characterisation of soft soils for deep water

CHARACTERISATION OF SOFT SOILS

FOR DEEP WATER DEVELOPMENTS

by

SHIN FUN CHUNG

B.E. (Hons)

This thesis is presented for the degree of

Doctor of Philosophy

at

The University of Western Australia

School of Civil and Resource Engineering

Jan 2005

ABSTRACT

i

ABSTRACT

This research has studied the penetration and extraction resistance profiles of different

types of penetrometers in soft clay. The penetrometers of interest include the cone,

T-bar, ball and plate. Effects of the surface roughness and aspect ratio of the T-bar

penetrometer on its resistance have also been investigated. Undrained shear strength, su,

profiles derived from the penetration tests are compared with the shear strengths

measured from field vane shear tests and laboratory (triaxial and simple shear) tests.

Both in situ and centrifuge model penetration tests were undertaken for the research. In

addition, ‘undisturbed’ tube samples were retrieved from both the field and the

centrifuge strongbox samples (after completion of the centrifuge tests) for laboratory

testing.

The in situ testing was carried out in Western Australia, at the Burswood site near Perth,

with tests including cone, T-bar, ball and plate penetrometer tests, and vane shear tests.

Interestingly, the T-bar, ball and plate (‘full-flow’) penetrometers showed a narrow

band of resistance profiles both during penetration and extraction, with a range of

around 15 % between the highest and lowest profiles and standard deviation of 15 %.

However, the cone penetrometer gave similar resistance at shallow depths but

increasingly higher penetration resistance at depths greater than 7 m – a phenomenon

that is also common in offshore results. During extraction, the cone penetrometer gave a

higher resistance profile than the full-flow penetrometers for much of the depth of

interest.

The su profile measured directly from the vane shear tests falls within the su profiles

derived from the penetration resistances of the full-flow penetrometers, using a single

bearing factor, N = 10.5 (the value originally suggested in the literature for a T-bar

penetration test). Again, the cone penetrometer demonstrated diverging results,

requiring two separate values for the cone factor, Nkt (10.5 initially increasing to 13 for

depths below 10 m) in order to give su similar to the vane shear tests. This highlights the

possible variability of the cone factor with depth.

Cyclic penetration and extraction tests were performed at specific depths for each full-

flow penetrometer. These tests comprised displacement cycles of ±0.5 m about the

relevant depth, recording the penetration and extraction resistances over five full cycles.

The results may be used to derive the remoulded strength and sensitivity of the soil.

ABSTRACT

ii

Laboratory tests such as triaxial and simple shear tests were performed on ‘undisturbed’

tube samples retrieved from the same site to evaluate the in situ shear strengths in the

laboratory. However, the resulting su data were rather scattered, much of which may be

attributed to variable sample quality due to the presence of frequent shell fragments and

occasional silt lenses within the test samples. In general, N factors for the full-flow

penetrometers, back-calculated using su values measured from the simple shear tests,

fell mainly in a range between 9.7 and 12.8 (between 10.4 and 12.2 for the standard size

T-bar (250 mm x 40 mm)).

Model penetrometer tests and hand vane tests were conducted in the centrifuge on

reconstituted samples of clay material collected from the same site. Two centrifuge

strongbox samples were prepared and tested in-flight. In the first sample, various types

of penetrometers (cone, T-bar, ball and plate) were tested, while in the second sample,

four different lengths (hence aspect ratios) of T-bars were investigated.

The model T-bar and ball penetrometers showed very similar penetration and extraction

resistance profiles, whereas the model cone and (unexpectedly) plate penetrometer

exhibited lower penetration resistance profiles. During extraction, all model

penetrometers show extremely similar resistance profiles. Of particular interest is the

rather gradual and smooth development of the full net cone resistance as the cone was

first extracted − a phenomenon that has also been observed in foundation tests.

The centrifuge tests also showed that the aspect ratio (varied between 4 and 8) of the

T-bar penetrometer does not have any obvious effect on its tip resistance.

T-bar factors, NT-bar, back-calculated using simple shear strengths for the model T-bars

are generally lower than for the field T-bars, perhaps implying greater strain rate effects

for the natural soil than for the reconstituted material.

Besides constant rate penetration tests, variable rate (‘twitch’) penetration tests were

also undertaken in the centrifuge. In the twitch tests, the penetration rate was

successively halved over 8 steps, with each step being triggered after the penetrometer

had advanced by 1 or 2 diameters (the interval being varied between tests). The

interpreted results, in conjunction with the results obtained by other researchers, have

been used to estimate the coefficient of consolidation, cv, of the clay. The cv values

estimated were similar to the cv values measured in oedometer tests (at high vertical

stress) and in Rowe cell tests.

ACKNOWLEDGEMENT

iii

ACKNOWLEDGEMENT

There are many people who have helped me throughout the course of my PhD research

and I would like to take this opportunity to thank them.

First, I would like to express my greatest gratitude to my supervisor, Prof. Mark

Randolph, for providing me with the opportunity to undertake this research, and for his

supervision and support for the research. He has guided me with great patience and

given me countless invaluable advice throughout the course.

Special thanks are due to Prof. Marcio Almeida for his assistance during his short visit

to the University of Western Australia (UWA). His kind words and encouragements are

always much appreciated.

I would also like to thank Dr. Hackmet Joer for his guidance in the early stage of my

study and Dr. Doug Stewart for his help with the field testing.

During my short visit to the Norwegian Geotechnical Institute (NGI) in Norway, I was

grateful to meet Tom Lunne and Knut Andersen, and have interesting discussions with

them. In this regard, I would also like to thank the Centre for Offshore Foundation

Systems (COFS), established and supported under the Australian Research Council’s

Research Centres Program, for funding my visit to NGI. In addition, particular thanks

are due to BP Exploration Operating Company, Statoil, Norsk Hydro, Woodside

Engineering, Petroleo Brasileiro and ConocoPhillips, who were sponsors of the joint

industry project: Characterization of Soft Soils in Deep Water by In Situ Tests,

undertaken jointly by NGI (project leader Tom Lunne) and COFS. The test results

obtained from the joint industry project have constituted the major results presented in

this thesis.

Many thanks must go to Binaya Bhattarai, Claire Bearman, Natalia Kroupnik and Alex

Duff for help in carrying out the field and laboratory tests, and Don Herley for

assistance with the centrifuge tests. In addition, I would like to extend my appreciation

to Nina Levy and Kervin Yeow for carrying out some of the laboratory tests and

assisting in the reports prepared for the joint industry project. The great skills possessed

by the staff of the Civil Engineering Workshop in UWA in manufacturing new

equipment required for the tests are also greatly acknowledged.

Thanks also to Monica Mackman, who is responsible for so much work ‘behind the

ACKNOWLEDGEMENT

iv

scene’. Her ability to organise basically ‘everything’ and make sure that students only

need to concentrate on their technical work is absolutely tremendous.

All my fellow colleagues in COFS have been very supportive and created a pleasant

working environment, for which I am truly grateful. Particularly, I wish to thank Dr.

George Vlahos and Dr. James Doherty, both of whom shared the same postgraduate

student room with me during their doctoral studies, for their laughter and the enormous

encouragement from time to time.

I also wish to thank Dr. Mostafa Ismail, Dr. Conleth O’Loughlin, Dr. Itai Einav and Dr.

Shambu Sharma for reviewing my thesis manuscript and giving constructive comments

on my writing. Additionally, Jing Yun Wong and Prof. Kit Po Wong, who have inspired

me in many ways, also helped to proof read some parts of my thesis. Their efforts are

much appreciated.

During the course of this research, I have received financial support through a

government scholarship (IPRS) and from COFS (in the final stage of the course). This

support is gratefully acknowledged.

Finally, I would like to express my sincere thanks to all of my friends who have helped

me and showed great understanding when my only focus was on the thesis. Also, I am

most indebted to my family and relatives for their unconditional love and support. I

could not have done it without them!

I certify that, except where specific reference is made in the text to the work of others,

the contents of this thesis are original and have not been submitted to any other

university.

Shin Fun Chung

December, 2004

TABLE OF CONTENT

v

TABLE OF CONTENT

ABSTRACT

ACKNOWLEDGEMENT

TABLE OF CONTENT

LIST OF SYMBOLS

LIST OF TABLES

LIST OF FIGURES

1 INTRODUCTION

1.1 Motivations ................................................................................................1-1

1.2 Aim of research..........................................................................................1-3

1.3 Outline of thesis .........................................................................................1-4

2 LITERATURE REVIEW

2.1 Vane shear test ...........................................................................................2-1

2.2 Cone penetration test..................................................................................2-5

2.3 ‘Full-flow’ penetrometer tests....................................................................2-9

2.4 Variable rate penetration test ...................................................................2-12

2.5 Summary ..................................................................................................2-15

3 BURSWOOD TEST SITE

3.1 Site background..........................................................................................3-1

3.2 Soil properties ............................................................................................3-2

4 FIELD TESTING

4.1 Field testing apparatus ...............................................................................4-1

4.1.1 Field penetrometers .......................................................................4-1

4.1.2 Shear vane......................................................................................4-2

4.1.3 Calibration details .........................................................................4-3

TABLE OF CONTENT

vi

4.2 Field testing procedure...............................................................................4-4

4.2.1 Field penetration tests....................................................................4-4

4.2.2 Vane shear tests .............................................................................4-4

4.3 Field test results .........................................................................................4-5

4.3.1 Assessment of penetrometer tip resistance.....................................4-5

4.3.2 Resistance profiles for various penetrometers...............................4-7

4.3.3 Field vane tests.............................................................................4-11

4.3.4 Assessment of undrained shear strength......................................4-11

4.3.5 Cyclic penetration and extraction tests........................................4-12

4.4 Summary for field testing ........................................................................4-13

5 LABORATORY TESTING

5.1 In situ tube samples....................................................................................5-1

5.2 Laboratory testing apparatus and procedure ..............................................5-2

5.2.1 Index tests.......................................................................................5-2

5.2.2 Constant rate of strain consolidation (CRSC) ...............................5-3

5.2.3 UU and CAU triaxial tests .............................................................5-3

5.2.4 CAU simple shear test....................................................................5-4

5.2.5 Model T-bar test in triaxial ............................................................5-5

5.2.6 Calibration for T-bar triaxial test ..................................................5-6

5.2.7 SHANSEP procedure .....................................................................5-6

5.3 Laboratory test results ................................................................................5-7

5.3.1 Index tests.......................................................................................5-7

5.3.2 CRSC test .......................................................................................5-8

5.3.3 UU and CAU triaxial tests ...........................................................5-11

5.3.4 CAU simple shear test..................................................................5-11

5.3.5 Model T-bar test in triaxial ..........................................................5-12

5.3.6 Triaxial and simple shear tests following SHANSEP ..................5-13

TABLE OF CONTENT

vii

5.4 Undrained shear strength profiles ............................................................5-14

5.5 Summary for laboratory testing ...............................................................5-16

6 CENTRIFUGE TESTING

6.1 Reconstituted sample properties ................................................................6-1

6.2 Centrifuge testing apparatus.......................................................................6-2

6.2.1 Model penetrometers and hand vane apparatus............................6-2

6.2.2 Calibration details .........................................................................6-3

6.3 Centrifuge testing procedure......................................................................6-3

6.3.1 Sample preparation........................................................................6-3

6.3.2 Penetration and hand vane tests ....................................................6-4

6.3.3 Laboratory testing on centrifuge samples......................................6-4

6.4 Centrifuge test results.................................................................................6-5

6.4.1 Consolidation in centrifuge............................................................6-5

6.4.2 Assessment of model penetrometer tip resistance..........................6-6

6.4.3 Resistance profiles for various model penetrometers ....................6-7

6.4.4 Hand vane tests ............................................................................6-13

6.4.5 Laboratory testing on centrifuge samples....................................6-13

6.4.6 Profiles of undrained shear strength ...........................................6-14

6.5 Summary for centrifuge testing................................................................6-16

7 CORRELATION OF PENETRATION RESISTANCE AND UNDRAINED

SHEAR STRENGTH

7.1 Cone factor, Nkt ..........................................................................................7-1

7.2 T-bar factor, NT-bar......................................................................................7-5

7.3 Ball factor, Nball..........................................................................................7-8

7.4 Plate factor, Nplate .......................................................................................7-9

7.5 Summary and recommendations..............................................................7-10

8 EFFECTS OF PARTIAL CONSOLIDATION AND VISCOSITY

TABLE OF CONTENT

viii

8.1 Resistance profiles of twitch tests ..............................................................8-1

8.2 Evaluation of consolidation coefficient from twitch tests..........................8-4

8.3 Effect of partial consolidation for various penetrometers..........................8-6

8.4 Effect of penetration rate in viscous (undrained) region............................8-8

8.4.1 Model twitch tests in centrifuge .....................................................8-9

8.4.2 Field twitch test results ................................................................8-12

8.5 Summary for effects of penetration rate ..................................................8-13

9 CONCLUSIONS

9.1 Findings of research ...................................................................................9-1

9.2 Recommendations for future work ............................................................9-6

REFERENCES

TABLES

FIGURES

APPENDIX A

APPENDIX B

APPENDIX C

APPENDIX D

APPENDIX E

APPENDIX F

APPENDIX G

LIST OF SYMBOLS

ix

LIST OF SYMBOLS

AN Inner cross-sectional area of cone head

Ap Projected area of penetrometer in plane normal to shaft

As Cross-sectional area of connection shaft in plane normal to shaft

AT Total cross-sectional area of cone head

δA Increment of shear surface area

B Breadth of rectangular footing

or ratio of change in pore pressure to change in cell pressure

Bq Ratio of excess pore pressure to net bearing pressure

cv Coefficient of consolidation

cvy Coefficient of consolidation determined at yield stress

Cc Compression index

COV Coefficient of variation

CAU Consolidated anisotropically undrained

CPT Cone penetration test

CPTU Cone penetration test with pore pressure measurement (or

piezocone test)

CRSC Constant rate of strain consolidation

d Diameter of penetrometer or shear vane

or length of drainage path

dE External diameter of sampling tube

e Void ratio

ei Initial void ratio

eo In situ void ratio

Δe Change in void ratio

fs Sleeve friction

g Acceleration due to gravity

G Shear modulus

Go Small strain shear modulus

Gs Specific gravity of soil particles

H Sample height

Hi Initial sample height

ΔH Change in sample height

LIST OF SYMBOLS

x

IP Plasticity index

Ir Rigidity index

k Coefficient of permeability

ko In situ coefficient of permeability

Ko Coefficient of earth pressure at rest

(Ko)nc Ko for normally consolidated soil

L Length of penetrometer or shear vane

n Number of data

or scaling ratio for centrifuge model (test acceleration level)

N Bearing factor for a penetrometer, defined as qnet/su

N Mean value of bearing factor data

Nav Bearing factor derived from the average laboratory strength (su, av)

Ni Empirical data of bearing factor

Nke Cone factor, defined as (qt − u2)/su

Nkt Cone factor, defined as (qt − σvo)/su

NΔu Cone factor, defined as Δu/su

OCR Overconsolidation ratio

OCR1 Actual OCR value in SHANSEP test

OCR2 In situ OCR derived from CRSC tests

p' Mean effective stress

q Deviator stress

or tip resistance

qc Measured cone resistance

qcnet Net cone resistance

qf Deviator stress at failure

qin Penetration resistance

qm Measured tip resistance

qnet Net tip resistance

qout Extraction resistance

qref Reference tip resistance

qt Total cone resistance

Qm Force applied (or measured by load cell)

Qnet Force exerted on penetrometer due to soil resistance

r Radius from centre of vane to shear surface

LIST OF SYMBOLS

xi

RL Reduced Level relative to datum sea level in Perth

s' Mean effective stress in plane strain conditions

su Undrained shear strength

su 1 Measured undrained shear strength in SHANSEP test

su 2 Adjusted undrained shear strength in SHANSEP test

su, av Average of triaxial compression, triaxial extension and simple

shear strengths

St Sensitivity

SD Standard deviation

SHANSEP Stress history and normalised soil engineering properties

SS Simple shear test

t Maximum shear stress in plane strain conditions

or real time

or thickness of vane blade

or wall thickness of sampling tube

t90 Real time to achieve 90 % degree of consolidation

Δt Change in time

T90 Time factor for 90 % degree of consolidation

Tq Torque required to shear soil

Tv Time factor for consolidation for surface footing

TXC, TXE Triaxial compression and extension tests

u Pore pressure

ub Pore pressure measured at sample base

uo Hydrostatic water pressure

u1 Pore pressure measured at conical face of cone tip

u2 Pore pressure measured immediately behind cone tip, at shoulder

position

u3 Pore pressure measured behind friction sleeve

Δu Excess pore pressure

Δuf Change in pore pressure at failure

Us Degree of consolidation (settlement) for surface footing

UU Unconsolidated undrained

v Rate of penetration

V Non-dimensional velocity

LIST OF SYMBOLS

xii

Vo Value of V for which the viscous effects start to decay

Vref Reference value of V

α Perimeter ratio

or unequal area ratio

αc Roughness coefficient for cone-soil interface

Δ In situ stress ratio

εa Axial strain

εaf Axial strain at failure

εv Volumetric strain

φ' Friction angle

φ'centrifuge Friction angle for reconstituted centrifuge sample

φ'field Friction angle for natural soil

γ Shear strain

or unit weight

γf Shear strain at failure

γ' Effective unit weight

γw Unit weight of water

λ Rate parameter

ξ95 Cumulative shear strain for 95 % of shear strength degradation

σh , σ'h Total and effective horizontal stresses

σho , σ'ho Total and effective in situ horizontal stresses

σv , σ'v Total and effective vertical (overburden) stresses

σvo , σ'vo Total and effective in situ vertical (overburden) stresses

σ'v1 , σ'v2 Effective vertical stresses applied in stage 1 (normal consolidation)

and stage 2 (swelling) in SHANSEP test

σ'yield Effective preconsolidation or yield stress

σ'yield 1 Actual yield stress experienced in SHANSEP test

σ'yield 2 In situ yield stress derived from CRSC tests

σ'1 , σ'3 Effective major and minor principal stresses

Δσ3 Change in cell pressure in triaxial and simple shear tests

τ Shear stress

ω Moisture content

LIST OF SYMBOLS

xiii

ωf Moisture content at the end of testing

ωi Initial moisture content

ωL Liquid limit

ωP Plastic limit

(su /σ'v)nc Ratio of su /σ'v for normally consolidated clay

(su /σ'v)oc Ratio of su /σ'v for overconsolidated clay

LIST OF TABLES

xiv

LIST OF TABLES

Table 2.1 Summary of constants derived for the backbone curves

Table 4.1 Calibration details for field test apparatus

Table 5.1 Dimensions of sampling tubes

Table 5.2 Summary of CRSC tests

Table 5.3 NGI’s criterion for sample quality

Table 5.4 Summary of UU tests

Table 5.5 Summary of CAU triaxial tests

Table 5.6 Summary of CAU simple shear tests

Table 5.7 Summary of T-bar in triaxial tests

Table 5.8 Summary of testing undergone SHANSEP procedure

Table 6.1 Soil properties for reconstituted Burswood clay

Table 6.2 Scaling relationships for centrifuge models

Table 6.3 Offset of tip resistance due to error generated from changes in

normal stress on the load cell

Table 6.4 Summary of CRSC tests on samples from centrifuge testing

Table 6.5 Summary of triaxial and simple shear tests on samples from

centrifuge testing

Table 7.1 Laboratory shear strengths adopted for calculating N values for the

field penetrometers

Table 7.2 Summary and comparison of cone factors (Nkt) for various clay

sites

Table 7.3 Summary and comparison of T-bar factors (NT-bar) for various clay

sites

Table 7.4 Summary of bearing factors (N) for the various shaped

penetrometers

LIST OF FIGURES

xv

LIST OF FIGURES

Figure 3.1 Map of test location at Burswood

Figure 4.1 (a) Test location layout: Location of main testing area

(b) Test location layout: Details of main testing area

Figure 4.2 Diagram of field cone penetrometer

Figure 4.3 Diagram of field T-bar penetrometer

Figure 4.4 Photograph of field cone, T-bar, ball and plate penetrometers

Figure 4.5 Diagram of shear vane

Figure 4.6 Photographs of truck and saturating a penetrometer

Figure 4.7 Photograph of frame for jacking the shear vane

Figure 4.8 Corrections of field cone and T-bar data

Figure 4.9 (a) Comparison of smooth and rough T-bar penetration resistances

(b) Comparison of smooth and rough T-bar extraction resistances

Figure 4.10 Ratios of average smooth to rough T-bar resistances

Figure 4.11 (a) Comparison of smaller and standard T-bar penetration

resistances

(b) Comparison of smaller and standard T-bar extraction

resistances

Figure 4.12 (a) Comparison of cone and T-bar penetration resistances

(b) Comparison of cone and T-bar extraction resistances

Figure 4.13 Comparison of friction ratios from cone tests

Figure 4.14 Comparison of Bq values from cone tests

Figure 4.15 (a) Comparison of ball and T-bar penetration resistances

(b) Comparison of ball and T-bar extraction resistances

Figure 4.16 (a) Comparison of plate and T-bar penetration resistances

(b) Comparison of plate and T-bar extraction resistances

Figure 4.17 Summary of tip resistance profiles for all various penetrometers

Figure 4.18 Comparison of ratios of extraction to penetration resistances

Figure 4.19 Peak and remoulded shear strengths from field vane tests

Figure 4.20 Sensitivity of clay from field vane tests

Figure 4.21 Comparison of undrained shear strength profiles

Figure 4.22 Cyclic penetration response for T-bar 1

Figure 4.23 (a) Cyclic penetration response for T-bar 3 (4 m and 9 m)

LIST OF FIGURES

xvi

(b) Cyclic penetration response for T-bar 3 (14 m and summary)


Figure 4.25 Cyclic penetration response for Smaller T-bar 1


Figure 4.27 Cyclic penetration response for Ball 1


Figure 4.29 Cyclic penetration response for Plate 1


Figure 4.31 Summary of degradation parameters for cyclic penetrometer tests

Figure 4.32 Smoothed degradation curves for cyclic penetrometer tests

Figure 5.1 X-ray of tube samples collected from the field

Figure 5.2 Schematic diagram of CRSC apparatus

Figure 5.3 Schematic diagram of triaxial apparatus

Figure 5.4 Schematic diagram of simple shear apparatus

Figure 5.5 Schematic diagram of T-bar in triaxial apparatus

Figure 5.6 Photographs of T-bar penetrometer for T-bar triaxial test

Figure 5.7 Natural water content, Atterberg limits and unit weight profiles

Figure 5.8 Grading curve for Burswood clay material

Figure 5.9 In situ vertical stress profiles

Figure 5.10 OCR and yield stress profiles

Figure 5.11 Compression index and compression ratio profiles

Figure 5.12 Void ratio and consolidation coefficient profiles

Figure 5.13 (a) Comparison of laboratory and field T-bar penetration

resistances

(b) Comparison of laboratory and field T-bar extraction resistances

(c) Extraction to penetration ratios for laboratory and field T-bars

Figure 5.14 (a) Undrained shear strength profiles from laboratory testing

(b) Undrained shear strength profiles including SHANSEP results

Figure 6.1 Photograph of centrifuge and actuator

Figure 6.2 (a) Photograph of model cone penetrometer

(b) Penetrometer rod attached with a model T-bar tip

Figure 6.3 Photographs of hand vane apparatus

Figure 6.4 Box 1: Details of testing layout


LIST OF FIGURES

xvii

Figure 6.6 Consolidation of Box 1: Settlement versus root time (minutes)


Figure 6.8 (a) Box 1: Comparison of penetration resistances for T-bars (20x5)

(b) Box 1: Comparison of normal and twitch tests for T-bars (20x5)

(c) Box 1: Comparison of extraction resistances for T-bars (20x5)

Figure 6.9 (a) Box 1: Comparison of cone and T-bar (20x5) penetration

resistances

(b) Box 1: Comparison of cone and T-bar (20x5) extraction

resistances

Figure 6.10 (a) Box 1: Comparison of ball and T-bar (20x5) penetration

resistances

(b) Box 1: Comparison of ball and T-bar (20x5) extraction

resistances

Figure 6.11 (a) Box 1: Comparison of plate and T-bar (20x5) penetration

resistances

(b) Box 1: Comparison of plate and T-bar (20x5) extraction

resistances

Figure 6.12 Box 1: Summary of tip resistances for all model penetrometers

Figure 6.13 Box 1: Ratios of extraction to penetration resistances for T-bars

(20x5)

Figure 6.14 Box 1: Summary of resistance ratios for all model penetrometers


(b) Box 2: Comparison of extraction resistances for T-bars (20x5)

Figure 6.16 (a) Box 2: Penetration resistances of T-bars 30x5 and 20x5

(b) Box 2: Extraction resistances of T-bars 30x5 and 20x5





Figure 6.19 Box 2: Summary of tip resistances for all various model T-bars

Figure 6.20 Box 2: Ratios of extraction to penetration resistances for T-bar

(20x5)

Figure 6.21 Box 2: Ratios of extraction to penetration resistances for all T-bars

Figure 6.22 Box 1: Results of hand vane tests

LIST OF FIGURES

xviii


Figure 6.24 Average results of hand vane tests for Boxes 1 and 2

Figure 6.25 Sensitivity of reconstituted clay sample (Box 2)

Figure 6.26 T-bar test in triaxial sample recovered from Box 1

Figure 6.27 Box 1: Summary of undrained shear strength profiles


Figure 6.29 Comparison of undrained shear strength profiles for Boxes 1 and 2

Figure 7.1 Best-fit trends for laboratory su data to evaluate the average shear

strength, su, av = (su, TXC + su, TXE + su, SS)/3

Figure 7.2 Cone factors, Nkt, calculated using different laboratory su values

Figure 7.3 Cone factors, Nkt, calculated using vane su values

Figure 7.4 T-bar factors, NT-bar, calculated using different laboratory su values

Figure 7.5 T-bar factors, NT-bar, calculated using vane su values

Figure 7.6 T-bar factors, NT-bar, calculated using simple shear strengths for

model T-bars tested in the centrifuge

Figure 7.7 T-bar factors, NT-bar, calculated using hand vane strengths for

model T-bars tested in the centrifuge

Figure 7.8 Ball factors, Nball, calculated using different laboratory su values

Figure 7.9 Ball factors, Nball, calculated using vane su values

Figure 7.10 Plate factors, Nplate, calculated using different laboratory su values

Figure 7.11 Plate factors, Nplate, calculated using vane su values

Figure 8.1 Resistance profiles of twitch tests for various shaped model

penetrometers

Figure 8.2 Resistance profiles of twitch tests for model T-bars with various

aspect ratios

Figure 8.3 Evaluation of consolidation coefficient using data from T-bar

twitch tests based on various backbone curves

Figure 8.4 Evaluation of consolidation coefficient using data from cone twitch

test

Figure 8.5 Non-dimensional plot for various shaped model penetrometers

Figure 8.6 Non-dimensional plot for different model T-bar penetrometers

Figure 8.7 Degree of consolidation (settlement), Us, versus time factor, Tv, for

surface footings − reproduced from Davis & Poulos (1972)

Figure 8.8 (a) Fitting results of model T-bar twitch test in Box 1: focus in

LIST OF FIGURES

xix

viscous region

(b) Fitting results of model T-bar twitch test in Box 1: overall best-

fit

Figure 8.9 (a) Fitting results of model cone twitch test in Box 1: focus in

viscous region

(b) Fitting results of model cone twitch test in Box 1: overall best-

fit

Figure 8.10 (a) Fitting results of various model T-bar twitch tests in Box 2,

using backbone curve from Watson & Suemasa (2000): focus in

viscous region

(b) Fitting results of various model T-bar twitch tests in Box 2,

using backbone curve from Watson & Suemasa (2000): overall

best-fit

Figure 8.11 (a) Fitting results of various model T-bar twitch tests in Box 2,

using backbone curve after Randolph & Hope (2004): focus in

viscous region

(b) Fitting results of various model T-bar twitch tests in Box 2,

using backbone curve after Randolph & Hope (2004): overall best-

fit

Figure 8.12 (a) Comparison of backbone curves from Watson & Suemasa

(2000) with different Vo

(b) Comparison of backbone curves from Randolph & Hope (2004)

with different Vo

Figure 8.13 (a) Fitting results of field T-bar twitch tests reported by Schneider

et al (2004), using backbone curve from Watson & Suemasa (2000)

(b) Fitting results of field T-bar twitch tests reported by Schneider

et al (2004), using backbone curve from Randolph & Hope (2004)

Figure 8.14 Fitting results of field cone twitch tests reported by Schneider et al

(2004), using the backbone curve from Randolph & Hope (2004)

INTRODUCTION

1-1

1 INTRODUCTION

1.1 Motivations

In deep water developments, where the water depths are greater than a few hundred

metres, sediments encountered are generally very soft, normally consolidated fine-

grained deposits. Accurate assessment of the geotechnical properties for such sediments

are essential in order to allow economic design of subsea facilities, pipelines and

foundation or anchoring systems.

