The Influence of Tetrahydrofuran Hydrate Veins on Fine ...

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The Influence of Tetrahydrofuran Hydrate Veins on

Fine-Grained Soil Behaviour

Smith, William

Smith, W. (2016). The Influence of Tetrahydrofuran Hydrate Veins on Fine-Grained Soil Behaviour

(Unpublished master's thesis). University of Calgary, Calgary, AB. doi:10.11575/PRISM/28218

http://hdl.handle.net/11023/2969

master thesis

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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

ii

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 fine-

grained 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.

iii

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.

iv

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, Jon2, 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 (1970-2016)

v

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 ..................................................................................1

1.1 Statement of Problem .................................................................................................1 1.2 Research Question .....................................................................................................3 1.3 Objectives of Thesis ...................................................................................................3

1.4 Scope of Thesis ..........................................................................................................4 1.5 Outline of Thesis ........................................................................................................4

CHAPTER TWO: LITERATURE REVIEW ......................................................................6

2.1 Introduction to Gas Hydrates .....................................................................................6 2.1.1 Formation and Stability Conditions ...................................................................6

2.1.2 Global Distribution ............................................................................................7 2.1.3 Significance .......................................................................................................8

2.2 Natural Gas Hydrate Formation and Morphology ...................................................12

2.2.1 Gas Availability and Migration .......................................................................12 2.2.2 Host Sediment and Hydrate Mode of Occurrence ...........................................13

2.3 Laboratory Formation Techniques of Hydrate-Bearing Sediment ..........................16

2.3.1 Dissolved Gas Method ....................................................................................17

2.3.2 Partial Water Saturation Method .....................................................................17 2.3.3 Hydrate Pre-mixing Method ............................................................................18

2.3.4 Analog Hydrate (Tetrahydrofuran) ..................................................................19 2.4 Previous Work on Geomechanical Behaviour of Hydrate-Bearing Sediment ........20

2.4.1 Strength Properties ..........................................................................................20

2.4.2 Consolidation Behaviour .................................................................................23 2.4.3 Dissociative Behaviour ....................................................................................26

2.5 Summary ..................................................................................................................27

CHAPTER THREE: EXPERIMENTAL PROCEDURE ..................................................42

3.1 Introduction ..............................................................................................................42

3.2 Materials ..................................................................................................................43

3.2.1 Fine-Grained Soil ............................................................................................43 3.2.2 Synthetic Hydrate ............................................................................................44

3.3 Specimen Preparation ..............................................................................................46 3.3.1 Soil Specimen Preparation ..............................................................................46 3.3.2 Hydrate Vein Formation within Soil ...............................................................48

3.3.2.1 Vein Void Formation .............................................................................48 3.3.2.2 In Situ Formation Method ......................................................................49 3.3.2.3 Transfer Method ....................................................................................50

3.3.2.4 Method Selection ...................................................................................51

vi

3.4 Baseline Geomechanical Testing on Fine-Grained Soil ..........................................52

3.4.1 Oedometer Consolidation Tests ......................................................................52 3.4.2 K0-Consolidation and Undrained (K0CU) Compression Tests .......................53

3.4.2.1 Geomechanical Testing Apparatus ........................................................53

3.4.2.2 Specimen Mounting and Cell Assembly ...............................................54 3.4.2.3 K0-Consolidation ...................................................................................55 3.4.2.4 Undrained Shear ....................................................................................57

3.5 Geomechanical Testing on Hydrate-Bearing Soil ...................................................58 3.5.1 Specimen Mounting and Cell Assembly .........................................................58

3.5.2 Consolidated Undrained (CU) Triaxial Compression Testing ........................59 3.5.3 Unconsolidated Undrained (UU) Triaxial Compression Testing ....................60

CHAPTER FOUR: LABORATORY RESULTS AND ANALYSIS ...............................77

4.1 Introduction ..............................................................................................................77 4.2 Baseline Geomechanical Testing on Fine-Grained Soil ..........................................77

4.2.1 Oedometer Consolidation Tests ......................................................................77

4.2.2 K0-Consolidation and Undrained (K0CU) Compression Tests .......................78 4.3 Consolidated Undrained (CU) Compression Testing ..............................................84

4.3.1 Isotropic Reconsolidation Results and Analysis .............................................85 4.3.2 Undrained Shear Compression Results and Analysis .....................................87 4.3.3 Issues Encountered ..........................................................................................90

4.4 Unconsolidated Undrained (UU) Triaxial Compression Testing ............................91 4.4.1 Pressurization Results and Analysis ................................................................91

4.4.2 Undrained Shear Compression Results and Analysis .....................................91 4.5 Summary ..................................................................................................................94

CHAPTER FIVE: DISCUSSION ....................................................................................114 5.1 Introduction ............................................................................................................114

5.2 Quantifying the Geomechanical Impact of THF Hydrate Veins on Specimens ....114 5.2.1 Quantifying the Hydrate Veins ......................................................................114

5.2.1.1 Hydrate Vein Saturation ......................................................................114

5.2.1.2 Area Ratio ............................................................................................115 5.2.1.3 Relationship between Hydrate Vein Saturation and Area Ratio ..........116

5.2.2 Quantifying the Impact of Hydrate Veins on Sediment Strength ..................117 5.2.2.1 Undrained Shear Strength Relationships .............................................117

5.2.2.2 Shear Strength Relationships from CU Test Results ...........................123 5.2.3 Quantifying the Impact of Hydrate Veins on Undrained Stiffness ...............126

5.2.3.1 Predicting the Stiffness of a Material using Hookean Springs ............126 5.2.3.2 Undrained Stiffness versus Area Ratio ................................................127 5.2.3.3 Undrained Stiffness versus Hydrate Vein Saturation ..........................129 5.2.3.4 Discussion ............................................................................................131

5.3 Theoretical Geomechanical Impact of Gas Hydrate Veins on Natural Sediment .132

5.3.1 Theoretical In-Situ Strength Behaviour .........................................................132 5.3.2 Theoretical In-Situ Consolidation Behaviour ................................................133 5.3.3 Theoretical In-Situ Dissociation Behaviour ..................................................135

5.4 Summary ................................................................................................................137

vii

CHAPTER SIX: SUMMARY AND CONCLUSIONS ..................................................148

6.1 Overview ................................................................................................................148 6.2 Summary of Laboratory Program ..........................................................................149 6.3 Conclusions ............................................................................................................150

6.4 Limitations .............................................................................................................152 6.5 Significance and Contributions ..............................................................................154 6.6 Recommendations and Future Work .....................................................................155

REFERENCES ................................................................................................................157

APPENDIX A: MATERIAL SPECIFICATION SHEETS .............................................168

APPENDIX B: OEDOMETER TEST RESULTS ..........................................................172

APPENDIX C: ANISOTROPIC CONSOLIDATION AND UNDRAINED SHEAR TEST

RESULTS ...............................................................................................................176

APPENDIX D: CONSOLIDATED UNDRAINED TRIAXIAL TEST RESULTS .......180

APPENDIX E: UNCONSOLIDATED UNDRAINED TRIAXIAL TEST RESULTS ..194

viii

List of Tables

Table 3.1: Characteristics of natural hydrate-bearing soils and prepared soil for this research ... 61

Table 3.2: Data from plastic limit determination on prepared soil using ASTM D4318 ............. 61

Table 3.3: Preliminary tests in the development of the THF hydrate formation procedure ......... 62

Table 3.4: Preliminary tests in the development of the in situ vein formation procedure ............ 63

Table 4.1: Summary of results from oedometer tests to 800 kPa vertical pressure on fine-

grained soil ............................................................................................................................ 96

Table 4.2: Summary of results from undrained shear tests on anisotropically consolidated and

isotropically reconsolidated fine-grained soil specimens ..................................................... 96

Table 4.3: Summary of results from consolidated undrained tests on soil specimen and

competent hydrate-vein-bearing specimens .......................................................................... 97

Table 4.4: Summary of results from consolidated undrained tests on non-competent hydrate-

vein-bearing specimens ......................................................................................................... 97

Table 4.5: Summary of results from unconsolidated undrained tests on soil specimen and

hydrate-vein-bearing specimens ........................................................................................... 98

Table B1: Oedometer consolidation test on Preconsolidated Soil 1 ........................................... 172



Table B4: Oedometer consolidation test on Slurried Soil .......................................................... 175

Table C1: Data from anisotropic consolidation and undrained shear of specimen .................... 176

Table C2: Data from K0-consolidation and undrained shear of specimen .................................. 178

Table D1: Data from CU test on specimen with no hydrate vein ............................................... 180

Table D2: Data from CU test on specimen with 0.75" diameter hydrate vein ........................... 182

Table D3: Data from CU test on specimen with 1" diameter hydrate vein ................................ 184




ix

Table D7: Data from CU Test on specimen with 0.75" diameter hydrate vein .......................... 192

Table E1: Data from UU test on specimen with no hydrate vein ............................................... 194

Table E2: Data from UU test on specimen with 0.25" diameter hydrate vein ........................... 195



Table E5: Data from UU test on specimen with 1" diameter hydrate vein ................................ 201

Table E6: Data from UU test on specimen with 1" diameter hydrate vein ................................ 203

x

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). .............. 30

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

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

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

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

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). ................................................................................................................... 35

