments in the KG Basin at around 20 m below seafloor. The soil slurry was formed by thoroughly mixing silica flour and kaolin with distilled, de-aired water at a water content of 55% (around 1.5 times the LL of the soil). Once mixed, a vacuum was applied to the slurry in a sealed bucket, to remove air introduced by the mixing process. The prepared slurry was carefully poured into a specially constructed consolidation cell to allow consolidation of the soil to 100 kpa vertical effective stress. The cell consisted of a cm internal diameter stainless steel tube held between two metal plates housing porous metal discs, 46

68 to allow for free drainage of the pore water during consolidation. A ram attached to a load cell and affixed to the moveable top plate was used to apply vertical stress using a 100kN load frame. O-rings were placed around the base and top plates and a Teflon wiper was installed around the top plate perimeter, in order to seal the soil within the cell. Filter papers were applied to the top and bottom porous metal discs to prevent the migration of fines during consolidation. Figure 3.5 shows the load frame and consolidation cell. The slurry was placed into the cell with a scoop and agitated regularly to prevent air entrapment and segregation. The top plate was brought into contact with the slurry, and a vertical load of 3.33kN (100kPa) was applied in stages. After full consolidation was achieved, the soil was unloaded in the presence of excess water to prevent the entry of air into the void space. Individual soil specimens were taken by slowly pushing 70 mm internal diameter cylindrical sampling tubes with sharpened edges into the soil using the load frame to minimize excess pore pressure development and soil structure disturbance. Specimens were stored within a sealed polyethylene bag, which in turn was stored within a sealed bucket to prevent moisture loss until ready for further specimen preparation procedures. When specimens were required for testing, they were extruded from the sample tubes using a vertically-mounted lever action hydraulic jack, shown in Figure 3.6, and placed on a steel dummy pedestal the same height as the triaxial base upstand. For specimens intended for reconsolidation in the triaxial apparatus, cm saturated filter paper radial drains were applied around the specimen to aid in reconsolidation. A latex rubber membrane was placed around the soil before being placed in a 70 mm internal diameter split mold to maintain the structural integrity of the specimen during subsequent vein formation. The specimen weight and dimensions were determined using an electronic scale with a precision of ±0.005g and a caliper 47

69 with a ±0.5 mm precision. Plastic wrap was applied to the top of the mold and placed in a refrigerator, where the soil was cooled to between 0 and 2⁰C prior to hydrate vein formation Hydrate Vein Formation within Soil Two hydrate vein formation processes were used in this research, and are outlined in Sections and However, the steps involved with forming a void in the soil specimen within which the hydrate vein is formed do not change and so are first detailed in Section Handling of open THF was carried out under a chemical fume hood using personal protective equipment, and THF liquid and hydrate were stored in sealed containers in dry, cool and wellventilated locations according to Material Safety Data Sheet specifications Vein Void Formation The mold containing the cooled soil was placed under a drill press and wood augers of 6.35 mm (¼"), 12.7 mm (½"), mm (¾") and 25.4 mm (1") diameter 1 were used to drill through the specimen to form a continuous cylindrical hole, as shown in Figure 3.7. The soil removed from the newly formed vein void was placed in a sealed bag to prevent moisture loss, and replaced in the refrigerator along with the specimen to cool them back into the hydrate stability field. Figure 3.8 shows the temperature changes of the specimen plotted versus time throughout the initial cooling to 2⁰C (which took 7 hours), the temperature increase during the drilling of the vein void (rose to 5.5⁰C over 15 minutes), followed by subsequent re-cooling of the specimen to 2⁰C (3 1 The vein sizes will henceforth be referred to in inches for simplicity. 48

70 hours). Prior to hydrate vein emplacement the void was re-cored, as minor soil migration and sloughing into the void occurred during re-cooling In Situ Formation Method The driving philosophy behind the in situ formation method is to replicate as best as possible natural formation conditions by allowing for the formation of THF hydrate in situ within the drilled vein void. A number of tests were conducted to refine this process, and the significant tests involved in the development of this method are outlined chronologically within Table 3.4. For this method, the cooled specimen containing the bored hole was re-weighed and placed under the fume hood. A circular latex membrane was placed between the dummy pedestal and the soil specimen, and a portion of the spoils from the drilling process were tamped into the bored hole to form a thin layer of soil at the base of the specimen. The purpose of the membrane was to ensure that when the THF-water mixture was placed in the vein void, it did not leak out while the hydrate was forming. The purpose of the tamped clay was to separate the hydrate from the base plate to prevent hydrate dissolution into the water-saturated porous stone during testing. As it was shown that the subsequent reformation of THF hydrate was faster due to the memory effect, a vial of THF hydrate was formed from the 1:15 THF-water mixture using methods outlined in Section 3.2.2, and then dissociated and poured into the vein void as shown in Figure 3.9. The triaxial top cap was placed on top prior to cooling the specimen to form THF hydrate. A number of cooling procedures were attempted to form hydrate within the specimen. Attempts to form hydrate by cooling the specimen and THF-water mixture to 2⁰C were unsuccessful 49

71 (Tests 1 & 2 in Table 3.4) due to the volatilization and infiltration of the THF solution into the soil matrix. Additionally, keeping the THF-water mixture in the open vein void throughout the time required for hydrate formation led to soil sloughing into the liquid, and slow deformation of the soil inwards due to hoop stress from the membrane. To reduce the formation time, the covered specimen and THF-water mixture was cooled to -20⁰C in a freezer to reducing the formation time to minutes depending on the vein size. The specimen was returned to the fume hood and a portion of the soil from the vein drilling process was tamped on top of the specimen to form a thin soil layer. The specimen was then covered with the top cap and returned to the refrigerator for storage prior to transfer into the geomechanical testing apparatus. The main issue with the in situ formation method, shown in Test 6, is that partial freezing of the soil specimen resulted in ice lensing due to the frost-susceptible nature of the soil (Clark and Phillips, 2003), which can be seen in Figure Indeed in Test 7, hydrate formation was found to occur in concert with soil freezing, meaning that some ice lensing could be expected during the formation of hydrate using this method. Since ice lenses can result in thaw-consolidation and affect the structure and behaviour of the soil (Nixon and Morgenstern, 1973), this method was determined to not be ideal for this study, and was only used when required, as discussed below Transfer Method To overcome the issues associated with the in situ formation method, an alternative method was developed in which a hydrate cylinder was formed independent of the soil and then transferred into the vein void. This method allowed for the competency of the hydrate cylinder to be evaluated prior to emplacement. 50

72 The hydrate cylinders were formed in cylindrical aluminium foil molds with internal diameters equivalent to the required vein size. A mold was constructed and sealed with vacuum grease, filled with 1:15 THF-water mixture, covered with a foil cap and then placed in the freezer to initiate hydrate formation. A fully formed 0.25" hydrate cylinder is shown in Figure The cooled specimen containing the vein void was moved from the refrigerator to the fume hood and weighed. A portion of the spoiled soil was tamped into the vein void, the hydrate cylinder was removed from the freezer, quickly unwrapped, and carefully inserted into the vein void. Cool soil was then tamped on the specimen, the top cap was emplaced and the specimen was placed in the freezer for 10 minutes to quickly reform any dissociated hydrate. This was done as it was discovered that during the insertion of the hydrate cylinder, minor dissociation occurred on the surface of the hydrate cylinder due to the heat generated by friction between the hydrate and soil. A preliminary test showed no freezing of soil occurred when cooled to -20⁰C for 10 minutes. The specimen was moved to the small refrigerator for storage until triaxial testing Method Selection The transfer method was found to work well for hydrate cylinders with 0.50", 0.75" and 1" diameters. However, any manipulation of 0.25" diameter hydrate cylinders caused them to fracture, negating any contribution they might have to the geomechanical behaviour of the specimen. Conversely, using the in situ formation method for the 0.25" vein required only 30 minutes in the freezer due to the small THF-water volume, resulting in minimal soil freezing. Thus the transfer method was used for 0.50", 0.75" and 1" inch veins, and the in situ method was used for 0.25" veins. The fragility of the 0.25" diameter hydrate cylinders should be noted, as this foreshadows their contribution to the behaviour of the fine-grained sediment. 51

73 3.4 Baseline Geomechanical Testing on Fine-Grained Soil Tests were first conducted on hydrate-free soil specimens using oedometer cells and the triaxial apparatus to determine its consolidation and shear behaviour at differing effective stresses Oedometer Consolidation Tests One-dimensional oedometer consolidation tests were carried out on both slurried and consolidated soil at room temperature, following ASTM Standard D2435. The slurry and consolidated soil were prepared as outlined in Section 3.3.1, with the consolidated specimens trimmed from soil consolidated to 100 kpa in the consolidation cell and the slurry spooned into the cell. In both cases, the soil was placed within a metal confining ring interposed between two saturated porous discs and filter papers, and then placed in the oedometer cell. The initial wet mass and the height of the specimen were determined. A load cap was seated on the top porous disc and the cell was placed within a pneumatic load frame capable of applying and maintaining pressures of kpa with a precision load regulator, set using a calibrated test gauge (0.25% accuracy). Change in height of the specimen was measured using a high resolution 25 mm LVDT compression gauge and logged continuously. The consolidation cell was saturated with distilled water, to ensure the soil remained saturated throughout the test. For the slurry, vertical pressures of 5, 10, 20, 50, 100, 200, 400 and 800 kpa were applied every 24 hours, with additional loading steps of 75, 125, 150 and 175 kpa applied to the consolidated soil to determine the pre-consolidation pressure (100 kpa). The slurry was unloaded by steps of 200, 50 and 5 kpa, as a stress decrement was skipped as per ASTM Standard. The consolidated soil was unloaded directly to 5 kpa, as the unloading behaviour was investigated in tests on the 52

74 slurry. The cell was dismantled quickly after releasing the final load, the specimen was removed and its mass, height and water content were determined K0-Consolidation and Undrained (K0CU) Compression Tests Standard triaxial tests involve consolidating soil isotropically while natural soils generally consolidate one-dimensionally due to lateral confinement by neighbouring soil. To mimic natural soil loading, a K0-consolidation can be conducted such that stresses are applied to a specimen so that radial deformation is prevented. As the objective of this research was to replicate natural conditions, initial studies were carried out to determine the feasibility of K0-consolidation in investigating the consolidation behaviour of hydrate-bearing fine-grained soils. While it was ultimately deemed overly time-consuming for use on soil containing hydrate veins given the time-instability of THF hydrate, K0-consolidation followed by undrained shear (CK0U) tests were carried out on non-hydrate-bearing soil specimens at 2⁰C using two different methods Geomechanical Testing Apparatus A double walled, computer-controlled triaxial system was used for this study as pictured in Figure 3.12 and shown in a schematic diagram in Figure The apparatus featured a 25 kn load frame with an external load cell (0.05% precision) that applied a vertical load with a convex loading piston housed within the top plate. Two clear, acrylic, hydraulically connected cell walls enabled specimen observation during testing and allowed confining pressures up to 2 MPa to be applied to the specimen. The specimen was placed between a 70 mm diameter stainless steel base pedestal and a top cap housing porous stones, allowing the specimen to be hydraulically connected to a computer servo-controlled hydraulic pump with an accuracy of ±1 kpa. The cell 53

75 pressure was controlled by a larger hydraulic pump with the same accuracy. Pressure transducers allowed independent measurements of both the cell and pore pressure to 0.1 kpa resolution. An electronic pore pressure transducer was used in the base plate with a low operating temperature and 0.25% accuracy, calibrated at 2⁰C. Axial displacement of the specimen was measured by an external LVDT on the load ram. Radial deformation was measured using a circumferential strain gauge mounted on the specimen, consisting of a Teflon roller assembly with a 0.2 μm resolution. The system was modified to enable testing at temperatures required for THF hydrate stability (<2⁰C). As shown in Figure 3.12, a refrigerated circulator was used to pump coolant fluid through a network of copper pipes submerged within the cell water, and a second circulator was connected to a copper pipe network inlayed within an aluminium plate and placed beneath the triaxial base plate. An insulation jacket was installed around the cell to help maintain the temperature, which was monitored by a thermocouple placed within the cell fluid Specimen Mounting and Cell Assembly The soil specimen was prepared using the procedure outlined in Section As the consolidated specimens were saturated, a wet mounting method was used following ASTM Standard D4767 involving saturating specimen drainage lines, the base pedestal and the top cap with de-aired water prior to specimen mounting. Saturated circular filter papers were placed on the top and bottom of the specimen, the specimen was mounted on the base pedestal, the top cap was installed and O-rings applied around the membrane to ensure a proper seal at the top cap and bottom pedestal. The circumferential strain gauge was mounted at mid-height on the specimen. The triaxial top platen was placed on the supporting bars, and the axial load bar was brought into 54

76 contact with the top cap, ensuring proper seating and alignment. The cell was sealed, insulated and filled with de-aired water and then cooled to 2⁰C using the refrigerated circulators K0-Consolidation Head and Epps (2014a) suggest that a virtually saturated soil does not require saturation procedures, but applying a back pressure on the drainage line is advantageous as air bubbles in the void space, between the membrane and in the back pressure system are forced into solution. Therefore, the specimen was isotropically reconsolidated to its preconsolidation pressure of 100 kpa by increasing the cell pressure to 500 kpa and setting the back pressure to 400 kpa. During reconsolidation, the specimen was drained through the top cap, aided by radial drainage through filter strips. The diameter of the specimen after isotropic reconsolidation was recorded and maintained during K0-consolidation, which was performed manually. A stress path suggested by Germaine & Ladd (1988) was followed, where the vertical stress (σ 1 ) was increased incrementally while adjusting the radial stress (σ 3 ) by changing cell pressure in response to specimen deformation to maintain the lateral strain equal to zero (ε r 0). The volume change of the specimen ( V) was approximated by fluid flow out of the specimen, calculated from the change in position ( P) of the back pressure piston in the pump, and the area of the piston (A p ): V = P A p (3.1) This was used along with the initial volume (V 0 ) to calculate the volumetric strain (ε v ): ε v = V V 0 100% (3.2) The axial strain (ε a ) was determined using the ram displacement ( H) and the initial height (H 0 ): 55

77 ε a = H H 0 100% (3.3) Therefore, assuming small deformations, the radial strain (ε r ) was calculated independent of the radial strain gauge according to: ε r = ε v ε a 2 (3.4) The circumferential strain gauge measured the change in circumference ( C) of the specimen, allowing a direct measurement of the radial strain (ε rgauge ) according to the following equation: ε rgauge = r r 0 = ( C/π)/2 (C 0 /π)/2 = C C 0 (3.5) One K0-consolidation was performed using Equation 3.4 and another using Equation 3.5, and both yielded significantly different results, which is discussed further in Section Vertical stress increments were kept relatively small ( σ v = 0.2σ v ) as suggested by Germaine and Ladd (1988) to minimize straining due to undrained shear deformation, which would have occurred if the specimen reached its yield envelope. Vertical stress was applied by lowering the axial load bar using a constant strain rate of mm/min until the desired stress value was reached. At this point the vertical stress was held constant, and the subsequent specimen deformation was measured. Consolidation during each stress increment application was considered complete when 95% of the excess pore pressure was dissipated, as suggested by Head and Epps (2014). The pore pressure distribution within the sample was assumed to be parabolic and the average pore pressure (u a ) was calculated using the pressure transducer readings at the base (u c ) and the top of the sample (u b ) using the following equation: 56

78 u a = 2 3 u c u b (3.6) After each stress increment and subsequent consolidation stage the radial strain was determined using Equation 3.4 or 3.5. As the soil expanded laterally under vertical loading, the cell pressure was increased incrementally until the specimen diameter returned to its original value within a certain tolerance. JGS Standard 0525 (Japanese Geotechnical Society, 2009) suggests that for K0 consolidation the tolerance should be: ε r < 0.05%. However, for the purposes of this research it was considered that a sample had been K0-consolidated if ε r < 0.5% due to the uncertainty associated with the radial strain measurements and calculations Undrained Shear Axial undrained compression tests were carried out once samples had been K0-consolidated. A strain rate of 0.07 mm/min was used (0.05%/min) that was slightly faster than rates suggested by the ASTM Standard and British Standard 1377: Part 8: 1990: 7 ( and mm/min respectively), but was compatible with the strain rate used for CU shear tests on hydrate-bearing specimens where the time-dependent hydrate stability required a faster testing time, as has been done in previous tests on hydrate-bearing soil (Yun et al., 2007). The shear stage was terminated at 15% axial strain as per ASTM Standard. Once the shear stage was completed, the axial load was removed, the cell and back pressure were reduced to zero, the cell was dismantled, the specimen was removed and its weight, height and water content determined. 57

79 3.5 Geomechanical Testing on Hydrate-Bearing Soil Following baseline testing, hydrate-bearing fine-grained soil specimens were subjected to undrained shear at different effective stress conditions. Procedures mostly follow ASTM standards, with some alterations to account for the presence of THF hydrate veins Specimen Mounting and Cell Assembly As THF hydrate dissociates if warmed above its stability conditions, a number of modifications were made to the wet mounting procedure adopted for CK0U testing to prevent this: The triaxial base plate was cooled prior to specimen mounting using the base cooling system, the refrigerated circulator for the upper cooling system was turned on, and the deaired cell water was cooled using ice cubes and ice packs to approximately 1⁰C. Specimen drainage lines, base pedestal and top cap were saturated with cooled cell water. Specimen transfer from refrigerator storage to the base pedestal was done as quickly as possible to minimize the time the specimen was exposed to room temperature. Ice cubes were placed around the specimen to maintain stability until the cell could be assembled and filled with the pre-cooled (1 o C) de-aired cell water. With these modifications, the specimen was only outside of the hydrate stability zone for 10 minutes with no significant hydrate dissociation being observed during this process, indicating that the heat capacity of the cooled soil was sufficiently high and its thermal conductivity sufficiently low to slow heat transfer from the surroundings to the hydrate vein. 58