For example, the required size of foundation and anchoring systems is intrinsically

linked to the shear strength of the seabed sediment. Accurate assessment of the strength

profile may allow significant savings by reducing the size of the foundation without loss

of reliability. For pipeline routes, the characteristics of the upper 1 or 2 m are critical for

determination of lateral stability, self-burial capability, or potential problems with

trenching. Thus, testing methods adopted must allow accurate determination of near-

surface strength profiles and detection of the presence (and horizontal connectivity) of

stronger layers. Reliable soil parameters are also essential for evaluation of natural

geohazards like submarine slides.

Although shear strength is the fundamental soil parameter for geotechnical design,

engineers always face the question: what is the ‘true’ value for the shear strength?

Unfortunately, due to strength anisotropy, shear strength does not have a unique value

(Wroth, 1984); rather, its value is dependent on several factors, such as the mode of

shearing, strain rate and strain softening of the soil. Consequently, results of shear

strength measurements vary depending on the testing methods adopted and engineers

are required to determine which strengths are best related to their particular design

problems.

In laboratory, measurements of shear strength can be obtained from triaxial and simple

shear tests conducted on samples collected from the field. In general, the various

laboratory strengths are found to have the following pattern: su, TXC > su, SS > su, TXE,

where su, TXC, su, SS and su, TXE are the shear strengths measured from triaxial

compression, simple shear and triaxial extension tests respectively (Wroth, 1984).

However, it is extremely difficult to obtain high quality samples, especially from deep

water sites, for laboratory tests. Disturbance is inevitable due to both sampling

(particularly where thick-walled gravity corers are used) and total stress reduction when

INTRODUCTION

1-2

the sample is recovered to the ground surface. In addition, the presence of gas in the

pore fluid may distort the response in undrained tests, or the gas may come out of

solution during stress reduction, leading to fracturing of the sample on horizontal planes

and possible formation of voids within the sample (Lunne et al, 2001). As a result, the

actual (maximum) strength of the sample may not be recovered in the laboratory and the

in situ strength will be underestimated.

The difficulties in obtaining high quality samples for laboratory tests have increased the

reliance on results from in situ tests for soil characterisation. Common in situ testing

tools used for clayey sites, both onshore and offshore, are the cone penetration test with

pore pressure measurement (CPTU) and the vane shear test. The latter test provides

direct measurements of the peak and remoulded undrained shear strengths; nevertheless,

the measurements can be affected by factors such as the waiting time before rotating the

vane, rate of rotation, aspect ratio of the vane, etc (Chandler, 1988), which often leads to

high degree of scatter in the data measured. Besides, the vane test can only be

performed at discrete depths, hence giving only discrete data for the strength profile.

In contrast, the CPTU gives much more consistent and continuous profiles of data of the

soil with depth. Yet, the primary disadvantage of the CPTU is that it does not measure

undrained shear strength directly, but the tip resistance exerted on the cone

penetrometer when it is being pushed into the soil. The resistance profile (after

appropriate corrections for effects of pore pressure and overburden pressure) is then

correlated to the shear strength profile by using a cone factor, Nkt. While extensive

theoretical solutions and empirical correlations have been published on the value of Nkt,

it is still customary in offshore practice to calibrate the value for each new site using

laboratory triaxial and simple shear test data, due to the high variability of Nkt

depending on soil properties such as the soil stiffness and the in situ stress ratio (Teh &

Houlsby, 1991).

A good illustration of the problem is given by Quirós & Little (2003) for lightly

overconsolidated clay in the Gulf of Mexico, where the cone factor was varied from

around 17 at shallow depths to 11 at 120 m in order to fit the profile of net cone

resistance to the design shear strength profile computed based on laboratory strengths.

In more extreme cases, Nkt values (based on shear strength from triaxial compression

tests) ranging from 8 to 29 were also reported (Rad & Lunne, 1988; Lunne et al, 1997b).

Of course one should be cautious that sample disturbance could severely distort the

INTRODUCTION

1-3

values of Nkt.

Besides the problem associated with the choice of Nkt, the need to correct the measured

cone resistance (before being correlated to the shear strength) for the effects of pore

pressure and overburden pressure is also unfavourable to interpretation, because errors

in the adjustments for the pore pressure and overburden pressure effects propagate into

the net cone resistance and the consequential shear strength.

In an attempt to avoid the uncertainties involved with interpretation of the CPTU in

clays, while preserving the advantage of a continuous profile of resistance, the T-bar

penetrometer was introduced (Stewart & Randolph, 1991). This was first tested in the

centrifuge and later used in the field, both onshore (Stewart & Randolph, 1994) and

offshore (Randolph et al, 1998). The principle behind the T-bar, which consists of a

cylindrical bar mounted at right angles to the push-rods, is that soil is able to flow

around the cylinder from front to back, causing a very localised plastic mechanism.

The relationship between net bearing resistance on the projected area of the T-bar and

the shear strength of the soil was originally based on the plasticity solution of Randolph

& Houlsby (1984). This, in principle, avoids the need for calibration against laboratory

strength data at each site. Besides, the T-bar penetrometer would give more reliable net

resistance compared to the cone, since the soil is able to flow around the T-bar, thereby

equilibrating the overburden pressure above and below the T-bar, except at the shaft;

however, since the projected area of the T-bar is considerably larger than the area of the

shaft, the correction for overburden stress is relatively insignificant.

Alternative ‘full-flow’ penetrometers consisting of either a spherical ball or circular

plate have also been investigated recently, mainly through centrifuge tests (Watson et al,

1998) or numerical analyses (Lu et al, 2000). The idea is that the axisymmetric shape of

these penetrometers would reduce the potential for the load cell to be subjected to

bending moments arising from non-symmetric resistance along the T-bar penetrometer.

1.2 Aim of research

The aim of this research is to provide an improved quantitative framework for the

characterisation of soft offshore sediments, with emphasis on in situ testing methods.

Specifically, the study aims to:

• Compare the penetration and extraction resistances of different types of

INTRODUCTION

1-4

penetrometers (cone, T-bar, ball and plate) in soft clay. This subsequently allows

correlation of net resistance profiles obtained for the different penetrometers with

undrained shear strength data measured from the vane and laboratory (triaxial and

simple shear) tests.

• Provide guidelines for the choice of bearing factors (N factors) for each

penetrometer (based on all available results in the database of empirical N factors)

in order to deduce the various shear strength profiles for lightly overconsolidated

clay.

• Explore responses of the full-flow penetrometers (T-bar, ball and plate) when

subjected to cyclic penetration and extraction tests. Data from the cyclic tests can

be used to compute sensitivity of the clay as the ratio of the fully remoulded tip

resistance to the initial tip resistance. The resulting sensitivities can then be

compared with values assessed from vane tests.

• Evaluate partial consolidation effects (predominant in partially drained conditions)

and viscous effects (predominant in fully undrained conditions) on the tip resistance

of each penetrometer.

• Validate the potential of using data interpreted from the variable rate (‘twitch’)

penetration tests to estimate the coefficient of consolidation of the clay.

1.3 Outline of thesis

The following chapter gives a review of the literature on relevant in situ testing

methods: first, the vane shear test; second, the cone penetration test; third, full-flow

penetration tests; and finally, the variable rate penetration test.

This is then followed by a chapter describing the site chosen for carrying out the in situ

testing and sampling for this research, in terms of location and geotechnical soil

properties of the site.

Chapter 4 explains the in situ penetration tests and the vane shear tests undertaken at the

site described in Chapter 3. Details of the penetrometers and the vane adopted, as well

as descriptions of the calibration and testing procedures are provided here. The in situ

test results are then presented: first, comparing the penetration and extraction resistances

for the various types of penetrometers; and second, comparing the shear strength

profiles deduced from the penetration test data with strength measurements obtained

from the vane tests. After that, results of the cyclic penetration and extraction tests are

INTRODUCTION

1-5

presented.

Chapter 5 subsequently explains the laboratory tests performed on the ‘undisturbed’

tube samples collected from the site described in Chapter 3. The laboratory tests carried

out include the constant rate of strain consolidation (CRSC) tests, triaxial (compression

and extension) tests, simple shear tests and model T-bar tests in triaxial conditions.

Triaxial (compression and extension) and simple shear tests following SHANSEP

procedure were also performed. Details of the testing procedures are given, followed by

the results obtained from these tests, which help to characterise the soil properties of the

site. Undrained shear strengths measured from the various laboratory tests are presented

for comparison and the strength data are used later for calibration of the empirical

bearing factors (N factors) presented in Chapter 7.

In Chapter 6, model penetration tests and hand vane tests conducted on reconstituted

clay samples prepared in the centrifuge are described. The bulk material used for

preparing the centrifuge samples was collected from the site described in Chapter 3.

Dimensions of the model penetrometers and testing procedures are described in detail.

Besides constant rate (normal) penetration tests, variable rate (twitch) penetration tests

were also conducted on the model penetrometers. However, the primary focus is given

to the results obtained from the normal penetration tests in this chapter. Interpretation

and discussion of the twitch tests are presented later in Chapter 8. Also, tube samples

were retrieved from the centrifuge samples for additional laboratory testing, so that

correlation of the tip resistances measured on the model penetrometers and the

laboratory strengths could be examined.

Chapter 7 presents correlation between the net penetration resistance for each

penetrometer and the undrained shear strength. The strength values used for calculation

of the N factors are those data measured from the triaxial compression tests, simple

shear tests, average values of the triaxial (compression and extension) and simple shear

tests, and the vane shear tests. Additional data of the empirical N factors for two other

clay sites (quoted from the reports of NGI-COFS (2004a, 2004b)) are presented for

comparison with the results obtained for the site described in Chapter 3. This chapter

also provides general guidelines for the choice of N factors for each penetrometer, based

on the available empirical data, in order to deduce the various shear strength profiles for

lightly overconsolidated clay from the penetration test data.

In Chapter 8, the effects of partial consolidation and viscosity on the tip resistances of

INTRODUCTION

1-6

the various shaped penetrometers are examined, in the light of the model twitch test

data. The characteristic of the tip resistance increasing with decreasing penetration rate

(when the condition is partially drained and partial consolidation effects are dominant)

is exploited to estimate the consolidation coefficient and the results are compared with

values measured from the oedometer tests and the Rowe cell tests. This chapter also

presents an attempt to quantify the viscous effects on the (normalised) tip resistance

when the condition is fully undrained, based on the model and field (reported by

Schneider et al (2004)) twitch test data.

Finally, Chapter 9 summarises the main outcomes of this research and provides

suggestions for future studies.

LITERATURE REVIEW

2-1

2 LITERATURE REVIEW

As has been pointed out in the previous chapter, the difficulties in obtaining high quality

samples for laboratory tests have increased the reliance on results from in situ tests for

soil characterisation. Traditionally, the vane shear test and the cone penetration test are

the common in situ tools used for clayey sites. Nevertheless, the T-bar penetrometer test

has recently started to be widely accepted and used as an alternative (additional) device

to measure soil parameters. The potential of other ‘full-flow’ penetrometers such as ball

and circular plate has also been recognised and studied recently, although mainly by

way of centrifuge tests or numerical analyses.

This chapter reviews the developments and associated challenges of the in situ testing

methods mentioned above: first, the vane shear test; second, the cone penetration test

with pore pressure measurement (CPTU); third, ‘full-flow’ (T-bar, ball and plate)

penetrometer tests; and finally the variable rate penetration test.

2.1 Vane shear test

The primary advantage of the vane shear test is that it gives direct in situ measurement

of the undrained shear strength profile of the soil. In addition to the peak value, the vane

test also measures the post-peak and remoulded strengths (thus the sensitivity) of the

soil at the same depth. It is the only in situ test with this capability at this time.

The vane shear test was originally used in Sweden in 1919 and has been employed

extensively around the world since the late 1940s. In 1947, the Swedish Geotechnical

Institute began an extensive study on the vane shear test. A standard test procedure was

then established and reported by Cadling & Odenstad (1950).

Detailed historical development of the vane test was given by Flodin & Broms (1981)

and Young et al (1988). This section, however, gives an overview of the discussions on

factors that may influence the results of the vane test. These factors include:

• disturbance from insertion of the vane;

• waiting time after insertion, but before rotating the vane;

• rate of rotation of the vane (strain rate effects);

• geometry of the failure surface;

• aspect ratio (length/diameter) of the vane due to anisotropy of strength on the

horizontal shear plane versus that on the vertical shear plane;

LITERATURE REVIEW

2-2

• errors caused by soil-rod and internal friction.

According to Chandler (1988), the ‘standard’ shear vane that seems to be globally

accepted consists of four rectangular blades with thickness, t, of 2 mm, set at right

angles. The standard vane has a length, L, and diameter, d, of 130 and 65 mm

respectively, giving a vane aspect ratio, L/d, of 2. Additionally, the commonly accepted

waiting time following insertion of the vane seems to be no more than 5 minutes, and

the rate of vane rotation to be 0.1°/s.

As the vane is inserted into the ground, soil particles are displaced and hence local

disturbance is inevitable in the soil adjacent to the vane. The amount of disturbance

caused by the vane insertion is often described using a vane ‘perimeter ratio’ suggested

by Cadling & Odenstad (1950). The perimeter ratio, α, is defined as:

dt4

π=α (2.1)

This ratio indicates that a small vane will in fact induce more soil disturbance than a

relatively larger vane, given that the blade thicknesses are the same. LaRochelle et al

(1973) showed that the measured undrained shear strength decreased almost linearly as

the perimeter ratio was reduced, and suggested that the ‘true’ undisturbed strength of the

soil might be inferred by back extrapolating the perimeter ratio to zero (zero blade

thickness). The extrapolated undisturbed strength was 15 % higher than the value

measured using the standard vane, which has a perimeter ratio of 4 %. Similar findings

were also reported by Cereto & Lutenegger (2004), for various sized vane tests

undertaken in varved clay. However, Roy & Leblanc (1988) demonstrated that the

disturbance due to vane insertion can be reduced by using tapered vane blades.

Besides soil displacement, insertion of the vane also generates excess pore pressure,

which leads to reduction of strength. The loss of strength may be recovered if the excess

pore pressure generated is allowed to dissipate, so that local consolidation of the soil

adjacent to the vane can occur prior to shearing. However, Chandler (1988) argued that,

if the waiting time before shearing was too long (for instance more than an hour), the

increase of strength resulting from the consolidation could be greater than the reduction

of strength caused by the vane insertion, resulting in overestimation of the undisturbed

strength. Besides, since the rate of increase of strength is associated with the

consolidation coefficient of the soil, the degree of strength increase for a given waiting

LITERATURE REVIEW

2-3

time can vary drastically between soil types, regardless of the fact that the final gain of

strength (if sufficient waiting time before shearing is allowed) could be similar (Roy &

Leblanc, 1988; Torstensson, 1977).

Significance of the effects of vane rotation rate on the measured shear strength has been

discussed by Chandler (1988). The effects are governed by two aspects: one is viscous

effects, which lead to increase of strength as the rotation rate is increased under

undrained conditions; the other aspect is consolidation effects, which dominate under

partially drained conditions and cause an increase of strength when the rotation rate is

reduced. These two conflicting effects are apparent in the vane test results presented by

Roy & Leblanc (1988), showing normalised shear strength increasing in both directions

with respect to a reference rate of 0.22°/s. A simplified analysis performed by Randolph

(2004) illustrated that the shear strain rate associated with the standard vane test (rotated

at 0.1°/s) is ~104 times the typical strain rates in laboratory tests and ~106 times the

strain rate associated with the base failure of a typical caisson. With such huge

differences in strain rate, one would not be surprised to see a variation of shear strengths

for the various problems or testing methods.

In order to take into consideration the differences in strain rates and to allow

comparison of the vane strength with that associated with foundation failure, Bjerrum

(1973) introduced vane correction factors that ranged from 0.6 (high plasticity) to 1

(low plasticity). Aas et al (1986) also presented a detailed review of vane correction

factors and concluded that the best correlation was with the strength ratio (vane strength

to effective overburden pressure), with the correction factor decreasing from unity for a

strength ratio of 0.2, down to 0.6 at a ratio of 0.6. Nonetheless, in offshore practice, the

vane shear strengths are rarely adjusted, even for strength ratios in excess of 0.4

(Randolph, 2004). A partial justification for this practice is the inevitable disturbance

due to insertion of the vane.

Veneman & Edil (1988) found that the actual failure surface is not a sharply defined

cylindrical surface, but rather a shear zone, and this failure surface may or may not have

fully developed when the measured torque reached a maximum value. Therefore, the

undrained shear strength computed based on the assumption of a fully developed

cylindrical surface may underestimate the actual soil strength. Nevertheless, it is

probably too difficult and not worthwhile to determine the actual surface mobilised

when the maximum torque is measured; common practice is to adopt the fully

LITERATURE REVIEW

2-4

developed cylindrical shear surface for assessing the shear strength.

Also, Skempton (1948) found that the diameter of the shear surface is slightly larger

than the vane diameter and suggested an effective diameter of 1.05 times the vane

diameter used for the shear surface. However, this suggested diameter may be

considerably larger for organic soils or peat, as Åhnberg et al (2004) pointed out that the

fibre content in the organic soils or peat would cause the shear zone to extend far

outside the perimeter of the vane. This may result in the opposite trend that small vanes

yield higher shear strengths than larger vanes (with the same blade thickness), albeit

according to the perimeter ratio, the small vanes will induce more disturbance during

insertion.

The distribution of shear stress around the periphery of the rotated vane has been found

to be highly non-uniform on the top and bottom horizontal planes of the vane, but

reasonably close to uniform on the vertical plane of the cylindrical shear surface

(Donald et al, 1977; Menzies & Merrifield, 1980). Wroth (1984) and Chandler (1988)

demonstrated that, for a standard vane with L/d of 2, the vertical plane of the shear

surface will contribute more than 90 % of the resistance to the total torque during

shearing of the vane, even if strength anisotropy has been taken into account. However,

the portion of contribution from the vertical plane will decrease with reduced vane

aspect ratio and vice versa. The aspect ratio effects are clearly illustrated in the results

reported by Watson et al (2000), who performed extensive model vane tests in-flight in

the centrifuge, on both overconsolidated and normally consolidated fine-grained

materials. Their results indicated that peak shear strength measured from the vane could

increase up to 30 % with L/d increasing from 0.33 to 1.5.

Errors in strength measurements caused by friction between the soil and push-rod

interface, and friction within the vane apparatus system itself were examined by Ortigão

& Collet (1988). Their results concluded that the soil-rod and internal friction could

significantly affect the strength measurements and is difficult to eliminate. Åhnberg et

al (2004) also noted that, for the field vane tests conducted using vane apparatus with a

slip coupling (but without casing), there is a small peak in the soil-rod friction and that

it decreases with further rotation. Consequently, it is impossible to determine the soil-

rod friction accurately. This can result in substantial error in the evaluated shear strength

if the friction constitutes a relatively large portion (possibly > 50 % for small vane) of

the total torque measured.

LITERATURE REVIEW

2-5

In conclusion, although the vane shear test provides valuable measurements of the peak,

post-peak and remoulded undrained shear strengths, the data can be affected

significantly by each of the factor discussed previously. Besides, the vane test can only

be performed at discrete intervals of depth, and therefore, it does not give a continuous

profile of the shear strength. Consequently, it is often used in companion with other

testing devices such as the cone penetration test.

2.2 Cone penetration test

The cone penetration test (CPT) has been used since the early 1930s, but originally with

a mechanical cone penetrometer. According to Broms & Flodin (1988), the first electric

cone penetrometer was probably developed during the Second World War. It was not

until the mid 1970s that electric piezometer probes were developed by Torstensson

(1975) in Sweden and by Wissa et al (1975) in USA, after Schmertmann (1974)

recognised the importance of pore pressure measurement for the interpretation of CPT

data. The first publication of the results from cone penetration tests with pore pressure

measurement (also referred to as piezocone tests or CPTU) was given by Roy et al

(1980). A detailed historical development of the CPT/CPTU has been given by Lunne et

al (1997b).

The cone penetrometer most widely used today is that with an apex angle of 60° and a

projected area of 1000 mm2 (diameter of 35.7 mm), as specified in the International

Reference Test Procedure (ISSMFE IRTP, 1999). The rate of penetration used in the

IRTP is 20 mm/s. The main advantage of the CPT/CPTU is that it provides very

consistent and continuous profiles of data of the soil with depth. With the cone

resistance and pore pressure being measured concurrently and continuously, the CPTU

is extremely useful in determining the soil stratigraphy and in detecting interbedded

layers (Baligh et al, 1981). Methods of interpreting CPTU data have been discussed

extensively, and correlations developed from the three measurements (cone resistance,

pore pressure and sleeve friction) to help identify the stratigraphic sequence (Robertson

& Campanella, 1983; Lunne et al, 1997b).

However, the CPTU does not give direct measurement of the undrained shear strength;

instead it measures the tip resistance exerted on the cone penetrometer when it is being

pushed into the soil. The measured cone resistance needs to be corrected for pore

pressure and overburden pressure effects, before being divided by a cone factor, Nkt, in

order to deduce the shear strength. In practice, errors in estimating the pore pressure and

LITERATURE REVIEW

2-6

overburden pressure effects, and the value of Nkt, may lead to large inaccuracies in the

consequential shear strength deduced, particularly for very soft sediments.

The pore pressure effect (also referred to as unequal area effect) arises owing to the pore

pressure acting on the shoulder area behind the cone. This effect was first identified in

deep water investigations, where the resistance readings before the cone penetrated into

the soil was found not to equal the water pressure (Lunne et al, 1997b). Therefore, the

measured cone resistance must be corrected for such effect using the following

relationship (Baligh et al, 1981; Campanella et al, 1982; Campanella, 1995):

)1(uqq 2ct α−+= (2.2)

where: qt = total cone resistance;

qc = measured cone resistance;

u2 = measured pore pressure immediately behind the cone tip;

α = unequal area ratio (i.e. ratio of ‘inner’ area to total area of the cone).

The importance of the correction for pore pressure effect was illustrated by Lunne et al

(1997b), based on results from the CPTUs performed in Bothkennar clay. Three

different piezocones with unequal area ratios ranging from 0.59 to very close to 1 were

used in the tests and the variations observed in the qc profiles disappeared once they

were corrected for the pore pressure effect to qt profiles.

Ideally, the unequal area ratio should be very close to unity, but in reality it varies with

each different cone design and ranges typically from 0.55 to 0.9 (Lunne et al, 1997b). In

very soft sediments, the pore pressure generated during penetration can be very large

relative to the cone resistance, hence the lower the unequal area ratio the greater

contribution of the pore pressure correction to qt and thus greater potential error in qt.

The necessity to correct for the pore pressure effect has led to development in the

industry to favour location for the pore pressure measurement immediately behind the

cone, at the shoulder position or most commonly referred to as the u2 position (Lunne et

al, 1997b). However, by comparing the pore pressure data measured at the u2 position

with that measured on the conical face of the cone head (u1 position; Lunne et al,

1997b), Randolph & Hope (2004) showed that the pore pressure response at the u2

position is much less sensitive than at the u1 position for stratification purpose. Another

location for the pore pressure measurement used in some piezocones is behind the

LITERATURE REVIEW

2-7

friction sleeve (u3 position; Lunne et al, 1997b), although this position is generally not

favoured due to the distance between the pressure transducer and the cone tip;

significant pore pressure will have dissipated when the transducer reaches the depth that

the cone tip has passed.

In clays, the undrained shear strength can be estimated from the net cone resistance,

qcnet, using the following expression (Lunne et al, 1997b):

kt

vot

kt

cnetu N

qNqs σ−

== (2.3)

where σvo is the in situ overburden pressure and Nkt the cone factor. Field test results

reported by Chung & Randolph (2004; also presented in Chapter 4) showed that the

correction for overburden pressure effects caused an average reduction in resistance of

31 % from qt to qcnet, or 19 % from qc to qcnet. The results imply that accurate estimation

of the overburden pressure is essential in order to give reliable assessment for the shear

strength.

Perhaps the most extensively discussed parameter in the interpretation of the CPTU

results, and consequently correlations developed, is the cone factor. Nevertheless, it is

worth mentioning that there are mainly three different cone factors used for correlations

with the shear strength. Besides the Nkt factor used in Equation 2.3, the other two cone

factors are denoted as Nke and NΔu (Lunne et al, 1997b). The Nke factor is defined as the

ratio of (qt – u2) to su, whilst the NΔu factor is the ratio of excess pore pressure (Δu) to su.

However, only the Nkt factor is considered in the thesis, since this factor allows

comparison with the bearing factors, N, calculated for other penetrometers of interest in

the study.

A number of theoretical solutions for the Nkt factor have been developed, based on

simple elastic-perfectly plastic idealisations of soil response (Teh & Houlsby, 1991; Yu,

2000; Lu et al, 2004). All theoretical solutions have shown that, even for a simple

Tresca soil model, the theoretical Nkt is affected by the rigidity index, Ir = G/su, where G

is the shear modulus, by the in situ stress ratio, Δ = (σvo – σho)/2su, where σvo and σho

are the in situ vertical and horizontal stresses, and by the roughness coefficient for the

cone-soil interface, αc.

Using large strain finite element analysis, Lu et al (2004) derived the following

approximate expression:

LITERATURE REVIEW

2-8

( ) crkt 3.19.1Iln6.14.3N α+Δ−+≈ (2.4)

which has a similar form to that obtained by Teh & Houlsby (1991), based on the strain

path method incorporated with small strain finite element analysis. Lu et al (2004) also

compared the ranges of Nkt from fully smooth to fully rough cone-soil interface (αc = 0

to 1), obtained from different theoretical solutions (Yu, 2000; Teh & Houlsby, 1991;

and Van den Berg, 1994) and found that the Nkt computed using Equation 2.4 has the

narrowest range, between 9.6 (fully smooth) and 14.5 (fully rough) for 50 ≤ Ir ≤ 500 and

Δ = 0. Nevertheless, for a typical roughness coefficient of around 0.3, Equation 2.4

gives Nkt varying from 10.2 to 13.9 for 100 ≤ Ir ≤ 300 (typical Ir range) and

−0.5 ≤ Δ ≤ 0.5.

However, these theoretical solutions have not taken into account the effects of strain

rate, strain softening and strength anisotropy of the natural soils. The consequence is

that empirical values of Nkt calibrated using strength data measured from laboratory

tests and the vane shear tests are often found to vary much more widely than the

theoretical values.

Aas et al (1986) proposed that the laboratory strengths computed from average values of

triaxial (compression and extension) and simple shear strengths be used as the reference

su to calibrate Nkt. Their results gave Nkt ranging from 11 to 20. The geometry of the

cone penetrometer has led to the general belief that the soil response around the cone tip

during penetration most resembles triaxial compression conditions (Baligh, 1986;

Lunne, 2001). Therefore, the triaxial compression strength is often favoured and used as

the reference su to calibrate Nkt.

Based on the triaxial compression strengths, Powell & Quarterman (1988) obtained Nkt

varying from 10 to 20, depending on the plasticity index (Ip). A similar range of Nkt, but

with slightly lower values, from 8 to 16, was also noted in the results from Aas et al

(1986), quoted by Lunne et al (1997b). However, a relatively wider range of Nkt,

between 8 and 29 was presented by Rad & Lunne (1988).

Interestingly, Tanaka & Tanaka (2004) showed that the strength profiles deduced from

the CPTU data using Nkt of 8 to 10 fit reasonably well to the UCS strength data for

normally consolidated to lightly overconsolidated clays. For moderately to heavily

overconsolidated clays (OCR from 2 to 5.8), the Nkt varied from 12 to 20 in order to fit

the laboratory data (Tanaka & Tanaka, 2004).

LITERATURE REVIEW

2-9

Having seen the wide range of Nkt, even in a single site, it is not surprising that the

practice of calibrating Nkt for each new site using strength data obtained from the

laboratory and vane shear tests is still undertaken by geotechnical engineers

(particularly offshore) even though extensive empirical correlations and theoretical

solutions have been developed.

Due to strength anisotropy, it is essential to state the testing methods used for obtaining

the reference su in the empirical correlations for the Nkt factors (Aas et al, 1986; Lunne

et al, 1997b; Tanaka & Tanaka, 2004). Furthermore, since the effects of sample

disturbance can severely distort the results and subsequently affect interpretation of Nkt,

the utmost importance of ensuring only high quality shear strength data be used in the

correlations is strongly emphasised (Randolph, 2004; Sandven & Black, 2004).

2.3 ‘Full-flow’ penetrometer tests

A ‘full-flow’ penetrometer refers to a penetrometer for which soil is able to flow

around, instead of being displaced by the penetrometer during penetration. The

rationales behind introduction and development of the full-flow penetrometers are that:

• Since soil is able to flow around the full-flow penetrometers, the overburden

pressure is equilibrated above and below the penetrometers, except at the shaft.