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

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). ................ 36

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

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

xi

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

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). .................................................................................................................................... 39

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

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

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

Figure 3.1: Flowchart summarizing the testing procedure adopted for this research program

including specimen preparation, baseline testing and geomechanical testing program. ...... 64

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

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

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

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

xii

Figure 3.6: Hydraulic jack used to extrude cylindrical consolidated soil specimens from 70

mm internal diameter sampling tube (left). ........................................................................... 69

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

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. .................................................................... 71

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

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

Figure 3.11: Aluminium foil mold containing a 0.25" hydrate cylinder, which proved

impossible to unwrap without fracturing into segments. ...................................................... 74

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

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

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). ..................................... 100

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

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

Figure 4.5: (a) Plot of deviatoric stress versus strain for the anisotropically consolidated and

isotropically reconsolidated specimens. (b) Similar 𝐴𝑓 values are observed for the

xiii

isotropically reconsolidated (to 100 kPa) and 𝐾0.75 specimens, with a lower value for the

𝐾0.38 specimen. ................................................................................................................... 103

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

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

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

Figure 4.9: (a) Excess pore pressure and (b) pore pressure coefficient versus axial strain. A

decrease in 𝐴𝑓 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

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

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

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

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 4.12. ............................................................................. 111

Figure 4.14: Stress-strain plots from unconsolidated undrained compression tests on

specimens containing hydrate veins of different diameters. ............................................... 112

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

xiv

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

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

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

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. ........................................... 141

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

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

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

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

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. .. 146

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

xv

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

Figure A1: Specification Sheet for EPK Kaolin ......................................................................... 169

Figure A2: Specification Sheet for Sil Industrial Minerals Ground Silica Flour 325 Mesh Size171

xvi

List of Symbols, Abbreviations and Nomenclature

𝐴 Pore pressure coefficient

𝐴𝑓 Pore pressure coefficient at failure

𝐴𝑝 Area of piston

𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) Threshold hydrate vein area ratio

𝐴𝑟 Hydrate vein area ratio

𝐴𝑠𝑜𝑖𝑙 Cross-sectional area of soil

𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 Cross-sectional area of specimen

𝐴𝑣𝑒𝑖𝑛 Cross-sectional area of hydrate vein

𝐶 Circumference

𝐶′ Effective cohesion

𝐶0 Initial circumference

𝐶𝑐 Compression index

𝐶𝑟 Recompression index

𝑐𝑢 Undrained shear strength

C4H8O Tetrahydrofuran

CH4 Methane

CK0U K0-consolidated undrained

CO2 Carbon dioxide

CU Consolidated undrained

𝐸 Young’s modulus

𝐸0.5% Secant Young’s modulus to 0.5% Axial Strain

xvii

𝐸50 Young’s modulus at 50% of failure stress

𝐸ℎ Young’s modulus of hydrate

𝐸𝑒𝑞 Equivalent Young’s modulus

𝐸𝑖 Initial tangent Young’s modulus

𝐸𝑠𝑒𝑐 Secant Young’s modulus

𝐸𝑠𝑜𝑖𝑙 Young’s modulus of soil

𝐸𝑢(𝑠𝑜𝑖𝑙) Undrained stiffness of soil

𝐸𝑢 Undrained elastic modulus

𝑒 Void ratio

𝑒0 Initial void ratio

𝑒1𝑘𝑃𝑎 Void ratio at 1 kPa on critical state line

𝑒𝑐𝑠 Critical state void ratio

𝑒𝑓 Void ratio after each consolidation stage

𝑒𝑠𝑜𝑖𝑙 Void ratio of hydrate-bearing soil component

𝑒𝑣𝑒𝑖𝑛 Void ratio of vein void

𝐹 Force

𝐹𝑒𝑞 Equivalent force

𝐹𝑚𝑎𝑥 Maximum axial load on the specimen

𝐹𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 Axial load on the specimen

GIP Gas-in-place

𝐻 Height

𝐻0 Initial height

xviii

H2S Hydrogen sulfide

JGS Japanese Geotechnical Society

𝐾 Stress ratio

𝐾0 Coefficient of lateral earth pressure at rest

𝐾0(𝑁𝐶) Coefficient of lateral earth pressure at rest for normally

consolidated soil

𝑘 Spring constant

𝑘ℎ Spring constant of hydrate

𝑘𝑒𝑞 Equivalent spring constant

𝑘𝑠𝑜𝑖𝑙 Spring constant of soil

KG Krishna-Godavari

𝐿 Length

𝐿𝑒𝑞 Equivalent length

𝐿𝑠𝑜𝑖𝑙 Length of soil

𝐿𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 Length of specimen

𝐿𝑣𝑒𝑖𝑛 Length of hydrate vein

LL Liquid limit

LVDT Linear voltage displacement transducer

𝑀 Slope of critical state line in 𝑞-𝑝′ space

𝑀𝑔 Methane gas concentration

𝑀𝑔𝑠𝑙 Methane gas solubility limit

𝑚 Slope of best fit line

Ma Million years ago

xix

𝑛0 Initial porosity

N2 Nitrogen

𝑟0 Initial radius

𝑂𝐶𝑅 Overconsolidation ratio

𝑃 Position

𝑝′ Mean effective stress

PI Plasticity index

PL Plastic limit

𝑞 Deviatoric stress

𝑟 Radius

𝑆ℎ Hydrate saturation

𝑆𝑢 Undrained shear strength

𝑆𝑢(𝑠𝑜𝑖𝑙) Undrained shear strength of soil

𝑆𝑣ℎ(𝑡ℎ𝑟𝑒𝑠ℎ) Threshold hydrate vein saturation

𝑆𝑣ℎ Hydrate vein saturation

THF Tetrahydrofuran

𝑢 Pore pressure

𝑢0 Initial pore pressure

𝑢𝑎 Average pore pressure

𝑢𝑏 Pore pressure at top of sample

𝑢𝑐 Pore pressure at base of sample

𝑢𝑓 Pore pressure at failure

xx

USCS Unified Soil Classification System

UU Unconsolidated undrained

𝑉 Volume

𝑉0 Initial volume

𝑉ℎ Hydrate volume

𝑉𝑇(𝑠𝑜𝑖𝑙) Total soil volume of the specimen

𝑉𝑠(𝑠𝑜𝑖𝑙) Volume of solids within host soil

𝑉𝑠 Volume of soil solids

𝑉𝑣(𝑠𝑜𝑖𝑙) Volume of voids within host soil

𝑉𝑣 Volume of voids

𝑉𝑣𝑒𝑖𝑛 Vein volume

𝜀1 Major principal strain

𝜀3 Minor principal strain

𝜀𝑎 Axial strain

𝜀𝑟 Radial strain

𝜀𝑟𝑔𝑎𝑢𝑔𝑒 Radial strain measured using circumferential strain gauge

𝜀𝑣 Volumetric strain

𝜀𝑉𝑙𝑒𝑛𝑠 Volumetric strain due to hydrate structure collapse

𝜀𝑉𝑝𝑠

Volumetric strain due to effective stress changes involved

with the depressurization production method

𝜀𝑉𝑡ℎ𝑎𝑤 Volumetric strain due to hydrate dissociation

𝜀𝑉𝑡𝑜𝑡 Total volumetric strain due to hydrate dissociation

𝜋 Pi

xxi

𝜎 Total stress

𝜎′ Effective stress

𝜎1 Total major principal stress

𝜎1′ Effective major principal stress

𝜎3 Total minor principal stress

𝜎3′ Effective minor principal stress

𝜎𝑐ℎ Compressive strength of THF hydrate

𝜎′𝑣 Vertical effective stress

𝜎′𝑣𝑐 Past maximum vertical effective stress or

preconsolidation pressure

𝜎𝑣𝑒𝑖𝑛 Vein stress

𝜎𝑣𝑒𝑖𝑛(𝑚𝑎𝑥) Maximum vein stress

𝜎′𝑣𝑜 Current vertical effective stress

(𝜎1 − 𝜎3)𝑚𝑎𝑥 Maximum deviatoric stress

𝜏′ Effective shear strength at failure

𝜙′ Effective friction angle

𝜙′𝑐𝑠

Critical state friction angle

𝜑 Angle of dilation

1

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).

2

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 fine-

grained 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 fine-

grained sediments. Studies on the geomechanical properties of natural gas hydrates within fine-

grained 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-

3

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.

4

1.4 Scope of Thesis

The research question will be addressed by carrying out a laboratory investigation of artificially-

formed 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 hydrate-

bearing 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.

5

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.

6

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).

2.1.1 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 (𝑀𝑔) exceeds the solubility limit (𝑀𝑔𝑠𝑙), 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 (𝑀𝑔 < 𝑀𝑔𝑠𝑙) then dissolution

of the hydrate crystal occurs, involving only a small volume increase (Lu et al., 2008; Sultan et

al., 2004).

7

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 (𝑀𝑔 > 𝑀𝑔𝑠𝑙). If 𝑀𝑔 falls

below 𝑀𝑔𝑠𝑙, hydrate dissolution will occur until the gas concentration increases such that 𝑀𝑔 =

𝑀𝑔𝑠𝑙, 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).

2.1.2 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

8

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.

2.1.3 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 1.2×1015 to 1.2×1017

m3 (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

9

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

54.95 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

10

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

11

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

12

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

2.2.1 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

13

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.