80 3.5.2 Consolidated Undrained (CU) Triaxial Compression Testing The steps taken during the reconsolidation phase of the CU tests were identical to those adopted for the reconsolidation portion of the K0 consolidation tests (outlined in Section ). Specimens with and without hydrate veins were reconsolidated to 100 kpa effective confining stress by applying a cell pressure of 500 kpa and a back pressure of 400 kpa, allowing drainage through the top cap and aided by radial drains. During consolidation, the axial load piston was brought into contact with the specimen cap while ensuring that an axial load of 0.5% of the estimated axial load at failure was not exceeded. Volume change was calculated using Equation 3.1 along with the change in height of the specimen. The consolidation stage was terminated when at least 95% of the pore pressure was dissipated, calculated using Equation 3.6. Once reconsolidated, the specimen was isolated from the back pressure line so that no free water was available to the specimen during the undrained shear stage. Shear compression was carried out at a strain rate of 0.07 mm/min (0.05%/min) for reasons outlined in Section until 15% axial strain was reached. Once this was completed the cell was dismantled and the specimen was removed in less than 10 minutes to prevent hydrate dissociation. The specimen was transferred to the fume hood where its dimensions and weight were determined, it was cut open so that the hydrate vein could be photographed, and the vein was removed and its weight determined. The moisture content was determined at the bottom, middle and top of the specimen followed by safe disposal of the dissociated THF-water mixture. 59

81 3.5.3 Unconsolidated Undrained (UU) Triaxial Compression Testing Unconsolidated undrained (UU) shear tests were carried out on specimens, following ASTM Standard D2850 with some modifications. The specimen preparation and cell assembly for UU testing was identical to that for CU testing, however no radial drains were applied to UU samples. After assembly, a cell pressure of 200 kpa was applied to ensure that any air within the sample or the drainage lines was forced into solution. The shear stage involved an axial strain rate of 0.3%/min for all specimens, as suggested for brittle materials by the ASTM Standard, which was expected to be the case for hydrate-bearing specimens. Axial loading was continued until 15% axial strain. After shear, the same procedure was used to dismantle the apparatus and determine specimen properties as described previously in Section for CU tests. 60

82 Table 3.1: Characteristics of natural hydrate-bearing soils and prepared soil for this research Characteristics Krishna- Godavari Basin 1 Ulleung Basin 2 Gulf of Mexico 1 Prepared Fine-Grained Soil Average Sand (% by weight) Average Silt (% by weight) Average Clay (% by weight) Liquid limit range N/A 34 Plastic limit range N/A 18 1 (Winters, 2011) 2 (Lee et al., 2011) Table 3.2: Data from plastic limit determination on prepared soil using ASTM D4318 Trial Number Water Contents of 3 mm Diameter Soil Threads (%) Average Plastic Limit 18 61

83 Table 3.3: Preliminary tests in the development of the THF hydrate formation procedure Test 1 Objective Define volatilization rate of pure THF liquid Hydrate Molar Ratio Conclusions THF:H2O THF Volatilization Test Pure THF: 4.45 ml to 3.4 ml in 18 hours at room temperature, little change over short term Hydrate Formation Tests 2 Formation at 2.5⁰C 1:15 Formation Time: 48 hours 3 Formation at 0.5⁰C 1:15 Formation Time: 45 hours Pyramidal crystal formation upward from bottom of test tube 4 Formation at 2⁰C after precooling THF/water 1:15 Formation Time: 23.5 hours 5 Formation at 2⁰C with some clay 1:15 Formation Time: 16.5 hours 6 Formation at 2⁰C, then re-pour mixture into 2 nd test tube 1:15 Formation Time: 16 hours 7 Formation at -20⁰C 1:15 Formation Time: 1.5 hours 8 Volume loss after storing hydrate in fridge Hydrate Storage Test Storage Time: 720 hours 1:15 No volume loss, completely solid Hydrate Dissociation Tests 9 Dissociation at 25⁰C 1:17 Dissociation Time: 20 minutes 10 Dissociation at 25⁰C 1:15 4 mins: No liquid apparent; 4.5 mins: Liquid development in cracks; 6 mins: 1 ml liquid; 12.5 mins: 2 ml of liquid; 20 mins: 5 ml of liquid; 24 mins: 10 ml of liquid; 30 mins: dissociated Hydrate Dissolution Tests 11 Dissolution at 2⁰C into 100 ml Complete dissolution of 5 ml hydrate after 20 1:15 THF/Water hours 12 Dissolution at 2⁰C into 1 ml 18 ml hydrate becomes 20 ml hydrate after 36 1:13 Water hours ml hydrate dissociates into water then reforms Dissolution at 2⁰C into 5 ml 1:15 at top of cylinder, 9 ml THF at bottom and 7 at Water top with 7 ml of water between after 336 hours Hydrate Strength Tests 14 Qualitative Compression of Veins at 25⁰C 1:15 1:16 1:17 1:15 Vein: Fractured into 2 pieces, very stiff 1:16 Vein: Fractured more easily into 4 pieces 1:17 Vein: Splintered very easily into pieces 62

84 Table 3.4: Preliminary tests in the development of the in situ vein formation procedure Test Objective Formation Time (hours) Observations/Conclusions 1 Form 0.50" Vein at 2⁰C Failed THF/H 2O escaped vein through bottom of mold 2 Form 0.50" Vein at 2⁰C Failed No formation after 24 hours, liquid still present in vein 3 Form 0.50" Vein at -20⁰C Form 0.50" Vein at -20⁰C Storage test over 6 days Form 0.50" Vein at -20⁰C Observe hydrate vein competency Form 0.50" Vein at -20⁰C Check for ice lensing after full vein formation Hydrate ~1 cm below top of sample, refilled with THFwater mixture and reformed fully competent vein 3 days no change 6 days - hydrate disappeared Could be due to dissolution into porewater or fridge warming to above 4⁰C Water around base after overnight storage in fridge Vein fairly intact (See Figure 3.9). Cut open after freezing: Bottom ~5 cm (in contact with upstand) and top ~5 cm (in contact with top cap) show ice lensing & freezing (See Figure 3.10) 7 Form 0.50" Vein at -20⁰C Observe ice lensing after partial vein formation 0.75 Bottom ~0.2 cm of clay frozen, hydrate only present in this interval indicating they are formed together. 63

85 Testing Program Experimental Procedure Baseline Testing Slurry Fine-Grained Soil Mixture Oedometer on Slurry Consolidate Soil to 100 kpa Oedometer on Consolidated Soil Core Soil Specimens CU/UU Tests on Non-Hydrate-Bearing Specimens Extrude Specimen K 0-Consolidation and Undrained Shear Tests Drill 0.50, 0.75, 1" Vein Void in Soil Drill 0.25" Vein Void in Soil Hydrate Vein Transfer Method In Situ Hydrate Vein Formation Method Modified Specimen Mounting and Cell Assembly (10 minutes) Undrained Shear (UU) Pressurize Specimen Undrained Shear (CU) Reconsolidate Specimen Figure 3.1: Flowchart summarizing the testing procedure adopted for this research program including specimen preparation, baseline testing and geomechanical testing program. 64

86 Figure 3.2: Grain size distribution curve of the prepared fine-grained soil compared to formerly gas-hydrate-bearing soil recovered from the KG Basin (after Clayton et al., 2008) and the Gulf of Mexico (after Winters, 2011), as well as basin averages from the KG Basin (after Winters, 2011) and Ulleung Basin (after Lee et al., 2011). 65

87 Figure 3.3: Liquid limit determined from fall cone penetrometer results. The liquid limit of the soil (~34%) is defined as the water content when penetration depth is equal to 20 mm. 66

88 A B C D Figure 3.4: THF hydrate cylindrical vein before dissociation (a) and during dissociation (b, c, d) with veins breaking into distinct segments along planes of weakness. 67

89 Top Plate Teflon Wiper Porous Metal Plate Consolidation Cell Drainage Lines Bottom Plate Figure 3.5: The specially constructed consolidation cell mounted in a load frame, with the aluminium top plate connected by ram to the load cell and porous discs fitted to the top and base plate allowing for the drainage of excess pore water during consolidation. 68

90 Figure 3.6: Hydraulic jack used to extrude cylindrical consolidated soil specimens from 70 mm internal diameter sampling tube (left). 69

91 Figure 3.7: Vein void installation in specimen using 0.50" wood auger hooked up to drill press. Excessive specimen deformation was prevented by confining the specimen within a latex rubber membrane, stainless steel split mold and steel dummy pedestal. 70

92 Figure 3.8: Specimen temperature as measured throughout the vein drilling procedure, showing the initial cooling after extrusion, warming during the vein drilling process, and specimen recooling before hydrate formation. 71

93 A B Figure 3.9: In situ hydrate vein formation method with (a) the THF-water mixture poured into the vein void and (b) the specimen after overnight storage within the hydrate stability field. 72

94 B A C Figure 3.10: Preliminary Test 6 described in Table 3.4 showing (a) ice lenses, (b) full hydrate vein formation, (c) de-structured soil after melting of ice lenses. 73

95 Figure 3.11: Aluminium foil mold containing a 0.25" hydrate cylinder, which proved impossible to unwrap without fracturing into segments. 74

96 Cooling Systems A B C Figure 3.12: Triaxial system showing (a) upper and lower cooling systems, (b) with double wall cells and (c) with insulation, hooked up to refrigerated circulators. 75

97 Legend - Plastic piping (water) - Insulated piping (glycol) - Copper piping (glycol) - Refrigerated circulator - Thermocouple - Manual ball valve - Automatic ball valve - Pressure transducer - Pressure gauge - De-aired water reservoir - Volume change device - Pressure intensifier Load Load Frame LVDT Load Cell Outer Cell Wall Inner Cell Wall Specimen Hydrate Vein Upper Cooling Tubes Porous Stone Base Plate Cooler 1⁰C Pressure Control Cabinet 25⁰C Figure 3.13: Schematic illustration of triaxial system showing modifications made to maintain specimen at 2⁰C, including refrigerated circulators pumping coolant through copper piping within cell fluid and below the base plate, and water reservoir containing water cooled to 1⁰C. 76

98 Chapter Four: Laboratory Results and Analysis 4.1 Introduction This chapter presents results from laboratory tests conducted with the goal of addressing the second objective of this thesis, to determine the impact of differing hydrate vein sizes on the geomechanical behaviour of fine-grained specimens under different effective stress conditions. Results from the baseline testing program on fine-grained soil are first presented, and then compared with results on THF hydrate vein-bearing specimens under effective and zero effective stress conditions. 4.2 Baseline Geomechanical Testing on Fine-Grained Soil Oedometer Consolidation Tests Oedometer tests were carried out on both slurry and preconsolidated soil (to 100 kpa) to study the consolidation behaviour of the soil. Load-deformation data from the oedometer tests are detailed in Appendix B and summarized in Table 4.1. Figure 4.1a shows the change in void ratio with vertical stress for all tests, highlighting that void ratios vary by ±0.02 for tests on preconsolidated soil, which is sufficiently repeatable given that the transitional behaviour of clayey silt can lead to a larger variance in void ratios (±0.05) (Ponzoni et al., 2014). The slurry and preconsolidated soil show a significant difference in void ratio at 100 kpa vertical stress, which is assumed to be related to hysteresis in the unloading/reloading of the preconsolidated soil, as the difference decreases with increasing vertical stress. Figure 4.1b shows the graphical Casagrande Method (Casagrande, 1936) that was used to confirm that soil specimens prepared in the consolidation cell using the load frame were indeed consolidated to a vertical stress of approximately 100 kpa. 77

99 Figure 4.2 shows the determination of the compression (C c ) and recompression (C r ) indices of the soil (simplified as the slope of the unloading line). The slurry was slightly more compressible (C c = 0.22) than the preconsolidated soil (C c = 0.19), which may be because the slurry had a higher initial water content than the liquid limit (LL) which can increase the compressibility, or due to hysteresis in the unloading/reloading curves (Head and Epps, 2014b). Equation 4.1 shows a relationship put forward by Skempton (1957) relating the LL of a soil to its C c : C c = 0.009(LL 10%) (4.1) By Equation 4.1, the C c should be approximately 0.216, similar to values calculated from results. In all cases the C r was found to be around These results can be compared to consolidation in the triaxial apparatus, although it must be noted that oedometer testing was carried out at room temperature while consolidation in the triaxial apparatus was at 2⁰C. A decrease in temperature from 25⁰C (oedometer) to 2⁰C (triaxial) should not lead to differences in primary compression, but can decrease the consolidation rate and the amount of secondary compression (Head and Epps, 2014b), both of which were not considered important for this study K0-Consolidation and Undrained (K0CU) Compression Tests Several K0-consolidation tests followed by undrained shear were carried out on preconsolidated soil specimens in the triaxial apparatus to determine the soil behaviour under different stress conditions and to consider the feasibility of K0-consolidation in investigating hydrate-bearing fine-grained soils. While the time taken to follow a K0 stress path was ultimately deemed to be too long for hydrate-bearing specimens given the time instability of THF hydrates, discussed further in Section 4.3.3, results are presented on non-hydrate-bearing specimens. 78

100 K0-Consolidation Results and Analysis In the K0-consolidation tests conducted, the radial strain was maintained within a certain threshold (ε r < 0.5% ) by applying small axial stress increments after each cell pressure increment. Initially the specimen diameter was determined using the circumferential strain gauge placed around the specimen, however after the first test it became apparent that it was not sufficiently accurate in measuring radial changes to carry out a true K0-consolidation. This resulted in anisotropic consolidation of the specimen, with an estimated K value (σ 3 /σ 1 ) of approximately 0.75 (K 0.75 ). A second test was conducted where the radial strain was calculated from axial and volumetric strains, and was considered to follow more closely a K0 stress path given the accuracy of the strain measurements, with a K value of approximately 0.38 (K 0.38 ). The data from K0-consolidation tests are detailed in Appendix C and summarized in Table 4.2. Calculated axial stress values were corrected for piston friction and changes in cross-sectional area. No filter strip correction was used, as Watabe et al. (2003) saw no difference in K0 test results with or without filter strips. The stress paths followed for both tests are shown in Figure 4.3, highlighting the stages in the successful K0-consolidation test at which the soil was considered to be K0-consolidated, giving a relatively constant K value of approximately 0.38 at higher stress values, which is taken as the normally consolidated K 0(NC) value (Watabe et al., 2003). Jaky (1948) relates the K 0(NC) value to the friction angle by: K 0(NC) = 1 sin φ (4.2) By this the friction angle is expected to be 38⁰, which can be verified in the following section. 79

101 The void ratio versus logarithm of vertical effective stress for the K0-consolidation is plotted in Figure 4.4. The void ratio after each consolidation stage (e f ) is calculated assuming the soil is saturated and volume change is due to pore water expulsion ( V = V v ): e f = e 0 e = e 0 V v V s (4.3) The C r of the soil appears to be greater for the isotropic reconsolidation of the triaxial specimen than the one-dimensional reconsolidation in the oedometer test, possibly due to all-round stress application to a one-dimensionally consolidated specimen. After reconsolidating the soil to its preconsolidation stress, further K0-consolidation of the specimen is theoretically onedimensional and results should resemble those from the oedometer tests. However, the C c of the K0-consolidated specimen is lower than observed in oedometer tests (0.14 as compared to 0.19), indicating it is slightly less compressible, which may be due to the presence of the membrane and radial drains or the result of imprecise strain measurements. The average initial void ratio of the triaxial specimens (0.67) was lower than oedometer samples (0.72), potentially due to sampling differences, since the triaxial specimens were obtained by pushing sample tubes into consolidated soil potentially leading to soil disturbance and densification while oedometer samples were cut from consolidated soil, potentially suffering less disturbance. Undrained Shear Results and Analysis Data Analysis Techniques Undrained shear tests were carried out on the two specimens that followed different anisotropic stress paths. Initial sample dimensions after consolidation were determined using strain relationships outlined previously. Following ASTM standard, corrections for vertical filter strips 80

102 and the rubber membrane (using a typical membrane stiffness value) were applied, as well as an increase in the cross-sectional area of the specimen with increasing axial strain. The corrected deviatoric stress (q) was used along with confining stress (σ 3 ) and pore pressure (u) measurements in the calculation of principal stresses (σ 1, σ 3 ) allowing the mean effective stress (p ) to be determined using: p = (σ 1 + 2σ 3 )/3 (4.4) Additionally, the pore pressure coefficient (A) which in a saturated soil defines the change in pore pressure with deviatoric stress, was calculated by: A = u u 0 σ 1 σ 3 (4.5) Graphical representation of the data taken throughout each shear test are shown in Appendix C. Using these plots, soil failure characteristics can be determined. In this thesis, two definitions of failure are used to describe the soil strength, the peak deviatoric stress and the deviatoric stress at critical state. Critical state is defined as when the soil continuously deforms at constant volume (constant pore pressure in undrained tests) under constant effective stress, whereby: du dq dp = 0, = 0, dε dε dε = 0 (4.6) Critical state strength is a fundamental soil property dependent only on effective stress while peak strength can depend on soil density. The failure criterion used for each test is identified. Material stiffness can be described by the Young s modulus of elasticity (E = σ/ ε), but applies only when strains are perfectly recoverable, which is generally only true at infinitesimal strains for soils. Therefore, the initial tangent slope (E i ) of the stress-strain curve of a soil through the origin represents its modulus of elasticity. In this study, axial strain is measured 81

103 externally, so the initial tangent modulus is unlikely to be accurate due to seating/bedding effects. To remedy this, the secant modulus joining the origin to the point on the stress-strain curve at half of the peak deviatoric stress (E 50 ) is typically used, although at times the secant modulus to 0.5% strain (E 0.5% ) or over some other strain interval, is more appropriate. Stress-Strain Response Figure 4.5a shows the deviatoric stress versus axial strain for the two anisotropically consolidated specimens (K 0.38 and K 0.75 ), along with the response of a specimen isotropically reconsolidated (K 1 ) to 100 kpa and submitted to undrained shear (discussed further in Section 4.3.2). The more heavily consolidated K 0.38 specimen exhibited a peak axial stress with post peak softening to critical state. The less consolidated K 0.75 and K 1 specimens showed no appreciable peaks, and the K 0.75 specimen had a critical state strength slightly lower than that of the K 0.38 specimen. The sharp rise in deviatoric stress with axial strain for the K 0.38 specimen gave rise to a significantly higher stiffness (E 50 ), while the E 0.5% values calculated over a larger strain interval for both anisotropically consolidated samples are comparable. Pore Pressure Response Figure 4.5b presents pore pressure coefficient (A) versus strain for the K 0.38, K 0.75 and K 1 specimens. The K 0.75 and K 1 specimens display early peak A coefficients, indicating initial compression of the specimens. This is followed by a decrease to the same coefficient at failure (A f = 0.35), indicating an increase in the specimen volume implying a dilatant tendency, although not sufficient to result in a total volumetric expansion (indicated by a negative A f ). 82