However, since the projected area of the full-flow penetrometers is considerably

larger than the area of the shaft, the corrections to the measured resistance for pore

pressure and overburden pressure effects to obtain net resistance will be minimal

and can be neglected.

• Improved resolution is obtained by measuring higher penetration force due to the

larger size of the full-flow penetrometers compared to the cone. This also reduces

the sensitivity to any load cell drift. Consequently, better accuracy is achieved in

the measurement of resistance.

• Accurate plasticity solutions exist, within certain idealisations, in the form of a

bearing factor, N, defined as the ratio of net penetration resistance to undrained

shear strength. Therefore, in principle, the full-flow penetrometers do not need

calibration against laboratory strength data for each new site.

• Tip resistance is also measured during extraction for the full-flow penetrometers.

This provides valuable additional data for the remoulded resistance and hence

sensitivity of the soil, which the traditional CPTU does not provide.

LITERATURE REVIEW

2-10

At present, three types of full-flow penetrometers have been developed, namely the

T-bar, ball and circular plate. These alternative devices can be deployed simply with the

existing CPTU equipment, and are generally performed at the same penetration rate of

20 mm/s as the CPTU (Stewart & Randolph, 1994).

The first full-flow penetrometer developed was the T-bar penetrometer, which consists

of a cylindrical bar mounted at right angles to the push-rods. It was originally developed

as a laboratory device to give improved definition of the shear strength profile, but later

scaled up for field use (Stewart & Randolph, 1991, 1994). According to Randolph

(2004), the first offshore T-bar penetrometer tests were conducted in Australian waters

in late 1997 with a purpose built device incorporating its own load cell and also two

pore pressure transducers within the cylindrical bar (Randolph et al, 1998). This device

was used 9 months later, again in Australian waters (Hefer & Neubecker, 1999).

The bearing factor for the T-bar penetrometer, NT-bar, was based originally on the

plasticity solution of Randolph & Houlsby (1984). This gave a bearing factor that

depended only on the surface roughness of the cylinder, varying between 9.1 (fully

smooth) and 11.9 (fully rough). Although an error was later found in the upper bound

calculation of Randolph & Houlsby, which after correction gave a bearing factor about

9 % higher than the lower bound solution for fully smooth cylinder (Randolph et al,

2000), an improved upper bound of 9.2 was subsequently obtained (Einav & Randolph,

2005). A value of 10.5 (mid value of the range, which was recommended by Randolph

& Houlsby (1984) for general use) has been used extensively to compute the undrained

shear strength (su) profile and the results showed generally good agreements with su

profiles obtained using other testing methods (Stewart & Randolph, 1991, 1994;

Randolph et al, 1998; Chung & Randolph, 2004; Oung et al, 2004).

The axisymmetric ball and circular plate penetrometers were subsequently introduced,

in order to reduce the potential for the load cell to be subjected to bending moments

arising from non-symmetric resistance along the T-bar. However, the plate penetrometer

tends to carry soil with it both during penetration and extraction, particularly in highly

non-homogeneous soils where softer soil may be trapped beneath the plate, or be carried

down covering the whole top side of the plate, thus affecting the resistance

measurements. This is supported by the numerical study presented by Lu et al (2001).

Therefore, the plate is often considered less attractive than the ball geometry for soil

LITERATURE REVIEW

2-11

profiling purposes.

The theoretical bearing factor for the ball penetrometer, Nball, was derived based on the

upper and lower bound approaches similar to the T-bar, with the flow mechanism

adapted to axisymmetric flow for the ball (Randolph et al, 2000; Randolph, 2004). The

resulting upper bound solution gave Nball factors of 11.6 for a fully smooth ball and 15.3

for a fully rough ball, while the lower bound solution yielded corresponding Nball factors

ranging from 11.0 to 15.1. These values are 19 to 29 % higher than the corresponding

theoretical NT-bar factors; these differences are, however, contrary to the experimental

findings, where the T-bar and ball penetrometers demonstrated essentially identical

resistances (hence bearing factors) in materials ranging from reconstituted clays to

calcareous silts (Watson et al, 1998). Furthermore, in situ penetration tests undertaken

in highly layered (varved) clay showed that the ball penetration resistance was 28 %

lower than the T-bar resistance (DeJong et al, 2004). The discrepancy between the

experimental and theoretical bearing factors is intriguing and is believed to be caused by

factors such as strength anisotropy, strain rate, strain softening, sensitivity etc, which are

not accounted for in the theoretical solutions. These factors may be enhancing or

compensating each other and the extent of each factor varies depending on the clay

types (NGI-COFS, 2004c). Therefore, the overall effects are difficult to gauge.

However, it should be noted that FE analyses presented by Lu et al (2000) show that the

soil rigidity index does not have any effect on the bearing factors or limiting resistances

for both the T-bar and ball penetrometers, but it does change the displacement needed to

reach the limiting resistances. The higher the rigidity, the smaller is the displacement to

reach the limiting resistances.

Randolph (2000) integrated the effects of strength anisotropy into the upper bound

analyses for the T-bar and ball, and verified that the ball is relatively more sensitive to

the strength anisotropy ratio than the T-bar. Numerical analysis presented in the report

of NGI-COFS (2004c) also showed that the NT-bar value for anisotropic shear strength is

close to that for isotropic strength, provided that the average strength is close to the

simple shear strength and there is no strain softening. According to Randolph’s analysis,

for typical strength anisotropy ratio (extension to compression strengths) of 0.5 to 0.7,

Nball would be about 7 to 10 % higher than NT-bar – hence the discrepancy has been

greatly reduced from the original 19 to 29 % for isotropic strength.

Based on the combined upper bound and strain path method, Einav & Randolph (2005)

LITERATURE REVIEW

2-12

incorporated the characteristics of strain softening and strain rate into their theoretical

soil models. A degradation parameter, ξ95, and rate parameter, λ, are used to quantify

the strain rate and strain softening effects. The consequent analyses illustrate that the

compensating effects of the strain softening and strain rate are stronger for the T-bar

than for the ball. For typical rate parameters in the range 0.1 to 0.2, NT-bar varied from

10.2 to 12.5 for rapidly softening soil, and 11.7 to 14.3 for very gradually softening soil

(Einav & Randolph, 2005; Randolph, 2004).

Besides the experimental and theoretical studies on the bearing factors for deducing

undrained shear strength, the potential for full-flow penetrometers to provide additional

information such as sensitivity and consolidation characteristics of the soil has also

started to be explored and the results published in the literature.

Since soil is able to flow around the full-flow penetrometers, measurement of the

resistance is also recorded during extraction of the penetrometer. This gives a measure

of the disturbance caused by the initial penetration. Fully remoulded resistance and

sensitivity of the soil can be determined by performing several cycles of penetration and

extraction tests over a short interval of depth (Chung & Randolph, 2004; DeJong et al,

2004; Long & Gudjonsson, 2004). Furthermore, such cyclic tests can provide valuable

insight to the strain softening characteristics of the soil and the data used to calibrate

theoretical models incorporating strain softening effects (Einav & Randolph, 2005;

Randolph, 2004).

2.4 Variable rate penetration test

The variable rate penetration test (also referred to as ‘twitch’ test) was proposed by

Randolph & House (2001) as an alternative means to assess the consolidation

coefficient, cv, of the soil. It involves pushing a penetrometer into the soil with the

penetration rate being successively halved, and the penetrometer advanced by 1 to 2

diameters of the probe at each stage. As discussed by Randolph & House (2001),

discrete steps for the penetration rate reduction are recommended, rather than a

continuous and smooth decrease in penetration rate, so as to overcome the effects of

creep and to allow the penetration resistance to fully develop and achieve its steady

state.

The idea of a twitch test originated from the fact that, penetration resistance increases as

the rate of penetration is increased when the conditions around the advancing

LITERATURE REVIEW

2-13

penetrometer are undrained, due to viscous effects. On the other hand, at low

penetration rates where the conditions become partially drained, the tip resistance

increases as the penetration rate is reduced due to partial consolidation effects and local

strengthening of the soil around the penetrometer. Therefore, there is a transition point

from undrained to partially drained response where the viscous and partial consolidation

effects balance out, leading to a minimum resistance. This phenomenon is clearly

illustrated in the results presented by Bemben & Myers (1974), who carried out in situ

(mechanical) cone penetration tests with penetration rates ranging from 0.2 to 200 mm/s

in varved clay, and by Roy et al (1982), who performed CPTU with rates varying from

0.5 to 40 mm/s in sensitive clay. A brief summary of the literature published on effects

of penetration rate on cone resistance has been given by Lunne et al (1997b).

A simple non-dimensional analysis shows that the drainage conditions depend not only

on the penetration rate, v, but also the diameter of the advancing probe, d, and the

consolidation coefficient, cv, of the soil. This led to the penetration rate being

normalised to a non-dimensional velocity (Randolph & House, 2001; Finnie &

Randolph, 1994):

vc

vdV = (2.5)

Based on the results from constant rate of penetration tests on shallow circular

foundations, Finnie & Randolph (1994) suggested transition points of V < 0.01 for

drained response and V > 30 for undrained response.

In order to verify the transition point from undrained to partially drained response, a

series of constant rate of T-bar penetration tests at various penetration rates covering 3

orders of magnitude, were conducted in the centrifuge independently by Watson &

Suemasa (2000, unpublished), House et al (2001) and Randolph & Hope (2004), who

also carried out similar tests for the cone penetrometer. The results obtained from these

penetration tests have been presented in the form of normalised resistance against non-

dimensional velocity, V. House et al (2001) suggested a curve of the form below to fit

the normalised data in the partially drained region:

mref cV1

baqq

++= (2.6)

where: q = penetration resistance at any rate;

LITERATURE REVIEW

2-14

qref = reference (undrained) resistance;

a, b, c, m = constants for the ‘backbone’ curve.

Note that m is used in Equation 2.6 rather than d originally used by House et al (2001),

in order to avoid confusion with the diameter of the penetrometer. The values of cv

required for computing V were obtained from Rowe cell tests. A summary of the

constants a, b, c and m published in the literature is given in Table 2.1. It should be

pointed out that the constants first presented by Randolph & Hope (2004) have been

affected by a correction factor for viscous effects applied to the right hand side of

Equation 2.6. However, this correction factor was later found inappropriate. Therefore,

a different set of constants has been derived to fit the T-bar and cone test data reported

by Randolph & Hope (2004), the values of which are also given in Table 2.1. The

‘backbone’ curves will be presented later in Chapter 8, although highlights of the curves

are briefly provided as follows.

The backbone curve obtained by Watson & Suemasa (2000) shows that the transition

point from undrained to partially drained conditions for the T-bar corresponds to V ~ 20

and that the T-bar resistance doubles its reference undrained value by V ~ 2. However,

the curves presented by House et al (2001) and by (after) Randolph & Hope (2004)

demonstrate a lower transition point of V ~ 10, but the curve from the latter shows

normalised resistance increasing more rapidly and doubling within one order of

magnitude by V ~ 1, compared to that from the former, which shows normalised

resistance doubling by V ~ 0.5. Randolph (2004) later commented that the T-bar tests

performed in the drum centrifuge by House et al (2001) may have been affected by

small vibrations transmitted from the central turntable of the drum centrifuge. These

could have led to additional excess pore pressures generated at the T-bar and delayed

the effects of consolidation, hence giving a lower rate of increase in resistance as the

penetration rate was reduced.

The backbone curve for the cone derived after Randolph & Hope (2004) shows a

gradual transition from undrained to partially drained for V < 30, with the normalised

cone resistance doubling by V ~ 1.

Although more data are required to confirm the backbone curves for the cone and T-bar

penetrometers, once the curves are established, they can be utilised either to assess

whether a given penetration test in, say, a silt is partially drained or not, or to deduce a

value for the cv of a particular soil.

LITERATURE REVIEW

2-15

The value of cv can be deduced by first normalising the (developed) penetration

resistances from each stage of the twitch test by the reference (undrained) values from

the normal penetration test. Values of V are then required. Since the penetration rates

and diameter of the penetrometer are known, the only variable (or unknown) is cv. It is

adjusted to fit the normalised twitch test data onto the backbone curve and the value that

gives the best-fit result corresponds to cv of the soil.

In an attempt to integrate viscous effects for the undrained zone (V > 10), Randolph &

Hope (2004) introduced a hyperbolic function multiplied to the right hand side of

Equation 2.6. The resulting expression is:

[ ]⎭⎬⎫

⎩⎨⎧

−⋅λ

+⋅⎟⎠⎞

⎜⎝⎛

++= −− )V/V(sinh)V/V(sinh

)10(n1

cV1ba

qq

oref1

o1

mref l

(2.7)

where: λ = rate parameter (varies between 0.1 and 0.2);

Vo = value of V for which the viscous effects start to decay;

Vref = reference V where the hyperbolic function term passes through unity.

The values of Vo and Vref were taken as 10 and 100 respectively in the interpretation by

Randolph & Hope (2004). However, as already mentioned, the correction factor for

viscous effects shown in Equation 2.7 was found inappropriate. A similar, yet different

correction factor has been attempted in this research. This is presented in Chapter 8.

2.5 Summary

This chapter has presented developments of the in situ testing methods and the

challenges faced in interpreting the data of such tests.

The primary advantage of the vane shear test is that it gives direct in situ measurements

of the undrained shear strength in peak, post-peak and remoulded modes, and hence

measurement for the soil sensitivity. Nevertheless, the measurements can be severely

influenced by many factors such as disturbance from the vane insertion, waiting time

before rotating the vane, rate of rotation etc, not to mention that it can only give shear

strength data at discrete intervals of depth.

On the other hand, the CPTU gives very consistent and continuous profiles of data of

the soil with depth. With the pore pressure and cone resistance measured concurrently,

the CPTU has been proven to be extremely useful in determining the soil stratigraphy

and in detecting interbedded layers. However, before interpreting the CPTU data, the

LITERATURE REVIEW

2-16

measured cone resistance needs to be corrected to net values for pore pressure and

overburden pressure effects. Such corrections may constitute a major fraction of the net

cone resistance, particularly in soft sediments, and thus imparting potential inaccuracies

in the subsequent interpretation. Additionally, the cone factor, Nkt, defined as the ratio

of net cone resistance to undrained shear strength, has been shown to vary widely from

one site to another, based on the experimental results published in the literature. This is

due to the dependency of the Nkt factor on soil parameters such as soil stiffness (or

rigidity index), in situ stress ratio, strength anisotropy, strain rate and strain softening.

Alternative, full-flow penetrometers (T-bar, ball and plate) have been introduced in

order to avoid the uncertainties involved with interpretation of the CPTU data in clays.

Upper and lower bound plasticity solutions gave bearing factors varying within a

relatively narrow range for each full-flow penetrometer. These, in principle, avoid the

need for calibration against laboratory strength data at each new site. A value of 10.5 for

the T-bar factor has been used extensively to compute the average strength profile from

the T-bar resistance, and the results indicated generally good agreements with the

strength data obtained from other testing methods. Nevertheless, the contradictory

findings between the theoretical and experimental bearing factors for the T-bar and ball

penetrometers emphasise the importance of strength anisotropy, strain rate and strain

softening effects, which should be accounted for in the analysis.

Another major advantage of the full-flow penetrometer tests over the CPTU is that tip

resistance is also measured during extraction of the penetrometer. This gives a measure

of the disturbance caused by the initial penetration. Fully remoulded resistance and

sensitivity of the soil can be determined by performing several cycles of penetration and

extraction tests over a short interval of depth. The results can also be used to calibrate

theoretical models that incorporate strain softening effects.

A variable rate penetration test (‘twitch’ test) was proposed as an alternative means to

assess the consolidation coefficient, cv, of the soil. It involves pushing a penetrometer

into the soil with the penetration rate being successively halved, and the penetrometer

advanced by 1 to 2 diameters of the probe at each stage. Interpretation of cv from the

twitch test data requires a ‘backbone’ curve to be established, in the form of normalised

penetration resistance against non-dimensional velocity, V = vd/cv, where v is the

penetration rate, d is the penetrometer’s diameter and cv the consolidation coefficient.

BURSWOOD TEST SITE

3-1

3 BURSWOOD TEST SITE

In the previous chapter, an overview was given of the developments in in situ testing

methods: the vane shear test, the cone penetration test and ‘full-flow’ (T-bar, ball and

plate) penetrometer tests, and the challenges associated with them. In order to

investigate and compare these testing methods, in situ tests were carried out at a local

site and laboratory tests performed on ‘undisturbed’ samples retrieved from the same

site.

In this chapter, a short background is given on the site used in this research (Burswood

Peninsula, which is referred to as the Burswood site), along with a description of the

soil properties at the site.

3.1 Site background

The Burswood site is situated on an inside meander of the Swan River, some few

kilometres upstream from the centre of Perth, Western Australia (see Figure 0). The soil

sediments at this site consist of estuarine alluvium deposited in recent geological time.

Churchill (1959) showed that, in 8000 BC, the sea level was 21 m lower than at the

present. At that time, the level would have been close to the base of the soft clay now

present at the Burswood site. Based upon the hypothesis, it was estimated that the soft

clay deposit would be less than 10,000 years old (Cray, 1988; Lee Goh, 1994).

The test region for this research is essentially level, with RL (Reduced Level relative to

datum sea level in Perth) of +0.95 m. The water table is within the top 1 to 2 m,

although previous fluctuations have probably occurred, so that the clay deposit is lightly

overconsolidated in the upper region. The stratigraphy of the site comprises a weathered

crust about 3 m thick, underlain by a layer of soft silty clay about 15 m thick,

underneath which is a layer of stiff, fine sand. Above a depth of 12 m, the soil contains

frequent shell fragments and silt lenses; below this depth, tiny shell fragments also exist

occasionally. Desiccated weeds and plants are generally found at shallow depths above

7 m.

Due to recent construction of the Graham Farmer Freeway, approximately 90 m south

of the test area for this research (see Figure 3.1), extensive site investigations have been

undertaken at the vicinity of the freeway by the state road authority, Main Roads

Western Australia (MRWA). In addition, Lee Goh (1994) also carried out some field

BURSWOOD TEST SITE

3-2

testing and sampling at the same vicinity for his PhD research before construction of the

freeway, but the tests were mainly concentrated at depths between 4 and 7.5 m.

3.2 Soil properties

The clay layer between depths of 4 and 17 m is the main interest for this research. Most

values of soil properties presented here were estimated from laboratory tests performed

for this research, testing details for which are presented in Chapter 5. Additionally, data

obtained from other sources are also presented for the purpose of soil characterisation.

The Burswood clay is lightly overconsolidated with overconsolidation ratio (OCR) of

about 2 at a depth of 4 m, decreasing gradually to about 1.65 at depths below 10 m. The

saturated unit weight of the clay is between 14 and 14.5 kN/m3 at depths above 6 m,

increasing to 16 kN/m3 at 13 to 14 m depth before reverting to around 14.5 kN/m3

below that depth.

Soil particle densities quoted from Lee Goh (1994) range from 2.60 to 2.64 Mg/m3 for

depth range 4.5 to 7 m. Particle size, based on grading test data at 6 m depth (presented

in Chapter 5), lies mainly in the categories of fine and medium silt fractions. The clay

content (particles finer than 2 μm) deduced from the grading curve is about 10 %.

However, Lee Goh (1994) found that the clay content increased with depth, from 17 to

51 % for depths 4.5 to 7 m. These results suggest variation of stratigraphy of the

Burswood site, both horizontally and vertically.

The mineral composition of the clay was determined using X-ray diffraction analysis.

Two samples collected from different locations were analysed and the results showed

high content of quartz (27 %) and sillimanite (44 %) in the first sample; the second

sample comprised mainly quartz (33%), albite (26 %) and kaolinite (26 %). Other

minerals such as calcite, gypsum, hematite etc were also present in small portions in

either one or both of the samples. Although the exact locations of the samples used in

the X-ray diffraction analyses were not specified, these results indicate that the mineral

composition of the Burswood clay may vary rather significantly.

The in situ water content is relatively high in the shallow depth region above 7 m, with

values ranging from 80 to 100 %. Below this depth, the values are mostly between 50

and 70 %. The liquid limits are mainly between 60 and 80 %, whereas the plastic limits

are more constant, with an average of 30 %. However, Lee Goh (1994) reported

generally much higher water content (99 and 110 %) and Atterberg limits at depth range

BURSWOOD TEST SITE

3-3

5 to 7 m, with liquid limits in the range 105 to 126 % and plasticity indices between 72

and 95 %; only one depth (at 4.5 m) recorded lower water content of 65 % and

Atterberg limits of 46 % (liquid limit) and 28 % (plasticity index). The discrepancies

between results from Lee Goh and those measured in this research may be attributed to

stratigraphic changes at different test regions and variation of the clay content within the

material taken from the sample for testing. Due to the presence of silt and other

impurities within the clay, it was extremely difficult to obtain small portion of materials

that are ‘absolutely’ representative.

Sensitivities measured by the field vane tests are mostly between 4 and 9 at shallow

depths and reduce to between 2 and 4 below a depth of 7 m. These are consistent with

the higher water content for depths above 7 m than that below this depth.

The compression index, Cc, lies between 1 and 1.5 for depths above 8 m, and between

0.5 and 1 below this depth. The in situ void ratio, eo, is found to be approximately 2.6

above 6 m, below which its values range from 1.2 to 1.7.

The consolidation coefficient determined at preconsolidation or yield stress, cvy, is

between 1 and 1.25 m2/year for depth range 6 to 12 m. Outside this depth range (both

above and below), the cvy values measured from the oedometer tests are somewhat

higher, particularly at a depth of 18.4 m, where a cvy of 4 m2/year was measured. The

high values of cvy in some tests may be caused by the presence of silt lenses within the

test samples. The cvy values reported by Lee Goh (1994) ranged between 0.33 and

0.89 m2/year over depths of 5 to 7 m. This reported cvy range was more similar to the

values of 0.5 to 1 m2/year obtained at high stress levels in the oedometer tests of this

research.

The in situ permeability, ko, deduced from oedometer tests is very low, in the order of

10-10 m/s.

The (peak) friction angle of the clay measured from the triaxial compression tests

(φ'TXC) was found to be very high, approximately 42°. The extension tests gave angles

(φ'TXE) from 22 to 31°. These give ratios φ'TXE/φ'TXC in the range 0.52 to 0.74, which are

rather low compared with the ratios reported by Kulhawy & Mayne (1990), which

ranged typically from 1.1 to 1.3. However, recent extension tests performed on the

Burswood clay in an industry project consistently gave friction angles as high as 43°

(Ismail, 2004: private communication). The simple shear tests gave friction angles of

BURSWOOD TEST SITE

3-4

around 32 to 42°.

The undrained shear strength (su) profile measured from the field vane tests gradually

increases from around 18 kPa at a depth of 6 m to 28 kPa at 13 m. However, laboratory

tests (details provided later in Chapter 5) gave an average strength of about 15 kPa at

6 m depth, increasing to 38 kPa at 17 m depth. The strength anisotropy ratio for triaxial

extension to compression strengths is approximately 0.61 for depth range 4 to 17 m,

while the ratio for simple shear to triaxial compression strengths is about 0.72 for the

same depth range.

The small strain shear modulus (Go) profile deduced from seismic cone test data is

approximately 2 MPa at depth of 3 m and increases to around 13 to 15 MPa at 17 m

(Schneider et al, 2004). The ratio, Go/su, computed based on field vane strengths ranges

from 230 to 420, whilst using average laboratory strengths, Go/su lies mostly between

300 and 400.

FIELD TESTING

4-1

4 FIELD TESTING

This chapter describes the experimental procedures and results of the in situ testing

carried out at the Burswood site, for which conditions and an overview of geotechnical

parameters have been described in Chapter 3.

The in situ tests were concentrated in an area of 20 m × 15 m, located about 40 m from

the waterfront of the Swan River, except for one cone penetration test that was

performed outside this area. The layout of the in situ tests is shown in Figure 4.1. The

testing programme comprised penetration tests with different types of penetrometers,

including the cone, T-bar, ball and plate, and vane shear tests. For the penetration tests,

in addition to monotonic tests, cyclic penetration and extraction tests were also

performed at specific depths, except for the cone penetrometer.

The following sections describe the equipment details, calibration and testing procedure,

followed by the results. When discussing the results, the focus is first on comparing the

tip resistances for the various types of penetrometer during penetration and extraction.

Then, using a provisional bearing factor (conventional values reported in the literature,

see Chapter 2), the undrained shear strength profiles are derived and compared with the

measurements from the vane shear tests. The discussion then turns to the cyclic test

results, with particular interest placed on the rate of degradation of the resistance and its

implications.

4.1 Field testing apparatus

4.1.1 Field penetrometers

Penetrometers tested in the field included cone, T-bar, ball and plate. A Hogentogler

cone, details of which are shown in Figure 4.2, was adopted. It has an apex angle of 60°

and a diameter of 35.7 mm, giving a projected area of 10 cm2 (ISSMFE IRTP ,1999).

The inner diameter of the cone is 25.1 mm and hence gives an inner area (AN) of about

4.95 cm2. Pore water pressure is measured at the shoulder of the cone behind the tip (u2

position; Lunne et al, 1997b). The cone has been calibrated by direct loading and also

by applying water pressure, in order to assess the effective area ratio, as is detailed in a

later section.

The T-bar penetrometer was manufactured by welding a cylindrical bar to the tip of an

identical cone described above. Hence, the T-bar has the same inner area (AN) and uses

FIELD TESTING

4-2

the same pore pressure measurement arrangement as the cone. Two different sizes of

T-bar penetrometer were tested, both of 40 mm nominal diameter. The first T-bar was

250 mm long (aspect ratio, L/d = 6.25), giving a projected area of 100 cm2 (Randolph et

al, 1998), while the second T-bar was 160 mm long (L/d = 4), giving projected area of

64 cm2. The longer T-bar was tested with both smooth and lightly sand-blasted

conditions for the cylindrical surface, whereas the shorter (smaller) T-bar was tested

only with the sand-blasted condition. A schematic diagram of the 250 mm × 40 mm

T-bar is shown in Figure 4.3. Note that the actual diameter measured for the standard

T-bar is 38.9 mm (as indicated in the figure), while that measured for the smaller T-bar

is 39.9 mm. The measured diameters have been adopted in calculating the T-bar tip

resistance.

The plate penetrometer consists of a thin circular plate with diameter of 113 mm (hence

projected area of 100 cm2) and thickness of 6 mm attached to the cone head with its tip

cut off.

The ball penetrometer was manufactured in one piece in the UWA workshop. The

sphere also has a diameter of 113 mm, and the connection head was ensured to be the

same as the cone head in order to have the same unequal area ratio and same pore

pressure measurement conditions.

Both the plate and ball penetrometers were only tested with the lightly sand blasted

surface condition. Figure 4.4 shows a photograph of plate and ball, together with other

penetrometers.

4.1.2 Shear vane

The shear vane used in the field was manufactured in the Civil Workshop of the

University of Western Australia. Both the vane and rods are unprotected during

insertion into the soil, with a slip coupling incorporated immediately above the vane.

The vane has a length of 176.3 mm and width (diameter) of 60 mm (Figure 4.5). The

thickness of the blades is 2.1 mm, giving a perimeter ratio (see Equation 2.1) and area

ratio (Chandler, 1988) of around 4.5 and 8.9 % respectively. This perimeter ratio is

marginally higher than that for the ‘standard’ vane referred to by Chandler (3.9 %), but

the area ratio complies well with the generally accepted standard of < 12 % (Chandler,

1988; Geise et al, 1988). Such ratios ensure disturbance induced during insertion of the

vane is kept to an acceptable level. The vane was also designed deliberately with cut

FIELD TESTING

4-3

away corners to help further reduce such disturbance.

Assuming isotropic shear strength distributed uniformly along all edges of the vane

shown in Figure 4.5, the torque, Tq, required to shear the soil is:

∫ δ⋅⋅= ArsT uq (4.1)

where: su = undrained shear strength along shear surface;

r = radius from centre of vane to shear surface;

δA = increment of shear surface area.

From Equation 4.1, it can be shown that a calibration factor of 1.054 is required to

convert the torque measured in Nm to the undrained shear strength of the soil in kPa.

However, a factor of 1 kPa/Nm was adopted for the interpretation in the thesis. This

may be justified by the potential uncertainties arising from the internal friction of the

apparatus itself. In fact, relatively high rod friction was observed in some measurements

of the vane tests. These friction errors are difficult to eliminate completely even though

a slip coupling has been used (Åhnberg et al, 2004). The consequence of this is that the

torque measured is higher than its ‘true’ value, thus overstating the value of shear

strength. Although, it is common to apply a reduction factor to the strength data

measured from the vane tests performed onshore for high plasticity clays (Aas et al,

1986; Bjerrum, 1973), no correction has been applied to the data reported here.