2.2.2 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

14

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%

15

(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 fine-

grained 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

16

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 best-

documented 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.

17

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.

2.3.2 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

18

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.

2.3.3 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.

19

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.

20

2.4 Previous Work on Geomechanical Behaviour of Hydrate-Bearing Sediment

2.4.1 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).

21

Previous Work on Hydrate-Bearing Coarse-Grained Sediments

Numerous studies have been carried out on laboratory-formed and naturally-occurring hydrate-

bearing 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 2.11. However, the effect of the distribution habit decreases with

increasing hydrate saturation.

22

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.

23

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-hydrate-

bearing 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.

2.4.2 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 one-

dimensional 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 (𝐶𝑐) 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

24

vertical strain to be recovered. The slope of the swelling line in 𝑒 versus log𝜎′𝑣 space is termed

the recompression index (𝐶𝑟), 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

25

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 well-

developed 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.

26

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 (𝜀𝑉𝑡𝑜𝑡) when small shear strains

are expected (e.g. around production wells on level ground), dependent on the volumetric strain

27

due to hydrate dissociation (𝜀𝑉𝑡ℎ𝑎𝑤), hydrate structure collapse (𝜀𝑉

𝑙𝑒𝑛𝑠) and effective stress

changes due to the depressurization production method (𝜀𝑉𝑝𝑠

):

𝜀𝑉𝑡𝑜𝑡 = 𝜀𝑉

𝑡ℎ𝑎𝑤 + 𝜀𝑉𝑙𝑒𝑛𝑠 + 𝜀𝑉

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:

𝜀𝑉𝑡𝑜𝑡 =

𝑒𝑐𝑠 − 𝑒0

1 + 𝑒0=

(𝑒1𝑘𝑃𝑎 − 𝜆 log𝑝′

1𝑘𝑃𝑎) − 𝑒0

1 + 𝑒0 (2.3)

Where 𝑒𝑐𝑠 is the critical state void ratio after dissociation, 𝑒0 is the initial void ratio, 𝑒1𝑘𝑃𝑎 and 𝜆

are critical state parameters and 𝑝′ is the final mean effective stress after dissociation.

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.

28

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%,

29

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.

30

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).

A

B

31

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).

32

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).

33

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).

34

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).

35

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).

A B C

A B

C D

36

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).

A

B C

37

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).

A

B

38

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).

A

B

39

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).

A B

A B

40

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).

41

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).

A

B

C

42

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 near-

vertical 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

(𝜎3) axisymmetrically (𝜀2 = 𝜀3 = 𝜀𝑟), 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 3.1.

43

3.2 Materials

3.2.1 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 hydrate-

bearing 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

44

(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.

3.2.2 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.

45

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 3.4.

7) 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.

46

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).

3.3.1 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 20.55 cm

internal diameter stainless steel tube held between two metal plates housing porous metal discs,

47

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, 8 14×1cm 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

48

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.

3.3.2 Hydrate Vein Formation within Soil

Two hydrate vein formation processes were used in this research, and are outlined in Sections

3.3.2.2 and 3.3.2.3. 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 3.3.2.1.

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 well-

ventilated locations according to Material Safety Data Sheet specifications.

3.3.2.1 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"), 19.05 mm (¾") and 25.4 mm (1") diameter1 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.

49

hours). Prior to hydrate vein emplacement the void was re-cored, as minor soil migration and

sloughing into the void occurred during re-cooling.

3.3.2.2 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

50

(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 30-75 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 3.10. 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.

3.3.2.3 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.

51

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 3.11. 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.

3.3.2.4 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.

52

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.

3.4.1 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 5-3400 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

53

slurry. The cell was dismantled quickly after releasing the final load, the specimen was removed

and its mass, height and water content were determined.

3.4.2 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.

3.4.2.1 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 3.13. 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

54

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.

3.4.2.2 Specimen Mounting and Cell Assembly

The soil specimen was prepared using the procedure outlined in Section 3.3.1. 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

55

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.

3.4.2.3 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 (𝜎1) was increased

incrementally while adjusting the radial stress (𝜎3) by changing cell pressure in response to

specimen deformation to maintain the lateral strain equal to zero (𝜀𝑟 ≈ 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 (𝐴𝑝):

∆𝑉 = ∆𝑃 × 𝐴𝑝 (3.1)

This was used along with the initial volume (𝑉0) to calculate the volumetric strain (𝜀𝑣):

𝜀𝑣 =

∆𝑉

𝑉0× 100% (3.2)

The axial strain (𝜀𝑎) was determined using the ram displacement (∆𝐻) and the initial height (𝐻0):

56

𝜀𝑎 =

∆𝐻

𝐻0× 100% (3.3)

Therefore, assuming small deformations, the radial strain (𝜀𝑟) was calculated independent of the

radial strain gauge according to:

𝜀𝑟 = 𝜀𝑣 − 𝜀𝑎

2 (3.4)

The circumferential strain gauge measured the change in circumference (∆𝐶) of the specimen,

allowing a direct measurement of the radial strain (𝜀𝑟𝑔𝑎𝑢𝑔𝑒) according to the following equation:

𝜀𝑟𝑔𝑎𝑢𝑔𝑒 =

∆𝑟

𝑟0=

(∆𝐶/𝜋)/2

(𝐶0/𝜋)/2=

∆𝐶

𝐶0 (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 4.2.2

Vertical stress increments were kept relatively small (∆𝜎𝑣 = 0.2𝜎𝑣′) 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 0.001 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 (𝑢𝑎) was calculated using the pressure transducer readings at the

base (𝑢𝑐) and the top of the sample (𝑢𝑏) using the following equation:

57

𝑢𝑎 =

2

3𝑢𝑐 +

1

3𝑢𝑏 (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: |𝜀𝑟| < 0.05%. However, for the purposes of this research

it was considered that a sample had been K0-consolidated if |𝜀𝑟| < 0.5% due to the uncertainty

associated with the radial strain measurements and calculations.

3.4.2.4 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 (0.0166 and 0.0134 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.

58

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.

3.5.1 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 de-

aired 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 (1oC) 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.

59

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 3.4.2.3).

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 3.4.2.4 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.

60

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 3.5.2 for CU tests.

61

Table 3.1: Characteristics of natural hydrate-bearing soils and prepared soil for this research

Characteristics

Krishna-

Godavari

Basin1

Ulleung

Basin2

Gulf

of Mexico1

Prepared

Fine-Grained

Soil

Average Sand (% by weight) 5 0 1 1

Average Silt (% by weight) 55 80 22 64

Average Clay (% by weight) 40 20 77 35

Liquid limit range 70-98 12-129 N/A 34

Plastic limit range 33-49 17-88 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 (%)

1 18.2

2 18.2

3 17.6

4 16.0

Average Plastic Limit 18

62

Table 3.3: Preliminary tests in the development of the THF hydrate formation procedure

Test Objective

Hydrate Molar

Ratio

THF:H2O

Conclusions

THF Volatilization Test

1 Define volatilization rate of

pure THF liquid

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 pre-

cooling 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 2nd test tube 1:15 Formation Time: 16 hours

7 Formation at -20⁰C 1:15 Formation Time: 1.5 hours

Hydrate Storage Test

8 Volume loss after storing

hydrate in fridge 1:15

Storage Time: 720 hours

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

THF/Water 1:15

Complete dissolution of 5 ml hydrate after 20

hours


Water 1:13

18 ml hydrate becomes 20 ml hydrate after 36

hours


Water 1:15

18 ml hydrate dissociates into water then reforms

at top of cylinder, 9 ml THF at bottom and 7 at

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

63

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/H2O 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 1 Hydrate ~1 cm below top of sample, refilled with THF-

water mixture and reformed fully competent vein

4 Form 0.50" Vein at -20⁰C

Storage test over 6 days 1.17

3 days – no change

6 days - hydrate disappeared

Could be due to dissolution into porewater or fridge

warming to above 4⁰C

5

Form 0.50" Vein at -20⁰C

Observe hydrate vein

competency

1.25 Water around base after overnight storage in fridge

Vein fairly intact (See Figure 3.9).

6

Form 0.50" Vein at -20⁰C

Check for ice lensing after full

vein formation

1.25

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.

64

Slurry Fine-Grained Soil Mixture Oedometer on Slurry

Consolidate Soil to 100 kPa

Core Soil Specimens

Oedometer on

Consolidated Soil

Extrude Specimen K0-Consolidation and

Undrained Shear Tests

Experimental Procedure Baseline Testing Testing Program

Drill 0.25" Vein

Void in Soil

In Situ Hydrate Vein

Formation Method

Drill 0.50, 0.75, 1"

Vein Void in Soil

Hydrate Vein

Transfer Method

CU/UU Tests on

Non-Hydrate-Bearing

Specimens

Modified Specimen Mounting and

Cell Assembly (10 minutes)

Pressurize Specimen

Reconsolidate Specimen

Undrained Shear (UU)

Undrained Shear (CU)

Figure 3.1: Flowchart summarizing the testing procedure adopted for this research program including

specimen preparation, baseline testing and geomechanical testing program.

65

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).

66

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.

67

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.

A B

C D

68

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.

Consolidation

Cell

Top Plate

Teflon Wiper

Porous Metal

Plate

Bottom Plate

Drainage Lines

69

Figure 3.6: Hydraulic jack used to extrude cylindrical consolidated soil specimens from 70 mm

internal diameter sampling tube (left).