104 However the K 0.38 specimen displays no peak and a lower A f value (A f = 0.18), indicating a more muted pore pressure response to deviatoric stress with no dilatant tendency during shear. Effective Stress Paths The difference between the responses is illustrated further by the effective stress paths shown in Figure 4.6. The K 0.75 and K 1 specimens initially follow a contractive stress path typical of normally consolidated clay until they approach the critical state line, at which point a change in behaviour occurs and the material begins to exhibit dilative behaviour (similar to dense sand) due to a pore pressure reduction with increasing deviatoric stress. This behaviour has been observed for triaxial tests on a similar soil (25% kaolin and 75% Sil-Co-Sil silt), as shown in Figure 4.6, which was postulated to be due to the high silica silt content resulting in a behaviour transitional between that of clay and sand (Dayarathne and Hawlader, 2015). However, the K 0.38 specimen exhibits only contraction towards the critical state line as expected for a normally consolidated clay. The difference in behaviour is most likely related to the difference in axial stress applied to each specimen, with increasing axial stress seemingly preventing the dilative tendency of the silt particles within the soil skeleton. For normally consolidated samples the c u /p ratio is generally assumed to be constant, where c u is the undrained shear strength. However, the c u /p ratio of the two anistropically consolidated specimens are 0.61 (K 0.38 ) and 0.70 (K 0.75 ), while that of the isotropically consolidated specimen (K 1 ) is 0.72, showing a decrease in undrained shear strength with a decrease in stress ratio, welldocumented in literature (Bishop and Henkel, 1962). The variability in undrained strength with different consolidation stress conditions leads to some scatter around the best fit critical state line 83

105 as shown in Figure 4.6. The slope of the critical state line (M) is determined to be 1.48 from the three shear tests, giving a critical state friction angle of 36⁰ (using Equation 4.7) which is higher than typical clayey silt, but lower than the value of 38⁰ predicted using Equation 4.2: sin φ cs = 3M 6 + M (4.7) 4.3 Consolidated Undrained (CU) Compression Testing Consolidated undrained (CU) triaxial testing was carried out on hydrate-bearing soil specimens to address one of the key objectives of this thesis, determining the impact of hydrate veins on soil behaviour. Isotropic reconsolidation followed by undrained shear was carried out on specimens with hydrate vein diameters of 0.25", 0.50", 0.75", 1", and soil specimens with no vein. Results are presented in Appendix D and summarized in Tables 4.3 and 4.4. The CU compression tests on specimens with hydrate veins of 0.75" and 1" diameter were found to strengthen and stiffen the fine-grained soil, however significant issues were encountered with the majority of specimens containing smaller hydrate veins. Therefore, the isotropic reconsolidation and undrained shear results from the two successful specimens (0.75" and 1" diameter hydrate veins) and the baseline specimen are presented in Sections and and summarized in Table 4.3. The results from the tests on the unsuccessful specimens are summarized in Table 4.4, and discussed in Section along with the reasons for the issues encountered. 84

106 4.3.1 Isotropic Reconsolidation Results and Analysis Development of Consolidation Parameters Due to the hydrate vein presence within the soil during reconsolidation, standard consolidation parameters must be adapted to allow for data analysis. The drilling of the vein decreases the total soil volume of the specimen (V T(soil) ), which can be calculated from the initial specimen volume (V 0 ) and the vein volume (V vein ) as given in Equation 4.8. This results in an increase in the void ratio of the specimen as a soil cylinder is removed and replaced with a hydrate vein: V T(soil) = V 0 V vein (4.8) The volume of voids of the host soil (V v(soil) ) can be determined using the porosity of the preconsolidated soil (n 0 ) assuming no volume changes in the surrounding soil during the vein drilling procedure (Equation 4.9), and the volume of the soil solids (V s(soil) ) can be determined using Equation 4.10: V v(soil) = V soil n 0 V v(soil) = V soil e e 0 (4.9) V s(soil) = V T(soil) V v(soil) (4.10) The void ratio of the soil component of the specimen (e soil ), which should be equal to the initial void ratio of the preconsolidated soil specimen (e 0 = e soil ) can be expressed as: e soil = V v(soil) V s(soil) (4.11) The void ratio due only to the vein void can be expressed separately from the soil: e vein = V vein V s(soil) (4.12) 85

107 The equation for the total void ratio (e) can then be developed from the above equations and given as Equation 4.13: e = V v(total) V s(total) = V v(soil) + V vein V s(soil) e = e soil + e vein (4.13) To normalize void ratio results for specimens with different vein sizes, the void ratio of the soil component (e soil ) is used rather than the total void ratio of the specimen (e). However in natural samples, unless the total vein volume (V vein ) can be estimated then void ratios of the soil (e soil ) and vein components (e vein ) cannot be determined or used. Consolidation Results and Analysis The volumetric strain versus time during isotropic reconsolidation to 100 kpa for the nonhydrate-bearing specimen and the 0.75" and 1" diameter vein-bearing specimens are shown in Figure 4.7. The calculated volumetric strain is greater for the two vein-bearing specimens than the hydrate-free specimen. This is counterintuitive as there is less compressible soil surrounding the relatively incompressible hydrate veins in these specimens. However, as the volume change is estimated using the amount of liquid removed from the specimen during reconsolidation, this suggests that the hydrate veins may undergo a degree of dissolution/dissociation, leading to the production of excess THF liquid and water which is then drained from the specimen along with the pore water. In other words, the assumption that the hydrate vein does not change in volume is likely incorrect ( e vein 0) and may result from dissociation/dissolution of the hydrate vein. This is discussed in greater detail in Section

108 4.3.2 Undrained Shear Compression Results and Analysis Stress-Strain Response Changes in deviatoric stress with increasing axial strain during shear are shown in Figure 4.8, with results clearly showing a significant increase in the stiffness and peak deviatoric stress for hydrate-bearing specimens. The E 50 value is used for the stiffness of the non-hydrate-bearing and 0.75" vein-bearing specimens, while the E sec from 1.2% to 1.9% axial strain is used for the 1"-vein-bearing specimen. This is because the stress-strain curve of the 1"-vein-bearing specimen initially shows a lower stiffness than is seen at higher axial strain, which may be due to a misalignment of the top cap and load ram relative to the vertical hydrate vein, leading to a delayed stress response. Failure was defined at the point of maximum deviatoric stress for all specimens. The stress-strain curve of the specimen with the 1" diameter vein exhibits a significant peak after which strain softening occurs, while a similar peak is not observed for the 0.75" diameter vein-bearing specimen. A stiff, brittle material like THF hydrate hosted within softer, elastoplastic soil would be expected to contribute dramatically to strength and stiffness until the vein structure is fractured or otherwise structurally compromised, leading to a drop in strength as the load is transferred to the soil skeleton, as seen in the 1" diameter vein. The behaviour of the 0.75" diameter vein-bearing specimen challenges this model, indicating that the mechanism by which it contributes to the strength and stiffness may differ. 87

109 Pore Pressure Response The excess pore pressure and pore pressure coefficient (A) versus strain for each test are shown in Figure 4.9a and b respectively. In all three tests the excess pore pressure at failure (u f ) was similar, decreasing from a peak pore pressure prior to failure, indicating a dilative tendency leading up to peak strength, potentially indicating that the soil is mobilized prior to peak strength. The pore pressure parameter at failure (A f ) decreases with increasing vein size, suggesting that the deviatoric stress is not fully felt by the soil and is mainly carried by the vertical hydrate vein prior to failure. Effective Stress Paths Effective stress paths followed by the non-hydrate-bearing and two vein-bearing specimens are shown along with the critical state line obtained from K0CU tests in Figure The presence of the hydrate vein allows the host soil to withstand stress conditions that exceed its critical state. After peak deviatoric stress, the 1"-diameter vein-bearing specimen falls back towards the soil s critical state line, however neither specimen returns to it, implying that the post-peak specimen behaviour is influenced by the hydrate vein. The hydrate-vein-bearing soil behaviour cannot be quantified in terms of effective stress parameters such as the effective friction angle and cohesion due to a lack of test results at higher stress levels. However, there is some indication that both specimens follow increasingly steep stress paths within q-p space with increasing vein diameter before failing, and falling to similar stress conditions above the critical state line of the soil. 88

110 Failure Modes and Post-Shear Analysis Figure 4.11 presents images of the two hydrate-bearing specimens before and after being cut open during post-shear analysis. Comparing the two specimens, it appears that different failure modes may have occurred during shearing, which may explain the difference in behaviour. The 1" diameter vein appears to have fractured horizontally, leading to rotation of the specimen about the fracture point, resulting in shear through the surrounding fine-grained soil as the two vein segments were unable to slide past one another due to the horizontal geometry of the fracture coupled with the THF hydrate strength. This failure mode likely led to a peak deviatoric stress followed by strain softening. In contrast, the 0.75" diameter vein appears to have fractured diagonally, allowing a shear plane to be developed through the soil and hydrate vein, so that the two hydrate segments could slide past one another, resulting in a higher stiffness and peak deviatoric stress than the baseline soil condition, but no distinct peak deviatoric stress. Post-shear analysis of the hydrate vein showed that for the 0.75" and 1"-vein-bearing specimens, 76% and 78% of the THF hydrate vein remained by weight respectively, supporting observations made during reconsolidation that the hydrate underwent dissociation/dissolution, discussed in greater detail in the following section. This is further supported by considering Figure 4.11 where the veins are thinner at the bottom of the specimen than at the top. To account for the change in vein geometry, the volume was calculated by assuming that the vein remained a cylinder and its average diameter was determined by re-measuring the vein at three locations after shearing. For the 0.75" and 1" diameter veins, the average diameters were approximately 16.7 mm (0.67") and 22 mm (0.87") respectively. 89

111 4.3.3 Issues Encountered While the 0.75" and 1"-vein-bearing specimens saw an increase in strength and stiffness due to the hydrate vein presence, specimens with smaller veins (including an additional test with a 0.75" diameter vein) did not. Figure 4.12 shows the deviatoric stress plotted versus axial strain for numerous specimens with hydrate vein diameters of 0.25", 0.50" and 0.75" along with a nonhydrate-bearing specimen. In all cases the soil stiffness is relatively unchanged and the peak deviatoric stress is similar or lower. Potential reasons for this can be determined by examining post-shear images of the specimens in Figure The 0.25" diameter hydrate vein completely disappeared. One 0.50" vein had a single shear band through its bottom third perhaps through a fracture formed prior to shear, while the second 0.50" vein and the 0.75" vein had a more distributed shear zone, and the hydrate had disintegrated into granular pieces. Hydrate dissociation is unlikely to have occurred as specimens were measured to have remained within the hydrate stability field throughout testing (below 2⁰C), therefore the most likely explanation for the hydrate disintegration is its dissolution into the pore water of the soil. If dissolution had occurred, this would have resulted in excess water and THF liquid in the pore space, which would have been drained during the reconsolidation stage (35 to 45 hours). However during shear, dissolution would lead to a weakening of the vein structure allowing the hydrate to fracture more easily, while also leading to an increase in excess pore pressure (as no drainage was allowed), which may have led to a decrease in the peak strength of the soil. Although dissolution was observed in the larger diameter veins (0.75" and 1"), the greater size of these veins seemed to have prevented the integrity of the vein from being significantly affected, and therefore they led to an impact on the strength and stiffness of the soil. 90

112 4.4 Unconsolidated Undrained (UU) Triaxial Compression Testing Due to hydrate dissolution encountered during CU tests, a series of unconsolidated undrained (UU) compression tests were carried out on soil specimens hosting various vein sizes. The goal of these tests was to minimize the time the THF hydrate veins spent within saturated soil specimens thereby limiting hydrate dissolution. UU test results, detailed in Appendix E and summarized in Table 4.5, are used to determine properties such as the undrained shear strength (c u ) and the undrained elastic modulus (E u ) of the hydrate-bearing specimens Pressurization Results and Analysis All specimens were subjected to a cell pressure of 200 kpa that gave rise to an equal rise in pore pressure. Theoretically, when unloading a preconsolidated clayey soil to zero total stress, the pore pressure should become negative (suction), such that the effective stress applied during consolidation is maintained (Head and Epps, 2014a). However, a rise in pore pressures equal to the applied confining stress implies that the effective stress is equal to 0 (u = σ 3 ). This may have occurred because the pore pressure was measured in the base pedestal rather than in the specimen. It is also possible that the dilatant tendency of the soil resulted in stress relief around the outside of the specimen during coring and storage such that the pore pressure response at this boundary was measured rather than the pore pressure within the majority of the specimen Undrained Shear Compression Results and Analysis Data Analysis Techniques The undrained shear strength (c u ) of a specimen can be determined from the peak deviatoric stress (σ 1 σ 3 ) f applied during axial compression, using Equation 4.14: 91

113 c u = (σ 1 σ 3 ) f 2 (4.14) The undrained elastic modulus was approximated by determining the E 50. However, for specimens containing the 0.50" and 0.75" diameter veins the initial stiffness was seen to be relatively low, with an increase in stiffness occurring at approximately 2% and 1% axial strain respectively. It is postulated that this may have been due to misalignment of the top cap on the specimen as observed in CU tests. Therefore, the secant moduli (E sec ) from 2.5% to 3.7%, and 1% to 1.9% axial strain are used as representative values for the elastic moduli for the 0.5" and 0.75" diameter vein-bearing specimens respectively. Stress-Strain Results and Analysis The stress-strain response from UU compression tests on specimens with different hydrate vein diameters are shown in Figure The results show a general increase in c u and stiffness with increasing hydrate vein diameter. The exception to this trend is the 0.25" diameter hydratebearing specimen which has a slightly lower c u and stiffness compared to other specimens. It is hypothesized that this was due to the structural weakness of the 0.25" diameter THF hydrate cylinders, which during specimen preparation proved to be incapable of withstanding the removal of the aluminium foil without fracturing along macroscopic structural defects. This phenomena along with slight hydrate dissolution may have led to a weakening of the specimen. The c u of the non-hydrate-bearing soil specimen (18.5 kpa) can be compared with values from established relationships. Skempton (1957) suggested that for normally consolidated, saturated 92

114 clays, the c u can be related to the effective vertical preconsolidation stress and the plasticity index (PI) using Equation This relationship predicts the c u of the soil to be 17 kpa: c u σ v = (PI) (4.15) A more simple relationship for clay proposed by Mesri (1989), shown as Equation 4.16, predicts the c u of the soil to be 22 kpa: c u σ v = 0.22 (4.16) The undrained shear strength of the non-hydrate-bearing soil (18.5 kpa) matches the two predicted values relatively closely, implying that the effective stress on the specimen was equivalent to the preconsolidation pressure as suggested in Sections and When investigating the strength of saturated normally consolidated clays, the Mohr circles typically give a horizontal failure envelope (φ = 0), and the cohesion intercept is the undrained shear strength. Therefore, an increase in the undrained shear strength with increasing vein size implies the apparent cohesion will increase with increasing vein diameter. Post-Shear Analysis of Failure Modes Figure 4.15 shows images of the exposed hydrate veins in the specimens, highlighting their failure modes. The 0.25" diameter vein is more destructured than the other veins, which is likely due to fracturing along structural asperities and slight hydrate dissolution. The 0.50" and 0.75" diameter veins fractured horizontally in their top and bottom quarters respectively. It is suggested that with increasing axial strain, the top of the specimen rotated about the fracture leading to soil deformation, as the two vein segments were unable to slide past one another due to the horizontal geometry of the fracture coupled with the strength of the THF hydrate. The stronger specimen 93

115 containing the 1" diameter vein (E in Figure 4.15) displayed a similar horizontal rupture, while the weaker 1" diameter vein (D in Figure 4.15) fractured diagonally leading to shear plane development through the fracture, allowing the two vein segments to translate past one another, possibly explaining why their stiffness values are the same but their peak strengths are not. This suggests that the orientation of hydrate vein fractures (horizontal versus inclined) can have a significant impact on the measured undrained shear strength of the specimen, while the location of the fracture is unimportant. Fracture orientations are difficult to predict, as they may occur along the random structural asperities in the THF hydrate cylinders. 4.5 Summary Baseline Geomechanical Testing on Fine-Grained Soil Oedometer consolidation tests led to the determination of the fine-grained soil s consolidation properties. Two specimens were anisotropically consolidated in the triaxial apparatus to the same confining stress (800kPa) and different K values of 0.75 and 0.38, the latter representing K0- consolidation. Undrained shear tests revealed that increasing the axial consolidation stress may inhibit the dilative tendency of the clayey silt. The critical state line of the soil was determined from the K 0.75 and K 0.38 specimens and an isotropically reconsolidated specimen. Summary of CU Triaxial Compression Testing CU compression tests performed on two specimens hosting THF hydrate veins of 0.75" and 1" diameter indicated an increase in the strength and stiffness as compared to non-hydrate-bearing specimens. The vertical THF hydrate veins allowed the soil to withstand stresses exceeding its critical state, and altering the post-peak soil behaviour. The 1" diameter vein had a more 94

116 significant impact on strength than the 0.75" vein, potentially due to the difference in failure mode. Horizontal rupture of the 1" vein may have led to a distinct peak in deviatoric stress and strain softening as the specimen rotated around the fracture, while the 0.75" vein fractured diagonally, exhibiting no peak strength. The pore pressure parameter at failure decreased with increasing vein size, indicating the axial stress was mainly carried by the vertical hydrate vein and not fully felt by the soil. Tests attempted on specimens with smaller vein sizes indicated that hydrate dissolution throughout reconsolidation and shear led to disintegration of the hydrate vein, resulting in little change in sediment stiffness and a reduction in peak strength. Summary of UU Triaxial Compression Testing UU triaxial compression tests were successful in maintaining the structural integrity of the hydrate by minimizing hydrate dissolution into the pore water. Results showed that an increase in hydrate vein diameter resulted in an increase in the undrained strength and stiffness of the specimen. The exception was the 0.25" diameter vein which had no impact on the soil behaviour, likely due to fracturing and minor dissolution of the hydrate. It is suggested that the fracture orientation of hydrate veins can affect the undrained shear strength, with no effect on the stiffness. Horizontally ruptured veins resulted in the highest observed impact on the specimen s undrained shear strength, due to the rotation of the specimen around the fracture. The orientation of apparently randomly occurring asperities observed within the veins may control where the fracture forms, making the undrained shear strength difficult to predict. 95