4.1.3 Calibration details

The load cell and pore pressure transducer used in the field penetration tests were

calibrated in the laboratory prior to use. The load cell was calibrated against a reference

standard load cell. It was loaded axially at appropriate load intervals, and both the load

cell readings and the reference readings were recorded. Readings were also recorded

while unloading.

The pore pressure transducer was calibrated in a calibration chamber filled with water.

It was essential to ensure the chamber was properly sealed, with no water leakage even

at high water pressure. Water pressure was applied to the chamber, with the applied

pressure read from an external Bourdon gauge. Readings of the pore pressure transducer

were recorded, as well as the readings of the load cell in order to assess the unequal area

ratio of the cone.

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Calibration factors for the load cell and pore pressure transducer, and the unequal area

ratio calibrated are summarised in Table 4.1. It may be seen that the calibrated unequal

area ratio of 0.699 is significantly larger than the intrinsic value of 0.495 (AN divided by

the total area, AT).

4.2 Field testing procedure

4.2.1 Field penetration tests

The penetrometers were pushed into the soil hydraulically from a truck (Figure 4.6).

Before commencing each test, the penetrometer was saturated with machine oil and

connected to the measuring devices. A hole was dug to a depth of about 0.6 m at the

location where the test was to be performed. Then the truck was positioned to locate the

penetrometer above the hole, and the truck was levelled to ensure verticality of the

penetrometer. After the depth reading was zeroed at the ground surface, the

penetrometer was lowered and submersed completely in the hole filled with water for

approximately 5 minutes, in order to achieve temperature equilibrium for the electronic

devices. The site is essentially level, with RL (Reduced Level relative to datum sea level

in Perth) of +0.95 m, and the depths plotted are direct values measured from the ground

surface. The penetrometer was penetrated to a depth of at least 18 m below the ground

surface before being extracted. In compliance with ISSMFE IRTP (1999), the rate of

penetration and extraction was 20 mm/s, with a push length of 1 m before additional rod

was added (for penetration) or removed. The data were logged at 50 mm depth intervals.

Cyclic penetration and extraction tests were carried out at specific depths for each

penetrometer, except for the cone. These tests comprised displacement cycles of ±0.5 m

about the mean depth, recording the penetration and extraction resistance over five full

cycles. The cyclic test was generally carried out during the initial penetration, except for

T-bar Tests 1 and 3, where it was conducted during extraction.

4.2.2 Vane shear tests

The shear vane was pushed into the soil manually using the frame shown in Figure 4.7.

The frame was anchored to provide the reaction forces required while pushing the vane

and while performing the vane shear test. Verticality of the vane was checked before

pushing it into the soil. Nevertheless, there was no means to prevent the vane from

deflecting laterally after it went into the soil. The vane was pushed into the soil at a rate

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of approximately 30 mm/s. Once the vane had reached the required depth, there was a

waiting period of one minute before performing the test.

The vane was turned using a torque wrench at an approximate rate of 5 revolutions per

minute (30°/s), and both the peak and post-peak torque readings were recorded. It

should be noted that this rate is much faster than the standard rotation rate of 0.1°/s

(Chandler, 1988). Assuming a variation of strength measurement of 10 % per log cycle

of rotation rate (Roy & Leblanc, 1988; Leroueil & Marques, 1996), the vane test

performed at such high rate of rotation could potentially give strength measurement

15 % higher than the ‘standard’ vane test.

After the initial peak and post-peak shear strengths had been measured, the vane was

turned rapidly for 10 rotations (ASTM, 2000a; Geise et al, 1988), and the same test

procedure was performed to measure the fully remoulded strength of the soil.

4.3 Field test results

The field testing conducted for the research included:

• four cone penetration tests;

• two smooth and two lightly sand blasted T-bar penetration tests using the 250 mm

× 40 mm bar;

• two lightly sand blasted T-bar penetration tests using the 160 mm × 40 mm bar;

• two lightly sand blasted plate penetration tests;

• two lightly sand blasted ball penetration tests;

• two vane shear tests.

4.3.1 Assessment of penetrometer tip resistance

As has been mentioned in Chapter 2, the measured cone resistance, qc, will be

influenced by the unequal pore pressure and the overburden pressure effects, hence it

should be corrected appropriately before proceeding to interpretation of the data from

the cone tests. The corrections are performed by first adjusting the measured cone

resistance for the unequal pore pressure effect, to total tip resistance, qt, using the

following relationship (Baligh et al, 1981; Campanella et al, 1982; Campanella, 1995):

)1(uqq 2ct α−+= (4.2)

where: u2 = measured pore pressure immediately behind the cone tip (shoulder

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4-6

position);

α = unequal area ratio (in this case 0.699).

The net tip resistance is then calculated as below:

votcnet qq σ−= (4.3)

where: σvo = total in situ overburden stress.

Similarly, the tip resistances of the T-bar and other ‘full-flow’ penetrometers (ball and

plate) are corrected for the unequal pore pressure and overburden pressure effects. Since

the soil is able to flow around the full-flow penetrometer, the pore pressure and

overburden pressure are equilibrated above and below the penetrometer, except at the

shaft. By force equilibrium, the following equation can be obtained:

sovomnet A)]1(u[QQ ⋅α−−σ−= (4.4)

where: Qnet = force exerted on the penetrometer due to soil resistance;

Qm = force applied (or measured by load cell);

uo = estimated hydrostatic water pressure;

As = cross-sectional area of connection shaft in plane normal to shaft;

Dividing Equation 4.4 by the projected area of the penetrometer, the equation becomes:

[ ] psovomnet A/A)1(uqq ⋅α−−σ−= (4.5)

where: qnet = net tip resistance of the penetrometer;

qm = measured tip resistance;

Ap = projected area of the penetrometer in plane normal to shaft.

Therefore, the measured tip resistance of a full-flow penetrometer can be corrected to its

net value using Equation 4.5.

The effects of the corrections for unequal pore pressure and overburden stress on the tip

resistance are best illustrated in Figure 4.8, which shows the measured resistance and

the corrected resistance profiles for a field cone and (100 cm2) T-bar penetration tests.

As may be seen, the adjustment of the cone data is very significant, with an average

reduction from measured to net values of 19 % over the depth range 3 to 18 m,

compared with a corresponding reduction of only 4 % for the T-bar. This is mainly

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because of the As/Ap term in Equation 4.5, which is relatively small for the T-bar

(generally around 0.1), and thus reduces the weight of the bracketed term.

Consequently, any uncertainty in estimating the unequal area ratio and overburden

stress would have much less impact for the T-bar than for the cone.

Note that, unless otherwise stated, all penetrometer resistance profiles presented in the

following sections are net values.

4.3.2 Resistance profiles for various penetrometers

Results of the field penetration tests are presented in Figures 4.9 to 4.18. The focus here

has been on comparing the net tip resistance measured by the different penetrometers,

and little attention has been paid to sleeve friction or excess pore pressure

measurements. These latter quantities are presented in terms of friction ratio (fs/qcnet)

and Bq value (Δu/qcnet) for the cone tests only. The upper 2 to 3 m includes some sand

seams and desiccated material, and further sand seams start below 18 m. The main zone

of interest is between 4 and 17 m, and the plots have been presented to show this region

in detail.

(a) Effect of T-bar surface

Figure 4.9 shows the penetration and extraction resistances for the smooth T-bars (Tests

1 and 2) and the lightly sand-blasted (rough) T-bars (Tests 3 and 4). All four tests show

very similar resistance profiles, although the smooth T-bars tend to show penetration

resistances slightly lower than the rough T-bars over the depth range 4 to 17 m. Also

presented is the T-bar profile (BTRT01) reported by Schneider et al (2004). This test

only recorded penetration resistances up to 14 m depth, and the resistance profile

obtained is lower than the other T-bar tests, although these T-bar tests were all

performed at the same location at the Burswood site.

During extraction, the smooth T-bars give slightly higher resistances than the rough

T-bars (Figure 4.9 (b)), as oppose to the results observed for the penetration. It should

be noted that the sudden reductions in T-bar extraction resistances observed at depths of

9 m for T-bar 3, and depths of 4 and 14 m for T-bar 4 are where cycles of penetration

and extraction were applied during the initial penetration test. These will be presented

later in the section dealing with the cyclic tests.

Comparison of resistances for the smooth and rough T-bars is best illustrated in

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Figure 4.10, which plots the ratios of the average smooth to rough T-bar resistances.

During penetration, it may be seen that, over the depth range 4 to 17 m, the ratio

fluctuates about an average value of 0.95. Such difference agrees with the theoretical

T-bar resistance (Randolph & Houlsby, 1984), which varies by 5 to 7 % as the interface

friction ratio increases from 0.2 to 0.4, or 0.3 to 0.5.

However, it is evident in Figure 4.10 that the smooth T-bars show higher resistances

than the rough T-bars during extraction. The discrepancy between the two ratio profiles

is possibly, at least partly, due to the smooth surface causing less ‘damage’ or softening

(remoulding) of the soil locally during the first penetration. It is also thought that the

extraction resistances for the rough T-bars may have been affected by the cyclic tests

performed during their initial penetrations. In order to minimise any such effect, only

the T-bar 3 result has been used to compute the ratios of smooth to rough T-bar

resistances during extraction (since only one cyclic test was conducted for T-bar 3), yet

the ratio profile (for extraction) fluctuates generally between unity and 1.1 in the region

below the cyclic test performed at 9 m depth.

(b) Effect of T-bar aspect ratio

Effect of the T-bar aspect ratio (L/d) has been explored by conducting two tests on a

(roughened) T-bar of length 160 mm (L/d = 4) compared with the standard length of

250 mm (L/d = 6.25). Figure 4.11 shows the penetration and extraction resistance

profiles of these two T-bar tests. It is clear that, during penetration, the resistances for

smaller T-bars are marginally lower than those for the standard T-bars, but during

extraction the smaller T-bars become relatively higher, particularly for Test 2 of the

smaller T-bar. Over the depth range 4 to 17 m, the average penetration resistance of the

standard T-bar is 6 % higher than the average penetration resistance of the smaller

T-bar, which is probably not statistically significant. However, the extraction resistance

of the standard T-bar is 14 % lower than the smaller T-bar. This suggests that the degree

of remoulding is lower for the smaller T-bar.

(c) Cone and T-bar

Figure 4.12 compares the net cone resistance profiles with the standard (100 cm2) T-bar

resistance profiles. A cone profile obtained from a seismic cone test (BSCT01)

performed at a nearby location (Schneider et al, 2004) is also included in the figure. The

cone tests have shown significant variation in the penetration resistance profiles, as

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illustrated in Figure 4.12 (a). It may be seen that the penetration resistances for Cones 3

and 4 are very similar, and these profiles are embraced by the results from Cones 1 and

2. It is suspected that the deviating results from Cones 1 and 2 may have been due to not

zeroing the load cell properly before commencing the tests. Since the load cell (with

maximum capacity of 10 MPa) was measuring load (generally < 0.6 MPa) at the lower

end of its resolution limit throughout the tests, slight offset of the zero reading may

cause a relatively significant shift in the measurement. Unfortunately, no data were

recorded before commencement of these two initial cone tests; therefore, it is not

possible to validate the zero reading of the load cell for these tests. The seismic cone

(SCone) shows slightly higher resistances than Cone 1 for depths above 14 m, but

below this depth, the latter cone profile becomes higher.

It can be seen that Cones 3 and 4 exhibit penetration resistances increasing more

strongly with depth and becoming higher than the T-bar resistances below 7 m. The

cone resistance is about 35 % higher than the T-bar resistance by a depth of 16 m.

Interestingly, the seismic cone result is very similar to the T-bar profiles, particularly

below the depth of 10 m.

During extraction, the resistance profiles for Cones 3 and 4 are well above the T-bar

results over the depth range of interest (Figure 4.12 (b)), but the difference appears to be

less between 15 and 17 m. Data were not recorded during extraction for Cones 1 and 2,

and the seismic cone, so extraction data are not available for these cone tests.

Data of sleeve friction, fs, for the cone tests are presented in Figure 4.13 in term of

friction ratio (fs/qcnet). Again, the friction ratios for Cones 3 and 4 are very consistent

both during penetration and extraction, while Cones 1 and 2 show significant variation.

Nevertheless, the values obtained appear suspect, especially for the results from Cones 3

and 4, for which the ratios seem to be extremely low (< 1 % during penetration and 1 –

2 % during extraction). In general, it is very difficult to obtain reliable values for the

sleeve friction for clay materials, as the magnitude is usually too low for the measuring

system in a CPT to measure accurately. Lee Goh (1994) also reported low friction ratios

(mostly 2 – 4 % with a few depth intervals recording zero value) for a CPT carried out

at a location approximately 90 m from where the other cone tests were performed.

Figure 4.14 presents the Bq values (Δu/qcnet) for the cone tests. Cones 3 and 4 show Bq

values fluctuate between 0.4 and 0.55, giving an average of about 0.45 during

penetration, while during extraction the values seem to vary somewhat irregularly

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between 0.1 and 0.25 (absolute), giving an average of about 0.17. Note that, these

average values were used in the correction of cone resistance for centrifuge model cone

tests, which will be described in Chapter 6 later.

(d) T-bar and ball

The penetration and extraction resistances of the standard T-bar and ball penetrometers

are plotted in Figure 4.15. The resistance profile, reported by Schneider et al (2004), for

a ball penetration test (BBTR01) conducted to a depth of 14 m, is also included.

Interestingly, all three tests for the ball penetrometer demonstrate extremely similar

resistance profiles, despite the previous observation that the T-bar and seismic cone

tests from Schneider et al (2004) showed lower resistances than the corresponding tests

performed for this research.

The ball resistances appear to be slightly lower than the T-bar resistances in the upper

14 m during penetration (Figure 4.15 (a)), but marginally higher over the same region

during extraction (Figure 4.15 (b)). Over the depths of interest (4 to 17 m) average

ratios of T-bar to ball resistances are 1.07 and 0.98 for penetration and extraction

respectively. Such differences are deemed to be statistically insignificant. Again, the

periodic decreases in the extraction resistance profiles (Figure 4.15 (b)) are where cyclic

tests were carried out.

(e) T-bar and plate

Comparisons of the penetration and extraction resistances of the standard T-bar

penetrometer and a circular plate are shown in Figure 4.16. The penetration resistance

of Plate 2 is very similar to that measured in the T-bar tests, while that from Plate 1 is

noticeably higher. During extraction, both plate tests show higher resistances than the

T-bar tests. Over the depth range 4 to 17 m, the average T-bar resistance is 0.95 times

the plate resistance during penetration and 0.91 times the plate resistance during

extraction.

(f) Summary of resistance profiles for field penetrometers

A summary plot of average penetration and extraction resistances for all the

penetrometers is shown in Figure 4.17. Note that, as mentioned earlier that, since Cones

1 and 2 have shown resistance profiles that deviate significantly from the profiles of

Cones 3 and 4, which gave extremely consistent resistances, only the latter two cone

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results are taken into account for averaging.

During penetration there is a very tight band of measured resistances, with the plate

penetrometer showing marginally higher resistance over the depth range of interest (4 to

17 m) and the cone penetrometer showing the highest resistance below about 7 m.

During extraction, the cone penetrometer shows the highest resistance in the upper

15 m, followed by the smaller T-bar penetrometer over the same depth range. The latter

penetrometer gives the highest extraction resistance below 15 m, possibly reflecting less

disturbance, or remoulding of the clay, from the smaller penetrometer. The differences

between the penetration and extraction resistances are illustrated more clearly in

Figure 4.18, which shows the ratios of extraction to penetration resistances for the

various penetrometers. It may be seen that, over the depth range of interest, the smaller

T-bar penetrometer has the highest ratio profile, with an average of around 0.72, while

the standard T-bar demonstrates the lowest ratio with an average of 0.55. The ball and

plate penetrometers have similar ratio profiles, just slightly above the standard T-bar. It

is, however, interesting to find that the cone resistance ratio decreases from about 0.9 to

below 0.5 from depths 4 to 17 m.

4.3.3 Field vane tests

Results from two profiles of vane shear tests are shown in Figure 4.19. The two sets of

tests give broadly similar peak shear strength values in the upper parts of the profile, but

show increasing scatter and some divergence with depth, particularly below 10 m. The

remoulded shear strengths range from as low as 8 to 12 % of the peak values, over the

depth range 3 to 5 m, but are more typically 25 to 30 % of the peak values at greater

depth. The corresponding sensitivities are 8 to 12 in the shallower range, decreasing to

between 3 and 4 (see Figure 4.20).

4.3.4 Assessment of undrained shear strength

Estimates of the undrained shear strength may be made from the measured penetration

resistances through the following relationship (Campanella, 1995):

N/qs netu = (4.6)

where: qnet = corrected net tip resistance of the penetrometer;

N = bearing factor for the penetrometer.

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The resulting profiles of undrained shear strength are illustrated in Figure 4.21, using a

single N value of 10.5 for all the penetrometers. The profiles are compared with the

average shear strengths deduced from the two profiles of field vane shear tests, without

application of any correction factor. Interestingly, the peak su profile obtained from the

field vane tests is embraced within the band of su profiles derived from the penetration

tests, except for the cone penetrometer below 7 m, where it starts to exhibit higher shear

strength. However, using N = 13 for the cone penetrometer for depth below 10 m, its su

profile will be comparable to those of the other penetrometers. This may be an

indication of the variation of N value for the cone penetrometer with the soil properties.

Further interpretation of the relationship between the shear strength and tip resistance

(N value) will be discussed in Chapter 7 later, with the addition of shear strength values

measured from the laboratory testing.

4.3.5 Cyclic penetration and extraction tests

As mentioned earlier, cyclic penetration and extraction tests were performed for each

penetrometer (except for the cone) at specific depths. Figure 4.22 shows the tip

resistance response from a single cyclic test carried out at depth of about 9.4 m for

T-bar Test 1. It may be seen that both penetration and extraction resistances continue to

degrade through the 5 cycles, but at a reducing rate, with the resistance stabilizing at a

fully remoulded value.

The degree of degradation in resistance may be measured using a degradation factor,

which is calculated by taking the mean (absolute) value of resistance during the half

cycle of each 1 m stroke divided by the mean value for the initial penetration. For the

cyclic resistance curve shown in Figure 4.22, the resistance during initial penetration is

about 0.222 MPa, while that during the first extraction cycle is around −0.137 MPa, thus

giving a degradation factor of 0.137/0.222 = 0.62 for the first half cycle.

Generally, each half cycle shows further degradation, although there is a slight tendency

in the test shown for the extraction resistance to be lower than the subsequent

penetration resistance, giving a saw-tooth effect in the degradation curve. The cyclic

resistance appears to stabilise after a few cycles, to just under 30 % of the initial

penetration resistance.

In T-bar Test 3, cyclic tests were carried out at a depth of 9.3 m during initial

penetration, and depths of 4.3 m and 14.3 m during extraction. The results of these tests

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are shown in Figure 4.23. The two deeper tests show more asymmetric cyclic resistance

curves, with a more marked saw-tooth pattern in the degradation factor with each half

cycle (see Figure 4.23 (b)). Again, the cyclic resistance stabilises at a value between 25

and 30 % of the initial penetration resistance.

Subsequent cyclic tests were all carried out during initial penetration, at depths of 4.3

and 14.3 m. The results are shown for each test in Figures 4.24 to 4.30, and the

degradation factors are summarised in Figure 4.31. The ball penetrometer shows the

most symmetric pattern of resistance during penetration and extraction, as does the

deeper cyclic test for each type of penetrometer.

A clearer pattern of degradation is shown in Figure 4.32, where the cyclic resistance

curves for each test have been ‘centred’ using a suitable offset, to give a relatively

smooth degradation curve. The plate penetrometer is found to demonstrate the most

rapid degradation, whereas the smaller T-bar (ST-bar) shows the most gradual

degradation, which is consistent with its higher extraction resistance compared to other

penetrometers. At the shallower depth (4.3 m), the final degradation factors for all

penetrometers range between 0.18 and 0.29, while at the greater depth (14.3 m), the

degradation curves are more tightly bunched and converge to a final value between 0.23

and 0.28 with an average of 0.25. Interestingly, the average degradation factor of 0.25 is

consistent with the sensitivity of 3 to 4 from the field vane tests. Therefore, it appears

that the cyclic test results may be used to measure the remoulded strength and

sensitivity of the soil.

4.4 Summary for field testing

This chapter has presented the procedures and results of the in situ testing carried out at

the Burswood site. The findings and observations from the field testing results are

summarised below.

First, it has been illustrated that the measured cone tip resistance is very sensitive to the

corrections for unequal pore pressure and overburden stress effects. As a result, errors in

estimating the quantities of these two parameters could lead to large inaccuracy in the

net values for the corrected cone resistance, and hence in the derived values for the

undrained shear strength. On the other hand, the ‘full-flow’ penetrometers (T-bar, ball

and plate) are relatively insensitive to the unequal pore pressure and overburden

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pressure effects, due to their much larger projected area compared to the cross-sectional

area of rod shaft (= projected area of cone). Therefore, the effects of corrections on the

measured resistance for a full-flow penetrometer are generally negligible.

Also, it has been shown that the difference in the T-bar resistance due to variation of its

cylindrical surface condition (smooth to lightly sand-blasted) appears to agree with the

variation predicted by the theoretical T-bar resistance (Randolph & Houlsby, 1984), as

the interface friction ratio is slightly altered. However, this was not the case during

extraction. One possible reason may be due to the smooth surface causing less

remoulding of the soil locally, which was reflected by the slightly lower penetration

resistance for the smooth T-bar. Therefore, the extraction resistance for the smooth

T-bar was in turn slightly higher than the rough T-bar.

Interestingly, all penetrometers, apart from the cone, have demonstrated very consistent

tip resistances both during penetration and extraction, although the plate tended to show

marginally higher penetration resistance, whereas the smaller T-bar was found give

slightly higher extraction resistance. However, the cone penetrometer generally gave the

highest penetration and extraction resistances.

In addition, the undrained shear strength, su, profiles derived from the penetration tip

resistances using a single value of N = 10.5 for all penetrometers match reasonably well

with the peak su measured from the field vane tests, although again, the cone gave the

highest su profile, diverging with increasing depth. Using N = 13 for the cone

penetrometer at depths below 10 m will bring its su profile comparable to those of the

other penetrometers. This may be an indication of variation of N value for the cone

penetrometer with the soil properties.

In cyclic penetration and extraction tests, the plate penetrometer was found to show the

most rapid degradation of tip resistance, while the smaller T-bar appeared to

demonstrate the most gradual resistance degradation. Nevertheless, after 5 or 6

complete cycles of penetration and extraction, all penetrometers show resistances

degrading to about 25 % of the initial values. The implied sensitivity of around 4 is

consistent with that observed in the vane shear tests.

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5 LABORATORY TESTING

In the previous chapter, results from the in situ testing conducted at the Burswood site

were presented. This chapter presents results from the laboratory testing carried out on

tube samples collected from the same site where the in situ testing was conducted.

The primary aim of the laboratory testing was to evaluate the in situ undrained shear

strength of the clay, in addition to assessing its other geotechnical parameters such as

consolidation coefficient, stress history, index properties etc. Since the shear strength is

not a unique value, but varies depending on the type of testing (Wroth, 1984), different

tests including UU compression, CAU triaxial compression and extension and CAU

simple shear tests were undertaken to deduce the various shear strength values.

The subsequent sections provide descriptions of the tube samples, test apparatus, testing

procedures and equations used for assessing the geotechnical parameters. The testing

results are then presented, which form part of the geotechnical background for the

Burswood site.

5.1 In situ tube samples

The tube samples used for the laboratory testing were collected from two boreholes,

referred to as BH1 and BH2, the locations of which are shown in Figure 4.1. The former

borehole had a nominal diameter of 72 mm, while that for the latter borehole was

100 mm. Continuous sampling was carried out, with sampling interval of approximately

700 mm, from the ground surface to depths of 19.06 and 18.85 m for boreholes BH1

and BH2 respectively.

Dimensions of the sampling tubes used for BH1 and BH2 are summarised in Table 5.1.

All the tubes had a length of 750 mm. Plastic and stainless steel tubes were used in

order to avoid or minimise corrosion. The plastic tubes had a relatively large wall

thickness, giving an external diameter to wall thickness ratio, dE/t, of around 31, while

the dE/t for the stainless steel tubes was approximately 51. The smaller diameter tubes

for BH1 were tapered to an outside cutting edge angle of ~15°, while the large tubes for

BH2 were tapered to ~7 to 9°, and the inside wall was completely flush (i.e. zero inside

clearance) for all tubes.

Clayton et al (1998) showed that sampling tubes with a sharp outside cutting edge and

zero inside clearance should give the best quality samples for soft clays. A 5° outside

LABORATORY TESTING

5-2

cutting edge for use in practice was subsequently suggested by Hight & Leroueil (2003).

Also, Ladd & DeGroot (2003) recommended that the sampling tube should have a

minimum external diameter of 76 mm and a wall thickness such that dE/t > 45.

As may be seen, the sampling tubes used for the research (similar to local practice) did

not always comply with the requirements recommended in the literature, particularly the

plastic tubes. Therefore, one would expect that these sampling tubes would generate a

higher degree of sample disturbance during insertion of the tube, compared to if an

‘ideal’ sampling tube had been used.

Before extrusion for any testing, all the tube samples were X-rayed to identify existence

of any cracks, shells or other abnormalities. For both boreholes, the X-rays suggested

that shells were encountered frequently to a depth of approximately 12 m below the

ground surface. Below this depth, occasional shells and shell fragments were observed

in some samples, but no clusters of large shell pieces were encountered. Figure 5.1

includes (a) X-ray of sample from between 8.40 and 9.15 m depth showing several

cracks and shells; and (b) X-ray of sample from greater depth between 12.90 and

13.65 m, in which only a few tiny shells were found. Where possible, portions

containing visible cracks and shells were avoided when selecting the section of a sample

for any laboratory testing.

5.2 Laboratory testing apparatus and procedure

5.2.1 Index tests

Soil characterisation tests have been carried out to determine the water content,

Atterberg limits, in situ unit weight and the particle size distribution.

The in situ water content profile against depth was estimated using trimmings from

samples prepared for laboratory testing. The trimmings were left air-dried in a

temperature controlled room at 20 °C to determine the water content at the

corresponding depth.

Atterberg limits were performed in accordance with Australian Standard AS 1289

(1991, 1995a), using the four-point determination method.

A grading test for the soil particle size was completed on a bulk sample of material

collected from a depth of approximately 6 m, at a location near the main testing area

(approximately 30 m East from BH1). The test consisted of two stages in which the first

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5-3

stage was to sieve particles with sizes ranging from 0.075 to 4.75 mm, in accordance

with AS 1289 (1995b) and the second stage was to determine the size distribution of the

particles smaller than 0.075 mm with the aid of a MASTERSIZER MAF 5000 sizing

machine, manufactured by Malvern Instruments. Details of the testing procedure with

the sizing machine have been provided by Levy et al (2002).

5.2.2 Constant rate of strain consolidation (CRSC)

The test apparatus used in the constant rate of strain consolidation (CRSC) tests is

shown schematically in Figure 5.2. It consists of a base cell with an oedometer cutting

ring, a piston and a load cell. A cutting ring with nominal internal diameter 71 mm was

used for tests CRSC 4 to CRSC 10, but for tests CRSC 11 to CRSC 18 a smaller

diameter (61 mm) cutting ring was used. The height of both cutting rings was 25 mm.

The rings were placed inside the oedometer cell with the soil sample inside it. The

piston and the load cell were then placed on top of the sample. A linear variable

differential transformer (LVDT) was placed on top of the piston to measure the sample

displacement during testing.

The sample was loaded at a rate of 0.003 mm/min (strain rate of 2 x 10-6 s-1), which rate

was found suitable for the range of samples tested to keep the excess pore pressure

below 20 % of the total applied vertical stress. The sample was left for compression for

at least 48 hours.

5.2.3 UU and CAU triaxial tests

The triaxial apparatus used for an unconsolidated undrained (UU) and a consolidated

anisotropically undrained (CAU) triaxial tests is illustrated schematically in Figure 5.3.

Further details of the apparatus have been provided by Levy et al (2002). Both types of

tests were undertaken on a sample with nominal diameter and length of 72 and 150 mm

respectively.

For a UU compression test, prior to loading the sample, a cell pressure of 300 kPa was

applied in order to eliminate any negative pore pressure. The sample was left for about

an hour to allow it to stabilise under the applied cell pressure. Then, it was compressed

at a rate of 1.5 mm/min (strain rate of 167 x 10-6 s-1) in an undrained condition. The test

was terminated after more than 15 % axial strain had been reached.