70

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.

71

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.

72

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.

A B

73

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.

A

B

C

74

Figure 3.11: Aluminium foil mold containing a 0.25" hydrate cylinder, which proved impossible

to unwrap without fracturing into segments.

75

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.

A B

C

Cooling Systems

76

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

1⁰C

25⁰C

Outer Cell Wall

Specimen

Hydrate Vein

Upper Cooling

Tubes

Porous Stone

Base Plate

Cooler

Inner Cell Wall

Pressure

Control Cabinet

Load FrameLoad

Load Cell

LVDT

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.

77

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

4.2.1 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.

78

Figure 4.2 shows the determination of the compression (𝐶𝑐) and recompression (𝐶𝑟) indices of

the soil (simplified as the slope of the unloading line). The slurry was slightly more compressible

(𝐶𝑐 = 0.22) than the preconsolidated soil (𝐶𝑐 = 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 𝐶𝑐:

𝐶𝑐 = 0.009(𝐿𝐿 − 10%) (4.1)

By Equation 4.1, the 𝐶𝑐 should be approximately 0.216, similar to values calculated from results.

In all cases the 𝐶𝑟 was found to be around 0.03. 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.

4.2.2 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.

79

K0-Consolidation Results and Analysis

In the K0-consolidation tests conducted, the radial strain was maintained within a certain

threshold (𝜀𝑟 < |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 (𝜎3′/𝜎1

′) of

approximately 0.75 (𝐾0.75). 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 (𝐾0.38).

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 𝐾0(𝑁𝐶) value (Watabe et al.,

2003). Jaky (1948) relates the 𝐾0(𝑁𝐶) value to the friction angle by:

𝐾0(𝑁𝐶) = 1 − sin 𝜙′ (4.2)

By this the friction angle is expected to be 38⁰, which can be verified in the following section.

80

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 (𝑒𝑓) is calculated assuming the soil is

saturated and volume change is due to pore water expulsion (∆𝑉 = ∆𝑉𝑣):

𝑒𝑓 = 𝑒0 − ∆𝑒 = 𝑒0 −

∆𝑉𝑣

𝑉𝑠 (4.3)

The 𝐶𝑟 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 one-

dimensional and results should resemble those from the oedometer tests. However, the 𝐶𝑐 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

81

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 (𝜎3) and pore pressure (𝑢)

measurements in the calculation of principal stresses (𝜎1′, 𝜎3

′) allowing the mean effective stress

(𝑝′) to be determined using:

𝑝′ = (𝜎1′ + 2𝜎3

′)/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 − 𝑢0

𝜎1 − 𝜎3 (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

𝑑𝜀= 0,

𝑑𝑞

𝑑𝜀= 0,

𝑑𝑝′

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 (𝐸𝑖) of the stress-strain curve of a soil

through the origin represents its modulus of elasticity. In this study, axial strain is measured

82

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 (𝐸50) is typically used, although at times the secant

modulus to 0.5% strain (𝐸0.5%) 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 (𝐾0.38 and 𝐾0.75), along with the response of a specimen isotropically

reconsolidated (𝐾1) to 100 kPa and submitted to undrained shear (discussed further in Section

4.3.2). The more heavily consolidated 𝐾0.38 specimen exhibited a peak axial stress with post

peak softening to critical state. The less consolidated 𝐾0.75 and 𝐾1 specimens showed no

appreciable peaks, and the 𝐾0.75 specimen had a critical state strength slightly lower than that of

the 𝐾0.38 specimen. The sharp rise in deviatoric stress with axial strain for the 𝐾0.38 specimen

gave rise to a significantly higher stiffness (𝐸50), while the 𝐸0.5% 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 𝐾0.38, 𝐾0.75 and 𝐾1

specimens. The 𝐾0.75 and 𝐾1 specimens display early peak 𝐴 coefficients, indicating initial

compression of the specimens. This is followed by a decrease to the same coefficient at failure

(𝐴𝑓 = 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 𝐴𝑓).

83

However the 𝐾0.38 specimen displays no peak and a lower 𝐴𝑓 value (𝐴𝑓 = 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 𝐾0.75 and 𝐾1 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 𝐾0.38 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 𝑐𝑢/𝑝′ ratio is generally assumed to be constant, where 𝑐𝑢

is the undrained shear strength. However, the 𝑐𝑢/𝑝′ ratio of the two anistropically consolidated

specimens are 0.61 (𝐾0.38) and 0.70 (𝐾0.75), while that of the isotropically consolidated specimen

(𝐾1) is 0.72, showing a decrease in undrained shear strength with a decrease in stress ratio, well-

documented 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

84

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 4.3.1 and 4.3.2 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 4.3.3 along with the reasons for the issues

encountered.

85

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 (𝑉𝑇(𝑠𝑜𝑖𝑙)), which can be calculated from the initial specimen volume

(𝑉0) and the vein volume (𝑉𝑣𝑒𝑖𝑛) 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:

𝑉𝑇(𝑠𝑜𝑖𝑙) = 𝑉0 − 𝑉𝑣𝑒𝑖𝑛 (4.8)

The volume of voids of the host soil (𝑉𝑣(𝑠𝑜𝑖𝑙)) can be determined using the porosity of the

preconsolidated soil (𝑛0) assuming no volume changes in the surrounding soil during the vein

drilling procedure (Equation 4.9), and the volume of the soil solids (𝑉𝑠(𝑠𝑜𝑖𝑙)) can be determined

using Equation 4.10:

𝑉𝑣(𝑠𝑜𝑖𝑙) = 𝑉𝑠𝑜𝑖𝑙 × 𝑛0

𝑉𝑣(𝑠𝑜𝑖𝑙) = 𝑉𝑠𝑜𝑖𝑙 ×𝑒0

1 + 𝑒0 (4.9)

𝑉𝑠(𝑠𝑜𝑖𝑙) = 𝑉𝑇(𝑠𝑜𝑖𝑙) − 𝑉𝑣(𝑠𝑜𝑖𝑙) (4.10)

The void ratio of the soil component of the specimen (𝑒𝑠𝑜𝑖𝑙), which should be equal to the initial

void ratio of the preconsolidated soil specimen (𝑒0 = 𝑒𝑠𝑜𝑖𝑙) can be expressed as:

𝑒𝑠𝑜𝑖𝑙 =

𝑉𝑣(𝑠𝑜𝑖𝑙)

𝑉𝑠(𝑠𝑜𝑖𝑙) (4.11)

The void ratio due only to the vein void can be expressed separately from the soil:

𝑒𝑣𝑒𝑖𝑛 =

𝑉𝑣𝑒𝑖𝑛

𝑉𝑠(𝑠𝑜𝑖𝑙) (4.12)

86

The equation for the total void ratio (𝑒) can then be developed from the above equations and

given as Equation 4.13:

𝑒 =

𝑉𝑣(𝑡𝑜𝑡𝑎𝑙)

𝑉𝑠(𝑡𝑜𝑡𝑎𝑙)=

𝑉𝑣(𝑠𝑜𝑖𝑙) + 𝑉𝑣𝑒𝑖𝑛

𝑉𝑠(𝑠𝑜𝑖𝑙)

𝑒 = 𝑒𝑠𝑜𝑖𝑙 + 𝑒𝑣𝑒𝑖𝑛 (4.13)

To normalize void ratio results for specimens with different vein sizes, the void ratio of the soil

component (𝑒𝑠𝑜𝑖𝑙) is used rather than the total void ratio of the specimen (𝑒). However in natural

samples, unless the total vein volume (𝑉𝑣𝑒𝑖𝑛) can be estimated then void ratios of the soil (𝑒𝑠𝑜𝑖𝑙)

and vein components (𝑒𝑣𝑒𝑖𝑛) cannot be determined or used.

Consolidation Results and Analysis

The volumetric strain versus time during isotropic reconsolidation to 100 kPa for the non-

hydrate-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 (∆𝑒𝑣𝑒𝑖𝑛 ≠ 0) and may result from dissociation/dissolution of the hydrate vein.

This is discussed in greater detail in Section 4.3.3.

87

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 𝐸50 value is used for the stiffness of the non-hydrate-bearing

and 0.75" vein-bearing specimens, while the 𝐸𝑠𝑒𝑐 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.

88

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 (𝑢𝑓) 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 (𝐴𝑓) 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 4.10. 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.

89

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.

90

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 non-

hydrate-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 4.13. 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.

91

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

(𝑐𝑢) and the undrained elastic modulus (𝐸𝑢) of the hydrate-bearing specimens.

4.4.1 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 = 𝜎3). 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.