117 Table 4.1: Summary of results from oedometer tests to 800 kpa vertical pressure on fine-grained soil Sample Name Slurried Soil Preconsolidated Soil 1 Preconsolidated Soil 2 Preconsolidated Soil 3 Soil Type Slurried to over 1.5 LL Consolidated to 100 kpa Consolidated to 100 kpa Consolidated to 100 kpa Initial Void Ratio, e 0 Final Void Ratio, e f Final Saturation, S (%) Compression Index, C c Recompression Index, C r Table 4.2: Summary of results from undrained shear tests on anisotropically consolidated and isotropically reconsolidated fine-grained soil specimens Sample Name Major Effective Stress, σ 1 (kpa) Final Consolidation Results Minor Effective Stress, σ 3 (kpa) K Value, σ 3 / σ 1 Void Ratio, e f Undrained Shear Data at Failure Failure Criterion: Maximum Deviatoric Stress Axial Strain, ε af (%) Deviatoric Stress, q f (kpa) Pore Pressure Parameter, A f Undrained Stiffness, E 50u (MPa) K = K K =

118 Table 4.3: Summary of results from consolidated undrained tests on soil specimen and competent hydrate-vein-bearing specimens Initial Hydrate Vein Diameter (mm/in) After Reconsolidation Undrained Shear Data at Failure and Post-Shear Failure Mode Failure Criterion: Maximum Deviatoric Stress Void Hydrate Axial Pore Undrained Ratio Area Deviatoric Vein Strain, Pressure Vein Stiffness, of Ratio, Stress, q Sat., S Soil, AR vh ε f Failure af Parameter E (kpa) 50u or Mode (%) e (%), A f E secu (MPa) soil 0/ N/A Competent Hydrate Vein-Bearing Specimens 19.05/ / Diagonal Rupture with Shear Band Horizontal Rupture Table 4.4: Summary of results from consolidated undrained tests on non-competent hydratevein-bearing specimens Initial Hydrate Vein Diameter (mm/in) Undrained Shear Data at Failure and Post-Shear Failure Mode Failure Criterion: Maximum Deviatoric Stress Axial Strain, ε af (%) Deviatoric Stress, q f (kpa) Pore Pressure Parameter, A f Undrained Stiffness, E 50u or E secu (MPa) 6.35/ / / / Vein Failure Mode Vein Disappeared Diagonal Rupture with Shear Band Distributed Shear Zone (Granulated Hydrate) Distributed Shear Zone (Granulated Hydrate) 97

119 Table 4.5: Summary of results from unconsolidated undrained tests on soil specimen and hydrate-vein-bearing specimens Hydrate Vein Diameter (mm/in) Specimen Properties Void Ratio of Soil, e soil Area Ratio, AR Hydrate Vein Sat., S vh (%) Undrained Shear Data at Failure and Post-Shear Failure Mode Failure Criterion: Maximum Deviatoric Stress Axial Strain, ε af (%) Deviatoric Stress, q f (kpa) Undrained Shear Strength, c u (kpa) Undrained Stiffness, E 50u or E secu (MPa) Vein Failure Mode 0/ N/A 6.35/ / / / / Vein disintegrated Horizontal Rupture Horizontal Rupture Horizontal Rupture Diagonal Rupture 98

120 A B Figure 4.1: (a) Consolidation data from one oedometer test on slurry and three tests on preconsolidated soil. (b) Data from Preconsolidated Soil 1 test used to verify the preconsolidation pressure (~100 kpa) using the Casagrande Method (Casagrande, 1936). 99

121 A B C D Figure 4.2: Determination of compression and recompression indices from oedometer tests on slurried soil (a) and preconsolidated soil samples (b, c and d). 100

122 Figure 4.3: Effective stress paths followed during anisotropic consolidation tests showing the stress increments applied for K=0.38 and K=0.75 anisotropic consolidations, along with stress levels at which the specimen returned to its original diameter, indicating a K0 value of approximately 0.38 for the soil. 101

123 Figure 4.4: Void ratio versus logarithm of vertical effective stress for oedometer and K0 consolidation tests. The recompression slope during isotropic reconsolidation is greater than seen in oedometer test results, however the soil appears to be less compressible once virgin compression is initiated. 102

124 A B Figure 4.5: (a) Plot of deviatoric stress versus strain for the anisotropically consolidated and isotropically reconsolidated specimens. (b) Similar A f values are observed for the isotropically reconsolidated (to 100 kpa) and K 0.75 specimens, with a lower value for the K 0.38 specimen. 103

125 A B Figure 4.6: (a) Effective stress paths from undrained shear tests on the isotropically reconsolidated specimen and two anisotropically consolidated specimens at the same effective confining pressure (800 kpa), along with derived critical state line. (b) Effective stress paths for undrained shear tests on similar clayey silt (75% Sil-Co-Sil silt and 25% kaolin) on isotropically reconsolidated (T5 and T8) and overconsolidated (T6 and T7) specimens, showing similar dilatant behaviour (Dayarathne and Hawlader, 2015). 104

126 Figure 4.7: Plot of volumetric strain versus square root of time during isotropic reconsolidation of specimens to 100 kpa effective stress. Greater volumetric strain is observed in vein-bearing specimens, which is counterintuitive as these specimens contain less compressible soil, implying the change in volume is due to the dissolution of the THF hydrate vein in addition to soil consolidation. 105

127 Figure 4.8: Deviatoric stress versus axial strain for three soil specimens with two different hydrate vein diameters (0.75" and 1"). The maximum deviatoric strength is chosen as the failure criteria. Specimens display an increase in peak strength and stiffness with increasing hydrate vein diameter. 106

128 A B Figure 4.9: (a) Excess pore pressure and (b) pore pressure coefficient versus axial strain. A decrease in A f is seen with increasing vein diameter. The soil exhibits a dilatant tendency with decreasing pore pressure coefficient after peak, but since the coefficient is never negative the specimen volume does not increase from its original volume. 107

129 Figure 4.10: Deviatoric stress versus mean effective stress, showing the presence of hydrate veins enhances the strength and allows the soil to exceed its critical state. 108

130 A B C D Figure 4.11: Images of 1" (a & b) and 0.75" (c & d) diameter hydrate-vein-bearing specimens post-shear (before and after being cut open) illustrating the differences in their failure modes (blue), the remaining THF hydrate (red) and the disappearance of THF hydrate at the base of the specimens. 109

131 Figure 4.12: Deviatoric stress versus axial strain for hydrate-vein-bearing specimens with diameters of 0.25", 0.50" and 0.75" showing similar stiffness and similar or lower peak deviatoric stress than non-hydrate-bearing soil. 110

132 A B C D Figure 4.13: Post-shear images of exposed hydrate veins for hydrate-vein-bearing specimens with diameters of 0.25" (a), 0.50" (b & c) and 0.75" (d) shown outlined with colours used in stress-strain plot in Figure

133 Figure 4.14: Stress-strain plots from unconsolidated undrained compression tests on specimens containing hydrate veins of different diameters. 112

134 A B C D E Figure 4.15: Images of specimens cut open after compression showing different failure modes. Hydrate veins of 0.25" (a), 0.50" (b), 0.75" (c) and 1" (d & e) diameter shown outlined with colours used in stress-strain plot shown as Figure 4.14, and the shear band through the 1" vein (d) shown in blue. 113

135 Chapter Five: Discussion 5.1 Introduction Analysis of triaxial test results on competent, vertical, cylindrical THF hydrate veins within finegrained specimens presented in the previous chapter led to the general conclusion that hydrate veins increase the strength and stiffness of specimens. This chapter presents relationships based on laboratory results that quantify the influence that simplified THF hydrate veins can have on soil behaviour, and discusses their applicability to determining the impact that gas hydrate veins may have within natural fine-grained sediment. 5.2 Quantifying the Geomechanical Impact of THF Hydrate Veins on Specimens Quantifying the Hydrate Veins To determine the potential relationship between hydrate vein size and specimen behaviour, two different methods of quantifying the hydrate vein size relative to the specimen dimensions are considered, each with merits and limitations Hydrate Vein Saturation Typically the hydrate content of a soil is quantified by the pore space hydrate saturation (S h ), which is the ratio of the hydrate volume (V h ) within the void space of the soil (V v ): S h = V h V v 100% (5.1) However, this definition is typically associated with hydrate that is homogeneously distributed within the pore space. Using this definition of hydrate saturation may lead to confusion, since the tests conducted in this research program were on concentrated, vertical, cylindrical hydrate 114

136 veins. Therefore the saturation of hydrate veins (S vh ) is suggested as an alternative definition. This is calculated by substituting the volume of the vein (V vein ) (equal in this case to the hydrate volume) and the volume of voids in the surrounding soil (V v(soil) ) for the total volume of voids (V v ) into Equation 5.1, as shown in Equation 5.2: S vh = V h V vein + V v(soil) 100% (5.2) The hydrate distribution of fracture-hosted deposits has been seen to be predominantly concentrated within vein structures, with little appreciable hydrate within the void space of the host sediment (Rees et al., 2011). Therefore in this laboratory study the hydrate was entirely concentrated in vein structures, with no hydrate in the surrounding soil. However, this may not necessarily be true for all natural fracture-hosted hydrate deposits Area Ratio As the hydrate veins created for this research are cylinders of constant diameter, the ratio of the cross-sectional vein area (A vein ) to the specimen area (A specimen ) is constant and can be used as a method of defining hydrate content. The calculation for the area ratio (A r ) is: A r = A vein A specimen (5.3) The benefit of this definition is that it is a simple relationship describing the relative areal extent of hydrate veins within the soil, which ranges from 0 in hydrate-free sediment to 1 if the specimen is entirely hydrate. Therefore, it can be used for natural samples where the hydrate volume and/or the soil s void ratio are unknown or variable, but the relative areal proportion of veins can be estimated. Additionally, as hydrate veins appear to dominate the geomechanical 115

137 behaviour, the void ratio of the surrounding soil may lose relevance in relation to strength and stiffness of the specimen. Areal relationships have been successfully employed in defining the contribution of competent cylindrical bodies to a fine-grained soil s geomechanical behaviour, for example stone columns (Barksdale and Bachus, 1983; Priebe, 1995) Relationship between Hydrate Vein Saturation and Area Ratio Since both methods outlined above quantify the hydrate volume within soil, they can be related. The hydrate vein saturation can be expressed in terms of the area ratio for specimens formed in this research as given by Equation 5.4, where n is the soil porosity, the hydrate volume is equal in this case to the vein volume, and the soil and hydrate height are equal (H) as the veins are pervasive: S vh = V h V vein + V v(soil) 100% S vh = A vein H A vein H + n H (A specimen A vein ) 100% 100% S vh = A vein + n A specimen n A vein A vein A vein A vein S vh 100% = 1 n A + 1 n r (5.4) Expressed in terms of the area ratio: A r = 100% S vh n (1 n) (5.5) The relationship between the two methods of defining hydrate saturation may not be applicable if hydrate forms in appreciable quantities within the soil pore space in addition to within veins. 116

138 5.2.2 Quantifying the Impact of Hydrate Veins on Sediment Strength As noted in Chapter 4, the sediment shear strength depends on the failure mechanism of the THF hydrate vein. Of the two mechanisms observed in UU and CU tests, the horizontal rupture of the hydrate vein followed by specimen rotation about the fracture point resulted in the most significant increase in shear strength, and was the most commonly observed failure mode. Therefore, strength relationships are developed in this section that apply to specimens with horizontally fractured THF hydrate veins Undrained Shear Strength Relationships The undrained shear strength of specimens determined from UU tests are shown plotted versus the area ratio and hydrate vein saturation in Figure 5.1a and b respectively. The hydrate veinbearing specimens can be said to follow two distinct behaviours, namely vein sizes that contribute to the specimen strength and those that do not. The solid blue line is a line of best fit for specimens with horizontally fractured veins that led to a significant impact on undrained shear strength, while the red line is drawn through specimens where the veins had no appreciable impact on the shear strength, thus representing the undrained shear strength of the fine-grained soil. The shear strength of the specimen with the diagonally fractured vein falls below the best fit line for the horizontally fractured veins. Undrained Shear Strength in terms of Area Ratio Empirical Relationship Figure 5.1a shows that a linear relationship with a slope of 1350 kpa can be used to empirically relate the undrained shear strength to the area ratio for specimens with horizontally fractured 117

139 hydrate veins that contributed to the specimen strength, while a horizontal line passing through the hydrate-free specimen describes specimens where veins had no impact on the shear strength. Extrapolating these lines from the data points to which they apply (dashed lines in Figure 5.1a) gives an intercept at approximately 0.014, corresponding to a vein diameter of 0.83 cm (0.33"). This value can be conceptualized as a threshold area ratio (A r(thresh) ), below which the hydrate vein has no impact on the undrained shear strength. Presenting this relationship mathematically: If A r (A r(thresh) = 0.014), S u = 18.5kPa If A r > (A r(thresh) = 0.014), S u = 1350kPa A r (5.6) Theoretical Relationship The empirically-derived relationship can be explored from a theoretical perspective to understand its physical meaning. Two different phases of material behaviour are apparent. Below the A r(thresh), the undrained shear strength can be generalized as constant and approximately equal to the undrained shear strength of the soil (S u(soil) ), implying the specimen shear strength is entirely dependent on the soil and the veins provide no shear resistance. This is generalized as: If A r A r(thresh), S u (A r ) = S u(soil) (5.7) For vein sizes above the predicted A r(thresh) the relationship is linear in S u versus A r space, the function for which can be generalized, with m as the slope and the y-intercept equal to 0 (b = 0): If A r > A r(thresh), S u (A r ) = m A r (5.8) If this equation is assumed to apply below the A r(thresh), a y-intercept of zero implies that when no hydrate is present (A vein = 0) the specimen would have no strength (S u = 0), implying that by this equation the specimen strength is entirely a function of the hydrate vein. The assumption 118

140 that the soil has no impact on specimen strength above the threshold area ratio can be tested by normalizing the axial load on the specimen (F specimen ) in terms of only the vein area by setting the soil area equal to zero (A soil = 0), a concept expressed as the vein stress (σ vein ): σ vein = F specimen A vein (5.9) The vein stresses for each UU test where the hydrate vein ruptured horizontally are shown plotted versus axial strain in Figure 5.2. It can be seen that the maximum vein stress is relatively constant, although the 0.50" diameter vein has a slightly higher peak due to seating error and hydrate vein eccentricity relative to the top cap, discussed previously in Section This implies that the load response is entirely controlled by the vertical hydrate vein, as normalizing for the vein area only gives a constant stress value, reinforcing the previous assumption. If the specimen strength is assumed to equal the hydrate vein strength, the maximum vein stress can be assumed to be equal the compressive strength of THF hydrate vein (σ vein(max) = σ ch ). By this, the compressive strength of THF hydrate is estimated to be around 2.7 MPa, averaged from maximum vein stresses. Bending tests carried out on THF hydrate indicate the range of strength values is MPa (Ohmura et al., 2002), broadly including 2.7 MPa. Axial compression tests carried out on identical THF hydrate cylinders of 0.50", 0.75" and 1" diameter gave a very similar peak strength of 2.8 MPa (Wu, personal communication, 2016). The physical meaning of the slope (m) of the undrained shear strength relationship can be further developed using previous relationships. The S u is equal to half the maximum deviatoric stress on the specimen (0.5(σ 1 σ 3 ) max ) so we can substitute this in Equation 5.8 and rearrange: 119

141 0.5(σ 1 σ 3 ) max A specimen = m A vein The maximum deviatoric stress, multiplied by the area on which it acts, is equal to the maximum load on the specimen (F max ) so substituting this into Equation 5.9 we get Equation 5.10: 0.5F max = m A vein m = 0.5F max A vein (5.10) The term to which the slope is equal can be replaced by the vein stress as shown in Equation 5.9, which in turn can be replaced by the compressive strength of the hydrate (σ ch ): m = 0.5σ vein(max) m = 0.5σ ch (5.11) Therefore for this soil, the slope is equal to half the compressive strength of THF hydrate (~1350 kpa). From this, a generalized relationship for the undrained shear strength at area ratios above the predicted threshold area ratio can be expressed as Equation 5.12: If A r > A r(thresh), S u (A r ) = 0.5σ ch A r (5.12) The predicted threshold area ratio value (A r(thresh) ) is calculated as the intercept of the functions defined as Equations 5.12 and 5.7: A r(thresh) = S u(soil) 0.5σ ch (5.13) Therefore, the empirical relationship can be generalized using Equations 5.7, 5.12 and 5.13: If A r S u(soil) 0.5σ ch, S u = S u(soil) (5.14) 120

142 If A r > S u(soil) 0.5σ ch, S u = 0.5σ ch A r Undrained Shear Strength in terms of Hydrate Vein Saturation Theoretical Relationship A theoretical relationship between the undrained shear strength and the hydrate vein saturation can be determined by substituting the equation relating the area ratio to the hydrate vein saturation (Equation 5.5) into the area ratio relationship (Equation 5.14), to give Equation 5.15 which relates the undrained shear strength to the hydrate vein saturation (S vh ) and soil porosity (n), based on a threshold hydrate vein saturation value (S vh(thresh) ): If If S vh 100% 1 n 0.5σ ch S + (1 n) u(soil) S vh 100% > 1 n 0.5σ ch S + (1 n) u(soil), S u = S u(soil) 0.5σ ch n, S u = 100% S (1 n) vh (5.15) Experimental Verification The theoretically derived relationship (Equation 5.15) is applied to experimental specimens using values for the average soil porosity (0.40), the soil s undrained strength (18.5 kpa) and the estimated compressive hydrate strength (2.7 MPa): If S vh 3.36%, S u = 18.5 If S vh > 3.36%, S u = % S vh 0.6 (5.16) 121