In a CAU triaxial test, before consolidating the sample, cell and back pressures of 1010

LABORATORY TESTING

5-4

and 1000 kPa respectively were applied. The sample was then allowed to saturate with

water from the system under the small pressure difference. A B-value was measured to

determine the degree of saturation. This was done by increasing the cell pressure to

1060 kPa, giving a change of 50 kPa. The corresponding increase in pore pressure was

noted and the B-value was calculated as:

3

uBσΔ

Δ= (5.1)

where: Δu = change in pore pressure;

Δσ3 = change in cell pressure.

The degree of saturation was considered to be satisfactory when a minimum B-value of

0.95 was achieved.

After that, the sample was consolidated anisotropically to the effective in situ vertical

stress, σ'vo, and effective in situ horizontal stress, σ'ho. Note that the value of σ'vo was

determined from the best-fit curve of the unit weight derived from the bulk densities

measured for UU triaxial tests, due to data from other testing being unavailable at the

time. The σ'v profile used for the testing will be presented later in Section 5.3.1.

A Ko value of 0.8 was adopted for σ'ho prior to results being available from the CRSC

tests. This value was deemed to be suitable for clay expected to have an

overconsolidation ratio, OCR, of 1.5 to 2. Thus, assuming a normally consolidated Ko of

around 0.5 to 0.6 and following the Mayne & Kulhawy (1982) correlation of:

( ) 65.0ncoo OCRKK ⋅= (5.2)

gives a Ko range of 0.65 to 0.94 for OCRs between 1 and 2, with an average of 0.8.

After primary consolidation was completed, the sample was loaded in an undrained

condition at a rate of 0.2 mm/min (strain rate of 23 x 10-6 s-1) to a minimum axial strain

of 15 % for both compression and extension tests.

5.2.4 CAU simple shear test

A schematic diagram of the simple shear cell is shown in Figure 5.4. Again, further

details have been provided by Levy et al (2002). The test sample was 29 mm high, with

a nominal diameter of 72 mm. To avoid slippage occurring between the sample and the

base pedestal, and between the sample and the top cap, two circular steel plates with

LABORATORY TESTING

5-5

pins on one face were placed on both ends of the sample, with the pins carefully

inserted into the sample. The pins were 1 mm in diameter and alternately 3 and 5 mm in

height, spanning over the plate on a 5 mm square grid.

Similar to the CAU triaxial test, the sample was saturated with water from the system at

cell and back pressures of 410 and 400 kPa respectively. After a minimum B-value of

0.95 was achieved for the saturation, the sample was consolidated anisotropically to the

required σ'v, with a Ko = 0.8 for the σ'h (as for the CAU triaxial test).

The consolidated sample was sheared at a rate of horizontal displacement of

0.1 mm/min (shear strain rate of 57 x 10-6 s-1) to a minimum shear strain of 30 %, with

the sample height and total vertical stress maintained constant.

5.2.5 Model T-bar test in triaxial

Figure 5.5 shows the diagram of a miniature T-bar test carried out on a sample

consolidated in the triaxial cell (further details are given by Levy et al (2002)). The

miniature T-bar penetrometer, photographs of which are shown in Figure 5.6, consists

of a 380 mm long penetrometer rod with a diameter of 6 mm and a T-bar tip with a

length and diameter of 29 and 6 mm respectively. The model T-bar tip was lightly sand-

blasted.

The T-bar test was performed on a sample with a nominal diameter and height of 100

and 190 mm respectively. Again, before consolidation, the sample was saturated at cell

and back pressures of 320 and 300 kPa respectively, but a slightly lower minimum B-

value of 0.92 was targeted.

However, with this apparatus set up, only isotropic consolidation was allowed.

Therefore, the sample was consolidated at a mean effective stress, p', calculated as:

3

'2''p hv σ+σ= (5.3)

where: σ'v, σ'h = Effective vertical and horizontal stresses respectively.

Similar to the CAU triaxial test, the value of σ'v was determined from the best-fit curve

of the unit weight derived from the bulk densities measured for UU triaxial tests (see

later in Section 5.3.1), and a Ko value of 0.8 was adopted. The sample was left for

consolidation for about 3 to 4 days.

Then, the model T-bar was pushed into the sample at a penetration rate of 50 mm/min

LABORATORY TESTING

5-6

for a distance of approximately 85 mm, before being extracted at the same rate.

5.2.6 Calibration for T-bar triaxial test

Two calibration tests were performed on the load cell at the tip of the penetrometer rod.

First, it was calibrated using a proving ring up to 470 N, where the capacity of the load

cell was 540 N. A calibration factor of 0.6672 N/bit was obtained, which corresponded

to 0.00384 MPa/bit for the T-bar penetrometer.

Then, the load cell was calibrated against pore pressure in a triaxial cell. A water

pressure was applied inside the cell and the corresponding response of the load cell was

recorded in bits. A calibration factor of −0.0509 MPa/bit was obtained, the negative

number per bit indicated that an apparent tension was exerted on the load cell when the

water pressure was applied, although the true reason for this is because the load cell was

sensitive to changes in normal stress acting directly on the load cell strain gauges.

Using the above calibration factors obtained from direct and water pressure loadings, it

may be shown that a change in normal stress of 50 kPa would give rise to an error in the

T-bar resistance of around 12 kPa (and consequential error in the deduced shear strength

of ~1.1 kPa, i.e. about 1 to 4 % error for typical strength values deduced from laboratory

T-bar tests). This error is considered insignificant.

5.2.7 SHANSEP procedure

SHANSEP (Stress history and normalised soil engineering properties) is a procedure

followed before shearing, in the anticipation to eliminate any sample disturbance

induced during sampling and during total stress reduction when the sample is recovered

and exposed to atmospheric pressure (Ladd & Foott, 1974). The method is first to

normally consolidate the sample to an effective vertical stress higher than the

preconsolidation, or yield stress (σ'yield); then, the vertical stress is reduced to a level that

gives the same overconsolidation ratio, OCR, as in the field. After the swelling is

completed, the sample is loaded in a similar manner to standard CAU triaxial and

simple shear tests.

In the triaxial and simple shear tests following the SHANSEP procedure, the sample

was first consolidated to an effective vertical stress (σ'v1) 20 % higher than its yield

stress, which value was estimated using the Casagrande method (presented later in

Section 5.3.2). The (Ko)nc value was back estimated using Equation 5.2, with a constant

LABORATORY TESTING

5-7

value of 0.8 for the overconsolidated Ko and the OCR value estimated from the results

of CRSC tests (see later in Section 5.3.2). During swelling, the effective vertical stress

(σ'v2) was reduced to the level that produced the same OCR and a Ko = 0.8 was used for

the horizontal stress.

5.3 Laboratory test results

5.3.1 Index tests

The in situ water content, unit weight and Atterberg limits profiles against depth are

summarised in Figure 5.7, while the soil grading is shown in Figure 5.8.

A best-fit trend is derived for the in situ unit weights measured from all laboratory test

samples, as illustrated in Figure 5.7. The derived unit weight is uniform at around

14 kN/m3 between depths of 4 and 6 m. It then increases at a gradient of 0.3 kN/m3 per

metre until 14.5 m, where it drops back to around 14.6 kN/m3. The trend of the unit

weight is consistent with the in situ water contents measured from all samples before

testing.

The water content is generally higher at shallow depths, with an average of about 90 %

between depths of 4 and 6 m, then decreasing to below 60 % from 6 to 14.5 m, before

increasing again to an average value of 70 %. The liquid limit, ωL, is mainly between 60

and 80 %, with a slight trend to decrease with increasing depth between 6 to 14.5 m.

In situ water content is generally found to be very close to ωL, with some values up to

20 % higher than ωL between 2 and 6 m. This is consistent with the high sensitivity of

the clay at this depth range, with the soil strength reducing significantly on remoulding.

The plastic limit, ωP, is relatively constant with depth, with a range of 24 to 37 % and an

average value of 30 %. The plasticity index (ωL - ωP) lies between 28 and 49 %, with an

average value of 38 %.

However, the liquid limit and plasticity index data obtained here are significantly lower

than the values reported by Lee Goh (1994), who performed classification tests on

material collected from a borehole located within a proximity of 90 m from borehole

BH1. Lee Goh reports ωL values of 105 to 126 % and plasticity indices of 72 to 95 %

over the depth range 5 to 7 m.

One possible reason for the lower ωL values obtained here could be due to the existence

LABORATORY TESTING

5-8

of shell fragments and tiny organic fragments, which cannot be removed completely by

hand from the samples. These organic fragments may affect the fall cone penetration

reading in the procedures following AS 1289 (1991). They may also affect the water

content during the determination of ωL, since only a small amount of material from the

sample is used for determining the value of ωL.

A total in situ vertical stress, σvo, profile was deduced using the best-fit trend of unit

weight shown in Figure 5.7. The water table was measured at the site and found to be

1 m below the ground surface. Using this information, the effective in situ vertical

stress, σ'vo, profile was computed and the stress profiles are illustrated in Figure 5.9. It

can be seen that the σ'vo profile increases rather uniformly with depth to about 100 kPa

at 20 m.

5.3.2 CRSC test

The CRSC test results are summarised in Table 5.2 and are shown for the different tests

in Appendix A. Also, it can be noted in Table 5.2 that quality of the test sample is rated

based on NGI’s criterion (Lunne et al, 1997a), as shown in Table 5.3.

The results from the CRSC tests are calculated as follows. The initial void ratio is first

determined as (postulating 100 % saturation):

sii Ge ω= (5.4)

where: ωi = initial water content;

Gs = specific gravity of the soil particles, taken as 2.62 (quoted from Lee

Goh (1994)).

The current void ratio at any stage during the test is then:

⎟⎟⎠

⎞⎜⎜⎝

⎛ +⋅Δ−=

i

ii H

e1Hee (5.5)

where: Hi = initial sample height;

ΔH = change in sample height (positive indicates decrease in height).

The average effective vertical stress in the sample is estimated assuming a parabolic

isochrone, as:

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5-9

bvv u32' −σ=σ (5.6)

where: σv = total vertical stress applied at the top of the sample (equal to the

effective stress, since the pore pressure is zero there);

ub = pore pressure measured at the sample base.

The coefficient of consolidation is then estimated as (Wissa et al, 1971):

tu2'Hc

b

v2

v ΔσΔ

= (5.7)

where: H = current sample height;

Δσ'v/Δt = current rate of effective stress increases.

The permeability may then be deduced as (Wissa et al, 1971):

tu2HHk

b

w

ΔΔγ

= (5.8)

where: γw = unit weight of water;

ΔH/Δt = compression rate adopted for the CRSC test.

(a) Compressibility data

The relationship between the void ratio, e, and the vertical effective stress, σ'v, shows

that the virgin consolidation lines are in most cases straight lines in the standard e

versus log (σ'v) plot (see Appendix A). The preconsolidation (yield) stress, σ'yield,

estimated using the Casagrande method and the overconsolidation ratio, OCR, are

plotted against depth in Figure 5.10. Also, a curve has been fitted through the OCR data

points, with OCR of around 3.5 at a depth of 2 m, decreasing parabolically to about 1.65

at 10 m, below which it remains almost constant.

The compression index, Cc, is derived as the magnitude of the gradient of the virgin

consolidation line in the e versus log (σ'v) plot. The compression ratio is then deduced

as the ratio of Cc/(1+eo), where eo is the estimated in situ void ratio. The plots of the

compression index and compression ratio against depth are shown in Figure 5.11. The

Cc data seem to show a trend decreasing with increasing depth, with values lying within

1 and 1.5 above 8 m, while below this depth, the Cc values reduce to between 0.5 and 1.

Nevertheless, the compression ratio remains relatively constant with depth, with an

LABORATORY TESTING

5-10

average value of around 0.34.

Figure 5.12 plots the profiles for the initial (ei) and in situ (eo) void ratios, both of which

seem to decrease as the depth increases. The eo values are generally around 2.6 for

depths above 6 m, but reduce to between 1.2 and 1.7 below this depth.

(b) Coefficient of consolidation

Most tests showed extremely high initial values of cv (see Appendix A), owing to the

very low measured pore pressure (~zero for some cases) with relatively high rate of

increase of the vertical stress at the beginning of testing (see Equation 5.7). This may be

partly caused by no back pressure being used in the CRSC tests; it is common to apply

back pressure to saturate the CRSC test samples (Larsson & Sällfors, 1986; ASTM,

2000b).

As the effective stress increased beyond the yield stress, the cv values in most cases

decreased markedly and stabilised at very high vertical stress levels (~3 to 5 times

σ'yield). For this reason, cv values were obtained at two stress levels: (a) at σ'yield; and (b)

at 5 x σ'yield after the cv stabilises. The two sets of cv values obtained are presented in

Figure 5.12, together with the profiles of initial and in situ void ratios.

The values of consolidation coefficient determined at σ'yield (cvy) range generally

between 1 and 1.25 m2/year at depths of 6 to 12 m, apart from the occasional high

values. However, higher cvy values are found at both the shallower and deeper depth

ranges, particularly at the depth of 18.4 m, where cvy of 4 m2/year was measured. The

relatively high values of cv measured in some tests are mainly attributed to the presence

of silt lenses in the samples. A similar trend is found for cv data obtained at 5 x σ'yield,

with the values in most cases lying between 0.5 and 1 m2/year. This range is broadly

consistent with results from Lee Goh (1994), who reported a (normally consolidated) cv

range of 0.33 to 0.89 m2/year over depths of 5 to 7 m.

(c) Coefficient of permeability

The relationship of void ratio, e, versus coefficient of permeability, k, is plotted for each

CRSC test in Appendix A. The e – log(k) curves in most tests are curled at high void

ratios, and then become linear as the void ratio decreases, which is common for soft

clays (Tavenas et al, 1983). The in situ values of k deduced for the Burswood clay are

very low, ranging from 3.5 x 10-10 to 2.4 x 10-9 m/s.

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5-11

(d) Pore pressure ratio

The measured pore pressure divided by the applied vertical stress, ub/σ'v, depends on the

strain rate adopted for the test. High strain rates, with correspondingly high pore

pressure ratios, result in viscous effects (Leroueil & Marques, 1996), which increase the

soil yield stress. On the other hand very low strain rates, and resulting low pore pressure

ratios, affect the accuracy of the computations of cv and k. (see Equations 5.7 and 5.8).

ASTM (2000b) recommends that pore pressure ratios should be in the range 3 to 20 %.

For Scandinavian clays, ub/σ'v ratios of 7 % (Larsson & Sällfors, 1986) and 15 %

(Sandbaekken et al, 1986) have been suggested. Almeida et al (1995) reported ub/σ'v in

the range 10 to 32 % and obtained good agreement between conventional oedometer

and CRSC tests.

Results of the Burswood clay CRSC tests in Appendix A show that the data on pore

pressure ratio plotted versus the effective vertical stress are in most tests high early in

the test and then decrease or stabilise, as the stress increases. Pore pressure ratios are in

the range 4 to 20 % for most tests, apart from occasionally higher ub/σ'v values observed

earlier in some tests.

5.3.3 UU and CAU triaxial tests

Results from the UU and CAU triaxial tests are summarised in Tables 5.4 and 5.5

respectively. Individual CAU triaxial test results from consolidation to loading stage are

also presented in Appendix B.

Most of the triaxial test samples showed only moderate consolidation strains, with Δe/ei

values around 0.05. However, two samples (TXC4 and TXE1) showed higher strains,

implying lower quality of the sample. The latter sample also showed extremely high

shear strength in comparison with other data and a strangely high stress ratio at failure.

In general, the compression results consistently show a high friction angle, with final

stress ratios of q/p' in the region of 1.75 (φ' ~ 42°). The extension tests on samples from

greater depths give q/p' of around 0.88, implying a mobilised friction angle of φ' ~ 31°

in extension, while that from shallower depth (excluding TXE1) shows a lower q/p' ~

0.68 (φ' ~ 22°).

5.3.4 CAU simple shear test

The simple shear test results are summarised in Table 5.6, while the individual test

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5-12

results are shown in Appendix C.

The samples all show high strains during consolidation. However, it is believed that the

apparent axial strains were caused by penetration of the pinned end platens into the

samples. Stress ratios, t/s', mobilised during testing range from 0.52 to 0.66

(corresponding to φ' ~ 32 to 42°), apart from tests SS3 and SS6, where the former test

gives extremely low ratio (t/s' ~ 0.32 or φ' ~ 19°), while the opposite is the case for the

latter test t/s' ~ 1 or φ' ~ 90°).

5.3.5 Model T-bar test in triaxial

Table 5.7 and Appendix D present the summary and individual test results respectively

for the T-bar in triaxial tests.

The samples showed consolidation strains consistent with those in the CAU triaxial

tests, despite the T-bar test samples being consolidated isotropically. The fluctuations

and irregularities observed in the resistance profiles may be attributed to the T-bar

interacting with shells or layer boundaries of the samples. To confirm this, one of the

samples was cut open after testing and some shells were found near the penetration path

at the locations of these fluctuations in resistance.

The average T-bar penetration (qin) and extraction resistances (qout) and the resistance

ratio of qout/qin (absolute values) are plotted against depth of the corresponding test

sample in Figure 5.13. The resistance profiles obtained from the in situ T-bar testing are

also included for comparison. It is obvious that the penetration resistance for the

laboratory T-bar is usually significantly higher than that for the field T-bars, but the

opposite is true for the extraction resistance, although the difference is less than that for

penetration. It should be reminded that cyclic penetration and extraction tests were

performed at depths of 4.3, 9.3 and 14.3 m for the field standard T-bar (see Chapter 4),

resulting in local reductions in resistance at those depths during extraction. Therefore,

although it can be seen in Figure 5.13 (b) that the extraction data for the laboratory

T-bar fall onto the profile for the field standard T-bar at depths of 4.3 and 9.8 m, these

laboratory data would have still been lower than the field standard T-bar results, if the

cyclic tests have not been performed.

Since the laboratory T-bar tests give significantly higher penetration resistances, but

lower extraction resistances than the field T-bar tests, the consequence is the resistance

(absolute) ratios of qout/qin for the laboratory T-bar become considerably lower (between

LABORATORY TESTING

5-13

0.20 and 0.31, apart from Test 4, see Table 5.7) compared to the field results (average

ratios of 0.55 and 0.72 for the field standard and smaller T-bar tests respectively), as

shown in Figure 5.13 (c).

Several factors may have contributed to the discrepancy between the laboratory and

field T-bar resistances. First, as already mentioned, the resistance for the laboratory

T-bar tests were influenced by the existence of shells and layering of the samples, which

would have had a much less severe influence on the tip resistance for the field T-bar due

to its relatively larger size.

Secondly, it is speculated that the mode of penetration may have been different in the

laboratory T-bar test, compared to the field T-bar test. The soil may not have flowed

back over the T-bar in the laboratory test, allowing the T-bar to be extracted essentially

within an open ‘slot’. This is supported by the observation of slurry-like material found

within the T-bar slot when the sample was dismantled after testing. As a result, the

extraction resistance measured is comparatively low.

Furthermore, since the samples were consolidated isotropically at the mean effective

stress, they would experience a lower effective vertical stress, but higher effective

horizontal stress than the stress levels experienced in the field. The effect of this on the

measured T-bar resistance is difficult to gauge, but might have partly contributed to the

higher resistance than the field results.

5.3.6 Triaxial and simple shear tests following SHANSEP

The results from triaxial and simple shear tests undertaken using the SHANSEP

procedure are summarised in Table 5.8, with the individual test results provided in

Appendix E.

The values of su measured in the tests are those corresponding to the higher vertical

stresses of SHANSEP and need to be adjusted to correspond to the in situ stress levels.

The adjustment can be made using an expression derived from the following correlation

of Ladd et al (1974, 1977) and Wroth (1984):

8.0

ncvu

ocvu OCR)'/s()'/s(

=σσ (5.9)

where: (su /σ'v)oc = ratio of su /σ'v for overconsolidated clay;

(su /σ'v)nc = ratio of su /σ'v for normally consolidated clay.

LABORATORY TESTING

5-14

Then, it can be shown that:

2.0

1

2

1yield

2yield1u2u OCR

OCR''

ss−

⎟⎟⎠

⎞⎜⎜⎝

⎛×

σσ

×= (5.10)

where: su 2 = adjusted undrained shear strength;

su 1 = measured undrained shear strength after SHANSEP is followed;

σ'yield 2 = in situ yield stress, which profile is shown in Figure 5.10;

σ'yield 1 = actual yield stress experienced in testing;

OCR2 = in situ OCR, which profile is shown in Figure 5.10;

OCR1 = actual OCR value in testing.

The adjusted su values are from 8 to 14 % lower than the measured values.

5.4 Undrained shear strength profiles

Figure 5.14 (a) presents the undrained shear strength, su, profiles measured from the

various laboratory testing. The strength data from the laboratory T-bar tests are derived

from the average penetration resistance using a bearing factor, N = 10.5. The field

strength profiles estimated from the cone and T-bar penetration tests are also included

for comparison.

It may be seen that there is a large scatter in the laboratory test data in both the shallow

and deep regions, much of which may be attributed to variable sample quality due to the

presence of frequent shell fragments and occasional silt lenses. Where shell fragments

were evident on the surface of the sample, they were removed and the voids filled with

material from trimming before testing. This may have weakened the sample. However,

shell fragments inside the sample could not be removed and this may have caused

higher strength measured for that sample.

As has been mentioned earlier, the presence of even small shell fragments near the

passage of the laboratory T-bar would have a relatively large influence on the measured

tip resistance (hence the deduced su). This is believed to be the main reason for the

laboratory T-bar tests generally showing the highest su values.

In the shallow depth region above 7 m, su values measured from the UU tests seem to be

the average for the data from other tests. However, the UU test data are found to be the

lowest between depths of 8 and 13 m, but the highest below 14 m (apart from the

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5-15

laboratory T-bar data). The erratic behaviour of the UU strengths was also noted by

Ladd & DeGroot (2003), who commented that the su values measured from UU tests

can easily be 25 to 50 % too high or too low, depending on the compensating effects

between the sampling disturbance and the relatively high shearing rate for the UU tests.

Nevertheless, it may be seen in Figure 5.14 (a) that the UU strength profile is broadly

similar to the field strength profile derived from the T-bar penetration test, although a

relatively high degree of scatter is observed for the UU tests.

For the CAU triaxial tests, the compression tests (TXC) gave an average su of about

19.6 kPa at 6 m depth, increasing to about 48.3 kPa at 17 m, while the extension tests

(TXE) gave an su of about 10 kPa at 5.3 m depth (ignoring the first TXE test at 5.4 m

depth, which showed a strangely high stress ratio, q/p'), increasing to about 31 kPa at

17 m. The TXE strength at the depth of 6 m was computed by interpolation, and this

consequently gives anisotropic strength ratio (su, TXE/su, TXC) of about 0.58 at 6 m depth,

increasing to 0.64 at 17 m. These ratios are consistent with (or marginally lower than)

the ratios estimated based on plasticity index (Ladd, 1991). A similar finding was also

noted for the Onsøy clay in Norway, with su, TXE/su, TXC ranging from 0.4 to 0.5 (Lunne

et al, 2003).

For the simple shear tests (SS), if the data points from tests SS3 at 5.8 m and SS6 at

4.8 m are discarded (since the former test gave extremely low stress ratio, t/s', compared

to others, whereas the latter gave a strangely high t/s'), then the simple shear data appear

to lie between the TXC and TXE strengths, although closer to the TXE results. The

strength ratio, su, SS/su, TXC is approximately 0.73 at depth of 6 m, decreasing to around

0.7 at 17 m (the su, SS values at 6 and 17 m depths are calculated by interpolation and

extrapolation respectively). The trend of su, SS/su, TXC decreasing with decreasing

overconsolidation ratio (OCR) was also noted for James Bay clay (a highly structured

and cemented clay), as quoted by Ladd & DeGroot (2003).

The TXC data follow the field cone profile very well, while the simple shear data agree

well with the field T-bar profile, particularly in the intermediate depth range 9 to 13 m.

It can be seen in Figure 5.14 (b) that all the tests with SHANSEP showed higher shear

strengths than the corresponding tests with standard re-consolidation to in situ stresses.

The simple shear and TXC tests using the SHANSEP procedure show consistent and

higher strength values than those for the TXE tests using the SHANSEP procedure, for

which the strength data fall directly on the field T-bar strength profile.

LABORATORY TESTING

5-16

Further interpretation between the shear strengths measured from the various laboratory

testing and measured tip resistances from the various in situ penetration tests will be

presented in Chapter 7.

5.5 Summary for laboratory testing

This chapter has presented data from the laboratory tests undertaken on tube samples

retrieved from the Burswood site. The overall quality of the samples was moderate to

poor by NGI’s criterion (Lunne et al, 1997a), and this has contributed to some scatter in

the data. Much of the difficulties have been associated with the occurrence of shell

fragments within the samples, mainly in the depth range 3 to 12 m.

The in situ unit weight was found to lie between 14 and 16 kN/m3, with data from the

depth range 11 to 15 m being closer to the upper value. The in situ water content is

higher in the shallow depth region above 7 m, with values ranging from 80 to 100 %.

Below this depth, it is mostly between 50 and 70 %. The liquid limit is mainly between

60 and 80 %, whereas the plastic limit is rather constant, with an average of 30 %.

The overconsolidation ratio data are in most cases between 1.3 and 2, with an average

around 1.65 below a depth of 10 m.

Data for the in situ consolidation coefficient are rather scattered, but range generally

between 1 and 1.25 m2/year in the depth range 6 to 12 m. The consolidation data

obtained at a higher reference stress (5 times the yield stress) lie generally between 0.5

and 1 m2/year over the same depth range.

The friction angle of the clay was found to be very high, approximately 42° measured

from the triaxial compression tests. The extension tests gave angles ranging from 22 to

31°, while those from the simple shear tests ranged from 31 to 42°.

It has been shown that scale effects can be of critical importance for the model T-bar

test in the laboratory, and this set up may not be appropriate for testing in situ samples

containing shell fragments.

In the shallow depth region above 7 m, shear strengths measured from the various

laboratory tests were mainly between 10 and 21 kPa, whereas at depths below 16 m,

strengths between 30 and 55 kPa were obtained.

All tests following the SHANSEP procedure gave higher shear strengths than the

corresponding standard tests, although the degree of strength increase varies for

LABORATORY TESTING

5-17

different test types.

CENTRIFUGE TESTING

6-1

6 CENTRIFUGE TESTING

The previous two chapters have described the in situ and laboratory tests carried out at

the Burswood site, in order to evaluate the relationship between the in situ and

laboratory estimated shear strengths. This chapter now presents the work and results for

the model tests conducted on reconstituted Burswood clay in the centrifuge.

Clay samples for the centrifuge testing were reconstituted using material collected from

the Burswood site at a depth of about 6 m, at a location about 30 m from where the in

situ tests were conducted. Two samples (centrifuge ‘strongboxes’) were prepared: the

first strongbox was tested with different types of model penetrometers and the second

strongbox with T-bar penetrometers of different aspect ratios. In addition, hand vane

tests were performed at 1 g at the end of the centrifuge tests, and several tube samples

were retrieved for laboratory testing, after the strongbox samples had been dismantled.

The following sections first describe the soil properties of the reconstituted samples, the

apparatus and the testing procedure. The results will then be presented with attention

focussed mainly on comparing the tip resistances and the deduced shear strengths from

the various model penetrometer tests.

6.1 Reconstituted sample properties

Soil properties for the centrifuge samples reconstituted from Burswood clay are

summarised in Table 6.1. These properties are generally based on results from

laboratory tests conducted on the tube samples retrieved after completion of the

centrifuge testing. It should be noted that the friction angles shown in Table 6.1 have

been measured from the triaxial compression tests. These angles (32 and 29°) are rather

low in comparison with corresponding values measured for the tube samples collected

from the field from depths of 5.6 to 17.1 m (φ'field ~ 42°).

Water contents of the centrifuge samples are similar to (or slightly lower than) the

moisture in field. The Atterberg limits listed in Table 6.1 were obtained from the bulk

material used for preparing the centrifuge samples. The liquid limit (ωL) and plasticity

index (PI) for this bulk material are relatively higher than the values for material

obtained from borehole BH1 at similar depth of 6 m, but are lower than values reported

by Lee Goh (1994).

The coefficients of consolidation, cv, listed in Table 6.1 have been estimated using data

CENTRIFUGE TESTING

6-2

from consolidation of the samples in the centrifuge at a centrifugal acceleration of

100 g. Other values of cv were also obtained using several different methods, which will

be presented and discussed in Chapter 8 later.

6.2 Centrifuge testing apparatus

Figure 6.1 shows a photograph of the beam centrifuge used for this research. A

complete description of the facility has been given by Randolph et al (1991). Also, full

discussion of the scaling relationships can be found in Schofield (1980) and Taylor

(1995). For convenience, a brief summary of the scaling relationships is shown in

Table 6.2. All tests reported here were conducted at an acceleration level of 100 g.