4.4.2 Undrained Shear Compression Results and Analysis

Data Analysis Techniques

The undrained shear strength (𝑐𝑢) of a specimen can be determined from the peak deviatoric

stress (𝜎1 − 𝜎3)𝑓 applied during axial compression, using Equation 4.14:

92

𝑐𝑢 =

(𝜎1 − 𝜎3)𝑓

2 (4.14)

The undrained elastic modulus was approximated by determining the 𝐸50. 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 (𝐸𝑠𝑒𝑐) 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 4.14. The results show a general increase in 𝑐𝑢 and stiffness with

increasing hydrate vein diameter. The exception to this trend is the 0.25" diameter hydrate-

bearing specimen which has a slightly lower 𝑐𝑢 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 𝑐𝑢 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

93

clays, the 𝑐𝑢 can be related to the effective vertical preconsolidation stress and the plasticity

index (𝑃𝐼) using Equation 4.15. This relationship predicts the 𝑐𝑢 of the soil to be 17 kPa:

𝑐𝑢

𝜎𝑣′

= 0.11 + 0.0037(𝑃𝐼) (4.15)

A more simple relationship for clay proposed by Mesri (1989), shown as Equation 4.16, predicts

the 𝑐𝑢 of the soil to be 22 kPa:

𝑐𝑢

𝜎𝑣′

= 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 4.2.1 and 4.4.1. 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

94

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 𝐾0.75 and 𝐾0.38 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

95

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.

96

Table 4.1: Summary of results from oedometer tests to 800 kPa vertical pressure on fine-grained

soil

Sample Name Soil Type

Initial

Void

Ratio,

𝒆𝟎

Final

Void

Ratio,

𝒆𝒇

Final

Saturation,

𝑺 (%)

Compression

Index, 𝑪𝒄

Recompression

Index, 𝑪𝒓

Slurried Soil Slurried to

over 1.5×LL 1.54 0.54 97.8 0.22 0.03

Preconsolidated Soil 1 Consolidated

to 100 kPa 0.73 0.54 99.1 0.20 0.03


to 100 kPa 0.73 0.52 101.9 0.193 0.03


to 100 kPa 0.72 0.53 100.1 0.187 0.03

Table 4.2: Summary of results from undrained shear tests on anisotropically consolidated and

isotropically reconsolidated fine-grained soil specimens

Sample

Name

Final Consolidation Results Undrained Shear Data at Failure

Failure Criterion: Maximum Deviatoric Stress

Major

Effective

Stress,

𝝈′𝟏

(kPa)

Minor

Effective

Stress,

𝝈′𝟑

(kPa)

K

Value,

𝝈′𝟑/𝝈′𝟏

Void

Ratio,

𝒆𝒇

Axial

Strain,

𝜺𝒂𝒇

(%)

Deviatoric

Stress, 𝒒𝒇

(kPa)

Pore

Pressure

Parameter,

𝑨𝒇

Undrained

Stiffness,

𝑬𝟓𝟎𝒖 (MPa)

K = 1 100 100 1 0.54 12 135 0.34 6.2

K ≈ 0.75 1100 800 0.72 0.46 8.5 1260 0.34 83.7

K = 0.38 2080 800 0.38 0.44 2.5 1495 0.16 944.6

97

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


Void

Ratio

of

Soil,

𝒆𝒔𝒐𝒊𝒍

Area

Ratio,

𝑨𝑹

Hydrate

Vein

Sat., 𝑺𝒗𝒉

(%)

Axial

Strain,

𝜺𝒂𝒇

(%)

Deviatoric

Stress, 𝒒𝒇

(kPa)

Pore

Pressure

Parameter

, 𝑨𝒇

Undrained

Stiffness,

𝑬𝟓𝟎𝒖 or

𝑬𝒔𝒆𝒄𝒖 (MPa)

Vein

Failure

Mode

0/0 0.54 0 0 12.0 136 0.34 6.2 N/A

Competent Hydrate Vein-Bearing Specimens

19.05/0.75 0.57 0.06 15.1 6.3 245 0.17 15.4

Diagonal

Rupture with

Shear Band

25.4/1.00 0.60 0.105 23.8 4.5 609 0.08 24.6 Horizontal

Rupture

Table 4.4: Summary of results from consolidated undrained tests on non-competent hydrate-

vein-bearing specimens

Initial

Hydrate

Vein

Diameter

(mm/in)

Undrained Shear Data at Failure and Post-Shear Failure Mode


Axial

Strain,

𝜺𝒂𝒇

(%)

Deviatoric

Stress, 𝒒𝒇

(kPa)

Pore

Pressure

Parameter

, 𝑨𝒇

Undrained

Stiffness,

𝑬𝟓𝟎𝒖 or


Vein

Failure

Mode

6.35/0.25 6.5 106 0.61 9.7 Vein

Disappeared

12.7/0.50 5.7 106 0.63 9.7

Diagonal

Rupture with

Shear Band

12.7/0.50 8.9 134 0.33 7.0

Distributed

Shear Zone

(Granulated

Hydrate)

19.05/0.75 4.2 107 0.49 5.0

Distributed

Shear Zone

(Granulated

Hydrate)

98

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 Undrained Shear Data at Failure and Post-Shear Failure Mode


Void

Ratio

of

Soil,

𝒆𝒔𝒐𝒊𝒍

Area

Ratio

, 𝑨𝑹

Hydrate

Vein

Sat., 𝑺𝒗𝒉

(%)

Axial

Strain,

𝜺𝒂𝒇

(%)

Deviatoric

Stress, 𝒒𝒇

(kPa)

Undrained

Shear

Strength, 𝒄𝒖

(kPa)

Undrained

Stiffness,

𝑬𝟓𝟎𝒖 or


Vein

Failure

Mode

0/0 0.67 0 0 12.0 37 18.5 3.6 N/A

6.35/0.25 0.70 0.008 2.0 14.2 33 16.5 3.2 Vein

disintegrated

12.7/0.50 0.68 0.033 7.8 4.6 105 52.5 3.9 Horizontal

Rupture

19.05/0.75 0.74 0.074 15.4 2.3 183 91.5 11.2 Horizontal

Rupture

25.4/1 0.74 0.132 26.0 1.5 360 180 25.1 Horizontal

Rupture

25.4/1 0.74 0.132 26.0 1.2 235 116.5 24.9 Diagonal

Rupture

99

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).

A

B

100

Figure 4.2: Determination of compression and recompression indices from oedometer tests on

slurried soil (a) and preconsolidated soil samples (b, c and d).

A B

C D

101

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.

102

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.

103

Figure 4.5: (a) Plot of deviatoric stress versus strain for the anisotropically consolidated and

isotropically reconsolidated specimens. (b) Similar 𝐴𝑓 values are observed for the isotropically

reconsolidated (to 100 kPa) and 𝐾0.75 specimens, with a lower value for the 𝐾0.38 specimen.

A

B

104

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).

A

B

105

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.

106

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.

107

Figure 4.9: (a) Excess pore pressure and (b) pore pressure coefficient versus axial strain. A

decrease in 𝐴𝑓 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.

A

B

108

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.

109

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.

A B

C D

110

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.

111

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 4.12.

A B

C D

112

Figure 4.14: Stress-strain plots from unconsolidated undrained compression tests on specimens

containing hydrate veins of different diameters.

113

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.

A B C

D E

114

Chapter Five: Discussion

5.1 Introduction

Analysis of triaxial test results on competent, vertical, cylindrical THF hydrate veins within fine-

grained 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

5.2.1 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.

5.2.1.1 Hydrate Vein Saturation

Typically the hydrate content of a soil is quantified by the pore space hydrate saturation (𝑆ℎ),

which is the ratio of the hydrate volume (𝑉ℎ) within the void space of the soil (𝑉𝑣):

𝑆ℎ =

𝑉ℎ

𝑉𝑣× 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

115

veins. Therefore the saturation of hydrate veins (𝑆𝑣ℎ) is suggested as an alternative definition.

This is calculated by substituting the volume of the vein (𝑉𝑣𝑒𝑖𝑛) (equal in this case to the hydrate

volume) and the volume of voids in the surrounding soil (𝑉𝑣(𝑆𝑜𝑖𝑙)) for the total volume of voids

(𝑉𝑣) into Equation 5.1, as shown in Equation 5.2:

𝑆𝑣ℎ =

𝑉ℎ

𝑉𝑣𝑒𝑖𝑛 + 𝑉𝑣(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.

5.2.1.2 Area Ratio

As the hydrate veins created for this research are cylinders of constant diameter, the ratio of the

cross-sectional vein area (𝐴𝑣𝑒𝑖𝑛) to the specimen area (𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛) is constant and can be used as

a method of defining hydrate content. The calculation for the area ratio (𝐴𝑟) is:

𝐴𝑟 =

𝐴𝑣𝑒𝑖𝑛

𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 (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

116

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).

5.2.1.3 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:

𝑆𝑣ℎ =

𝑉ℎ

𝑉𝑣𝑒𝑖𝑛 + 𝑉𝑣(𝑠𝑜𝑖𝑙)× 100%

𝑆𝑣ℎ =

𝐴𝑣𝑒𝑖𝑛 × 𝐻

𝐴𝑣𝑒𝑖𝑛 × 𝐻 + 𝑛 × 𝐻 × (𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 − 𝐴𝑣𝑒𝑖𝑛)× 100%

100%

𝑆𝑣ℎ=


𝐴𝑣𝑒𝑖𝑛+ 𝑛

𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛

𝐴𝑣𝑒𝑖𝑛− 𝑛



𝑆𝑣ℎ

100%=

1𝑛

𝐴𝑟+ 1 − 𝑛

(5.4)

Expressed in terms of the area ratio:

𝐴𝑟 = 𝑛

100%𝑆𝑣ℎ

− (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.

117

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.