143 This relationship fits experimental results fairly well, as shown in Figure 5.1b. The threshold hydrate vein saturation (S vh(thresh) ) of 3.36% corresponds to a hydrate vein diameter of cm (0.32"), similar to the vein diameter determined using the threshold area ratio. Discussion of Hydrate Vein Effect on Specimen Undrained Shear Strength The relationships between hydrate vein size and the undrained shear strength are developed in terms of a predicted threshold area ratio and hydrate vein saturation. However, the idea of a physical threshold vein size is postulated in light of the lack of data on vein sizes in the interval between 0.25" and 0.50" diameter hydrate veins, over which the behaviour is seen to transition from soil to hydrate vein controlled. The simplest method to generalize this is an absolute threshold value at the intercept of the extrapolation of the two best fit lines where the specimen behaviour is predicted to switch from soil to hydrate controlled behaviour. However, this may be an oversimplification, and the behaviour may transition more gradually within this zone. The physical reason for the transition in behaviour between soil and hydrate controlled strength is likely due to the nature of both the fine-grained soil and the hydrate veins, discussed below. The preconsolidated soil specimen has a relatively low undrained shear strength (18.5 kpa) as discussed in Section 4.3.2, due to its high silica silt content and low plasticity. It is suggested that due to the low relative shear strength of the soil as compared to the hydrate vein, and because the lateral effective confining stress of the soil on the vein was not further increased from when the veins were installed, the soil would provide little structural support to the hydrate vein. Therefore, the load applied to the specimen will be mostly carried by the vein until rupture (peak strength), after which it is transferred to the soil. 122

144 The predicted threshold hydrate vein diameter may indicate the size at which the hydrate veins transition from being too slender to support the applied axial stress given their macroscopic structural defects, to being stronger and stiffer than the soil in which they are hosted. Further testing on hydrate-bearing soil should confirm whether this relationship applies for different soil conditions. However, the relationships outlined within this section may serve as a reasonable prediction for the undrained shear strength of vertical, cylindrical THF hydrate veins hosted within soft fine-grained soil at the lateral effective confining stress at which they were formed Shear Strength Relationships from CU Test Results CU compression tests demonstrated that hydrate veins strengthen the soil, however the number of tests was insufficient to allow for the development of a rigorous relationship between effective strength parameters and hydrate vein size. Despite this, general hypotheses are suggested on the basis of one test on the horizontally ruptured 1" diameter hydrate vein. However, as the hydrate vein experienced significant dissolution when reconsolidated (~78% remaining by weight), the area ratio and hydrate vein saturation are calculated using the average diameter of the remaining hydrate vein cylinder (~0.87"). Impact on Deviatoric Stress at Failure The deviatoric stress at failure for CU tests was significantly higher than for UU tests on soil with similar vein sizes as shown in Figure 5.3. This is further demonstrated in Figure 5.4, which shows the increase in peak deviatoric stress with axial strain with different failure modes and at different effective confining stresses on the specimen. As hydrate veins are rigid solids, their compressive strength should not be appreciably enhanced with increasing effective stress. 123

145 However, the strength of the soil specimen is significantly increased due to isotropic reconsolidation, as seen by comparing UU and CU tests on specimens with no hydrate veins (A r = 0). This is because in UU tests, specimens were preconsolidated one-dimensionally under an effective vertical stress of 100 kpa, with the lateral stress determined by the soil s coefficient of earth pressure at rest (K 0 ), predicted to be around 38 kpa. However in CU tests, the specimen was isotropically reconsolidated to 100 kpa all-round effective stress, leading to further soil densification (decrease in void ratio) and a higher undrained shear strength as the soil was consolidated from 38 to 100 kpa effective confining stress in the lateral direction. Results from the CU test suggest that the isotropically consolidated soil provides structural support to the hydrate vein, leading to a higher peak strength than seen in UU tests in which the hydrate vein controls the specimen strength. This may be because the soil applies a greater effective confining pressure to the hydrate vein, frustrating its deformation and resulting in a greater resistance to rupture and relative displacement of the vein segments subsequent to rupture. Therefore, it is predicted that hydrate vein-bearing soil consolidated further laterally may exhibit a hybrid strength behaviour dependent on both the hydrate and soil due to increased bonding at the hydrate-soil interface. As a result, the hydrate-controlled relationships in terms of hydrate vein size developed from UU test results are not applicable to CU test results. The increase in specimen strength between the UU and CU tests could be related to the increase in total stress on the specimen and the hydrate vein. While the strength of a saturated soil is only affected by an increase in effective stress (as theoretically the pore pressure acts outwards on the grains to reduce the total stress on the grain contacts), a solid material such as the hydrate vein 124

146 may increase in strength under increasing total confining stress. The total confining stress imparted to the hydrate vein in the UU test was 200 kpa, while in the CU test the total confining stress on the vein was 500 kpa, the sum of the pore pressure (400 kpa) and the effective stress of the soil (100 kpa). Therefore, the difference in measured strength between UU and CU tests may result from the difference in total applied stresses. However, as the undrained shear strength of just the soil was seen to increase dramatically from 18.5 kpa in the UU test to 68 kpa in the CU test due to an increase in the lateral effective confining stress on the soil, this would likely have a more significant impact on the specimen strength than an increase in total stress on the vein. Potential Impact of Veins on Effective Friction Angle and Cohesion Mohr-Coulomb failure criteria is often used to determine a soil s effective friction angle and cohesion. Pore pressure measurements were used to generate Mohr circles in terms of effective stress (shown in Figure 5.5) for the CU test on the non-hydrate-bearing specimen, and UU and CU tests on the specimen containing the horizontally ruptured ~1" diameter hydrate vein. Previous studies outlined in Chapter 2 suggest that at low hydrate saturations the hydrate does not affect the friction angle, but increases the cohesion of the soil. An effective friction angle of 36⁰ was obtained from baseline tests on non-hydrate-bearing specimens. Applying the same friction angle to the CU test on the ~1" diameter vein would result in an effective cohesion of 138 kpa, as shown in Figure 5.5. The stress conditions at failure for the UU test on the ~1" diameter vein-bearing specimen are also shown, the undrained shear strength is 180 kpa. The effective cohesion is expected to increase with increasing vein size (area ratio/hydrate vein saturation) similar to the undrained shear strength, while the effect of hydrate veins on the 125

147 friction angle cannot be determined using data presented in this thesis. Relationships between the vein size and effective cohesion (and effective friction angle if appropriate) can be developed if CU tests are carried out at different effective stresses and vein sizes Quantifying the Impact of Hydrate Veins on Undrained Stiffness Laboratory results indicate that the undrained stiffness of hydrate-bearing specimens increases with vein diameter and is not dependent on the failure mode, so all test results on competent specimens from UU and CU tests are examined regardless of vein fracture orientation Predicting the Stiffness of a Material using Hookean Springs Hooke s law describes the force (F) required to compress an elastic spring of constant stiffness (k), by a small displacement (dl): F = kdl (5.17) It is assumed that the stiff hydrate vein and the soil respond elastically to small-strain deformation. Since the hydrate veins and soil are continuous over the height of the specimens they can be modelled as two Hookean springs in parallel, and therefore the behaviour of the hydrate-bearing soil can be predicted as one equivalent spring according to the following: k eq = k 1 + k 2 (5.18) dl eq = dl 1 = dl 2 (5.19) F eq = F 1 + F 2 (5.20) The spring constant (k) is related to the Young s modulus (E) of a material by: k = E Area L (5.21) 126

148 Substituting the spring constants for hydrate and soil (k h,soil ) into Equations 5.18 and 5.21, and knowing that the hydrate and soil are equal to the height of the specimen (L specimen = L vein = L soil ) allows the elastic modulus of the composite material to be expressed as: k eq = k h + k soil E eq A specimen L specimen = E ha vein L vein + E soila soil L soil E eq = E ha vein A specimen + E soila soil A specimen (5.22) This relationship assumes that hydrate veins remain sufficiently competent at small diameters to contribute to the stiffness, which was not true for the undrained shear strength of the specimens. While determining the stiffness of THF hydrate outside the scope of investigation for this research, Sloan (1998) estimated THF hydrate stiffness to be ~8.2 GPa, and Ohmura et al. (2002) evaluated the stiffness from bending tests to be GPa Undrained Stiffness versus Area Ratio The Hookean relationship (Equation 5.22) can be expressed in terms of the area ratio as follows: E eq = E ha vein A specimen + E soila soil A specimen E eq = E ha vein A specimen + E s(1 A r )A specimen A specimen E eq = E h A r + E s (1 A r ) E eq = (E h E s )A r + E s (5.23) 127

149 UU Test Results The undrained stiffness values from UU tests are plotted versus the area ratio in Figure 5.6a along with the theoretical relationship based on Hooke s Law. The overall trend is similar to that for undrained shear strength, in that the specimen stiffness is not immediately increased by the hydrate vein presence. The same methodology adopted for the undrained shear strength is applied, by which a threshold area ratio can be predicted by extending the best fit straight lines for the soil and hydrate-controlled stiffness values. The threshold area ratio is 0.020, translating to a hydrate vein diameter of 1 cm (0.4"). The empirically determined relationship between stiffness and area ratio can be presented as: If A r (A r(thresh) = 0.02), E u = 3700kPa If A r > (A r(thresh) = 0.02), E u = kPa A r (5.24) If an area ratio of 1 is substituted into Equation 5.24 (representing an entire specimen of THF hydrate), the undrained stiffness of the THF hydrate would be 185 MPa (equal to the slope). This value is lower than estimated by Sloan (1998) (~8.2 GPa) and slightly below the range determined by Ohmura et al. (2002) using small-strain bending tests ( GPa). However, axial compression tests on identical THF hydrate cylinders with diameters of 0.50", 0.75" and 1" led to the calculation of a very similar large-strain stiffness of approximately 0.23 GPa (Wu, personal communication, 2016). Therefore, 185 MPa is assumed to represent the stiffness of THF hydrate for the purposes of this research. The relationship between the undrained stiffness from UU tests and the area ratio is generalized using methods from Section as Equation 5.25, assuming above the threshold area ratio the stiffness depends entirely on the hydrate vein (using E h = 185MPa): 128

150 If A r E u(soil) E h, E u = E u(soil) If A r > E u(soil) E h, E u = E h A r (5.25) CU Test Results Undrained stiffness values from CU tests are plotted versus area ratio in Figure 5.6b for the two hydrate-bearing specimens along with UU test results and the relationships developed previously. Area ratios are corrected to account for hydrate vein dissolution and so are smaller than their UU test counterparts which began at the same vein size. The undrained soil stiffness (no hydrate) is higher in the CU test (6.2 MPa) than the UU test (3.7 MPa), as the soil has been isotropically consolidated to a lower void ratio as described in Section The CU test data appears to follow the Hookean relationship more closely than the hydrate-controlled relationship. This suggests that when the specimen is isotropically reconsolidated under effective stress, the stiffness can be predicted with reasonable accuracy by using the soil and hydrate stiffness in Equation 5.23 to generate Equation 5.26: E eq = AR (5.26) However, the lack of CU tests on specimens bearing smaller hydrate veins makes it difficult to predict whether this behaviour is followed by hydrate veins with smaller diameters Undrained Stiffness versus Hydrate Vein Saturation The relationship between the area ratio and the hydrate vein saturation previously presented (Equation 5.13) is substituted into the theoretical Hookean relationship for two parallel springs (Equation 5.25) as shown below, and can be compared to results from CU and UU tests: 129

151 If If S vh 100% E u(soil) E h n + E u(soil) (1 n) S vh 100% > E u(soil) E h n + E u(soil) (1 n), E u = E u(soil), E u = 100% S vh E h n (1 n) (5.27) UU Test Results Stiffness is shown with respect to hydrate vein saturation for UU tests in Figure 5.7a along with the two aforementioned relationships. The UU test data follows the threshold hydrate vein saturation theory, below which the specimen stiffness is controlled by the soil, and above it is controlled by the hydrate stiffness. The predicted threshold hydrate vein saturation value of 4.9% translates to a vein diameter of 1 cm (0.4"). Using an average porosity value of 0.40, the soil stiffness and estimated hydrate stiffness, the relationship is shown as Equation 5.28: If S vh 4.9%, E u = E u(soil) If S vh > 4.9%, E u = % S vh 0.6 (5.28) CU Test Results Stiffness is plotted versus the hydrate vein saturation for CU tests in Figure 5.7b along with the aforementioned relationships, predicted using an average reconsolidated porosity value of The stiffness follows the trend predicted by the Hookean springs in parallel. Substituting the averaged porosity value of for specimens and the soil and hydrate stiffness in Equation 5.27, the relationship can be described by Equation 5.29: 130

152 E eq = % S vh (5.29) Discussion The undrained stiffness results from UU tests are seen to follow a transitional soil to hydratecontrolled relationship, while those from CU tests follow the parallel spring theory. The increase in specimen stiffness in CU tests compared to UU tests is likely due to the increased confining pressure on the vein, as the surrounding soil is isotropically consolidated to a greater lateral effective confining stress (100 kpa) than was imparted during one-dimensional consolidation (~38 kpa), and a greater total stress is applied, leading to a hybrid material response as seen with the shear strength determined in CU tests. Conversely, the hydrate-controlled stiffness relationship applies to UU test results, indicating that forming the hydrate vein within soft soil without increasing the lateral confining stress may lead to the hydrate controlling the behaviour. Similar behaviour was observed for the undrained stiffness from UU tests as for the undrained shear strength, however the threshold area ratios and hydrate vein saturations were slightly higher in the undrained stiffness relationships. Theoretically, if the predicted threshold area ratio/hydrate vein saturation value represents an absolute vein size at which the hydrate veins become competent in terms of both stiffness and strength, then this value should be the same. This difference may be due to uncertainty in the experimental measurements, or it could be that the transition from soil to hydrate-controlled behaviour is defined by more of a gradual transition zone over the interval shown by dashed lines on Figures 5.1, 5.3, 5.6 and

153 5.3 Theoretical Geomechanical Impact of Gas Hydrate Veins on Natural Sediment The relationships presented in the previous section relating the undrained shear strength, undrained stiffness and effective strength parameters to hydrate vein size will be discussed with regards to their potential applicability to natural hydrate-bearing sediments Theoretical In-Situ Strength Behaviour The results presented within this thesis suggest that hydrate veins increase the in-situ strength and stiffness of the sediment in which they are hosted, parallel to the direction in which the hydrate veins are aligned, and the increase is directly dependent on the size of the hydrate veins. THF hydrate veins were created within specimens up to an area ratio of 0.13 and 26% hydrate saturation, while fine-grained fracture-hosted hydrate deposits have been seen to be present at average saturation values of 20-30% (Rees et al., 2011), falling in the middle of this range. If gas hydrates are hosted within sub-vertical fractures in soft marine soil (i.e. soil described by Priest et al. (2014)), and have not experienced significant consolidation in the lateral direction after vein formation, relationships in terms of the area ratio and hydrate vein saturation derived from UU test results may apply such that the undrained shear strength can be predicted using Equations 5.14 and 5.15 respectively, and the undrained stiffness using Equations 5.25 and 5.26 respectively. These equations involve the assumption that below a certain vein size the strength and stiffness are controlled by the soil, and above which they are controlled by the hydrate vein. If this proves true for natural sediments, then it is suggested that sediment consolidated onedimensionally to an effective stress of 100 kpa (~20 m below seafloor) with a gas hydrate 132

154 saturation as low as 5% may lead to an increase in undrained shear strength and stiffness, as predicted by the threshold hydrate saturation value. If the hydrate-bearing host soil experiences further isotropic consolidation after vein formation under the overburden pressure to a higher effective confining stress, strength and stiffness relationships derived from CU test results may be more likely to apply. The strength increase of the sediment due to the hydrate veins is expected to be greater than in soft unconsolidated soil due to confining pressure and structural support provided by the soil, requiring further testing to predict. The undrained stiffness may be estimated by using the parallel Hookean spring theory in terms of the area ratio and hydrate vein saturation by using Equations 5.23 and 5.27 respectively. The relationships developed from undrained shear results in terms of the strength and stiffness of sediment may not be applicable to long-term stress changes where the pore pressure can stabilize (e.g. natural slope stability and long-term marine foundation stability). While fracture-hosted hydrate deposits are generally characterized by sub-vertical vein networks, the strength and stiffness of a hydrate-bearing sediment may differ depending on the vein orientations relative to the direction of natural loading, so relationships developed on vein structures parallel to the applied stress may not be applicable to soil in which this is not the case Theoretical In-Situ Consolidation Behaviour The consolidation behaviour of hydrate-bearing fine-grained soil could not be investigated due to the time-sensitive nature of the THF hydrate. However, the geomechanical effect of hydrate veins on the soil determined from triaxial tests can be used to discuss their potential impact on 133

155 the in-situ consolidation behaviour. Marine soils consolidate under self-weight in onedimensional conditions according to their compressibility. Priest et al. (2014) suggest that if subvertical hydrate veins lead to an increase in the stiffness of the soil matrix, this might lead to a significant reduction in the sediment compressibility under vertical loading. Results presented in this thesis indicate that an increase in the undrained stiffness occurs at hydrate saturations of 8% parallel to the orientation of the veins (as low as 5% given the validity of the extrapolated threshold value), which could lead to a decrease in the compressibility of sediment if hydrate veins are formed in this direction within sediment one-dimensionally consolidated to an effective stress of 100 kpa (~20 m below seafloor). The consolidation of a layer of fine-grained soil hosting continuous, interconnected networks of vertical gas hydrate veins of sufficient size to provide an increase in stiffness will be discussed. Figure 5.8 illustrates this schematically, with veins assumed to have formed during continuous sedimentation of the seafloor. Figure 5.9 shows the theoretical one-dimensional stress path this submarine deposit might follow if it was to undergo vertical consolidation due to sedimentation, plotting the void ratio of the soil (excluding the voids hosting hydrate veins) versus pressure. Prior to hydrate formation, the soil follows the normal consolidation line from Point A to Point B as sedimentation leads to an increase in overburden pressure. Hydrate veins form when the soil enters the hydrate stability zone at Point B. During this process the void ratio of the soil (e soil ) is assumed to remain constant (given no porewater outflow to aid in gas hydrate formation), despite a potential increase in the total void ratio (e T ) due to the fracturing of soil and infill of gas hydrate. As sedimentation continues, the increasing vertical stress may now be partially carried 134