6.2.1 Model penetrometers and hand vane apparatus

Similar to the field testing, various model penetrometers were employed in the first

centrifuge sample, comprising the cone, ball, plate with respective diameters of 10,

11.9, 11.2 mm, and the T-bar with dimensions 20 mm x 5 mm. The thickness of the

model plate is 1 mm. Photographs of the model cone and T-bar penetrometers are

shown in Figure 6.2. For the model T-bar penetrometer, the T-bar tip can be unscrewed

to replace it with a different penetrometer tip (i.e. a ball or plate).

In the second strongbox, four T-bars of different lengths: 20, 30, 40 and 50 mm, all with

the same diameter of 5 mm, were tested.

Unfortunately, no pore pressure transducer was integrated in the penetrometer rod used

(due to the small diameter of the rod), thus the pore pressures generated by the model

penetrometers during testing could not be measured.

Figure 6.3 shows photographs of the hand vane apparatus. The vane has a length and

width (diameter) of 28.6 and 19.1 mm respectively. Four blades with a thickness of

0.9 mm are mounted at right angles to a push-rod of 3 mm diameter. The perimeter ratio

and area ratio of the hand vane are 6 and 12.6 % respectively, both of which are slightly

higher than the corresponding ratios of the ‘standard’ vane used at field (Chandler,

1988; Geise et al, 1988). The consequence is that a higher degree of disturbance due to

vane insertion will be expected for the hand vane compared to the standard vane.

The measurement is taken manually from readings on the scale during the vane test. The

data recorded are then multiplied by a calibration factor of 1.346 kPa per division,

which was provided by the vane manufacturer, to obtain the values of shear strength.

CENTRIFUGE TESTING

6-3

6.2.2 Calibration details

The unequal area ratio, α, of the model cone penetrometer was calibrated in a small

calibration chamber filled with water, and a value of 0.859 was obtained.

As for the laboratory T-bar tests performed in the triaxial samples, the penetrometer rod

used in the centrifuge was examined to assess whether its load cell would be influenced

significantly by changes in the surrounding pore pressure (normal stress). It was found

that a change in normal stress of 100 kPa would give rise to an under-register of the

model T-bar (20 mm x 5 mm) resistance by around 19.63 kPa, which represents ~13 %

of the measurements. The error is smaller for penetrometers with larger projected areas.

Since the errors generated from changes in surrounding pore pressure were rather

significant, the resistance profiles measured for the full-flow penetrometers must be

adjusted accordingly. This was done by estimating the errors based on changes in static

pore pressure, uo, then adding the estimated errors to the corresponding measured

resistances. The amount of offset per 100 kPa change in uo for each model penetrometer

is summarised in Table 6.3. Nevertheless, it should be pointed out that, although the

errors should be estimated technically based on the total horizontal stress, σh, exerted on

the load cell, this value can be difficult to determine after insertion of the penetrometer.

In addition, σh may not be much higher than uo, particularly after insertion of the

penetrometer. Therefore, it is considered that estimating the error in the measurement

based on uo is reasonable.

6.3 Centrifuge testing procedure

6.3.1 Sample preparation

The bulk material collected from the site was screened through a sieve of 2.4 mm in

diameter to eliminate any large shell fragments. The sieved material was then

reconstituted at a water content of approximately 125 %. In an attempt to model an

idealised soil stratigraphy for the Burswood site, with the upper few metres below the

surface being lightly overconsolidated, merging to normally consolidated conditions,

both samples were first consolidated in a press under a vertical stress of 35 kPa. The

first sample was consolidated in the press for 7 days, whilst the second sample was

consolidated for 58 days, before they were transferred to the centrifuge for

consolidation at a centrifugal acceleration of 100 g. A final sample thickness of between

200 to 230 mm was targeted, which is equivalent to a prototype depth of 20 to 23 m.

CENTRIFUGE TESTING

6-4

6.3.2 Penetration and hand vane tests

The centrifuge testing layouts are shown in Figures 6.4 and 6.5 for the first and second

samples respectively. Both constant rate (‘normal’) penetration and variable rate

(‘twitch’) penetration tests were conducted. In a normal penetration test, the

penetrometer was pushed-in and pulled-out at a rate of 1 mm/s.

A ‘twitch’ test commenced by first pushing the penetrometer through the upper lightly

overconsolidated layer at 1 mm/s down to a depth of 60 – 110 mm; the rate was then

successively halved over 8 steps to a final value of 0.0078 mm/s. Each succeeding step

was triggered after the penetrometer had advanced by either 1 or 2 diameters. During

completion of the final step, the penetrometer was either pulled-out immediately at a

rate of 1 mm/s, or pushed in further at a rate reverted to 1 mm/s before being extracted

at the same rate.

Hand vane tests were conducted immediately after the sample was removed from the

centrifuge, and completed within an hour. It is believed that the vane tests were

performed before significant swelling of the sample occurred after ramping down the

centrifuge (since the material has a permeability, k, in the order of 10-7 mm/s and

consolidation coefficient of 0.5 − 0.75 m2/year). The vane was rotated at an approximate

rate of 1 revolution per minute (6 °/s) to measure the peak undrained shear strength.

This rotation rate is much higher than the rotation rate of 0.1°/s for a ‘standard’ vane

test (Chandler, 1988). Therefore, it is believed that the hand vane test will give higher

shear strength measurement compared to the standard vane; Watson et al (2000) noted

in their results of vane tests performed in the centrifuge, that increasing the rotation rate

will increase the peak shear strength. However, the effect of rotation rate is

compensated to some degree by the effect of disturbance generated from the vane

insertion, due to higher perimeter ratio and area ratio for the hand vane compared to the

standard vane, as mentioned earlier.

Remoulded shear strengths were also obtained for some depths, by turning the vane for

ten rotations prior to recording the measurement (ASTM, 2000a; Geise et al, 1988).

6.3.3 Laboratory testing on centrifuge samples

Several tube samples with diameter of 100 mm and height of 200 mm were retrieved

from the strongbox samples after the centrifuge tests were completed (see Figures 6.4

and 6.5). These tube samples were used for laboratory testing to evaluate the soil

CENTRIFUGE TESTING

6-5

properties of the reconstituted clay samples. The laboratory tests undertaken included

the CRSC, CAU triaxial (compression and extension), CAU simple shear and T-bar in

triaxial tests.

In the CRSC tests, the samples were trimmed to 61 mm in diameter and 25 mm in

height, and then loaded at a constant displacement rate of 0.003 mm/min (strain rate of

2 x 10-6 s-1).

For the triaxial compression, triaxial extension and the simple shear tests, the samples

were cut into heights of 190, 175 and 40 mm respectively, all with a diameter of

100 mm. The samples were re-consolidated to a stress level similar to that experienced

at the base of the sample during the centrifuge tests. The vertical effective stress was

estimated using the effective unit weight listed in Table 6.1, multiplied by the prototype

depth of the sample, taking zero stress at the top surface of the sample. A Ko value of

0.8 was adopted to compute the horizontal effective stress. The triaxial test samples

were then loaded at 0.2 mm/min (strain rate of 18 − 20 x 10-6 s-1) in both compression

and extension tests, whereas the simple shear samples were sheared at 0.1 mm/min

(shear strain rate of 45 x 10-6 s-1).

The T-bar in triaxial test was only conducted on a sample retrieved from the first

centrifuge strongbox sample. The test sample had a height of 190 mm and diameter of

100 mm. Since only isotropic consolidation is allowed in this set up, the sample was

consolidated to the mean effective stress, p', calculated using Equation 5.3. Again, the

effective vertical stress used in the calculation was the estimated vertical stress

experienced in the centrifuge at the sample base. However, a Ko value of 0.6 was

adopted. After consolidation, the T-bar was pushed-in and pulled-out at a rate of

50 mm/min.

6.4 Centrifuge test results

6.4.1 Consolidation in centrifuge

Both strongbox samples were allowed to consolidate in the centrifuge for about 209

hours (almost 9 days), before model penetration tests were performed. Figures 6.6 and

6.7 plot the settlement versus root time in minutes during consolidation for the first and

second samples respectively. The time for 90 % degree of consolidation (t90) has been

estimated using Taylor’s root time method, as indicated in the figures. This gave t90

values of about 201.7 and 144.2 hours (or 8.4 and 6 days) for the first and second

CENTRIFUGE TESTING

6-6

samples respectively. Using these data, average values for the consolidation coefficient,

cv, for both reconstituted samples can then be estimated from:

90

290

v tdTc = (6.1)

where: T90 = time factor, taken as 0.848;

d = length of the drainage path.

These led to cv values of approximately 0.5 and 0.75 m2/year for the first and second

samples respectively, as noted in Table 6.1.

6.4.2 Assessment of model penetrometer tip resistance

Due to the inability to measure pore pressure on the model penetrometers during testing,

the measured cone resistance, qc, is corrected to its net value using the following

equation (Robertson & Campanella, 1983; Watson et al, 1998):

q

ovccnet B)1(1

)u'(qqα−−

α+σ−= (6.2)

where: σ'v = estimated effective vertical stress;

uo = estimated hydrostatic water pressure;

α = unequal area ratio (in this case 0.859);

Bq = ratio of the excess pore pressure to the net bearing pressure.

Since no pore pressure was measured here, the value of Bq is estimated from the field

cone testing. It may be seen in Figure 4.14 that the Bq values from the field cone tests

fluctuate at about 0.45 and 0.17 (absolute) during penetration and extraction

respectively. So, these values were adopted in the correction for the model cone

resistance.

However, in a twitch test for the cone penetrometer, the Bq value will be reducing as the

rate is decreased in each successive step, because the conditions are gradually becoming

more partially drained. Therefore, the value of Bq is assumed to decrease evenly in each

succeeding step, to 50 % of its initial value in the final step. This assumption is based on

the Bq results plotted against the non-dimensional velocity, V, shown by Randolph &

Hope (2004). During extraction, the same value of 0.17 was used. Note that, since the

unequal area ratio, α, is relatively high (quite close to unity), the net cone resistance is

CENTRIFUGE TESTING

6-7

in fact relatively insensitive to the Bq value adopted.

The tip resistances for the model T-bar, ball and plate (full-flow) penetrometers were

not corrected, since the areas of these penetrometers are much larger than the area of the

penetrometer rod, so that the correction will be insignificant.

6.4.3 Resistance profiles for various model penetrometers

Results of the penetration tests on the first strongbox sample (denoted as Box 1) are

presented in Figures 6.8 to 6.14, whilst results for the second strongbox sample (Box 2)

are shown in Figures 6.15 to 6.21. Note that the resistance profiles are plotted against

the equivalent prototype depth, where 1 mm in the centrifuge represents 0.1 m at

prototype scale.

(a) Box 1: Model T-bar tests

Figure 6.8 shows the tip resistances of all model T-bar (20 mm x 5 mm) tests performed

in the first sample. It can be seen in Figure 6.8 (a) that all normal T-bar tests gave

similar penetration resistances. After the initial increase in resistance, the results show a

nearly constant resistance, with a value of around 90 kPa, to a depth of about 6 m. This

is due to the sample having been pre-consolidated at 35 kPa before it was transferred to

the centrifuge, which created a lightly overconsolidated layer. Below 6 m, the

resistances increase almost linearly to around 330 kPa at a depth of about 18 m,

indicating normally consolidated material. It may also be noticed that there is a small

(local) peak in the penetration resistance at a depth of about 4 m, which is thought to be

due to a slight layering effect resulting from the topping up of clay slurry during the pre-

consolidation phase.

For clarity, the penetration resistance profile for the T-bar twitch test is shown

separately in Figure 6.8 (b), together with the comparison tests (T-bars 4 and 6). The

end of each step of the twitch test is also indicated. It may be seen that the twitch test

initially shows similar (but slightly lower) penetration resistance to the normal tests

until step 5, from where the twitch test starts to show resistance increasing more

strongly than the normal tests. At the end of step 8, resistance of the twitch test is

approximately 78 % higher than those of the normal tests.

Also, as the penetration rate of the twitch test reverted to 1 mm/s at completion of step

8, the resistance continued to increase over a short distance before reducing rapidly,

CENTRIFUGE TESTING

6-8

forming a peak of resistance between depths of 14 and 16 m. The resistance then slowly

merged with the normal test results. This phenomenon has also been reported by House

et al (2001).

Results for the extraction stages of each test, including the twitch test, are presented in

Figure 6.8 (c). Extraction resistance profiles for the normal tests are very consistent.

Similar features as mentioned above for the penetration profiles are also observed for

the extraction profiles.

For the twitch test, in addition to the peak similar to that seen in its penetration profile, a

trough is also seen in the extraction profile at a depth of around 12 m. This may be due

to a (partial) physical gap (water filled) forming above the penetrometer during slow

penetration of the T-bar. However, above this depth, the extraction resistance of the

twitch test stays above the resistance profiles of the normal tests. A possible reason is

that soil above the T-bar had gained some strength from local consolidation occurring

during slow penetration, or possibly that stronger soil (resulting from partial

consolidation) adhered to the T-bar during extraction.

Tip resistance profiles of model T-bars 4 and 6 will be used for comparison with

resistance profiles for the other penetrometers, because the two profiles are deemed

most representative of the average profiles of all the model T-bar tests.

(b) Box 1:Model cone and T-bar tests

The measured cone resistance and its net value (qc and qcnet respectively) are plotted in

Figure 6.9. As mentioned earlier, a Bq value of 0.45 was used for the correction during

penetration (Figure 6.9 (a)). It may be seen that the qc for both model Cones 1 and 3

have been greatly reduced, by about 50 %, after being corrected to qcnet. While Cone 3

shows a slightly lower resistance profile than Cone 1, both cones exhibit lower qcnet

profiles than the model T-bar profiles. Profiles of qcnet have also been assessed for

Bq = 1, with the intention of examining the sensitivity of the qcnet to the Bq value. The

results show that the qcnet values for both cones are increased by about 9 %, by changing

the Bq value from 0.45 to 1.

During extraction, the measured resistances of Cones 1 and 3 are very similar and

generally remain positive, because of the overburden and pore pressures acting upwards

on the cone tip, but after correction the net cone resistances follow closely the extraction

profiles of the T-bars for depths above 12 m. It is interesting to note that the net cone

CENTRIFUGE TESTING

6-9

resistances develop gradually and smoothly as the cones are first extracted, in contrast

with the very rapid development of full resistances for the T-bars. Similar observations

have also been noted by Watson et al (2000). Also, the extraction qcnet profiles

computed using Bq = 1 are approximately 14 % higher than those computed with

Bq = 0.17. These results confirm that the value of qcnet is relatively insensitive to

changes of the Bq value both during penetration and extraction.

The twitch test of the cone penetrometer demonstrated penetration resistance increasing

in a similar manner as shown in the T-bar twitch test, but during extraction, instead of a

trough as noted for the T-bar test, the cone resistance decreases gradually and remains

higher than the normal test profiles of Cones 1 and 3 until a depth of 6 m.

(c) Box 1: Model ball and T-bar tests

Results of the model ball penetration tests are plotted in Figure 6.10. It can be seen that

both the penetration and the extraction resistances of model Balls 1 and 3 follow the

model T-bar profiles closely, although Ball 1 exhibits a slightly lower penetration

resistance in the top 5 m. In addition, the ball resistance profiles show relatively less

obvious (local) peaks around the 5 m depth, as evident in the T-bar profiles. One

possible reason may be due to the greater size (diameter) of the ball compared to the

T-bar, thus smearing out the effects of any localised stronger soil at the interface.

For the twitch test, the ball was extracted immediately after the final step was

completed. Consequently, no peaks in either the penetration and extraction resistance

profiles were observed. Also, it may be noted that the extraction resistance for the ball

twitch test was not recorded between depths of 14.5 and 15.5 m. This was due to the

rate of data logging during slow penetration not being changed back to the rapid rate of

logging as the ball was first extracted at 1 mm/s.

(d) Box 1: model plate and T-bar tests

Figure 6.11 presents results for the model plate penetrometer. Plates 1 and 3 gave

penetration resistances up to 25 % lower than the model T-bar results (Figure 6.11 (a)).

This finding is unexpected and contrary to the field test results, where the field plate

showed higher resistance than the field T-bar (see Chapter 4). However, during

extraction (Figure 6.11 (b)), Plate 3 shows marginally higher resistance than the T-bars

for much of the depth, while Plate 1 gives similar or slightly lower extraction resistance

than the T-bars. Interestingly, the plates show the same local peak in resistance at 4 m

CENTRIFUGE TESTING

6-10

depth as the T-bars. This may be due to the fact that, even though the plate has a

diameter similar to the ball, the thickness of the plate is very small (even less than that

of the T-bar), and thus the plate may be more sensitive to the layering effect than the

ball.

For the twitch test, as was the case for the ball, the plate penetrometer was extracted

immediately after completion of the final step, and hence no peak of resistance is

observed.

(e) Summary of Box 1 penetration test results

A summary plot of the penetration and extraction resistance profiles for all model

penetrometers tested in the first strongbox sample is shown in Figure 6.12. The average

resistance profile of the field T-bar (250 mm x 40 mm) is also included for comparison.

It is clear that the model cone penetrometer exhibits the lowest net resistance, in

contrast to the field results. This finding may be affected by the uncertainty in

correcting the model cone resistance, but more probably reflects higher rigidity index

for the natural clay, leading to higher cone resistance for a given shear strength (Teh &

Houslby, 1991). It is also interesting that the model plate penetrometer shows relatively

lower penetration resistance compared to the model T-bar and ball penetrometers, while

the reverse is true in the field testing.

During extraction, all model penetrometers show very similar, but lower resistances

compared to the field T-bar test, although the difference is less than that during

penetration. This suggests that the rate of degradation in resistance (and possibly

sensitivity) is lower for the reconstituted sample than for the natural soil in the field.

Ratios of the extraction to penetration resistances, qout/qin, for all model T-bar (20 mm x

5 mm) tests are plotted in Figure 6.13, whilst a summary of the resistance ratios for the

various shaped model penetrometers is presented in Figure 6.14. It may be seen that the

individual model T-bar tests show some scatter at depths less than 8 m, below which the

ratios converge to an average value of around 0.65 at a depth of 17 m. The average

result of the model ball shows similar (but ~5 % higher in average) ratio profile to the

model T-bar result (averaged from T-bars 4 and 6). Nevertheless, the ratio profiles of

these two model penetrometers are somewhat higher than ratio profiles of the

corresponding field penetrometers, which fluctuate around 0.6 and 0.55 for the field ball

and T-bar penetrometers respectively.

CENTRIFUGE TESTING

6-11

The model plate shows a somewhat irregular curve of qout/qin for depths above 8 m and

arguably a plateau at around 0.88 between depths of 4 and 8 m. Below 8 m, the ratio

qout/qin decreases rather steadily to about 0.67 at a depth of 16 m.

Results obtained for the model cone penetrometer are very unusual, as the ratios exceed

unity between depths of 5 and 12 m.

(f) Box 2: Model T-bar (20 mm x 5 mm) tests

Figure 6.15 presents test results for all model 20 mm x 5 mm T-bars tested in the second

strongbox sample (Box 2). As may be seen, the sample shows a lightly overconsolidated

layer to a depth of 9 m, below which the material is normally consolidated. Note that,

this depth of transition from overconsolidated to normally consolidated is apparently

greater than that for the first strongbox sample (Box 1) observed earlier. Also, it may be

noted that the penetration resistance profiles for the T-bar tests are higher in Box 2 than

in Box 1. This can be seen more clearly when the results are compared directly in term

of shear strength profile, which will be presented later. These differences are due to the

much longer laboratory-floor consolidation period for Box 2 (58 days) than for Box 1 (7

days).

As was the case with the first sample, the resistance profiles also show consistent local

peaks at 8 m depth, which is thought to be due to a slight layering effect caused by

topping up the clay.

(g) Box 2: Effect of model T-bar aspect ratio

Tests in the second sample were aimed at exploring the effect of various aspect ratios,

L/d, of the T-bar penetrometer. A range of L/d values (4, 6, 8 and 10) were studied and

the results are presented in Figures 6.16 to 6.18.

Evidently, the overall results do not show any obvious effect of the various aspect ratios

on the model T-bar tip resistance. The penetration results for T-bars 30 mm x 5 mm and

40 mm x 5 mm (Figures 6.16 (a) and 6.17 (a)) show similar resistances to those for

T-bar 20 mm x 5 mm. Although T-bar 50 mm x 5 mm (Figure 6.18 (a)) does exhibit

slightly higher resistances than for T-bar 20 mm x 5 mm over much of the depth, this is

believed due to other effects, such as a locally stronger region of the soil sample or a

slight bending moment on the load cell.

During extraction, all different sized model T-bars show similar resistances (Figures

CENTRIFUGE TESTING

6-12

6.16 (b) to 6.18 (b)). It is interesting that there is much more scatter of T-bar results

during penetration than extraction, again possibly due to slight bending effects during

the former (results getting worse as the aspect ratio increases).

Note that twitch tests were also carried out for each of these various sized T-bars,

although their results are not included in the figures, because they show similar trends

of resistances as has been seen for the T-bar twitch test conducted in the first sample.

However, further interpretation of the twitch test results will be presented in Chapter 8.

(h) Summary of Box 2 penetration results

A summary plot of the penetration and extraction resistances for all model T-bars of

various aspect ratios is shown in Figure 6.19, including average resistance profile of the

field T-bar (250 mm x 40 mm) penetrometer. It is clear that the longest model T-bar

demonstrates slightly higher penetration resistance than other model T-bars, the reason

for which has been mentioned above, and all model T-bar penetration resistances are

lower than the field T-bar penetrometer, although the difference is smaller compared to

that observed in the first sample, due to higher strength of the second sample. Also, it

may be noted in Figure 6.19, that the difference between the model and field T-bar

resistances during penetration is less in the overconsolidated region (above 9 m), with

the model T-bar resistances being about 20 % lower than the field T-bar result,

compared to that in the transition region from overconsolidated to normally

consolidated, where the model T-bar resistances are up to 32 % lower than the field

resistance at 9.5 m. Below this depth, the model T-bar resistances increase slightly more

rapidly than the field T-bar resistance, and merge with the field result at a depth of

about 17 m.

During extraction, the centrifuge results follow very closely the field test profiles,

except for the model T-bar 50 mm x 5 mm showing marginally higher extraction

resistance at depths below 9 m.

Profiles of the ratio qout/qin of model T-bar 20 mm x 5 mm are plotted in Figure 6.20,

while profiles (average values) of the other model T-bars are presented in Figure 6.21.

Also included in Figure 6.21 are the average ratio profiles of the field standard and

smaller T-bars, and the model T-bar 20 mm x 5 mm from the first strongbox sample.

Interestingly, the model T-bar results from both strongbox samples are tightly bunched

(fluctuating between 0.62 and 0.72), and are embraced by the field results (but much

CENTRIFUGE TESTING

6-13

closer to the field smaller T-bar curve).

6.4.4 Hand vane tests

Figures 6.22 and 6.23 illustrate the results of hand vane tests in the first and second

strongbox samples respectively. In the second sample, the remoulded strength has also

been measured at some depths. The tests give repeatable results of both the peak and

remoulded shear strengths.

Average shear strength profiles from the hand vane tests are summarised in Figure 6.24,

and the sensitivity profile of the reconstituted clay is shown in Figure 6.25. In addition,

both the strength and sensitivity profiles of the field vane test are also included for

comparison in the figures. It can be seen that the first sample exhibits a lower peak

shear strength profile (20 – 40 %) than the second sample, but the peak strength profile

from the second sample is still slightly lower than that of the field tests, although the

two appear to merge below a depth of 14 m. This all seems consistent with the T-bar

results.

Interestingly, the remoulded strength results from the second sample seem to lie within

the zone of remoulded strength results from the field vane test, which is consistent with

the similarity between penetrometer extraction resistances from the centrifuge and field

tests.

It can be seen that the sensitivity data of the centrifuge samples are typically between

2.7 and 3. These values are generally lower than the sensitivity measured for the natural

clay (3 to 4). This is nevertheless consistent with the model penetrometers showing

higher ratios of qout/qin than the field penetrometers.

6.4.5 Laboratory testing on centrifuge samples

Results of the CRSC tests on samples retrieved from the centrifuge test samples are

summarised in Table 6.4 and the individual test results can be found in Appendix F. It

should be noted that the depth range shown in the table refers to the prototype depth

range of the sample in the centrifuge testing. Also, it is noteworthy that the depth range

is rather wide, since the sample height of 25 mm required for the CRSC test would

represent a prototype depth range of 2.5 m. For example, the sample for test

CF1 CRSC1 (see Table 6.4) has a prototype depth of 3.3 m at its top surface and 5.8 m

at the base. For this reason, ranges of estimated effective vertical stress and OCR in the

CENTRIFUGE TESTING

6-14

centrifuge at the top and bottom of the sample are shown in Table 6.4.

The consolidation coefficients obtained at the yield stress levels (cvy) lie mostly within 1

to 1.2 m2/year, while the cv values obtained at 5 times yield stress lie between 0.5 and

0.7 m2/year. The later cv values are consistent with those values derived from the

consolidation data of the strongbox samples in the centrifuge. Further comparison and

discussion of the values of cv will be presented in Chapter 8.

Table 6.5 presents a summary for the results of the triaxial (compression and extension)

and simple shear tests, while their individual test results are plotted in Appendix G.

Again, the prototype depth ranges are shown in the table. The consolidation vertical and

horizontal stresses refer to the estimated stresses experienced in the centrifuge at the

base of the sample. The measured undrained shear strengths will be plotted against the

prototype depths and compared with the strength profiles deduced from the model

penetration tests in the following section.

One T-bar test in a triaxial sample was carried out on a tube sample recovered from the

first centrifuge sample. The results from this test are illustrated in Figure 6.26. The

prototype depth at the base of the sample was about 19 m and the sample was

consolidated isotropically to its mean effective stress, p', of approximately 61 kPa. It

may be seen in Figure 6.26 (b) that the results give reasonably constant penetration and

extraction resistances of about 0.25 and –0.21 MPa respectively, hence giving a

resistance ratio (qout/qin) of about 0.84. Interestingly, this ratio is considerably higher

than the qout/qin ratios shown by the model T-bar tests performed in the centrifuge

(Figures 6.13 and 6.21), and very much higher than the ratios obtained from the similar

tests conducted on the in situ samples, which demonstrated very low ratios between 0.2

and 0.31, except for one test recording value of 0.72 (Table 5.7).

6.4.6 Profiles of undrained shear strength

Estimates of the undrained shear strength, su, may be made from the model penetration

test data using Equation 4.6 shown in Chapter 4. However, since correction for the

measured tip resistances of the full-flow penetrometers are deemed insignificant and

pore pressure data required for the correction are unavailable, the measured resistances

have been taken as qnet in Equation 4.6. As for the field penetrometers, a single value of

N = 10.5 is used for all model penetrometers and further discussion of the appropriate

value of N will be presented in Chapter 7.

CENTRIFUGE TESTING

6-15

The resulting profiles of undrained shear strength are illustrated in Figures 6.27 and

6.28 for the first and second centrifuge samples respectively. Results of the hand vane

tests and laboratory tests have also been included in the figures. It should be noted that

the laboratory test data are plotted against the prototype depths at the base levels of the

samples.

It is obvious in Figure 6.27 that the su data measured from the hand vane tests in Box 1

appear to agree well with the su profiles deduced from the model T-bar and ball

penetration tests. In Box 2, the hand vane tests and the T-bar penetration tests gave

similar su values in the lightly overconsolidated region (depths above 9 m), but in the

normally consolidated region, the su values of the former tests are about 10 % higher

than the strength profiles of the latter tests (Figure 6.28). The results from Box 2 are

somewhat consistent with the findings noted by Watson et al (1998, 2000), who

conducted vane tests in-flight in the centrifuge, on both overconsolidated and normally

consolidated fine-grained materials. They reported su values being about 20 % higher

for the vane with aspect ratio, L/d, of 1.5 (the same as the hand vane) compared to the

T-bars.

In Figure 6.27, it can be seen that the su value obtained from the simple shear test at a

depth of 10.8 m (Box 1) lies significantly above the su profiles for the model penetration

and hand vane tests. In fact, the result from this particular simple shear test is

considered suspect, perhaps due to a machine fault occurring during testing, since the

stress-strain curves for both simple shear tests from Box 1 showed significant

irregularities (see Appendix G, Figures G.5 and G.6). By contrast, other simple shear

tests from Box 2 show much smoother stress-strain curves and the strength values plot

close to the su profiles from the model penetration tests, as illustrated in Figure 6.28.

Results of the triaxial compression tests in both samples indicate slightly lower su values

than from the simple shear tests at the corresponding depths, while the extension tests

give the lowest su values. The low triaxial compression strengths may be partly due to

the limited consolidation period for the sample at the elevated stress level (dictated by

the previous stress experienced at the sample base). Typical consolidation periods were

just over 100 hours (4 days), compared with a centrifuge consolidation period

(including stopping and starting) in excess of 10 days. As such, less secondary

consolidation would have occurred for the triaxial samples compared to the simple shear

samples, as the latter samples were much smaller in height and hence the average

CENTRIFUGE TESTING

6-16

increase in consolidation stress was correspondingly much lower.