5.2.2.1 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 vein-

bearing 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

118

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 (𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ)), below which the hydrate

vein has no impact on the undrained shear strength. Presenting this relationship mathematically:

𝐼𝑓 𝐴𝑟 ≤ (𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) = 0.014) , 𝑆𝑢 = 18.5𝑘𝑃𝑎

𝐼𝑓 𝐴𝑟 > (𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) = 0.014) , 𝑆𝑢 = 1350𝑘𝑃𝑎 × 𝐴𝑟

(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 𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ), the undrained shear strength can be generalized as constant and approximately

equal to the undrained shear strength of the soil (𝑆𝑢(𝑠𝑜𝑖𝑙)), implying the specimen shear strength

is entirely dependent on the soil and the veins provide no shear resistance. This is generalized as:

𝐼𝑓 𝐴𝑟 ≤ 𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) , 𝑆𝑢(𝐴𝑟) = 𝑆𝑢(𝑠𝑜𝑖𝑙) (5.7)

For vein sizes above the predicted 𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) the relationship is linear in 𝑆𝑢 versus 𝐴𝑟 space, the

function for which can be generalized, with 𝑚 as the slope and the y-intercept equal to 0 (𝑏 = 0):

𝐼𝑓 𝐴𝑟 > 𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) , 𝑆𝑢(𝐴𝑟) = 𝑚 × 𝐴𝑟 (5.8)

If this equation is assumed to apply below the 𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ), a y-intercept of zero implies that when

no hydrate is present (𝐴𝑣𝑒𝑖𝑛 = 0) the specimen would have no strength (𝑆𝑢 = 0), implying that

by this equation the specimen strength is entirely a function of the hydrate vein. The assumption

119

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 (𝐹𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛) in terms of only the vein area by setting

the soil area equal to zero (𝐴𝑠𝑜𝑖𝑙 = 0), a concept expressed as the vein stress (𝜎𝑣𝑒𝑖𝑛):

𝜎𝑣𝑒𝑖𝑛 =

𝐹𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛

𝐴𝑣𝑒𝑖𝑛 (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 4.3.2. 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 (𝜎𝑣𝑒𝑖𝑛(𝑚𝑎𝑥) = 𝜎𝑐ℎ). 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 0.9-44MPa (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 𝑆𝑢 is equal to half the maximum deviatoric stress

on the specimen (0.5(𝜎1 − 𝜎3)𝑚𝑎𝑥) so we can substitute this in Equation 5.8 and rearrange:

120

0.5(𝜎1 − 𝜎3)𝑚𝑎𝑥 × 𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 = 𝑚 × 𝐴𝑣𝑒𝑖𝑛

The maximum deviatoric stress, multiplied by the area on which it acts, is equal to the maximum

load on the specimen (𝐹𝑚𝑎𝑥) so substituting this into Equation 5.9 we get Equation 5.10:

0.5𝐹𝑚𝑎𝑥 = 𝑚 × 𝐴𝑣𝑒𝑖𝑛

𝑚 =

0.5𝐹𝑚𝑎𝑥

𝐴𝑣𝑒𝑖𝑛 (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 (𝜎𝑐ℎ):

𝑚 = 0.5𝜎𝑣𝑒𝑖𝑛(𝑚𝑎𝑥)

𝑚 = 0.5𝜎𝑐ℎ

(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 𝐴𝑟 > 𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) , 𝑆𝑢(𝐴𝑟) = 0.5𝜎𝑐ℎ𝐴𝑟 (5.12)

The predicted threshold area ratio value (𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ)) is calculated as the intercept of the functions

defined as Equations 5.12 and 5.7:

𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) =

𝑆𝑢(𝑆𝑜𝑖𝑙)

0.5𝜎𝑐ℎ (5.13)

Therefore, the empirical relationship can be generalized using Equations 5.7, 5.12 and 5.13:

𝐼𝑓 𝐴𝑟 ≤

𝑆𝑢(𝑠𝑜𝑖𝑙)

0.5𝜎𝑐ℎ , 𝑆𝑢 = 𝑆𝑢(𝑠𝑜𝑖𝑙) (5.14)

121

𝐼𝑓 𝐴𝑟 >𝑆𝑢(𝑠𝑜𝑖𝑙)

0.5𝜎𝑐ℎ , 𝑆𝑢 = 0.5𝜎𝑐ℎ𝐴𝑟

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 (𝑆𝑣ℎ) and soil porosity

(𝑛), based on a threshold hydrate vein saturation value (𝑆𝑣ℎ(𝑡ℎ𝑟𝑒𝑠ℎ)):

𝐼𝑓

𝑆𝑣ℎ

100%≤

1

𝑛0.5𝜎𝑐ℎ

𝑆𝑢(𝑆𝑜𝑖𝑙)+ (1 − 𝑛)

, 𝑆𝑢 = 𝑆𝑢(𝑆𝑜𝑖𝑙)

𝐼𝑓 𝑆𝑣ℎ

100%>

1

𝑛0.5𝜎𝑐ℎ

𝑆𝑢(𝑆𝑜𝑖𝑙)+ (1 − 𝑛)

, 𝑆𝑢 =0.5𝜎𝑐ℎ𝑛

100%𝑆𝑣ℎ

− (1 − 𝑛)

(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 𝑆𝑣ℎ ≤ 3.36% , 𝑆𝑢 = 18.5

𝐼𝑓 𝑆𝑣ℎ > 3.36% , 𝑆𝑢 =540

100%𝑆𝑣ℎ

− 0.6

(5.16)

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This relationship fits experimental results fairly well, as shown in Figure 5.1b. The threshold

hydrate vein saturation (𝑆𝑣ℎ(𝑡ℎ𝑟𝑒𝑠ℎ)) of 3.36% corresponds to a hydrate vein diameter of 0.812

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.

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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.

5.2.2.2 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.

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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

(𝐴𝑟 = 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 (𝐾0), 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

125

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

126

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.

5.2.3 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.

5.2.3.1 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:

𝑘𝑒𝑞 = 𝑘1 + 𝑘2 (5.18)

𝑑𝐿𝑒𝑞 = 𝑑𝐿1 = 𝑑𝐿2 (5.19)

𝐹𝑒𝑞 = 𝐹1 + 𝐹2 (5.20)

The spring constant (𝑘) is related to the Young’s modulus (𝐸) of a material by:

𝑘 =

𝐸 × 𝐴𝑟𝑒𝑎

𝐿 (5.21)

127

Substituting the spring constants for hydrate and soil (𝑘ℎ,𝑠𝑜𝑖𝑙) into Equations 5.18 and 5.21, and

knowing that the hydrate and soil are equal to the height of the specimen (𝐿𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 = 𝐿𝑣𝑒𝑖𝑛 =

𝐿𝑠𝑜𝑖𝑙) allows the elastic modulus of the composite material to be expressed as:

𝑘𝑒𝑞 = 𝑘ℎ + 𝑘𝑠𝑜𝑖𝑙

𝐸𝑒𝑞𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛

𝐿𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛=

𝐸ℎ𝐴𝑣𝑒𝑖𝑛

𝐿𝑣𝑒𝑖𝑛+

𝐸𝑠𝑜𝑖𝑙𝐴𝑠𝑜𝑖𝑙

𝐿𝑠𝑜𝑖𝑙

𝐸𝑒𝑞 =


𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛+


𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛 (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 0.36-32 GPa.

5.2.3.2 Undrained Stiffness versus Area Ratio

The Hookean relationship (Equation 5.22) can be expressed in terms of the area ratio as follows:

𝐸𝑒𝑞 =





𝐸𝑒𝑞 =



𝐸𝑠(1 − 𝐴𝑟)𝐴𝑠𝑝𝑒𝑐𝑖𝑚𝑒𝑛


𝐸𝑒𝑞 = 𝐸ℎ𝐴𝑟 + 𝐸𝑠(1 − 𝐴𝑟)

𝐸𝑒𝑞 = (𝐸ℎ − 𝐸𝑠)𝐴𝑟 + 𝐸𝑠 (5.23)

128

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 𝐴𝑟 ≤ (𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) = 0.02) , 𝐸𝑢 = 3700𝑘𝑃𝑎

𝐼𝑓 𝐴𝑟 > (𝐴𝑟(𝑡ℎ𝑟𝑒𝑠ℎ) = 0.02) , 𝐸𝑢 = 185000𝑘𝑃𝑎 × 𝐴𝑟

(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 (0.36-32 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 5.2.2.1 as Equation 5.25, assuming above the threshold area ratio

the stiffness depends entirely on the hydrate vein (using 𝐸ℎ = 185𝑀𝑃𝑎):

129

𝐼𝑓 𝐴𝑟 ≤

𝐸𝑢(𝑠𝑜𝑖𝑙)

𝐸ℎ , 𝐸𝑢 = 𝐸𝑢(𝑠𝑜𝑖𝑙)

𝐼𝑓 𝐴𝑟 >𝐸𝑢(𝑠𝑜𝑖𝑙)

𝐸ℎ , 𝐸𝑢 = 𝐸ℎ𝐴𝑟

(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 5.2.2.2. 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:

𝐸𝑒𝑞 = 178800𝐴𝑅 + 6200 (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.