156 by the stiff vertical hydrate veins, leading to a reduction in compressibility such that the soil follows a less steep consolidation path from Point B to C. If this is the case, the soil containing vertical hydrate veins (Point C) will have a higher metastable void ratio relative to hydrate-free soil under the same effective stress (Point D); this would make the soil appear underconsolidated for the depth at which it is found, when in reality the soil is fully consolidated under the stress it experiences while the hydrate vein network carries some of the overburden pressure. While the observed underconsolidation of formerly hydrate-vein bearing fine-grained sediments has been attributed to rapid sedimentation of the continental margins or the natural structuration of the soil (Lee et al., 2013; Winters, 2011), the theory of sediment stiffening due to hydrate vein presence in the direction of the general orientation of the veins presents a potential alternative mechanism Theoretical In-Situ Dissociation Behaviour While the dissociation behaviour of hydrate-bearing fine-grained soil was not investigated, the potential in-situ dissociation behaviour can be discussed using the experimentally-determined influence of hydrate veins on the undrained shear behaviour and their postulated effect on the sediment s consolidation behaviour. If the hydrate vein-bearing deposit illustrated in Figure 5.8, is subject to an increase in temperature or decrease in pressure, hydrate dissociation will occur. Strength Change during Dissociation Hydrate vein dissociation will lead to the generation of excess pore pressures within the vein structures, and if the heat transport and/or pressure change processes are relatively fast compared 135

157 to pore pressure dissipation, this may lead to an effective stress reduction within the fracture networks which may present significant zones of weakness along which failure may be initiated. Volume Change after Dissociation The loss of hydrate vein structures due to dissociation will lead to a total void ratio change ( e T ) (Equation 5.30) due to both the collapse of the vein structure ( e vein ) and the loss of the structural support of the veins which results in a change in the soil structure ( e soil ): e T = e soil + e vein (5.30) The void ratio change due to vein void collapse ( e vein ) is not necessarily equal to the vein volume as it is the result of a complex series of events involving fluid volume expansion, cavity closure, and interaction with interconnected veins in the network (Lee et al., 2010). The change in void ratio of the soil ( e soil ) may occur due to the collapse of the soil from the metastable state (Point C) to the expected soil state given the in situ effective stress (Point D), as shown in Figure However, a sudden transfer of overburden stresses to the weak, underconsolidated soil may also induce high pore pressures and exceed the soil s shear strength, leading to further void ratio change as the soil tends towards its critical state. In this case, the void ratio of the soil will fall to the critical state void ratio ( e cs ) for the in situ effective stress after dissociation as shown in Figure 5.10 (Point E) and in the following equation for the void ratio change: e T = e soil + e vein + e cs (5.31) 136

158 Strength Change After Dissociation Following the dissipation of excess pore pressures generated during hydrate dissociation, the shear strength and stiffness of the sediment will be significantly reduced due to the disappearance of the hydrate veins, and will be controlled by in situ effective stress conditions. 5.4 Summary Relationships were established between hydrate vein size and geomechanical behaviour based on laboratory results. Two methods were developed to define the hydrate content of the veinbearing specimens, the area ratio and the hydrate vein saturation. The impact of hydrate veins on the undrained shear strength and stiffness from UU tests are generalized by relationships defined by a threshold area ratio/hydrate vein saturation, below which the undrained strength/stiffness is dependent on the soil and above which it is dependent on the hydrate vein. Data from the CU compression tests indicates the stiffness may follow the Hookean parallel spring theory, while strength data was insufficient in developing relationships between effective shear strength parameters and hydrate vein size. Despite this, it is postulated that the specimen strength is dependent on both the soil and hydrate vein when the soil is consolidated to greater effective confining stress, due to the increased confining pressure on the vein. It is expected that the effective cohesion increases with increasing vein size, while the effect on the effective friction angle is currently not understood. The relationships developed suggest that hydrate veins of increasing size will increase both the undrained shear strength and stiffness of sediment parallel to the direction in which they are aligned. The geomechanical behaviour of one-dimensionally consolidated sediment that has not 137

159 undergone an increase in lateral effective stress after vein formation may be predicted in terms of the area ratio and hydrate vein saturation, the undrained shear strength using Equations 5.14 and 5.15, and the undrained stiffness using Equations 5.25 and 5.26 respectively. If the hydrate-veinbearing soil is further laterally consolidated under overburden pressure, then the strength of the deposits due to the hydrate veins may be greater due to the increased effective confining pressure and structural support provided by the soil, and the undrained stiffness can be predicted using Equations 5.23 and As the compressibility is expected to be lower for soil containing stiff hydrate veins aligned in the direction of one-dimensional loading, its consolidation may result in the host soil having a higher, metastable void ratio than expected for the effective stress at which it is found. The dissociation of hydrate-vein-bearing sediment may lead to significant instability, and is expected to result in volume change due to a collapse of vein voids, a decrease in void ratio from the metastable state to the expected void ratio of the soil, and potentially a decrease in void ratio if the soil s critical state is reached by the transfer of overburden pressure from the vein network to the soil. A reduction in shear strength may occur after the sediment has stabilized postdissociation, due to the disappearance of the strong, stiff vein network. 138

160 A B Figure 5.1: Undrained shear strength from UU tests versus (a) area ratio and (b)hydrate vein saturation. The transition from soil controlled strength behaviour (red) to hydrate vein controlled behaviour (blue) is extrapolated (dashed lines) to predict a threshold value at which the two behaviours transition. 139

161 Figure 5.2: Vein stress (load on specimen divided by hydrate vein area) versus axial strain for horizontally fractured vein-bearing specimens. An approximately constant peak for the three different vein sizes suggests that the soil has little to no impact on the undrained shear strength in UU tests, and that their peaks represent the compressive strength of hydrate which controls the strength behaviour. 140

162 A B Figure 5.3: Deviatoric stress at failure versus (a) the area ratio and (b) hydrate vein saturation for CU and UU tests on specimens. The significant increase in deviatoric stress at failure for veinbearing CU specimens indicates that the strength in CU tests may be influenced by the interaction between the soil and hydrate vein strength. 141

163 Figure 5.4: Deviatoric stress versus axial strain for different tests on specimens with ~1" diameter hydrate veins. Different hydrate vein failure modes for UU tests give rise to differences in peak strength. A much higher peak strength is measured in the CU test, which exceeds the estimated compressive strength of the THF hydrate, indicating that the isotropically reconsolidated soil provides additional strength to the specimen. 142

164 Figure 5.5: Mohr circles of effective stress and Mohr-Coulomb failure envelopes for a CU test on a specimen with no hydrate vein (green) and for a UU test on a specimen with a 1" diameter hydrate vein (purple), as well as a tentative failure envelope for a CU test on a specimen with 1" diameter hydrate vein (dotted red). The failure envelope for the 1" diameter hydrate vein is defined assuming no change in the friction angle but an increase in cohesion. 143

165 A B Figure 5.6: Comparison of undrained stiffness versus area ratio for (a) UU and (b) CU compression tests, showing that UU results follow the hydrate-controlled stiffness relationship after a predicted threshold ratio, while the CU results follow the parallel Hookean spring theory. 144

166 A B Figure 5.7: Comparison of undrained stiffness versus hydrate vein saturation for (a) UU and (b) CU compression tests, showing that UU results follow the hydrate-controlled stiffness relationship after a predicted threshold value while the CU results follow the parallel Hookean spring theory. 145

167 Figure 5.8: Schematic illustration of a layer of fine-grained marine soil containing continuous vertical gas hydrate vein networks of sufficient size to provide an increase in stiffness. Figure 5.9: Theoretical consolidation behaviour of hydrate-bearing fine-grained soil before and after vein formation, resulting in the soil being at a higher metastable void ratio than would be expected at the same in situ effective stress state. 146

168 Figure 5.10: Potential void ratio change due to hydrate dissociation from its metastable state to its expected state given the effective stress conditions on the normal consolidation line (NCL), and potential further collapse to its critical state line (CSL) due to the transfer of overburden pressure from the hydrate vein network to the soil. 147

169 Chapter Six: Summary and Conclusions 6.1 Overview Gas hydrates are ice-like compounds found in deepwater marine sediments and beneath permafrost, strengthening and stiffening the soil in which they form. Hydrates may pose a geohazard during hydrate dissociation since this involves the release of free gas and liquid water into the sediment pore space, potentially leading to sediment failure. Gas hydrates are most abundant within fine-grained sediments, where they form as segregated lenses, nodules, and fracture-filling sub-vertical complex fibrous vein structures. The challenges in recovering intact samples and the difficulty in forming laboratory specimens has limited our understanding of fine-grained hydrate-bearing soils. Determining the geomechanical behaviour of hydrate-veinbearing fine-grained sediments that more closely mimic natural deposits is fundamental to understanding this potential marine geohazard. Therefore the research reported in this thesis set out to address the following question: How do gas hydrate veins influence the geomechanical behaviour of fine-grained sediment? In order to answer this question, the following research objectives were established: 1) establish a simple, repeatable procedure to enable the formation of simplified hydrate vein structures within finegrained soil that resemble naturally-occurring structures; 2) determine the impact of hydrate vein size on the geomechanical behaviour of a specimen under different effective stress conditions, and 3) establish a relationship between hydrate vein size and the resulting geomechanical behaviour of the fine-grained soil in which they are hosted. 148

170 6.2 Summary of Laboratory Program Soil specimens used in the laboratory testing program were prepared by extruding samples from a consolidated (to 100 kpa) mixture of silt-sized silica (65% by weight) and kaolin (35%). A procedure was adopted to form simplified vertical cylinders of tetrahydrofuran (THF) hydrate centred within the specimens, which involved drilling vertical, cylindrical voids within the soil sample and emplacing THF hydrate veins. Specimens were then placed in the triaxial apparatus, which was modified to maintain conditions conducive to THF hydrate stability (<2⁰C). Baseline material properties were established using non-hydrate-bearing soil specimens, including isotropic reconsolidation and anisotropic consolidation followed by undrained shear at different effective stress conditions. Consolidated undrained (CU) compression tests were attempted on specimens containing hydrate veins, however it became apparent that hydrate dissolution into the pore water compromised the structural integrity of the hydrate vein, which was a significant issue that could not be overcome. Therefore, unconsolidated undrained (UU) compression tests were carried out on specimens with different sized hydrate veins. The results from the testing were used to develop relationships quantifying the impact of THF hydrate veins on specimen behaviour, and their applicability to natural fine-grained sediment containing hydrate veins was discussed. 149

171 6.3 Conclusions The following conclusions can be drawn from the theoretical and experimental work carried out: 1. The formation of cylindrical THF hydrate veins within saturated, pre-consolidated finegrained soil specimens can be achieved through a simple, repeatable laboratory procedure that allows rapid geomechanical testing to be carried out. 2. UU compression tests on specimens show that the undrained shear strength and stiffness increase with increasing hydrate vein diameter, with the exception of the 0.25" diameter vein. The results led to the development of relationships that suggest that a threshold vein size exists where the undrained shear strength (of horizontally fractured hydrate veins) and the stiffness of soft soil were entirely soil-controlled below the threshold and transitioned to hydrate-controlled above this threshold. This is postulated to be due to the low soil strength and the low lateral effective confining stress that the soil applies to the hydrate vein, such that when load is applied it is taken up by the hydrate vein. The threshold vein size may either represent an absolute size at which hydrate veins are too slender to support the applied axial stress given their macroscopic structural defects, or may be the product of the limited range of vein sizes tested over a transitional region. 3. CU compression tests on specimens consolidated to an isotropic effective stress of 100 kpa exhibited higher peak strength and stiffness than measured in UU tests. This suggests that increasing the effective and total confining stress after vein formation leads to greater lateral stresses at the hydrate-soil interface and may allow the soil to provide support to the vein, possibly resisting vein deformation, fracture and/or relative vein motion after fracture. The undrained stiffness from CU tests can be predicted by a relationship derived from the 150

172 parallel Hookean spring theory, indicating that the specimen provides a hybrid material response under initial deformation. 4. The orientation of the fracture formed in the THF hydrate vein was found to influence the shear strength in both CU and UU tests, however it had no effect on the undrained stiffness. Veins that ruptured horizontally in UU tests led to the highest undrained shear strength and in the CU test led to a distinct peak strength and strain softening as the specimen rotated around the fracture point. Veins that ruptured diagonally resulted in a lower undrained shear strength. Macroscopic structural weaknesses observed within THF hydrate veins may control where the fracture forms, making the undrained shear strength difficult to predict. 5. It is suggested that hydrate veins of increasing size increase both the shear strength and stiffness of natural sediment parallel to the direction in which they are aligned at fairly low hydrate saturations, which might be predicted using the developed relationships for both one-dimensionally consolidated soil at the stress level at which the hydrate veins formed, as well as sediment consolidated further laterally under overburden pressure. 6. A hypothesis was developed to explain the apparent underconsolidation that has been observed in natural formerly hydrate-vein-bearing soil. It is suggested that the formation of hydrate veins which increase the sediment strength and stiffness would result in a reduction in the sediment compressibility parallel to their orientation, preventing the normal consolidation of the sediment under increasing overburden pressure. This would result in the host soil having a higher, metastable void ratio than expected given the in situ effective stress applied at a given burial depth within the sedimentary column. 7. It was also hypothesized that the dissociation of hydrate veins within natural hydrate-bearing sediments would result in a significant reduction in effective stress within the vein structures 151

173 and sediment, potentially leading to instability. The loss of the vein structures to complete dissociation is expected to lead to significant volume change due to vein void collapse, as well as the transfer of overburden pressure from the vein network to the soil resulting in a significant decrease in void ratio from the metastable state, which could result in further volume change if the critical state of the soil is reached. After dissociation and pore pressure stabilization, the overall strength of the sediment will be reduced due to the disappearance of the strong, stiff vein network. 6.4 Limitations Due to experimental difficulties encountered in the testing of hydrate-vein-bearing fine-grained soil, assumptions and simplifications were made leading to several limitations on the theoretical and experimental work presented in this thesis: Laboratory studies were limited to silty clay consisting of ground silt-sized silica and kaolin, which was of low plasticity (PI of 16), exhibited dilatant behaviour at high axial strains when isotropically reconsolidated, and exhibited a high critical state friction angle (36⁰) which are properties not typical of natural fine-grained marine soils. Soil properties may have led to the weak bonding with the hydrate vein when unconsolidated, such that relationships for unconsolidated soil may not be applicable to typical fine-grained soil. THF was used as the hydrate former, however it has been suggested that it may behave differently from natural gas hydrate. Therefore strength relationships developed in this thesis may not truly represent the behaviour of natural gas hydrate-bearing fine-grained sediments, as fracture orientation was seen to play a significant role in determining strength behaviour and vein fracture may differ for different hydrate vein types. 152

174 The measured strength and stiffness of the composite hydrate-soil material are a function of the testing apparatus and procedure. As the axial loading of the specimen was carried out with a rigid top cap, the material behaviour may significantly differ if vertical loading is applied via a flexible boundary, as the soil will deform more than the stiff hydrate vein. Additionally, since shear tests were strain-controlled, the material behaviour may be affected by the strain rate. Due to the simplification of complex natural vein structures to concentrated cylindrical, vertical veins centred in the middle of the specimen, relationships generated in terms of the area ratio/hydrate vein saturation may not apply to the geomechanical behaviour of samples with thin, dispersed veins of different shapes and sizes, but with the same hydrate volume. Due to the anisotropy of the artificial specimens, relationships are limited in applicability to when hydrate veins are parallel to the principal stress orientation. The stiffness and strength of hydrate-vein-bearing specimens may differ greatly depending on the dominant vein orientation relative to the principal stress direction. Significant difficulty was encountered in maintaining hydrate stability during CU testing due to THF hydrate dissolution into the pore water. Therefore, only a limited number of tests were conducted, making relationships derived from these results speculative in nature. The limited number of tests conducted on specimens with small hydrate vein diameters, coupled with the difficulty in maintaining the stability of small veins may have given rise to the transition zone between soil and hydrate-controlled strength and stiffness behaviour. Four different hydrate vein diameters (0.25", 0.50", 0.75" and 1") were tested in this research. Although these sizes represent the range of hydrate saturations seen in natural fracture-hosted deposits, the developed relationships may not necessarily apply to vein- 153

175 bearing soil with higher hydrate saturations. Additionally, the limited data between hydrate veins of 0.25" and 0.50" diameter means that the transition from soil controlled behaviour to hydrate vein controlled behaviour could not be thoroughly characterized, so was simplified by an extrapolated threshold value. The hydrate was concentrated within discrete veins, and not within the pores of the host soil. Natural hydrate bearing sediments where hydrate is also dispersed within the sediment pore space may increase the strength and stiffness of the surrounding soil, potentially altering the impact of hydrate veins. Geomechanical testing was limited to undrained shear of unconsolidated and reconsolidated specimens, meaning that discussion on the consolidation and dissociation behaviour of hydrate-vein-bearing soil is theoretical, and based on assumptions related to poorlyunderstood natural processes such as hydrate vein formation mechanisms and the nature of seafloor sedimentation processes relative to when hydrate veins are formed. 6.5 Significance and Contributions The fundamental purpose of this research was to investigate the influence of gas hydrate veins on the geomechanical behaviour of fine-grained sediment, which had never been attempted. Therefore, in light of the lack of previous work, an important contribution of this thesis is the development of laboratory procedures for hydrate vein formation within fine-grained sediment. Hydrate-vein-bearing specimens were created with the intention of developing an experimental basis upon which future studies can be undertaken. An innovative geomechanical testing program confirmed for the first time that hydrate veins of increasing size lead to an increase in the strength and stiffness of both unconsolidated and reconsolidated soil. This allowed for 154