The laboratory T-bar test from Box 1 yielded an undrained shear strength that is lower

than the model penetration and hand vane test results, which is in contrast with similar

tests conducted on the in situ tube samples (see Figure 5.14). This may also be

attributed to the limited consolidation period for the centrifuge sample (as mentioned

above for the triaxial tests) and the absence of shell fragments, since the material was

sieved before being reconstituted for the centrifuge testing. Nevertheless, the strength

value of the laboratory T-bar test appears to lie midway between that from the triaxial

extension and compression tests.

Figure 6.29 compares the su profiles deduced from the T-bar (20 mm x 5 mm) tests in

both centrifuge samples. The effect of ‘aging’ or secondary consolidation is evident, as

the second sample shows a greater strength than the first sample, particularly between 3

and 9 m, with a deeper transition depth from lightly overconsolidated to normally

consolidated. These differences are due to the much longer laboratory-floor

consolidation period for Box 2 (58 days compared with 7 days for Box 1), allowing

more secondary consolidation to occur in addition to full primary consolidation.

However, despite the longer consolidation period for Box 2, its strength profile is still

lower than the field strength profile due to the innate structure of the undisturbed

material in the field.

6.5 Summary for centrifuge testing

This chapter has presented results of the centrifuge testing on reconstituted samples of

clay collected from the Burswood site. The findings are summarised as follows.

The T-bar and ball penetrometers have exhibited similar tip resistances both during

penetration and extraction, despite the plane strain soil flow during passage of the T-bar

penetrometer and the axisymmetry soil flow for the ball penetrometer. Results from the

field tests also support this observed agreement. The model cone and (unexpectedly)

plate penetrometers demonstrated lower penetration resistances than the model T-bar

and ball. However, results of the cone penetrometer are abnormal and showed ratios

qout/qin that exceed unity at some depths.

The tests on model T-bars with aspect ratios between 4 and 10 do not seem to show any

obvious effect on the tip resistance. Therefore, any aspect ratio in the range 4 to 8 for

the T-bar penetrometer would be appropriate.

CENTRIFUGE TESTING

6-17

All variable rate (twitch) penetration tests gave higher penetration resistances (at low

penetration rates) and extraction resistances than the corresponding constant rate

penetration tests. Further discussion of the penetration rate effect is presented in

Chapter 8 later.

The undrained shear strength profile obtained from the hand vane tests in the first

strongbox sample are shown to be similar to the strength profiles derived from the T-bar

and ball penetration tests using a single value of N = 10.5, whereas the hand vane tests

in the second sample gave a higher shear strength profile compared to those deduced

from the T-bar tests, particularly in the lightly overconsolidated region. Results from the

second sample are somewhat consistent with the findings noted by Watson et al (1998,

2000), although the average difference in su values for the two testing methods appeared

to be less in Figure 6.28, compared to that reported by Watson et al (1998, 2000).

Laboratory tests conducted on samples recovered after the centrifuge testing may have

been affected by: (a) the limited period of consolidation for the triaxial samples; and (b)

mechanical problems in the simple shear results for the two samples from Box 1, which

led to irregular stress-strain curves. The most reliable results were obtained from the

four simple shear tests conducted on samples from Box 2, which gave shear strengths

that agree very well with the strength profiles derived from the model T-bar tests, using

the same N factor of 10.5.

It is noteworthy that the T-bar test conducted on a triaxial sample from Box 1 gave an

average shear strength that lies midway between that of the triaxial extension and

compression tests. This implies that, while the absolute strengths are low because of

limited secondary consolidation of the triaxial samples, the relative values support the

choice of N ~ 10.5 to obtain shear strengths from T-bar tests that are consistent with the

average laboratory strengths.

The importance of ‘aging’ or secondary consolidation effects has been shown, as the

second centrifuge sample showed a higher shear strength and deeper transition depth

(from lightly overconsolidated to normally consolidated) than the first sample, due to

the much longer laboratory-floor consolidation period for the second sample compared

with the first one.

CORRELATION OF PENETRATION RESISTANCE AND UNDRAINED SHEAR STRENGTH

7-1

7 CORRELATION OF PENETRATION RESISTANCE AND UNDRAINED

SHEAR STRENGTH

Results of the field, laboratory and centrifuge tests are presented successively in

Chapters 4, 5 and 6. Undrained shear strength (su) profiles have been derived from

penetration resistances of the various shaped penetrometers, using a single bearing

factor, N, of 10.5 (the value originally recommended by Stewart & Randolph (1991,

1994) for a T-bar penetration test) for comparison with strength data measured from the

laboratory and vane shear tests. This chapter in turn examines the ranges of N for the

various shaped penetrometers back-calculated using the laboratory and vane shear

strengths.

Due to strength anisotropy, it is necessary to specify which reference su is being used for

computing N (Aas et al, 1986; Lunne et al, 1997b; Tanaka & Tanaka, 2004). The

reference su used in the correlations includes the vane shear strength, su, vane, the CAU

triaxial compression strength, su, TXC, the CAU simple shear strength, su, SS and the

average laboratory strength, su, av = (su, TXC + su, TXE + su, SS)/3, where su, TXE is the CAU

triaxial extension strength. However, since the laboratory test samples were not

extracted from the same depth levels, a linear trendline is fitted to each set of the

compression, extension and simple shear strength data (excluding the suspect data

reported in Chapter 5), in order to compute the values of su, av at depths corresponding to

the compression strength data (Figure 7.1). For convenience, values of the laboratory

shear strengths used in the correlations are summarised in Table 7.1.

The following sections first investigate the bearing factors for the cone, then for the

‘full-flow’ penetrometers (T-bar, ball and plate). In addition to the bearing factors

derived for the Burswood clay, empirical T-bar and cone factors for two other soft clay

sites collected in a joint NGI-COFS project are also obtained for comparison.

Furthermore, the theoretical cone factor is evaluated for each site and the resulting

values are compared with the empirical data. Finally, recommendations for the choice of

N values, as a general guide, to be adopted for the various shaped penetrometers are

provided. Limitations and some factors that may cause the N values to fall outside the

suggested range are addressed.

7.1 Cone factor, Nkt

As has been pointed out in Chapter 2, there are three different cone factors often used


7-2

for correlations with the undrained shear strength, su. However, only Nkt, defined as net

cone resistance, qcnet, divided by su, is considered in the thesis (see Equation 2.3 in

Chapter 2).

Figure 7.2 presents the profiles of cone factor derived from the laboratory shear

strengths. Two sets of data of net cone resistance have been obtained for calculating the

cone factors: (a) from the average net resistance profile of Cones 3 and 4 presented in

Figure 4.17 (Chapter 4); and (b) from the net cone resistance profile (BSCT01) reported

by Schneider et al (2004). For ease of reference, the cone factors calculated from the

former cone resistance profile are denoted as Cone-(a), whereas the results deduced

from the latter profile are represented by Cone-(b).

Some scatter is clearly observed within each set of Nkt data deduced from the different

reference su, for both Cones-(a) and (b), particularly where the simple shear strength is

the reference strength. In order to assess and compare the dispersion and variability of

the empirical data, standard deviation and coefficient of variation are evaluated for each

set of the Nkt data. The standard deviation, SD, can be calculated as:

1n

)NN(SD

n

1i

2i

−

−=

∑= (7.1)

where: n = number of data;

Ni = data of bearing factor;

N = mean value of bearing factor data.

The coefficient of variation, COV, is the ratio of standard deviation to mean value,

expressed in percentage.

For Cone-(a), it may be seen in Figure 7.2 that the Nkt, TXC (= qcnet/su, TXC) values range

between 9.5 and 10.9, giving a mean value of 10.3. The standard deviation evaluated is

0.8 and the coefficient of variation is 7.6 %. The values of Nkt, av (= qcnet/su, av) lie

between 12.6 and 13.6, with a mean value of 13.2 (SD = 0.4 and COV = 3.4 %). The

data of Nkt, SS (= qcnet/su, SS) demonstrate the largest degree of scatter, between 13.2 and

16.5, with an average of 14.6 (SD = 1.5 and COV = 10.3 %), although three data points

between the depths of 9 and 13 m are very similar (~13.3).

The values of Nkt for Cone-(b) are lower than the corresponding values for Cone-(a),


7-3

because of lower cone resistances for Cone-(b) than for Cone-(a). Besides, the Nkt, TXC

and Nkt, av data for Cone-(b) are slightly more scattered than those for Cone-(a), with the

Nkt, TXC values for the former ranging from 7.6 to 10.0 (mean value being 8.5 ± SD = 1.1

and COV = 12.4 %), and the Nkt, av values lying between 10 and 12.4 (mean value being

10.9 ± SD = 1.1 and COV = 10.2 %). The Nkt, SS values for Cone-(b) vary from 10.1 to

12.8, with an average of 11.7 (SD = 1.1, COV = 9.3 %).

Also included in Figure 7.2 is the range of theoretical Nkt, represented by the dashed

lines, calculated using the following approximate expression (Lu et al, 2004):

( ) crkt 3.19.1Iln6.14.3N α+Δ−+≈ (7.2)

where: Ir = rigidity index;

Δ = in situ stress ratio;

αc = roughness coefficient for the cone-soil interface (varies from 0 to 1).

The rigidity index of the soil is the ratio of shear modulus to shear strength, Ir = G/su,

and the in situ stress ratio is defined as difference between in situ vertical and horizontal

stresses divided by two times the shear strength, Δ = (σvo − σho)/2su (Teh & Houlsby,

1991). Schneider et al (2004) reported, for Burswood, a small strain shear modulus (Go)

profile deduced from the seismic cone test data varies from about 4.5 MPa at a depth of

5 m to 13 MPa at 17 m. Although it may not be appropriate to calculate Ir using Go,

since this will give too high a value of Ir for large strain problems such as cone

penetration, it provides a consistent basis for comparison between sites. Based on the

reported Go and the average laboratory shear strengths, the resulting rigidity indices lie

mainly between 300 and 400. The in situ stress ratios, Δ, deduced based on a Ko value

of 0.8 and the average laboratory strengths, range from 0.22 to 0.24. Therefore, for a

typical roughness coefficient, αc, of 0.3 (Lu et al, 2004; Randolph, 2004), the theoretical

cone factor calculated using Equation 7.2 varies between 12.5 and 13.0 for the

Burswood clay.

It is apparent that the empirical Nkt data fall mostly outside the range of the theoretical

values, although the Nkt, av data for Cone-(a) are reasonably close, and three data points

of Nkt, SS from Cone-(a) plot marginally above the upper range of the theoretical Nkt.

Besides the effects of sample disturbance on the measurements of the shear strength, the

discrepancy between the empirical and theoretical cone factors is mainly due to the

theoretical solution not taking into account the effects of strain rate, strain softening and


7-4

strength anisotropy of the natural soil, each of which factors can significantly influence

the shear strength, and thus the cone factor. Therefore, a wider theoretical range is to be

expected once these factors are incorporated in the solution. At this stage, even though it

is not possible to isolate the effects of each factor, studying the ranges and mean values

of the different Nkt may, however, give some broad indications as to what combined

effects these factors would have on each set of the Nkt data. Of course, the effects of

sample disturbance can be, and should be, minimised by ensuring the test samples are of

high quality.

The cone factors calculated using the vane shear strengths, Nkt, vane (= qcnet/su, vane), are

presented in Figure 7.3. The Nkt, vane data obtained from the centrifuge testing are also

plotted in the figure. Interestingly, both field cones tend to show general trends of

Nkt, vane increasing with depth, while the centrifuge model cone shows Nkt, vane initially

decreasing with depth, before increasing, from a depth of 7 m. In addition, the profile of

Nkt, vane for the model cone plots well below the profiles of the field cone tests, with a

range of 5.5 to 8.6 and a mean value of 7.1 (SD = 1.1 and COV = 15.5 %). It should be

pointed out that although the empirical results show Nkt, vane varying with depth, there is

no particular basis for Nkt, vane to vary with depth (or vertical stress). Therefore, at this

stage, it has been assumed that Nkt, vane (and other N factors) does not vary with depth,

and the values of SD and COV are calculated based on this assumption.

In Figure 7.3, the Nkt, vane data for Cone-(a) vary from 10.0 to 14.6, giving a mean of

12.5 (SD = 1.5 and COV = 12.2 %). It can be noted that the theoretical range lies within

the mean ± one SD (11 − 14) of Cone-(a) results. The Nkt, vane values for Cone-(b) lie

between 8.6 and 11.8, with an average of 10.3 (SD = 1.0 and COV = 9.3 %).

For ease of comparison, the range and mean value (± one standard deviation) of each set

of Nkt are summarised in Table 7.2. In addition, summary of the empirical cone factors

for two other clay sites collected in the joint NGI-COFS project (NGI-COFS, 2004a) are

also quoted in the table. One of the sites is located onshore in Norway (Onsøy) and the

other site is offshore Australia (Laminaria). Both sites are underlain with lightly

overconsolidated clays with plasticity indices rather similar to the Burswood clay (IP

mainly between 30 and 50 %).

As for Burswood, theoretical Nkt values for the Onsøy and Laminaria clays are also

computed using Equation 7.2. The shear modulus for Onsøy was measured by bender

elements, but only one measurement, with a value of 22.1 MPa, was reported (NGI-


7-5

COFS, 2004b). Based on the triaxial compression strength (since corresponding

extension and simple shear strengths are not available), the rigidity index deduced is

around 630. For Laminaria, the profile of shear modulus derived from triaxial

compression tests increases from about 8 MPa at a depth of 2 m to around 30.5 MPa at

20 m (Woodside, 1997). Using the strength profile averaged from the penetration and

laboratory test results, the rigidity index profile varies from 550 to 660. Then, assuming

Δ = 0.24 for lightly overconsolidated clay for both sites (similar to Burswood), the

theoretical Nkt value computed for Onsøy is 13.6 and that for Laminaria varies from

13.4 to 13.7.

It may be noted in Table 7.2 that the empirical Nkt data for the three different sites, and

the reconstituted Burswood sample, show different ranges and mean values. It should be

pointed out that the laboratory tests for Onsøy were conducted on high quality block

samples, which were believed to have minimal sample disturbance. Yet, the Onsøy clay

appears to show the widest range (as well as highest COV) of Nkt, TXC. The range of

Nkt, av is also greater for Onsøy than for Burswood (for Cone-(a)). These may be an

indication of greater variations of soil properties at the Onsøy site. Additionally, all four

sets of Nkt for Onsøy are relatively higher than for Burswood. This is, however,

consistent with the theoretical estimations indicating values for the Onsøy clay being

higher than for the Burswood clay, although to a lesser extent compared to the empirical

results. On the other hand for the Laminaria clay, while its theoretical values are higher

than for Burswood, the empirical Nkt ranges and mean values are more similar or

slightly lower (except for Nkt, TXC) than the Burswood results (considering Cone-(a)).

Furthermore, the empirical data for Laminaria in most cases are considerably lower than

for Onsøy, even though the theoretical value for Onsøy lies within the theoretical range

for Laminaria.

7.2 T-bar factor, NT-bar

Profiles of bearing factors for the T-bar penetrometer (NT-bar) deduced from the

laboratory shear strengths are presented in Figure 7.4. Three resistance profiles of T-bar

penetrometer have been obtained for calculating the T-bar factors: first, the average

resistance profile of the standard T-bar (250 mm x 40 mm) presented in Figure 4.17;

second, the average resistance profile of the smaller T-bar (160 mm x 40 mm) also

presented in Figure 4.17; and finally the standard T-bar resistance profile (BTRT01)

reported by Schneider et al (2004). Again, for ease of reference, the T-bar factors


7-6

deduced from the first and second resistance profiles are referred to as T-bar-(a1) and

T-bar-(a2) respectively, whereas the results obtained based on the third resistance

profile are denoted as T-bar-(b). However, it should be noted that the T-bar penetration

test reported by Schneider et al (2004) was performed only to a depth of 14 m, and

hence the T-bar factors below this depth cannot be evaluated for T-bar-(b).

As may be observed in Figure 7.4, large scatter of the NT-bar data is notable at a depth of

around 6 m. This is due to the fact that more sets of data are available at this depth

(three sets of NT-bar for each of the three T-bars) and each data set shows a different

range of T-bar factor. In addition, it may be seen that each set of NT-bar seems to show

higher values at shallow depths (~6 m) than at greater depths (16 to 17 m), which

perhaps implies greater degree of sample disturbance at the shallow depths (due to

greater frequency of shells), resulting in significant underestimate in shear strengths and

overestimate in N. This trend is also true, although less apparent, for the cone factors

(Figure 7.2).

For T-bar-(a1), the NT-bar, TXC values range from 7.4 to 11.5, giving a mean value of 9.2

(SD = 1.8 and COV = 19.2 %), whilst the NT-bar, av data lie between 9.7 and 14.3, with

an average value of 11.8 (SD = 2.2 and COV = 18.6 %). The NT-bar, SS data are in the

range 10.4 to 15.1, with a mean value of 12.0 (SD = 1.7 and COV = 13.9 %). Note that,

although the NT-bar, SS data show the widest range compared to the previous two T-bar

factors, this is mainly attributed to the data point of NT-bar, SS at 6 m depth, which has a

much higher value than other NT-bar, SS data. Therefore, the coefficient of variation for

NT-bar, SS is in fact lower than for NT-bar, TXC and NT-bar, av. Also, it can be noticed that the

degree of scatter is much greater for the T-bar factors than for the corresponding cone

factors, as reflected by larger COV for the former.

T-bar-(a2) tends to give T-bar factors lower than T-bar-(a1) at shallow to intermediate

depths (6 to 13 m), but the data at greater depths (below 14 m) are extremely similar for

the two T-bars. The NT-bar, TXC values for T-bar-(a2) are within 7.4 and 10.3, with a

mean value of 8.6 (SD = 1.3 and COV = 15.1 %). The NT-bar, av and NT-bar, SS data range

from 9.7 to 12.8 and 9.8 to 13.7 respectively. Their respective mean values ± SD are

11.1± 1.6 (COV = 14.7 %) and 11.3 ± 1.4 (COV = 12.7 %).

T-bar-(b) gives the lowest values of NT-bar compared to the results of the previous two

T-bars at the corresponding depths. As already mentioned, the penetration test for

T-bar-(b) was terminated at a depth of 14 m, thus T-bar factors below this depth cannot


7-7

be evaluated. Only two data at about 6 m were obtained for both NT-bar, TXC and NT-bar, av,

for which the two values for NT-bar, TXC are 7.8 and 8.6, while the two NT-bar, av data have

a similar value of 10.7. The mean value of the former factor is 8.2, with SD of 0.6 and

COV of 6.8 %. However, larger SD and COV would be expected, if data from greater

depths had been available. This was the case for NT-bar, SS, where data are also available

at depths between 9 and 13 m. The data of NT-bar, SS range from 9.2 to 12.1, giving a

mean value of 10.2 (SD = 1.3 and COV = 13.0 %).

Figure 7.5 shows the profiles of T-bar factor derived from the vane shear strength

(NT-bar, vane). In contrast to the cone, all three T-bars show a general trend of constant

NT-bar, vane values against depth, and rather similar range of variations. Overall, the data

lie between 8 and 12, with T-bar-(a1) results being the highest, followed by T-bar-(a2)

results, and T-bar-(b) results being the lowest. The NT-bar, vane data for T-bar-(a1) lie in

the range 9.8 and 11.8, while the NT-bar, vane range for T-bar-(a2) is between 9.1 and 11.1.

The mean values for T-bar-(a1) and T-bar-(a2) are 10.9 (SD = 0.7 and COV = 6.1 %)

and 10.3 (SD = 0.6 and COV = 6.3 %) respectively. These values are very close to 10.5

− the value suggested by Stewart & Randolph (1991, 1994). For T-bar-(b), the NT-bar, vane

values range from 7.8 to 10.2, giving a mean value of 9.2 (SD = 0.8 and COV = 8.5 %).

Interestingly, the values of COV for all three sets of NT-bar, vane data are lower than the

COV for Nkt, vane, entailing lower degree of scatter for the T-bar factor, in contrast with

the case for the factors derived from laboratory strengths.

Figure 7.6 plots the profiles of NT-bar, SS for various model T-bar tests performed in the

centrifuge. As may be seen, all four T-bars show very consistent NT-bar, SS profiles. The

data range between 10.2 and 12.0, with a mean value of 11.0 (SD = 0.5 and 4.5 %). It is

also interesting to note that the NT-bar, SS values decrease initially from 5.4 to 8.3 m

depth, and then increase gradually at greater depths (below ~11 m). This trend is also

noted for the NT-bar, vane profiles for the model T-bars, which are presented in Figure 7.7.

However, this is not apparent for the field T-bar results, particularly for NT-bar, vane.

Perhaps, this is due to the NT-bar, vane profiles for the field T-bars being limited to the

depth range 5 to 14 m.

In Figure 7.7, the NT-bar, vane profile for the model T-bar test performed in the first

centrifuge sample (Box 1) is also plotted along with the results from the second sample

(Box 2) for comparison. As may be seen, the NT-bar, vane profile from Box 1 shows

somewhat higher values (~10.8) than other profiles from Box 2 (~9.0 to 10.3) for depth


7-8

range 9 to 17 m. Nonetheless, the total range of NT-bar, vane for the model T-bars is

between 8.7 and 11.1, and the average value is 9.8 (SD = 0.7 and COV = 7.0 %).

A summary of the T-bar factors, including the results obtained for Onsøy and Laminaria

clays (NGI-COFS, 2004a), is provided in Table 7.3. Note that the results for T-bar-(a1)

and T-bar-(a2) have been combined and are shown in the row labelled Burswood (a) in

the table. Also, as a reminder, the NT-bar, TXC and NT-bar, av values for T-bar-(b), labelled

Burswood (b) in the table, are based solely on two data points.

It may be seen that Burswood shows the widest ranges of T-bar factors (apart from

NT-bar, vane) compared to Onsøy and Laminaria, where the ranges fall within that for

Burswood. In addition, the mean values of T-bar factors for the different clays are more

similar, as opposed to the case for the cone factors, where the mean values for Onsøy

are much higher than for Burswood and Laminaria. Furthermore, aside from NT-bar, TXC

and the results of T-bar-(b) for Burswood, the mean values of T-bar factors are found to

be slightly higher (maximum of 19 %) than the suggested value of 10.5 for the T-bar

factor (Stewart & Randolph, 1991; 1994). It is also noteworthy that the mean values and

standard deviations for the reconstituted Burswood samples are lower than for the

natural soils, perhaps implying greater strain rate effects for the natural soils than for the

reconstituted soils, since theoretical studies show that the T-bar factor increases with

increasing dependency of the strength on the strain rate (Einav & Randolph, 2005).

7.3 Ball factor, Nball

Figure 7.8 presents the bearing factors for the ball penetrometer (Nball) derived from the

laboratory shear strengths. Again, the ball factors computed using the average ball

resistance profile shown in Figure 4.17 are referred to as Ball-(a), whereas those

obtained based on the ball resistance profile (BBTR01) reported by Schneider et al

(2004) are denoted as Ball-(b). As for T-bar-(b), the penetration test for Ball-(b) was

completed at a depth of about 14 m, so the ball factors below this depth cannot be

calculated for this test.

It may be seen that Ball-(a) and Ball-(b) both demonstrate very similar ball factors. For

Ball-(a), the Nball, TXC values vary between 7.0 and 10.0, giving an average of 8.5

(SD = 1.3 and COV = 15.0 %), while the Nball, av values lie between 9.2 and 12.5, with a

mean value of 10.9 (SD = 1.6 and COV = 14.7 %). The data of Nball, SS are in the range

9.7 to 13.7, with an average value of 11.2 (SD = 1.5 and COV = 13.0 %).


7-9

For Ball-(b), only two data at about 6 m depth are obtained for both Nball, TXC and

Nball, av, for which the values are 8.5 and 9.7 for the former, and 11.7 and 12.1 for the

latter. The Nball, SS values range from 9.8 to 13.2, giving a mean value of 11.0 (SD = 1.6

and COV = 14.4 %).

Figure 7.9 shows the ball factor profiles deduced from the vane shear strength. The

results obtained from the centrifuge test are also plotted in the figure. It is rather

difficult to infer the trends for the two sets of field ball results from the limited data

available. It may be argued that they show a slight tendency to increase with depth. The

total range of Nball, vane for the two field balls is between 8.8 and 11.0, and the overall

mean value is 10.1 (SD = 0.7 and COV = 7.1 %). However, it may be seen clearly that

the model ball shows Nball, vane values increasing slightly with depth and the data in

depth range 4.5 to 12 m lie within the scatter of the field results. The range of Nball, vane

for the model ball is between 8.3 and 11.3, and the mean value is 10.0 (SD = 1.1 and

COV = 10.9 %).

7.4 Plate factor, Nplate

There is only one resistance profile available for the plate penetrometer (from

Figure 4.17) to calculate the plate factors (Nplate). The profiles of plate factor derived

from the laboratory shear strengths are presented in Figure 7.10.

It may be seen that plate factors in most cases demonstrate larger ranges of variations

compared to the corresponding factors for other penetrometers discussed previously.

The Nplate, TXC values range from 7.7 to 11.8, with a mean value of 9.4 (SD = 1.9 and

COV = 20.3 %), whereas the Nplate, av values fall between 9.9 and 14.8, giving a mean

value of 12.1 (SD = 2.5 and COV = 20.2 %). The data of Nplate, SS lie within 10.8 and

15.7, with an average value of 12.5 (SD = 1.8 and COV = 14.1 %).

Profiles of Nplate, vane for the field and model plate penetrometers are presented in

Figure 7.11. The Nplate, vane values for the field plate range from 10.1 to 12.3, with an

average value of 11.4 (SD = 0.7 and COV = 6.2 %). The centrifuge model plate

penetrometer shows much lower Nplate, vane values, even if compared to the

corresponding bearing factors for the model T-bar and ball penetrometers. This is due to

the strangely low penetration resistance profile of the model plate presented in

Figure 6.12 (Chapter 6), where the reason causing the low plate resistances was not

identified. The range of Nplate, vane for the model plate is between 7.5 and 10.3, and the


7-10

mean value is 8.9 (SD = 0.9 and COV = 10.0 %).

7.5 Summary and recommendations

It has been shown that all bearing factors (N) for the various shaped penetrometers vary

rather considerably. The empirical N values are found to vary much greater than the

theoretical solutions (Lu et al, 2004; Randolph & Houlsby, 1984; Randolph et al, 2000).

The contradictory findings between the experiments and theories are mainly due to the

theoretical solutions not taking into account the effects of strain rate, strain softening

and strength anisotropy of the soil, each of which effects can significantly influence the

shear strength, and thus the N values (Randolph, 2000; Einav & Randolph, 2005). In

addition, the inevitable effects of sample disturbance also contribute to the scatter of N

data, and hence widen the empirical ranges.

Cone and T-bar factors from three different clay sites (Burswood, Onsøy and

Laminaria) have been collected for comparison. The cone factors are found to vary

more drastically from one site to another compared to the T-bar factors, the reasons for

which may be due to the cone resistance being more susceptible to changes in

stratigraphy and to corrections for the pore pressure and overburden pressure effects.

Errors in the corrections for the pore pressure and overburden pressure effects may give

consistent bias in the net cone resistance, but the empirical cone factors derived are

more likely to vary from one site to another.

A summary of the empirical ranges of the various N values is given in Table 7.4. The

empirical ranges for the cone and T-bar factors are the total ranges collected from the

three sites: Burswood, Onsøy and Laminaria. However, the Nvane data obtained from the

model penetration tests are not included in the total ranges, because the vane tests

performed in the centrifuge samples were deemed to measure much higher shear

strength compared to in the field, due to the relatively large size of the vane in the

centrifuge and potential strength anisotropy effects (Watson et al, 1998).

Recommendations for the choice of N values, as a general guide, to be adopted for the

various shaped penetrometers are given in the table. The recommendations are given in

the form of a recommended value ± two standard deviations (SD). The recommended

values are derived by averaging the mean values of N for the different sites. Assuming a

Gaussian (normal) distribution of the N data with respect to the mean value, there is

about 95 % probability that an empirical N data from a different site is within the


7-11

interval (mean − 2SD) to (mean + 2SD).