5.2.3.3 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:

130

𝐼𝑓

𝑆𝑣ℎ

100%≤

𝐸𝑢(𝑆𝑜𝑖𝑙)

𝐸ℎ𝑛 + 𝐸𝑢(𝑠𝑜𝑖𝑙)(1 − 𝑛) , 𝐸𝑢 = 𝐸𝑢(𝑠𝑜𝑖𝑙)

𝐼𝑓 𝑆𝑣ℎ

100%>

𝐸𝑢(𝑆𝑜𝑖𝑙)

𝐸ℎ𝑛 + 𝐸𝑢(𝑠𝑜𝑖𝑙)(1 − 𝑛) , 𝐸𝑢 =

𝐸ℎ𝑛

100%𝑆𝑣ℎ

− (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 𝑆𝑣ℎ ≤ 4.9% , 𝐸𝑢 = 𝐸𝑢(𝑠𝑜𝑖𝑙)

𝐼𝑓 𝑆𝑣ℎ > 4.9% , 𝐸𝑢 =74000

100%𝑆𝑣ℎ

− 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 0.363.

The stiffness follows the trend predicted by the Hookean springs in parallel. Substituting the

averaged porosity value of 0.363 for specimens and the soil and hydrate stiffness in Equation

5.27, the relationship can be described by Equation 5.29:

131

𝐸𝑒𝑞 =

65000

100%𝑆𝑣ℎ

− 0.637+ 6200

(5.29)

5.2.3.4 Discussion

The undrained stiffness results from UU tests are seen to follow a transitional soil to hydrate-

controlled 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 5.7.

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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.

5.3.1 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 one-

dimensionally to an effective stress of 100 kPa (~20 m below seafloor) with a gas hydrate

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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.

5.3.2 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

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the in-situ consolidation behaviour. Marine soils consolidate under self-weight in one-

dimensional conditions according to their compressibility. Priest et al. (2014) suggest that if sub-

vertical 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 (𝑒𝑠𝑜𝑖𝑙) is

assumed to remain constant (given no porewater outflow to aid in gas hydrate formation), despite

a potential increase in the total void ratio (𝑒𝑇) due to the fracturing of soil and infill of gas

hydrate. As sedimentation continues, the increasing vertical stress may now be partially carried

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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.

5.3.3 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

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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 (∆𝑒𝑇)

(Equation 5.30) due to both the collapse of the vein structure (∆𝑒𝑣𝑒𝑖𝑛) and the loss of the

structural support of the veins which results in a change in the soil structure (∆𝑒𝑠𝑜𝑖𝑙):

∆𝑒𝑇 = ∆𝑒𝑠𝑜𝑖𝑙 + ∆𝑒𝑣𝑒𝑖𝑛 (5.30)

The void ratio change due to vein void collapse (∆𝑒𝑣𝑒𝑖𝑛) 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 (∆𝑒𝑠𝑜𝑖𝑙) 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 5.10.

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 (∆𝑒𝑐𝑠) 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:

∆𝑒𝑇 = ∆𝑒𝑠𝑜𝑖𝑙 + ∆𝑒𝑣𝑒𝑖𝑛 + ∆𝑒𝑐𝑠 (5.31)

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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 vein-

bearing 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

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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-vein-

bearing 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 5.27.

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 post-

dissociation, due to the disappearance of the strong, stiff vein network.

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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.

A

B

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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.

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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.

A

B

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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.

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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.

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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.

A

B

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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.

A

B

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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.

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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.

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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-vein-

bearing 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 fine-

grained 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.

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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.

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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 fine-

grained 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

151

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

152

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.

153

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-

154

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 poorly-

understood 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

155

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.

156

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.

157

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168

Appendix A: Material Specification Sheets

169

Figure A1: Specification Sheet for EPK Kaolin

170

171

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

172

Appendix B: Oedometer Test Results

Table B1: Oedometer consolidation test on Preconsolidated Soil 1

Initial Specimen Properties

Cell Number 4 Specimen Density (g/cm3) 1.89

Moisture Content (%) 24 Dry Density (g/cm3) 1.53

Weight of Sample (g) 101.24 Initial Void Ratio 0.73

Specimen Height (cm) 1.69 Initial Saturation (%) 86.7

Specimen Volume (cm3) 53.52 Equivalent Height of Solids (cm) 0.98

Particle Density (g/cm3) 2.64

Final Specimen Properties

Moisture Content (%) 20 Final Volume (cm3) 47.50

Weight of Sample (g) 97.91 Final Density (g/cm3) 2.06

Final Height (cm) 1.50 Final Dry Density (g/cm3) 1.71

Overall Settlement (cm) 0.19 Final Void Ratio 0.54

Volume Change (cm3) 6.02 Final Saturation (%) 99.1

Load-Deformation Data

Pressure (kPa) Final Specimen Height (mm) Void Ratio Modulus of Compressibility (m2/MN)

0 16.90 0.731

4 16.70 0.710 3.12

10 16.62 0.702 0.72

20 16.53 0.693 0.61

50 16.31 0.670 0.45

75 16.18 0.657 0.33

100 16.06 0.644 0.30

125 15.94 0.632 0.29

150 15.83 0.621 0.27

175 15.75 0.613 0.22

200 15.62 0.599 0.32

400 15.01 0.537 0.20

800 14.45 0.479 0.10

5 15.01 0.537 0.05

173








Particle Density (g/cm3) 2.64 Final Specimen Properties








0 16.90 0.731

5 16.72 0.713 2.15

10 16.63 0.704 1.03

20 16.50 0.691 0.83

50 16.22 0.662 0.58

75 16.05 0.644 0.42

100 15.90 0.629 0.38

125 15.78 0.617 0.30

150 15.65 0.603 0.35

175 15.54 0.592 0.25

200 15.43 0.581 0.27

400 14.84 0.520 0.20

805 14.31 0.466 0.09

10 14.84 0.520 0.04

174








Particle Density (g/cm3) 2.64 Final Specimen Properties








0 16.90 0.721

5 16.71 0.701 2.68

10 16.62 0.692 1.05

20 16.47 0.677 0.96

50 16.23 0.652 0.50

75 16.10 0.639 0.32

100 15.98 0.627 0.29

125 15.88 0.617 0.23

150 15.77 0.606 0.27

180 15.70 0.598 0.18

205 15.60 0.588 0.24

405 15.04 0.531 0.19

810 14.50 0.476 0.09

10 15.04 0.531 0.04

175

Table B4: Oedometer consolidation test on Slurried Soil






Specimen Volume (cm3) 80.44 Equivalent Height of Solids

(cm) 1.00

Particle Density (g/cm3) 2.64 Final Specimen Properties (From Final Height)








0 25.40 1.542

50 17.61 0.763 9.39

100 16.87 0.689 0.85

200 16.14 0.615 0.45

400 15.48 0.549 0.21

800 14.84 0.485 0.11

200 15.00 0.501 0.02

50 15.18 0.519 0.08

5 15.38 0.539 0.29

176

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) 13.90 Final Height (cm) 13.54

Specimen Diameter (cm) 7.0 Final Diameter (cm) 6.94

Volume of Soil (cm³) 534.93 Cross-Sectional Area after

Reconsolidation (cm2) 37.79

Weight (g) N/A Isotropic Coefficient of

Compressibility (mvi) (m2/MN)

0.44

Wet Density (g/cm³) N/A Reconsolidated Void Ratio 0.60

Water Content (excess material) 25

Void Ratio 0.67

Dry Unit Weight (kN/m³) 15.08

Saturation (%) 98.4

Volume of Voids (cm³) 214.82

Volume of Solids (cm³) 320.12

Anisotropic Consolidation Data

Minor Effective Stress

(kPa)

Major Effective Stress

(kPa) Height (cm) Diameter (cm) Volume (cm3) Void Ratio

100 100 13.54 6.94 511.49 0.60

200 265 13.48 6.92 506.43 0.58

800 1100 13.28 6.71 466.83 0.46

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) 10.58

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) 0.46

Pore Pressure Parameter at Failure (Af) 0.34 Saturation (%) 97

Undrained Stiffness - (E50u) (kPa) 83700 Reconsolidated Dry Unit Weight

(kN/m³) 18.10

Undrained Stiffness – (E0.5%) (kPa) 93000

Notes Attempt at Ko-consolidation using circumferential strain gauge

177

Data Plots

178

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) 10.94

Shelby Tube Number 3 Change in Volume (%) 5.90

Specimen Height (cm) 14.10 Final Height (cm) 13.98

Specimen Diameter (cm) 7.0 Final Diameter (cm) 6.82

Volume of Soil (cm³) 542.63 Cross-Sectional Area after

Reconsolidation (cm²) 36.56

Weight (g) N/A Isotropic Coefficient of

Compressibility (mvi) (m2/MN)

0.57

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) 15.08

Saturation (%) 98.4

Volume of Voids (cm³) 217.91


Anisotropic Consolidation Data

Minor Principal Stress

(kPa)

Major Principal Stress

(kPa) Height (cm) Diameter (cm) Volume (cm3) Void Ratio

100 110 13.98 6.82 510.64 0.57

200 310 13.85 6.79 467.87 0.54

400 1040 13.35 6.79 448.52 0.48

800 2080 13.01 6.78 435.29 0.44


Failure Criterion: Maximum Deviatoric Stress Specimen Properties after Failure




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)