176 hypotheses to be developed regarding the potential influence of gas hydrate veins on the behaviour of natural, fine-grained marine sediment layers. Therefore, a significant contribution to understanding the behaviour of natural gas hydrate-bearing fine-grained sediments has been made, increasing our understanding of this important potential unconventional energy resource, potential natural climate change driver and potential geotechnical hazard. 6.6 Recommendations and Future Work The research initiative undertaken has shown that hydrate veins increase the strength and stiffness of the soil in which they are hosted. In addition, relationships were developed by which the geomechanical behaviour can be predicted for cylindrical, THF hydrate veins aligned in the principal stress direction within unconsolidated and reconsolidated soil. However, given the limitations associated with this research highlighted previously, knowledge gaps related to the effect of gas hydrate veins on fine-grained soil behaviour still exist. Therefore, the following recommendations are suggested for further studies: An experimental study on the macroscopic physical behaviour of THF hydrate, such that strength and stiffness properties estimated in this research can be validated and the relative contribution of the hydrate within the soil structure can be better understood. Improvements to the hydrate vein formation process are necessary to prevent dissolution and destructuration of the hydrate. This would allow for a more extensive CU testing program to be conducted at different effective confining stresses to investigate the effect of varying hydrate vein diameters on the effective cohesion and friction angle of the soil, and confirm the undrained stiffness relationship proposed in this research based on limited CU test data. 155

177 A more extensive UU testing program in which the preconsolidation stress of the specimens is increased would allow a greater understanding to be gained of the soil/hydrate-controlled transitional behaviour for both the strength and stiffness of the sediment. Additionally, by investigating hydrate vein sizes between 0.25" and 0.50" diameter, the concept of the threshold value could be further explored. Further triaxial testing programs in which vein orientation, shape, location and dispersion within soil specimens are varied are required to fully understand the impact that heterogeneous hydrate veins may have on the in-situ geomechanical behaviour of the sediment. One-dimensional consolidation testing is also suggested to test the theoretical consolidation behaviour postulated within this thesis on hydrate-vein-bearing specimens, either using zero lateral strain consolidation cells or through K0-consolidation in triaxial cells. Furthermore, dissociation of consolidated hydrate-vein-bearing specimens should be undertaken to better understand how volume and strength changes that may occur within natural deposits, although THF hydrate is not the ideal hydrate type for this form of investigation as it does not dissociate into free gas and water. 156

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189 Appendix A: Material Specification Sheets 168

190 Figure A1: Specification Sheet for EPK Kaolin 169

191 170

192 Figure A2: Specification Sheet for Sil Industrial Minerals Ground Silica Flour 325 Mesh Size 171

193 Appendix B: Oedometer Test Results Table B1: Oedometer consolidation test on Preconsolidated Soil 1 Initial Specimen Properties Cell Number 4 Specimen Density (g/cm 3 ) 1.89 Moisture Content (%) 24 Dry Density (g/cm 3 ) 1.53 Weight of Sample (g) Initial Void Ratio 0.73 Specimen Height (cm) 1.69 Initial Saturation (%) 86.7 Specimen Volume (cm 3 ) Equivalent Height of Solids (cm) 0.98 Particle Density (g/cm 3 ) 2.64 Final Specimen Properties Moisture Content (%) 20 Final Volume (cm 3 ) Weight of Sample (g) Final Density (g/cm 3 ) 2.06 Final Height (cm) 1.50 Final Dry Density (g/cm 3 ) 1.71 Overall Settlement (cm) 0.19 Final Void Ratio 0.54 Volume Change (cm 3 ) 6.02 Final Saturation (%) 99.1 Load-Deformation Data Pressure (kpa) Final Specimen Height (mm) Void Ratio Modulus of Compressibility (m 2 /MN)

194 Table B2: Oedometer consolidation test on Preconsolidated Soil 2 Initial Specimen Properties Cell Number 5 Specimen Density (g/cm 3 ) 1.89 Moisture Content (%) 24 Dry Density (g/cm 3 ) 1.52 Weight of Sample (g) Initial Void Ratio 0.73 Specimen Height (cm) 1.69 Initial Saturation (%) 86.6 Specimen Volume (cm 3 ) Equivalent Height of Solids (cm) 0.98 Particle Density (g/cm 3 ) 2.64 Final Specimen Properties Moisture Content (%) 20 Final Volume (cm 3 ) Weight of Sample (g) Final Density (g/cm 3 ) 2.09 Final Height (cm) 1.48 Final Dry Density (g/cm 3 ) 1.74 Overall Settlement (cm) 0.21 Final Void Ratio 0.52 Volume Change (cm 3 ) 6.52 Final Saturation (%) Load-Deformation Data Pressure (kpa) Final Specimen Height (mm) Void Ratio Modulus of Compressibility (m2/mn)

195 Table B3: Oedometer consolidation test on Preconsolidated Soil 3 Initial Specimen Properties Cell Number 3 Specimen Density (g/cm 3 ) 1.90 Moisture Content (%) 24 Dry Density (g/cm 3 ) 1.53 Weight of Sample (g) Initial Void Ratio 0.72 Specimen Height (cm) 1.69 Initial Saturation (%) 87.9 Specimen Volume (cm 3 ) Equivalent Height of Solids (cm) 0.98 Particle Density (g/cm 3 ) 2.64 Final Specimen Properties Moisture Content (%) 20 Final Volume (cm 3 ) Weight of Sample (g) Final Density (g/cm 3 ) 2.07 Final Height (cm) 1.50 Final Dry Density (g/cm 3 ) 1.73 Overall Settlement (cm) 0.19 Final Void Ratio 0.53 Volume Change (cm 3 ) 6.02 Final Saturation (%) Load-Deformation Data Pressure (kpa) Final Specimen Height (mm) Void Ratio Modulus of Compressibility (m 2 /MN)

196 Table B4: Oedometer consolidation test on Slurried Soil Initial Specimen Properties Cell Number 1 Specimen Density (g/cm 3 ) 1.62 Moisture Content (%) 56 Dry Density (g/cm 3 ) 1.04 Weight of Sample (g) Initial Void Ratio 1.54 Specimen Height (cm) 2.54 Initial Saturation (%) 95.5 Specimen Volume (cm 3 ) Particle Density (g/cm 3 ) 2.64 Equivalent Height of Solids (cm) Final Specimen Properties (From Final Height) Moisture Content (%) 20 Final Volume (cm3) Weight of Sample (g) Final Density (g/cm 3 ) 2.08 Final Height (cm) 1.54 Final Dry Density (g/cm 3 ) 1.74 Overall Settlement (cm) 1.00 Final Void Ratio 0.54 Volume Change (cm 3 ) Final Saturation (%) 97.8 Load-Deformation Data Pressure (kpa) Final Specimen Height (mm) Void Ratio Modulus of Compressibility (m 2 /MN)

197 Appendix C: Anisotropic Consolidation and Undrained Shear Test Results Table C1: Data from anisotropic consolidation and undrained shear of specimen Initial Properties Data from Initial Reconsolidation Consolidation Date (Batch) 31/10/2014 Effective Consolidation Stress (kpa) 100 Shelby Tube Number 2 Change in Volume (%) 4.38 Specimen Height (cm) Final Height (cm) Specimen Diameter (cm) 7.0 Final Diameter (cm) 6.94 Volume of Soil (cm³) Weight (g) Cross-Sectional Area after Reconsolidation (cm 2 ) Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) Wet Density (g/cm³) N/A Reconsolidated Void Ratio 0.60 Water Content (excess material) 25 N/A Void Ratio 0.67 Dry Unit Weight (kn/m³) Saturation (%) 98.4 Volume of Voids (cm³) Volume of Solids (cm³) Minor Effective Stress (kpa) Major Effective Stress (kpa) Anisotropic Consolidation Data Height (cm) Diameter (cm) Volume (cm 3 ) Void Ratio Undrained Shear and Post-Shear Data Failure Criterion: Maximum Deviatoric Stress/Critical State Specimen Properties after Failure Axial Strain (%) 8.5 Average Soil Height (cm) Deviatoric Stress (Corrected) (kpa) 1260 Average Soil Diameter (cm) 7.76 Induced Excess Porewater Pressure (kpa) 427 Weight of Specimen (g) N/A Major Principal Effective Stress (kpa) 1620 Weight of Soil (g) N/A Minor Principal Effective Stress (kpa) 360 Water Content (%) 17 Effective Principal Stress Ratio 4.50 Final Void Ratio (From Reconsolidation) Pore Pressure Parameter at Failure (A f) 0.34 Saturation (%) 97 Undrained Stiffness - (E 50u) (kpa) Undrained Stiffness (E 0.5%) (kpa) Notes Reconsolidated Dry Unit Weight (kn/m³) Attempt at K o-consolidation using circumferential strain gauge

198 Data Plots 177

199 Table C2: Data from K0-consolidation and undrained shear of specimen Initial Properties Data from Initial Reconsolidation Consolidation Date (Batch) 31/10/2014 Effective Consolidation Stress (kpa) Shelby Tube Number 3 Change in Volume (%) 5.90 Specimen Height (cm) Final Height (cm) Specimen Diameter (cm) 7.0 Final Diameter (cm) 6.82 Volume of Soil (cm³) Weight (g) N/A Cross-Sectional Area after Reconsolidation (cm²) Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) Wet Density (g/cm³) N/A Reconsolidated Void Ratio 0.57 Water Content (excess material) 25 Void Ratio 0.67 Dry Unit Weight (kn/m³) Saturation (%) 98.4 Volume of Voids (cm³) Volume of Solids (cm³) Minor Principal Stress (kpa) Major Principal Stress (kpa) Anisotropic Consolidation Data Height (cm) Diameter (cm) Volume (cm 3 ) Void Ratio Failure Criterion: Maximum Deviatoric Stress Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 2.5 Average Soil Height (cm) Deviatoric Stress (Corrected) (kpa) 1495 Average Soil Diameter (cm) 7.96 Induced Excess Porewater Pressure (kpa) 238 Weight of Specimen (g) N/A Major Principal Effective Stress (kpa) 2052 Weight of Soil (g) N/A Minor Principal Effective Stress (kpa) 557 Water Content (%) 16.5 Effective Principal Stress Ratio 3.7 Final Void Ratio (From Reconsolidation) Pore Pressure Parameter at Failure (A f) 0.16 Saturation (%) 98.2 Undrained Stiffness - (E 50u) (kpa) Undrained Stiffness (E 0.5%) (kpa) Notes Reconsolidated Dry Unit Weight (kn/m³) Failure criterion chosen to be maximum deviatoric stress, however critical state was also reached, as indicated on plots 178

200 Data Plots 179

201 Appendix D: Consolidated Undrained Triaxial Test Results Table D1: Data from CU test on specimen with no hydrate vein Prior to Hydrate Vein Formation Initial Specimen Properties After Hydrate Vein Formation Consolidation Date (Batch) 17/08/2015 Specimen Height (cm) N/A Shelby Tube Number 3 Specimen Diameter (cm) N/A Specimen Height (cm) Specimen Volume (cm 3 ) N/A Water Content (excess material from consolidation) 24 Weight of Soil (g) N/A Void Ratio 0.63 Weight of THF Hydrate (g) N/A Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) N/A Saturation (%) Volume of Voids in Soil (cm³) N/A Volume of Voids (cm³) Volume of Voids including vein (cm³) N/A Volume of Solids (cm³) Volume of Solids (cm³) N/A Consolidation Stage Initial Pore Pressure (kpa) Time to 100% Primary Reconsolidation (min.) Effective Consolidation Stress (kpa) Isotropic Coefficient of Consolidation (C vi) (m 2 /year) Change in Volume (%) 5.25 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.56 Final Height (cm) Reconsolidated Void Ratio 0.54 Final Diameter (cm) 6.94 Reconsolidated Saturation (from final water content) (%) 98.9 Cross-Sectional Area after Reconsolidation (cm²) Failure Criterion: Maximum Deviatoric Stress/Critical State Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 12 Average Soil Height (cm) N/A Deviatoric Stress (Corrected) (kpa) 136 Average Soil Diameter (cm) N/A Induced Excess Porewater Pressure (kpa) 47 Weight of Specimen (g) N/A Major Principal Effective Stress (kpa) 182 Weight of THF Hydrate (g) N/A Minor Principal Effective Stress (kpa) 46 Weight of Soil (g) N/A Effective Principal Stress Ratio 4 Water Content (%) 20 Pore Pressure Parameter at Failure (A f) 0.34 Final Void Ratio (From Reconsolidation) 0.54 Undrained Stiffness - (E 50u) (kpa) 6180 Reconsolidated Dry Unit Weight (kn/m³) Undrained Stiffness (E 0.5%) (kpa) Notes Some axial strain applied accidentally at start of isotropic reconsolidation, barreling failure mode 180

202 Data Plots and Post-Shear Pictures [No Pictures] 181

203 Table D2: Data from CU test on specimen with 0.75" diameter hydrate vein Prior to Hydrate Vein Formation Initial Specimen Properties After Hydrate Vein Formation Consolidation Date (Batch) 17/08/2015 Specimen Height (cm) Shelby Tube Number 3 Specimen Diameter (cm) 7.00 Specimen Height (cm) Specimen Volume (cm 3 ) Specimen Diameter (cm) 7.0 Vein Height (cm) Volume of Soil (cm³) Vein Diameter (cm) 1.91 Weight (g) 1100 Vein Volume (cm³) Wet Density (g/cm³) 2.00 Soil Volume (cm³) Water Content (excess material from consolidation) 25 Weight of Soil (g) Void Ratio 0.65 Weight of THF Hydrate (g) Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) Saturation (%) Volume of Voids in Soil (cm³) Volume of Voids (cm³) Volume of Voids including vein (cm³) Volume of Solids (cm³) Volume of Solids (cm³) Consolidation Stage Initial Pore Pressure (kpa) 498 Time to 100% Primary Reconsolidation (min.) Effective Consolidation Stress (kpa) 100 Isotropic Coefficient of Consolidation (C vi) (m 2 /year) Change in Volume due to hydrate dissolution and consolidation (%) 6.72 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.67 Final Height (cm) Reconsolidated Void Ratio 0.57 Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) 98.0 Cross-Sectional Area after Reconsolidation (cm²) Failure Criterion: Maximum Deviatoric Stress Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 6.3 Average Soil Height (cm) Deviatoric Stress (Corrected) (kpa) 245 Average Soil Diameter (cm) 7.30 Induced Excess Porewater Pressure (kpa) 41 Weight of Specimen (g) Major Principal Effective Stress (kpa) 297 Weight of THF Hydrate (g) Minor Principal Effective Stress (kpa) 53 Weight of Soil (g) Effective Principal Stress Ratio 5.7 Water Content (%) 21 Pore Pressure Parameter at Failure (A f) 0.17 Final Void Ratio (From Reconsolidation) 0.57 Undrained Stiffness - (E 50u) (kpa) Reconsolidated Dry Unit Weight (kn/m³) Undrained Stiffness (E 0.5%) (kpa) Notes Significant amount of hydrate dissolved at bottom (~76% by weight) so hydrate vein approximately 1.67 cm in diameter (leading to a reduced specimen area), failure mechanism appears to be diagonal shear plane through vein 182

204 Data Plots and Post-Shear Pictures 183

205 Table D3: Data from CU test on specimen with 1" diameter hydrate vein Prior to Hydrate Vein Formation Initial Specimen Properties After Hydrate Vein Formation Consolidation Date (Batch) 17/08/2015 Specimen Height (cm) Shelby Tube Number 4 Specimen Diameter (cm) 7.00 Specimen Height (cm) Specimen Volume (cm3) Specimen Diameter (cm) 7.0 Vein Height (cm) Volume of Soil (cm³) Vein Diameter (cm) 2.54 Weight (g) 1109 Vein Volume (cm³) Wet Density (g/cm³) 1.95 Soil Volume (cm³) Water Content (excess material from consolidation) 24 Weight of Soil (g) Void Ratio 0.69 Weight of THF Hydrate (g) Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) Saturation (%) 93.8 Volume of Voids in Soil (cm³) Volume of Voids (cm³) Volume of Voids including vein (cm³) Volume of Solids (cm³) Volume of Solids (cm³) Consolidation Stage Initial Pore Pressure (kpa) 497 Time to 100% Primary Reconsolidation (min.) Effective Consolidation Stress (kpa) 98 Isotropic Coefficient of Consolidation (C vi) (m 2 /year) Change in Volume due to hydrate dissolution and consolidation (%) 6.68 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.68 Final Height (cm) Reconsolidated Void Ratio 0.60 Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) 93 Cross-Sectional Area after Reconsolidation (cm²) Failure Criterion: Maximum Deviatoric Stress Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 4.5 Average Soil Height (cm) Deviatoric Stress (Corrected) (kpa) 609 Average Soil Diameter (cm) 7.55 Induced Excess Porewater Pressure (kpa) 51 Weight of Specimen (g) Major Principal Effective Stress (kpa) 653 Weight of THF Hydrate (g) Minor Principal Effective Stress (kpa) 44 Weight of Soil (g) Effective Principal Stress Ratio 14.9 Water Content (%) 21 Pore Pressure Parameter at Failure (A f) 0.08 Final Void Ratio (From Reconsolidation) 0.60 Undrained Stiffness - (E 50u) (kpa) Reconsolidated Dry Unit Weight (kn/m³) Secant Stiffness (E sec) (1.2% to 1.9% strain) (kpa) Notes Hydrate dissolved at bottom of vein (around 78% by weight remaining) so hydrate vein approximately 2.2 cm in diameter (leading to a reduced specimen area), took until 1.2% strain for hydrate strength to mobilize 184

206 Data Plots and Post-Shear Pictures 185

207 Table D4: Data from CU test on specimen with 0.25" diameter hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation Date (Batch) 11/09/2015 Specimen Height (cm) Shelby Tube Number 2 Specimen Diameter (cm) 7.00 Specimen Height (cm) Specimen Volume (cm3) Specimen Diameter (cm) 7.0 Vein Height (cm) Volume of Soil (cm³) Vein Diameter (cm) 0.64 Weight (g) 1106 Vein Volume (cm³) 4.54 Wet Density (g/cm³) 2.00 Soil Volume (cm³) Water Content (excess material from consolidation) 25 Weight of Soil (g) Void Ratio 0.65 Weight of THF Hydrate (g) 4.53 Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) Saturation (%) Volume of Voids in Soil (cm³) Volume of Voids (cm³) Volume of Voids including vein (cm³) Volume of Solids (cm³) Volume of Solids (cm³) Consolidation Stage Initial Pore Pressure (kpa) Time to 100% Primary Reconsolidation (min.) Effective Consolidation Stress (kpa) 105 Isotropic Coefficient of Consolidation (C vi) (m 2 /year) Change in Volume (%) 7.92 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.75 Final Height (cm) Reconsolidated Void Ratio 0.53 Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) N/A Cross-Sectional Area after Reconsolidation (cm²) Undrained Shear and Post-Shear Data Failure Criterion: Maximum Deviatoric Stress Specimen Properties after Failure Axial Strain (%) 6.5 Average Soil Height (cm) N/A Deviatoric Stress (Corrected) (kpa) 106 Average Soil Diameter (cm) N/A Induced Excess Porewater Pressure (kpa) 65 Weight of Specimen (g) N/A Major Principal Effective Stress (kpa) 130 Weight of THF Hydrate (g) N/A Minor Principal Effective Stress (kpa) 24 Weight of Soil (g) N/A Effective Principal Stress Ratio 5.4 Water Content (%) N/A Pore Pressure Parameter at Failure (A f) 0.61 Final Void Ratio (From Reconsolidation) 0.53 Undrained Stiffness - (E 50u) (kpa) 9640 Reconsolidated Dry Unit Weight (kn/m³) Undrained Stiffness (E 0.5%) (kpa) 9640 Notes No data after shear; significant cell pressure oscillation during shear. No hydrate left 186

208 Data Plots and Post-Shear Pictures 187

209 Table D5: Data from CU test on specimen with 0.50" diameter hydrate vein Prior to Hydrate Vein Formation Initial Specimen Properties After Hydrate Vein Formation Consolidation Date (Batch) 17/08/2015 Specimen Height (cm) Shelby Tube Number 1 Specimen Diameter (cm) 7.00 Specimen Height (cm) Specimen Volume (cm 3 ) Specimen Diameter (cm) 7.0 Vein Height (cm) Volume of Soil (cm³) Vein Diameter (cm) 1.27 Weight (g) 1132 Vein Volume (cm³) Wet Density (g/cm³) 1.99 Soil Volume (cm³) Water Content (excess material from consolidation) 24 Weight of Soil (g) Void Ratio 0.66 Weight of THF Hydrate (g) Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) Saturation (%) 98.3 Volume of Voids in Soil (cm³) Volume of Voids (cm³) Volume of Voids including vein (cm³) Volume of Solids (cm³) Volume of Solids (cm³) Consolidation Stage Initial Pore Pressure (kpa) 494 Time to 100% Primary Reconsolidation (min.) Effective Consolidation Stress (kpa) 97 Isotropic Coefficient of Consolidation (C vi) (m 2 /year) Change in Volume (%) 7.81 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.81 Final Height (cm) Reconsolidated Void Ratio 0.52 Final Diameter (cm) 6.81 Reconsolidated Saturation (from final water content) (%) 110 Cross-Sectional Area after Reconsolidation (cm²) Failure Criterion: Maximum Deviatoric Stress Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 5.7 Average Soil Height (cm) Deviatoric Stress (Corrected) (kpa) 106 Average Soil Diameter (cm) 7.67 Induced Excess Porewater Pressure (kpa) 67 Weight of Specimen (g) Major Principal Effective Stress (kpa) 134 Weight of THF Hydrate (g) 0 Minor Principal Effective Stress (kpa) 28 Weight of Soil (g) Effective Principal Stress Ratio 4.8 Water Content (%) 22 Pore Pressure Parameter at Failure (A f) 0.63 Final Void Ratio (From Reconsolidation) 0.52 Undrained Stiffness - (E 50u) (kpa) 9680 Reconsolidated Dry Unit Weight (kn/m³) Undrained Stiffness (E 0.5%) (kpa) 9680 Notes Volumetric change calculated from axial and radial strain (measured with gauge), as back piston position was not logged, so is an approximation. 188

210 Data Plots and Post-Shear Pictures 189

211 Table D6: Data from CU test on specimen with 0.50" diameter hydrate vein Prior to Hydrate Vein Formation Initial Specimen Properties After Hydrate Vein Formation Consolidation Date (Batch) 09/11/2015 Specimen Height (cm) Shelby Tube Number 4 Specimen Diameter (cm) 7.00 Specimen Height (cm) Specimen Volume (cm 3 ) Specimen Diameter (cm) 7.0 Vein Height (cm) Volume of Soil (cm³) Vein Diameter (cm) 1.27 Weight (g) Vein Volume (cm³) Wet Density (g/cm³) 2.02 Soil Volume (cm³) Water Content (excess material from consolidation) 25 Weight of Soil (g) Void Ratio 0.64 Weight of THF Hydrate (g) Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) Saturation (%) Volume of Voids in Soil (cm³) Volume of Voids (cm³) Volume of Voids including vein (cm³) Volume of Solids (cm³) Volume of Solids (cm³) Consolidation Stage Initial Pore Pressure (kpa) 500 Time to 100% Primary Reconsolidation (min.) Effective Consolidation Stress (kpa) 81 Isotropic Coefficient of Consolidation (C vi) (m 2 /year) Change in Volume (%) 7.04 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.87 Final Height (cm) Reconsolidated Void Ratio 0.52 Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) 110 Cross-Sectional Area after Reconsolidation (cm²) Failure Criterion: Maximum Deviatoric Stress/Critical State Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 8.9 Average Soil Height (cm) Deviatoric Stress (Corrected) (kpa) 134 Average Soil Diameter (cm) 7.35 Induced Excess Porewater Pressure (kpa) 45 Weight of Specimen (g) Major Principal Effective Stress (kpa) 159 Weight of THF Hydrate (g) Minor Principal Effective Stress (kpa) 25 Weight of Soil (g) Effective Principal Stress Ratio 6.36 Water Content (%) 22 Pore Pressure Parameter at Failure (A f) 0.33 Final Void Ratio (From Reconsolidation) 0.52 Undrained Stiffness - (E 50u) (kpa) 7000 Reconsolidated Dry Unit Weight (kn/m³) Undrained Stiffness (E 0.5%) (kpa) Notes Reconsolidation met with issues which is suspected to be due to hydrate vein dissolution leading to incomplete consolidation; Hydrate vein offered no strength 190

212 Data Plots and Post-Shear Pictures 191

213 Table D7: Data from CU Test on specimen with 0.75" diameter hydrate vein Prior to Hydrate Vein Formation Initial Specimen Properties After Hydrate Vein Formation Consolidation Date (Batch) 09/11/2015 Specimen Height (cm) Shelby Tube Number 1 Specimen Diameter (cm) 7.00 Specimen Height (cm) Specimen Volume (cm 3 ) Specimen Diameter (cm) 7.0 Vein Height (cm) Volume of Soil (cm³) Vein Diameter (cm) 1.91 Weight (g) 1097 Vein Volume (cm³) Wet Density (g/cm³) 1.99 Soil Volume (cm³) Water Content (excess material from consolidation) 25 Weight of Soil (g) Void Ratio 0.66 Weight of THF Hydrate (g) Dry Unit Weight (kn/m³) Weight of Soil and Hydrate (g) Saturation (%) Volume of Voids in Soil (cm³) Volume of Voids (cm³) Volume of Voids including vein (cm³) Volume of Solids (cm³) Volume of Solids (cm³) Consolidation Stage Initial Pore Pressure (kpa) 481 Time to 100% Primary Reconsolidation (min.) N/A Effective Consolidation Stress (kpa) 93 Isotropic Coefficient of Consolidation (C vi) (m 2 /year) N/A Change in Volume (%) 4.03 Isotropic Coefficient of Compressibility (m vi) (m 2 /MN) 0.43 Final Height (cm) Reconsolidated Void Ratio 0.59 Final Diameter (cm) 6.90 Reconsolidated Saturation (from final water content) (%) N/A Cross-Sectional Area after Reconsolidation (cm²) Failure Criterion: Maximum Deviatoric Stress/Critical State Undrained Shear and Post-Shear Data Specimen Properties after Failure Axial Strain (%) 4.2 Average Soil Height (cm) N/A Deviatoric Stress (Corrected) (kpa) 107 Average Soil Diameter (cm) N/A Induced Excess Porewater Pressure (kpa) 52 Weight of Specimen (g) N/A Major Principal Effective Stress (kpa) 129 Weight of THF Hydrate (g) N/A Minor Principal Effective Stress (kpa) 22 Weight of Soil (g) N/A Effective Principal Stress Ratio 5.9 Water Content (%) N/A Pore Pressure Parameter at Failure (A f) 0.49 Final Void Ratio (From Reconsolidation) 0.59 Undrained Stiffness - (E 50u) (kpa) 5050 Reconsolidated Dry Unit Weight (kn/m³) Undrained Stiffness (E 0.5%) (kpa) 6550 Notes No data taken after shear; reconsolidation not complete due to suspected hydrate dissociation 192

214 Data Plots and Post-Shear Pictures 193

215 Appendix E: Unconsolidated Undrained Triaxial Test Results Table E1: Data from UU test on specimen with no hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation 08/10/2015 Specimen Height (cm) N/A Specimen Height (cm) Specimen Diameter (cm) N/A Specimen Diameter (cm) 7.00 Specimen Volume (cm 3 ) N/A Volume of Soil (cm³) Vein Height (cm) N/A Void Ratio 0.67 Vein Diameter (cm) N/A Weight (g) 1139 Vein Volume (cm³) N/A Dry Unit Weight (kn/m³) Soil Volume (cm³) N/A Water Content (excess material) 26 Weight of Soil (g) N/A Saturation Weight of THF Hydrate (g) N/A Volume of Voids (cm³) Weight of Soil and Hydrate (g) N/A Volume of Solids (cm³) Volume of Voids in Soil (cm³) N/A Wet Density (g/cm³) 2.00 Volume of Voids including vein (cm³) N/A Volume of Solids (cm³) N/A Hydrate Saturation including vein (%) N/A Undrained Shear and Post-Shear Data At Specimen Failure Specimen Properties after Failure Axial Strain (%) 12.0 Average Soil Height (cm) 9.65 Deviatoric Stress (kpa) 37 Average Soil Diameter (cm) N/A Excess Porewater Pressure (kpa) -1 Vein Height (cm) Major Principal Total Stress (kpa) 237 Weight of Specimen (g) Minor Principal Total Stress (kpa) 200 Weight of THF Hydrate (g) N/A Undrained Stiffness - (E 50u) (kpa) 3600 Weight of Soil (g) Undrained Stiffness (E 0.5%) (kpa) 3700 Water Content (%) 26 Notes Baseline test Data Plots and Post-Shear Pictures 194

216 Table E2: Data from UU test on specimen with 0.25" diameter hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation 02/10/2015 Specimen Height (cm) Specimen Height (cm) Specimen Diameter (cm) 7.00 Specimen Diameter (cm) 7.0 Specimen Volume (cm 3 ) Volume of Soil (cm³) Vein Height (cm) Void Ratio 0.70 Vein Diameter (cm) 0.64 Weight (g) 1135 Vein Volume (cm³) 4.75 Dry Unit Weight (kn/m³) Soil Volume (cm³) Water Content (excess material) 26 Weight of Soil (g) Saturation 99.7 Weight of THF Hydrate (g) 4.74 Volume of Voids (cm³) Weight of Soil and Hydrate (g) Volume of Solids (cm³) Volume of Voids in Soil (cm³) Wet Density (g/cm³) 1.97 Volume of Voids including vein (cm³) Volume of Solids (cm³) Hydrate Saturation including vein (%) 1.98 Undrained Shear and Post-Shear Data At Specimen Failure Specimen Properties after Failure Axial Strain (%) 14.2 Average Soil Height (cm) Deviatoric Stress (kpa) 33 Average Soil Diameter (cm) N/A Excess Porewater Pressure (kpa) 2 Vein Height (cm) Major Principal Total Stress (kpa) 233 Weight of Specimen (g) Minor Principal Total Stress (kpa) 200 Weight of THF Hydrate (g) 4.64 Undrained Stiffness - (E 50u) (kpa) 3200 Weight of Soil (g) Undrained Stiffness (E 0.5%) (kpa) 3100 Water Content (%) 24 Notes 195

217 Data Plots and Post-Shear Pictures 196

218 Table E3: Data from UU test on specimen with 0.50" diameter hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation 02/10/2015 Specimen Height (cm) Specimen Height (cm) Specimen Diameter (cm) 7.00 Specimen Diameter (cm) 7.0 Specimen Volume (cm 3 ) Volume of Soil (cm³) Vein Height (cm) Void Ratio 0.68 Vein Diameter (cm) 1.27 Weight (g) 1145 Vein Volume (cm³) Dry Unit Weight (kn/m³) Soil Volume (cm³) Water Content (excess material) 26 Weight of Soil (g) Saturation Weight of THF Hydrate (g) Volume of Voids (cm³) Weight of Soil and Hydrate (g) Volume of Solids (cm³) Volume of Voids in Soil (cm³) Wet Density (g/cm³) 1.98 Volume of Voids including vein (cm³) Volume of Solids (cm³) Hydrate Saturation including vein (%) 7.8 Undrained Shear and Post-Shear Data At Specimen Failure Specimen Properties after Failure Axial Strain (%) 4.6 Average Soil Height (cm) Deviatoric Stress (kpa) 105 Average Soil Diameter (cm) N/A Excess Porewater Pressure (kpa) 1 Vein Height (cm) Major Principal Total Stress (kpa) 305 Weight of Specimen (g) Minor Principal Total Stress (kpa) 200 Weight of THF Hydrate (g) Undrained Stiffness - (E 50u) (kpa) 2000 Weight of Soil (g) Secant Stiffness (2.5% to 3.7%) (E sec) (kpa) 3900 Water Content (%) 24 Notes Secant tangent calculated as vein strength took until 2.5% strain to mobilize 197

219 Data Plots and Post-Shear Pictures 198

220 Table E4: Data from UU test on specimen with 0.75" diameter hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation 02/10/2015 Specimen Height (cm) Specimen Height (cm) Specimen Diameter (cm) 7.00 Specimen Diameter (cm) 7.0 Specimen Volume (cm 3 ) Volume of Soil (cm³) Vein Height (cm) Void Ratio 0.74 Vein Diameter (cm) 1.91 Weight (g) 1101 Vein Volume (cm³) Dry Unit Weight (kn/m³) Soil Volume (cm³) Water Content (excess material) 26 Weight of Soil (g) Saturation 94.1 Weight of THF Hydrate (g) Volume of Voids (cm³) Weight of Soil and Hydrate (g) Volume of Solids (cm³) Volume of Voids in Soil (cm³) Wet Density (g/cm³) 1.99 Volume of Voids including vein (cm³) Volume of Solids (cm³) Hydrate Saturation including vein 15.4 Undrained Shear and Post-Shear Data At Specimen Failure Specimen Properties after Failure Axial Strain (%) 2.3 Average Soil Height (cm) Deviatoric Stress (kpa) 183 Average Soil Diameter (cm) N/A Excess Porewater Pressure (kpa) 9 Vein Height (cm) Major Principal Total Stress (kpa) 384 Weight of Specimen (g) Minor Principal Total Stress (kpa) 200 Weight of THF Hydrate (g) Undrained Stiffness - (E 50u) (kpa) 8200 Weight of Soil (g) Secant Stiffness (1% to 1.9%) (E sec) (kpa) Water Content (%) 25 Notes Some seating issues at start of shear, secant modulus used instead 199

221 Data Plots and Post-Shear Pictures 200

222 Table E5: Data from UU test on specimen with 1" diameter hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation 02/10/2015 Specimen Height (cm) Specimen Height (cm) Specimen Diameter (cm) 7.00 Specimen Diameter (cm) 7.0 Specimen Volume (cm3) Volume of Soil (cm³) Vein Height (cm) Void Ratio 0.74 Vein Diameter (cm) 2.54 Weight (g) 1061 Vein Volume (cm³) Dry Unit Weight (kn/m³) Soil Volume (cm³) Water Content (excess material) 26 Weight of Soil (g) Saturation 94.5 Weight of THF Hydrate (g) Volume of Voids (cm³) Weight of Soil and Hydrate (g) Volume of Solids (cm³) Volume of Voids in Soil (cm³) Wet Density (g/cm³) 1.92 Volume of Voids including vein (cm³) Volume of Solids (cm³) Hydrate Saturation including vein (%) 26 Undrained Shear and Post-Shear Data At Specimen Failure Specimen Properties after Failure Axial Strain (%) 1.5 Average Soil Height (cm) Deviatoric Stress (kpa) 360 Average Soil Diameter (cm) N/A Excess Porewater Pressure (kpa) 5 Vein Height (cm) Major Principal Total Stress (kpa) 560 Weight of Specimen (g) Minor Principal Total Stress (kpa) 200 Weight of THF Hydrate (g) Undrained Stiffness - (E 50u) (kpa) Weight of Soil (g) Undrained Stiffness (E 0.5%) (kpa) Water Content (%) 25 Notes Vein strength took until around 0.5% strain to mobilize 201

223 Data Plots and Post-Shear Pictures 202

224 Table E6: Data from UU test on specimen with 1" diameter hydrate vein Initial Specimen Properties Prior to Hydrate Vein Formation After Hydrate Vein Formation Consolidation 02/10/2015 Specimen Height (cm) Specimen Height (cm) Specimen Diameter (cm) 7.00 Specimen Diameter (cm) 7.0 Specimen Volume (cm3) Volume of Soil (cm³) Vein Height (cm) Void Ratio 0.74 Vein Diameter (cm) 2.54 Weight (g) 1012 Vein Volume (cm³) Dry Unit Weight (kn/m³) Soil Volume (cm³) Water Content (excess material) 26 Weight of Soil (g) Saturation Weight of THF Hydrate (g) Volume of Voids (cm³) Weight of Soil and Hydrate (g) Volume of Solids (cm³) Volume of Voids in Soil (cm³) Wet Density (g/cm³) 1.94 Volume of Voids including vein (cm³) Volume of Solids (cm³) Hydrate Saturation including vein (%) 26 Undrained Shear and Post-Shear Data At Specimen Failure Specimen Properties after Failure Axial Strain (%) 1.2 Average Soil Height (cm) Deviatoric Stress (kpa) 235 Average Soil Diameter (cm) N/A Excess Porewater Pressure (kpa) 3 Vein Height (cm) Major Principal Total Stress (kpa) 435 Weight of Specimen (g) Minor Principal Total Stress (kpa) 200 Weight of THF Hydrate (g) Undrained Stiffness - (E 50u) (kpa) Weight of Soil (g) Undrained Stiffness (E 0.5%) (kpa) Water Content (%) 25 Notes Vein fractured diagonally 203

225 Data Plots and Post-Shear Pictures 204