0.44

Pore Pressure Parameter at Failure (Af) 0.16 Saturation (%) 98.2

Undrained Stiffness - (E50u) (kPa) 944600 Reconsolidated Dry Unit Weight

(kN/m³) 18.35


Notes Failure criterion chosen to be maximum deviatoric stress, however critical state

was also reached, as indicated on plots

179

Data Plots

180

Appendix D: Consolidated Undrained Triaxial Test Results

Table D1: Data from CU test on specimen with no hydrate vein


Prior to Hydrate Vein Formation 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) 14.48 Specimen Volume (cm3) 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) 15.93 Weight of Soil and Hydrate (g) N/A

Saturation (%) 103.1 Volume of Voids in Soil (cm³) N/A

Volume of Voids (cm³) 214.48 Volume of Voids including vein (cm³) N/A

Volume of Solids (cm³) 342.59 Volume of Solids (cm³) N/A

Consolidation Stage

Initial Pore Pressure (kPa) 494.00 Time to 100% Primary Reconsolidation (min.) 1157.48

Effective Consolidation Stress (kPa) 94.48 Isotropic Coefficient of Consolidation (Cvi) (m2/year) 0.086

Change in Volume (%) 5.25 Isotropic Coefficient of Compressibility (mvi) (m2/MN) 0.56

Final Height (cm) 13.97 Reconsolidated Void Ratio 0.54

Final Diameter (cm) 6.94 Reconsolidated Saturation (from final water content) (%) 98.9

Cross-Sectional Area after Reconsolidation (cm²) 37.82 Undrained Shear and Post-Shear Data


Axial Strain (%) 12 Average Soil Height (cm) N/A

Deviatoric Stress (Corrected) (kPa) 136 Average Soil Diameter (cm) 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 (Af) 0.34 Final Void Ratio (From Reconsolidation) 0.54

Undrained Stiffness - (E50u) (kPa) 6180 Reconsolidated Dry Unit Weight (kN/m³) 17.14


Notes Some axial strain applied accidentally at start of isotropic reconsolidation, barreling failure mode

181

Data Plots and Post-Shear Pictures

[No Pictures]

182

Table D2: Data from CU test on specimen with 0.75" diameter hydrate vein



Consolidation Date (Batch) 17/08/2015 Specimen Height (cm) 14.50

Shelby Tube Number 3 Specimen Diameter (cm) 7.00

Specimen Height (cm) 14.28 Specimen Volume (cm3) 558.03

Specimen Diameter (cm) 7.0 Vein Height (cm) 14.28

Volume of Soil (cm³) 549.37 Vein Diameter (cm) 1.91

Weight (g) 1100 Vein Volume (cm³) 40.69

Wet Density (g/cm³) 2.00 Soil Volume (cm³) 517.34

Water Content (excess material from consolidation) 25 Weight of Soil (g) 1051.33

Void Ratio 0.65 Weight of THF Hydrate (g) 40.00

Dry Unit Weight (kN/m³) 15.67 Weight of Soil and Hydrate (g) 1091.89

Saturation (%) 102.5 Volume of Voids in Soil (cm³) 204.24

Volume of Voids (cm³) 216.89 Volume of Voids including vein (cm³) 244.93

Volume of Solids (cm³) 332.48 Volume of Solids (cm³) 313.10

Consolidation Stage

Initial Pore Pressure (kPa) 498 Time to 100% Primary Reconsolidation (min.) 932.21

Effective Consolidation Stress (kPa) 100 Isotropic Coefficient of Consolidation (Cvi) (m2/year) 0.102

Change in Volume due to hydrate dissolution and

consolidation (%) 6.72 Isotropic Coefficient of Compressibility (mvi) (m

2/MN) 0.67


Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) 98.0

Cross-Sectional Area after Reconsolidation (cm²) 36.34





Induced Excess Porewater Pressure (kPa) 41 Weight of Specimen (g) 1053.00

Major Principal Effective Stress (kPa) 297 Weight of THF Hydrate (g) 30.36

Minor Principal Effective Stress (kPa) 53 Weight of Soil (g) 1022.64

Effective Principal Stress Ratio 5.7 Water Content (%) 21




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

183


184

Table D3: Data from CU test on specimen with 1" diameter hydrate vein
















Consolidation Stage



Change in Volume due to hydrate dissolution and

consolidation (%) 6.68 Isotropic Coefficient of Compressibility (mvi) (m

2/MN) 0.68


Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) 93












Secant Stiffness – (Esec) (1.2% to 1.9% strain) (kPa) 24600

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

185


186

















Consolidation Stage

Initial Pore Pressure (kPa) 500.00 Time to 100% Primary Reconsolidation (min.) 884.15




Final Diameter (cm) 6.80 Reconsolidated Saturation (from final water content) (%) N/A




Axial Strain (%) 6.5 Average Soil Height (cm) N/A





Effective Principal Stress Ratio 5.4 Water Content (%) N/A



Undrained Stiffness – (E0.5%) (kPa) 9640 Notes No data after shear; significant cell pressure oscillation during shear. No hydrate left

187


188

















Consolidation Stage












Major Principal Effective Stress (kPa) 134 Weight of THF Hydrate (g) 0






Notes Volumetric change calculated from axial and radial strain (measured with gauge), as back piston position was not logged, so is an approximation.

189


190









Weight (g) 1096.73 Vein Volume (cm³) 18.62








Consolidation Stage


















Notes Reconsolidation met with issues which is suspected to be due to hydrate vein dissolution leading to incomplete consolidation; Hydrate vein offered no strength

191


192

Table D7: Data from CU Test on specimen with 0.75" diameter hydrate vein
















Consolidation Stage

Initial Pore Pressure (kPa) 481 Time to 100% Primary Reconsolidation (min.) N/A

Effective Consolidation Stress (kPa) 93 Isotropic Coefficient of Consolidation (Cvi) (m2/year) N/A



Final Diameter (cm) 6.90 Reconsolidated Saturation (from final water content) (%) N/A




Axial Strain (%) 4.2 Average Soil Height (cm) N/A





Effective Principal Stress Ratio 5.9 Water Content (%) N/A




Notes No data taken after shear; reconsolidation not complete due to suspected hydrate dissociation

193


194

Appendix E: Unconsolidated Undrained Triaxial Test Results

Table E1: Data from UU test on specimen with no hydrate vein



Consolidation 08/10/2015 Specimen Height (cm) N/A

Specimen Height (cm) 14.83 Specimen Diameter (cm) N/A

Specimen Diameter (cm) 7.00 Specimen Volume (cm3) N/A

Volume of Soil (cm³) 570.53 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) 15.51 Soil Volume (cm³) N/A

Water Content (excess material) 26 Weight of Soil (g) N/A

Saturation 103.6 Weight of THF Hydrate (g) N/A

Volume of Voids (cm³) 228.84 Weight of Soil and Hydrate (g) N/A

Volume of Solids (cm³) 341.69 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


At Specimen Failure Specimen Properties after Failure


Deviatoric Stress (kPa) 37 Average Soil Diameter (cm) N/A

Excess Porewater Pressure (kPa) -1 Vein Height (cm) 14.40

Major Principal Total Stress (kPa) 237 Weight of Specimen (g) 1143.70

Minor Principal Total Stress (kPa) 200 Weight of THF Hydrate (g) N/A

Undrained Stiffness - (E50u) (kPa) 3600 Weight of Soil (g) 1143.70

Undrained Stiffness – (E0.5%) (kPa) 3700 Water Content (%) 26

Notes Baseline test


195

Table E2: Data from UU test on specimen with 0.25" diameter hydrate vein



Consolidation 02/10/2015 Specimen Height (cm) 15.00

Specimen Height (cm) 15.00 Specimen Diameter (cm) 7.00

Specimen Diameter (cm) 7.0 Specimen Volume (cm3) 577.27

Volume of Soil (cm³) 577.27 Vein Height (cm) 15.00

Void Ratio 0.70 Vein Diameter (cm) 0.64


Dry Unit Weight (kN/m³) 15.28 Soil Volume (cm³) 572.52

Water Content (excess material) 26 Weight of Soil (g) 1136.11

Saturation 99.7 Weight of THF Hydrate (g) 4.74

Volume of Voids (cm³) 236.78 Weight of Soil and Hydrate (g) 1140.85

Volume of Solids (cm³) 340.49 Volume of Voids in Soil (cm³) 234.83

Wet Density (g/cm³) 1.97 Volume of Voids including vein (cm³) 239.58


Hydrate Saturation including vein (%) 1.98





Excess Porewater Pressure (kPa) 2 Vein Height (cm) 14.40


Minor Principal Total Stress (kPa) 200 Weight of THF Hydrate (g) 4.64



Notes

196


197

















Hydrate Saturation including vein (%) 7.8









Secant Stiffness – (2.5% to 3.7%) (Esec) (kPa) 3900 Water Content (%) 24

Notes Secant tangent calculated as vein strength took until 2.5% strain to mobilize

198


199

















Hydrate Saturation including vein 15.4









Secant Stiffness – (1% to 1.9%) (Esec) (kPa) 11200 Water Content (%) 25

Notes Some seating issues at start of shear, secant modulus used instead

200


201

Table E5: Data from UU test on specimen with 1" diameter hydrate vein
















Hydrate Saturation including vein (%) 26










Notes Vein strength took until around 0.5% strain to mobilize

202


203

Table E6: Data from UU test on specimen with 1" diameter hydrate vein
















Hydrate Saturation including vein (%) 26










Notes Vein fractured diagonally

204


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