However, it should be noted that, since the NTXC and Nav data for T-bar-(b) and Ball-(b)

are limited to a depth of 6 m, they were not included in the evaluations of the

recommended values for NTXC and Nav for the T-bar and ball. Nevertheless, inclusion of

these data would only cause variations of about 3 − 4 % to the recommended values,

which are deemed insignificant, considering the variations are well within 2SD. Also,

since the N data for the ball and plate penetrometers are only available from a single site

(Burswood), the variation of N due to different clay sites cannot be assessed. Therefore,

a 10 % coefficient of variation (COV) is assumed for all the ball and plate factors,

which is the maximum COV calculated for the T-bar factors. The standard deviations

for the ball and plate factors given in the recommendations are estimated based on the

assumed value of COV. It is evident in Table 7.4 that the values of 2SD of N for the

full-flow penetrometers are significantly lower than for the cone, suggesting N values

for the former being much less erratic over different sites with rather similar soil

properties.

A final note for the recommendations in Table 7.4 is that they should be used only as a

guide to estimate the shear strength profiles from the results of a penetration test for

design purposes. One should bear in mind that the recommendations given in the table

are based solely on the results obtained from three sites − all consisting of lightly

overconsolidated clay deposits (OCR < 2) with rather similar soil properties. Therefore,

the empirical N values could differ from the recommended values for clay deposits with

much significantly different properties, such as with extremely high plasticity, or for

varved clays.

For example, DeJong et al (2004) reported an average NT-bar, vane value of 13.7, but 10.0

for both average values of Nball, vane and Nplate, vane for a lightly overconsolidated deposit

comprising alternating layers of clay and silt-fine sand (varved clay). The reported

NT-bar, vane value lies outside (higher than) the suggested range, but the Nball, vane value is

very similar to the recommended value given in Table 7.4. The much higher T-bar

factor compared to the ball and plate factors is because the T-bar penetration resistance

profile was about 38 % higher than the ball and plate resistance profiles, which were

very similar. This is in contrast with the results of both in situ and centrifuge model

penetration tests performed in the Burswood clay, where the T-bar and ball showed very

similar resistances. The contradictory findings in the varved clay may be attributed to


7-12

the high sensitivity and possibly rapid strain softening of the varved clay, since

theoretical studies indicate that the ball (possibly the plate as well) is more sensitive to

strain softening effects than the T-bar (Einav and Randolph, 2005). Besides, it is also

interesting to note that the upper layer crust of the varved clay deposit gave significantly

lower Nvane values: 9.7, 6.8 and 6.9 for the T-bar, ball and plate respectively (DeJong et

al, 2004).

In addition, Long & Gudjonsson (2004) also presented data of NT-bar, TXC collected from

three different clay sites. Their results showed that the NT-bar, TXC values for a laminated

clay deposit are on average close to 10.5, whereas the NT-bar, TXC values for the other two

clay deposits with plasticity index of about 22 % lie typically between 7 and 8.

However, in the upper layers of these deposits, which consisted of organic materials, the

NT-bar, TXC values obtained vary widely from 15 up to 25 (Long & Gudjonsson, 2004).

In conclusion, the average N values for the full-flow penetrometers appear to vary over

a smaller range compared to the cone, for clay deposits with rather similar

characteristics. Significant variations of N for different soil deposits generally imply

significant differences in the properties of the soil. Therefore, it may be possible to

evaluate additional soil parameters, in addition to shear strength, from parallel in situ

penetrometer tests conducted using different types of penetrometers.

At this stage, it is still necessary to develop empirical correlations for the full-flow

penetrometers for each site, as for the cone penetrometer, but ultimately, as more

empirical N values from different sites are accumulated in the database, soil

characteristics that cause the variations of N may be clearly identified and thus

quantified. For example, by conducting parallel cone and T-bar penetration tests, there

is the potential to interpret differences in the cone and T-bar factors, perhaps to indicate

differences in overconsolidation ratio or in rigidity index (Randolph, 2004). Also, by

performing parallel T-bar and ball penetration tests, relative values of the T-bar and ball

factors may help quantify the effects of strain softening or infer the anisotropic strength

ratio. However, in order to facilitate these, it is of utmost importance that the shear

strength data used for the empirical correlations are of high quality, in order to minimise

complications due to sample disturbance when evaluating laboratory strengths.

EFFECTS OF PARTIAL CONSOLIDATION AND VISCOSCITY

8-1

8 EFFECTS OF PARTIAL CONSOLIDATION AND VISCOSITY

The previous chapter explored the empirical ranges of bearing factors for the various

penetrometers. One of the primary parameters affecting the bearing factors is the rate of

penetration, and the effect of this on the penetrometer tip resistance is discussed in

further detail in this chapter, in the light of the variable rate model penetration tests

(referred to as ‘twitch’ tests) performed in the centrifuge.

As already mentioned in Chapter 2, the tip resistance increases as the penetration rate is

increased when the conditions around the advancing penetrometer are undrained, due to

viscous effects. On the other hand, when the conditions are partially drained, the tip

resistance will increase as the penetration rate is decreased, due to partial consolidation

effects and local strengthening of the soil around the penetrometer. Therefore, there is a

transition point from undrained to partially drained conditions where the tip resistance is

minimum. This is clearly illustrated in the results presented by Bemben & Myers

(1974), who carried out in situ mechanical cone penetration tests with penetration rates

ranging from 0.2 to 200 mm/s in varved clay, and by Roy et al (1982), who performed

piezocone tests with rates ranging from 0.5 to 40 mm/s in sensitive clay. A brief

summary of literature published on effects of penetration rate on the cone resistance has

been given by Lunne et al (1997b).

The effect of partial consolidation on the penetrometer tip resistance has been exploited

to estimate the consolidation coefficients of the centrifuge samples by performing a

series of twitch tests and interpreting the results in conjunction with the results obtained

by other researchers. The partial consolidation effects for the various shaped (cone,

T-bar, ball and plate) penetrometers and for the T-bars with different aspect ratios will

also be discussed. After that, attention is turned to the viscous zone, where the

conditions are undrained.

8.1 Resistance profiles of twitch tests

The testing procedure for a twitch test has been described in Chapter 6. It involved

pushing a penetrometer into the centrifuge box sample initially at a rate of 1 mm/s.

After the penetrometer had been advanced below the lightly overconsolidated layer of

the sample, the penetration rate was successively halved over 8 steps to a final value of

0.0078 mm/s, with the penetrometer advanced a set distance at each stage. During

completion of the final stage, the penetrometer was either pulled-out immediately at a


8-2

rate of 1 mm/s, or was penetrated further at the original rate of 1 mm/s before being

extracted at the same rate.

Figure 8.1 summarises the tip resistance profiles of the twitch tests performed for the

various shaped model penetrometers. The average resistance profile of the normal

(constant rate) T-bar test is also included in the figure as a broad reference for the tip

resistance under undrained conditions. Individual comparisons of each twitch test and

the corresponding normal test results for the different shaped model penetrometer have

been presented in Chapter 6 (Figures 6.8 to 6.11).

It may be seen in Figure 8.1 that the penetration resistance for the T-bar twitch test

starts to exceed that for the normal T-bar test from step 5 (where the rate is

0.0625 mm/s), and the increase in resistance becomes more significant in succeeding

steps. This is also true for the model ball and plate penetrometers (as has been presented

earlier in Chapter 6). These suggest that partial consolidation effects start to become

significant at the same penetration rate, regardless of the differences in shapes and

diameters for these particular penetrometers.

In addition, twitch tests for the model T-bar and ball penetrometers are found to give

extremely similar penetration resistances even at different penetration rates (at different

stages of the tests), implying that the two penetrometers experienced similar degrees of

local strengthening of the surrounding soil. The reasons may be attributed to the relative

size and flow pattern of the two penetrometers. Since the diameter of the model ball

(11.9 mm) is considerably larger than that of the model T-bar (5 mm), one would expect

the ball to experience a delay of partial consolidation effects compared to the T-bar,

owing to larger size of the former. However, the axisymmetric flow for the ball allows

more rapid dissipation of pore pressure compared to the plane strain flow for the T-bar,

hence accelerating the partial consolidation effects for the ball. The consequence is that

both of the penetrometers show similar tip resistances regardless of the penetration rate,

although this would only be the case for this particular relative size (or diameter ratio)

between the T-bar and ball penetrometers.

The model plate twitch test gave lower penetration resistances than the model T-bar and

ball twitch tests, whereas the resistance profile for the model cone twitch test appears to

be the lowest. The hierarchy of resistance profiles for the twitch tests is consistent with

that for the normal tests.

During extraction, the model penetrometers were all extracted at a rate of 1 mm/s, as for


8-3

the normal tests. It is evident in Figure 8.1 that the extraction resistances of the twitch

tests have been influenced by the effects of partial consolidation during penetration.

Several interesting features can be noted: first, a peak of resistance similar to that

observed during penetration is also noted in the extraction resistance profiles of the

T-bar and cone twitch tests. No peak is noted in the profiles of the twitch tests for the

model ball and plate, because they were extracted immediately (without further

penetration) at completion of the final stage of penetration. Also, it may be noted that

the extraction resistance for the ball twitch test was not recorded between depths of 14.5

and 15.5 m. This was due to the rate of data logging during slow penetration not being

changed back to the rapid rate of logging as the ball was first extracted at 1 mm/s.

Secondly, following the peak is a trough in the extraction resistance profile of the T-bar

twitch test. Similar troughs are also noted in the extraction results for the model ball and

plate. The extraction resistances at the troughs for the twitch tests are even lower than

that for the normal T-bar test at the corresponding depths, whereas at other depths, the

twitch tests gave consistently higher extraction resistances than the normal T-bar test. It

is suspected that the trough may have been caused by a (partial) physical gap (water

filled) forming above the model penetrometer, and thus resulting in reduction of

extraction resistance in the region. At depths above the troughs, the higher extraction

resistances for the twitch tests compared to the normal T-bar test may be owing to the

soil regaining strength from local consolidation occurring during slow penetration, or

stronger soil (resulting from partial consolidation during slow penetration) adhering to

the penetrometer.

In contrast with the full-flow penetrometer results, the cone twitch test shows a smooth

and steady decrease in extraction resistance from the peak value as the cone was

extracted.

Figure 8.2 presents the twitch test results for the model T-bar penetrometers with

different aspect ratios. It can be seen that the penetration and extraction resistance

profiles are tightly bunched. In addition, the interesting features noted previously in the

resistance profiles for the T-bar twitch test shown in Figure 8.1 are also noted for the

various T-bar twitch test results presented in Figure 8.2. Therefore, it can be concluded

that the aspect ratio (L/d) in the range 4 to 10 does not affect the T-bar tip resistance,

even at low penetration rates where the conditions are partially drained.


8-4

8.2 Evaluation of consolidation coefficient from twitch tests

Penetration test data measured from a twitch test can be used to interpret a value for the

consolidation coefficient, cv, of the soil (Randolph & House, 2001). To facilitate this, a

‘backbone’ curve that gives the variation of penetration resistance against the

penetration rate, in an appropriate non-dimensional form, is essential. A simple

dimensional analysis shows that the drainage conditions depend not only on the

penetration rate, v, but also on the diameter of the advancing probe, d, and the

coefficient of consolidation, cv. This led to the penetration rate being normalised to non-

dimensional velocity (Randolph & House, 2001; Finnie & Randolph, 1994):

vc

vdV = (8.1)

The penetration resistance at a particular rate can be normalised easily by the undrained

value from a ‘rapid’ penetration test (Randolph & House, 2001; House et al, 2001).

In an attempt to establish a backbone curve for the T-bar, a series of constant rate of

T-bar penetration tests at various penetration rates were performed in the centrifuge

independently by Watson & Suemasa (2000, unpublished), House et al (2001) and

Randolph & Hope (2004), who also performed similar tests for the cone penetrometer.

House et al (2001) suggested an expression of the form below to fit the data obtained

from the penetration tests, in order to derive the backbone curve:

mref cV1

baqq

++= (8.2)

where: q = penetration resistance at any rate;

qref = reference (undrained) resistance;

a, b, c, m = constants for the backbone curve.

Note that m is used in the equation above rather than d originally used by House et al

(2001), in order to avoid confusion with the diameter of the penetrometer. The values of

cv used for the normalisation of V were obtained from Rowe cell tests. A summary of

the constants a, b, c and m published in the literature is given in Table 2.1 (in

Chapter 2), and the resulting backbone curves for the T-bar and cone penetrometers are

presented in Figures 8.3 and 8.4 respectively. It should be noted that the backbone

curves shown in the figures have not incorporated viscous effects, and the curves

derived after Randolph & Hope (2004) are plotted using different sets of constants from


8-5

those originally presented by them, since the original constants were derived with a

correction factor for viscous effects being applied to Equation 8.2. This correction factor

was later found to be inappropriate, as it affected the ratio of partially drained to

undrained penetration resistances. Therefore, a different sets of constants for each of the

T-bar and the cone backbone curves have been derived. From hereinafter, the backbone

curve from Randolph & Hope (2004) refers to that plotted using the new derived

constants shown in Table 2.1, which still matches well the actual data points from that

study.

It may be seen in Figure 8.3 that the backbone curve for the T-bar derived by Watson &

Suemasa (2000) shows a transition point from undrained to partially drained response of

V ~ 20 and normalised resistance doubling within one order of magnitude by V ~ 2.

House et al (2001) derived a curve showing a later transition point of V ~ 10 and

resistance doubling by V ~ 0.5, while results from Randolph & Hope (2004) indicate a

transition point similar to that from House et al, but resistance increasing more strongly

and doubling by V ~ 1. Randolph (2004) commented that the T-bar tests performed in

the drum centrifuge by House et al (2001) may have been affected by small vibrations

transmitted from the central turntable of the drum centrifuge. These could have led to

additional excess pore pressures generated at the T-bar, delaying the effects of

consolidation, and hence showing a relatively low rate of increase in resistance with

reducing penetration rate.

Average values for the consolidation coefficients of the first (Box 1) and the second

(Box 2) centrifuge samples can be deduced by using the backbone curves mentioned

above in conjunction with the twitch test data, as illustrated in Figure 8.3. This has been

carried out based on the twitch test results of the model T-bar 20 mm x 5 mm. The

normalised resistances for the T-bar are computed using the (developed) resistances at

each stage of the twitch test normalised by the corresponding reference undrained

resistances from the normal T-bar test. Since the penetration rate and the probe’s

diameter are known, the only variable in V (= vd/cv) is cv. The value of cv is adjusted to

fit the twitch test data onto the backbone curve and the value that gives the best-fit result

corresponds to the average cv of the sample. This results in cv values of 0.45 and

0.4 m2/year for sample Boxes 1 and 2 respectively, based on the backbone curve derived

by Watson & Suemasa (2000). While fitting the normalised data to the curve from

Randolph & Hope (2004) gave respective cv of 1 and 0.9 m2/year for Boxes 1 and 2,


8-6

results deduced based on the curve from House et al (2001) illustrate the highest values,

corresponding to 1.8 and 1.6 m2/year for Boxes 1 and 2 respectively.

The backbone curve for the cone penetrometer derived by Randolph & Hope (2004) is

shown in Figure 8.4. It demonstrates an earlier but more gradual transition from

undrained to partially drained conditions (for V < 30) compared to that for the T-bar.

Similarly, normalised resistances from the cone twitch test (from Box 1 only) were

fitted to the backbone curve and this yields a cv value of 0.6 m2/year (Figure 8.4). This

is within the range of cv deduced from the interpretation of the T-bar twitch test data,

although closer to that deduced based on the backbone curve from Watson & Suemasa

(2000).

The consolidation coefficients deduced from the curve fitting methods, based on the

T-bar backbone curve from Watson & Suemasa (2000) and the cone backbone curve

from Randolph & Hope (2004), appear to be reasonably close to the cv values of 0.5 and

0.75 m2/year for Boxes 1 and 2 respectively, estimated using data from consolidation of

the centrifuge samples and Equation 6.1 (see Table 6.1 in Chapter 6). In addition, this is

also found to be true when compared to cv values measured from Rowe cell tests

conducted on Burswood clay material from slurry, which range between 0.35 and

0.54 m2/year.

However, it is interesting to note that oedometer tests performed on samples recovered

after the centrifuge tests gave cv values ranging generally between 1 and 1.2 m2/year at

the yield stress, σ'yield, but reducing to between 0.5 and 0.7 m2/year at stress levels of 5 x

σ'yield. It may be argued that the cv values obtained at stress levels greater than yield

from the oedometer tests reflect the ‘true’ cv for Burswood clay in normally

consolidated states, for which the values are consistent with those interpreted from the

twitch tests using the T-bar backbone curve from Watson & Suemasa (2000) and the

cone backbone curve from Randolph & Hope (2004).

8.3 Effect of partial consolidation for various penetrometers

The effect of partial consolidation for the different shaped penetrometers may be studied

by comparing their normalised resistances against the non-dimensional velocity, V, as

plotted in Figure 8.5. A single value of cv = 0.45 m2/year has been adopted for all

penetrometers for comparison.

All sets of data in Figure 8.5 exhibit similar trends of variation of normalised resistance


8-7

with non-dimensional velocity, V, but plot at different abscissae. It can be seen that the

non-dimensional profile of the ball plots to the right of the T-bar profile by a factor of

about 2.5, which implies a higher rate of local consolidation around the ball than around

the T-bar during penetration. The transition point for the ball inferred from the twitch

test data lies at V ~ 50, although more data are required to verify this. It is, nevertheless,

interesting to note that the diameter ratio of the ball to the T-bar (dball/dT-bar) is 2.38,

which is close to 2.5 (the horizontal position of the ball profile in relation to that of the

T-bar profile). Also, it has been noted previously that the ball and T-bar penetrometers

showed very similar penetration resistance profiles both in the normal and twitch tests.

These suggest that the relative rates of consolidation around the ball and T-bar are in a

ratio of about 2.5 : 1. Therefore, if the ball diameter is about 2 to 3 times the T-bar

diameter, the actual consolidation rates around the two penetrometers are expected to be

similar.

It is surprising to note that the plate penetrometer demonstrates very similar normalised

resistance profile to the ball, even though it has shown ‘actual’ penetration resistances

lower than the T-bar and ball both in the normal and twitch tests. This suggests that the

ball and plate penetrometers have similar rates of local consolidation around them,

despite the round base for the ball compared to the flat base for the plate penetrometer.

The non-dimensional profile for the cone lies between the T-bar and ball (and the plate)

profiles. This suggests that the local consolidation rate around the cone is higher than

that around the T-bar (consistent with that noted in Section 8.2), but lower than for the

ball and plate penetrometers.

Figure 8.6 plots the normalised resistance against the non-dimensional velocity for the

T-bars with various aspect ratios tested in sample Box 2. The curves are tightly bunched

and fit reasonably well on the backbone curve from Watson & Suemasa (2000) using a

single value of cv = 0.4 m2/year. Again, no obvious effect on the normalised T-bar

resistances can be identified due to the various aspect ratios. Nevertheless, it can be seen

in the figure that some T-bars show normalised resistances slightly lower than unity

near the transition point. These are due to reduced viscous effects relative to the

reference undrained resistances.

It is intriguing in the previous discussion that the consolidation rates for plane strain and

axisymmetric flows inferred from the relative positions between the non-dimensional

profiles of the T-bar and ball penetrometers, are in a ratio of about 1 : 2.5. There is as


8-8

yet (to the author’s knowledge) no analysis undertaken to study consolidation rates for

the penetration problems. Nonetheless, in an attempt to assess the relative rates of

consolidation for plane strain and axisymmetric flows, a surface footing analogy is

presented as follows.

Consolidation around the T-bar and ball penetrometers may be analogous to

consolidation under rectangular and circular footings respectively. Based on Biot

theory, Gibson & McNamee (1957, 1963) derived theoretical curves for the degree of

consolidation (settlement, Us) for rectangular footings with length to breadth ratios, L/B,

ranging from 1 to 5, on an infinitely deep layer. Similarly, De Jong (1957) and

McNamee & Gibson (1960) derived theoretical consolidation curves for circular

footing. Davis & Poulos (1972) replotted these curves in term of Us against time factor,

Tv = cvt/(area of footing), where t is the real time. The resulting curves are reproduced in

Figure 8.7 (a) for the square (L/B = 1) and rectangular (L/B = 5) footings; and (b) for

the square and circular footings with the same area. It can be seen in Figure 8.7 (a) that

the curves for the square and rectangular footings are very similar, which is also true in

the case for the square and circular footings, as seen in Figure 8.7 (b). The findings

suggested that, for an infinitely deep layer, the rate of consolidation is determined by the

area of the footing and is virtually independent of the shape, at least for L/B up to 5 for

the rectangular footing (Davis & Poulos, 1972).

The model T-bar tested in the first centrifuge sample (Box 1) had an aspect ratio, L/d, of

4 and projected area of 100 mm2, which was rather similar to that of the model ball

(111.2 mm2). As seen earlier, the actual consolidation rates around the model T-bar and

ball penetrometers were similar, which is in agreement with the theoretical findings

obtained by Davis & Poulos (1972). However, the results of model T-bar twitch tests,

with L/d ranging from 4 to 10 and projected area from 100 to 250 mm2, performed in

the second centrifuge sample (Box 2) were contrary to the suggestion of Davis &

Poulos, as the actual consolidation rates around the different model T-bars did not seem

to vary with the projected area. Further investigation is required to verify this.

8.4 Effect of penetration rate in viscous (undrained) region

Randolph & Hope (2004) incorporated viscous effects in their ‘original’ backbone

curves, by multiplying a hyperbolic function to the right hand side of Equation 8.2 (see

Equation 2.7 in Chapter 2). However, as already mentioned, the hyperbolic function

originally adopted by Randolph & Hope (2004) was found to be inappropriate. In order


8-9

to comply with the hypothetical relationship of shear strength and strain rate proposed

by Mitchell (1976), an alternative correction factor for viscous effects, which is

expressed as a ratio of hyperbolic functions, is attempted. The resulting equation is:

⎟⎟⎟⎟

⎠

⎞

⎜⎜⎜⎜

⎝

⎛

⋅λ

+

⋅λ

+⋅⎟

⎠⎞

⎜⎝⎛

++=

−

−

)V/V(sinh)10(n

1

)V/V(sinh)10(n

1

cV1ba

qq

oref1

o1

mref

l

l (8.3)

where: λ = rate parameter (typically between 0.1 and 0.2);

Vo = value of V for which the viscous effects start to decay;

Vref = reference V where the hyperbolic function term passes through unity.

This equation has been adopted to fit the twitch test data, with attention given to the

viscous region. In addition to twitch test data of the model T-bar and cone, in situ T-bar

and cone twitch test results reported by Schneider et al (2004) were also obtained for

curve fitting purposes. The primary objective is to assess the values of λ and Vo.

8.4.1 Model twitch tests in centrifuge

Figure 8.8 (a) shows the fitted results of the model T-bar twitch test performed in

sample Box 1. The twitch test data are first fitted to the backbone curves without

viscous effects; as already presented in Section 8.2, these result in cv values of

0.45 m2/year based on the backbone curve from Watson & Suemasa (2000), and

1 m2/year based on that from Randolph & Hope (2004). These set the values of Vref in

Equation 8.3, since Vref is the value of V corresponding to the reference (undrained)

resistance (qref) and is determined by the reference penetration rate and the value of cv.

Hence, it is only required to adjust the values for Vo and λ, to fit the ‘modified’

backbone curves to the twitch test data. Nevertheless, it should be stated that the

following curve fitting analysis performed for the model twitch test data has been based

on visual assessment. This is because the number of data in the viscous region is limited

(since Vref is close to the transition point from undrained to partially drained), so that

undertaking the analysis following a statistical procedure tended to produce a curve

without viscous effects (λ tending to zero).

It is found that with Vo = 300 and λ = 0.15, the modified backbone curve of Watson &

Suemasa (2000) fits the data reasonably well in the viscous zone, while Vo = 150 and


8-10

λ = 0.15 are suitable for the case of the backbone curve from Randolph & Hope (2004),

as shown in Figure 8.8 (a). However, it can be seen clearly that both of the backbone

curves have been shifted slightly to the left from their original positions, due to the

influence of the multiplier for viscous effects extending to the partially drained portion

of the curves.

If the three parameters (Vref, Vo and λ) are adjusted to fit the twitch test data, a

combination of Vo = 200, λ = 0.15 and Vref = 263 (for cv = 0.6 m2/year) for the

backbone curve from Watson & Suemasa (2000), and Vo = 100, λ = 0.15 and Vref = 121

(for cv = 1.3 m2/year) for that from Randolph & Hope (2004) produce rather good fits

with the test data, as illustrated in Figure 8.8 (b). The values of cv are, however, slightly

greater than those deduced previously.

Similar procedures of curve fitting were carried out for the cone twitch test data and the

results are presented in Figure 8.9. Based on the cv value of 0.6 m2/year (from

Section 8.2), Vo and λ with values of 100 and 0.1 respectively are found to give the

best-fit curve to the cone test data (Figure 8.9 (a)). However, it may be seen that the

modified backbone curve increases somewhat rapidly in the viscous region, particularly

beyond the point Vref. In contrast, the cone twitch test gave normalised resistances

remaining close to unity in the viscous region. This was in fact also observed in the

cone test data reported by Randolph & Hope (2004). From these data, a trend of rather

gradual increase in cone resistance due to viscous effects may be expected. Therefore, a

relatively low λ of 0.05 is attempted, and the best-fit curve is found for Vo = 50 and

Vref = 394 (for cv = 0.8 m2/year), as shown in Figure 8.9 (b).

Since there is no backbone curve established for either the ball and plate penetrometers,

similar curve fitting procedures cannot be performed for these penetrometers.

Nevertheless, based on the limited data from the model ball and plate twitch tests

presented in Figure 8.5, it may be seen that the normalised resistances for these

penetrometers remain close to unity in the viscous region, which suggest relatively

minor decrease in viscous effects as the (reference) rate was reduced by a factor of ~10.

For clarity, fitted results of the various T-bar twitch tests performed in sample Box 2,

based on two different backbone curves, are presented separately in Figures 8.10 and

8.11. In Figure 8.10, the backbone curve from Watson & Suemasa (2000) is adopted.

Using the cv value deduced in Section 8.2 (0.4 m2/year), the best-fit result for the


8-11

viscous region is obtained for Vo = 200 and λ = 0.15 (Figure 8.10 (a)), while the overall

best-fit curve is obtained for a combination of Vref = 315 (for cv = 0.5 m2/year),

Vo = 150 and λ = 0.15 (Figure 8.10 (b)). It can be noted that the model T-bars show

some scatter in the normalised resistances, with T-bar Tests 32 (40 mm x 5 mm) and 42

(50 mm x 5 mm) showing greater reductions in resistances due to reduced viscous

effects relative to the reference values, while T-bar Tests 12 and 52 (both 20 mm x

5 mm) showed more uniform normalised resistances in the viscous region. The

backbone curves give an average fit to the data in this region.

For curve fitting with the modified backbone curve of Randolph & Hope (2004), with

the cv value of 0.9 m2/year obtained from Section 8.2, Vo = 100 and λ = 0.15 are

required to produce a curve in good agreement with the test data in the viscous zone

(Figure 8.11 (a)). Alternatively, a cv of 1 m2/year (i.e. Vref = 158) coupled with Vo = 70

and λ = 0.15 give the overall best-fit curve, as illustrated in Figure 8.11 (b).

It is interesting to note that a single value of rate parameter, λ, of about 0.15 was

obtained from all fitted results of T-bar twitch tests presented above. This suggests

similar viscous effects in both of the centrifuge samples. The value is also within the

typical range of λ (between 0.1 and 0.2) suggested by Randolph (2004).

A range of Vo was obtained from fitted results of the different cases. When curve fitting

procedures were carried out using the backbone curve from Watson & Suemasa (2000),

Vo varied from 150 to 300, but if the backbone curve from Randolph & Hope (2004)

was adopted, it varied from 70 to 150. Since Vo is where the viscous effects start to

decay, it is believed to be unique for each penetrometer (although varies for different

backbone curves). Although more data, perhaps from higher values of V, are required to

assess Vo accurately, the influence of Vo on the ‘modified’ backbone curves is examined

in Figure 8.12. It can be seen that, for the range of Vo considered (which is obtained

from the fitted results), the modified backbone curves are very similar.

Also, as already mentioned, the effect of the correction factor for viscous effects

extends to the partially drained portion of the backbone curves, causing the curves to

shift slightly to the left from their original positions. As a result, values of cv deduced

using backbone curves with correction for viscous effects are slightly greater (about 10

to 30 %) than the values obtained using backbone curves without the correction. Ideally,

the constants a, b, c and m in Equation 8.2 should be derived along with Vo and λ in the


8-12

correction factor for viscous effects (as in Equation 8.3), rather than deriving them

separately.

It is worth pointing out that, in the field, a penetrometer such as the T-bar has a diameter

(d) of 40 mm and the standard (reference) rate of penetration (v) is 20 mm/s, both of

which are much greater than those for the centrifuge tests. The product (vd)field for a

field T-bar test is 800 mm2/s, which is 160 times greater than (vd)centrifuge (= 5 mm2/s)

for the model T-bar test. Hence, field tests have 2 orders of magnitude greater

‘undrained’ range than centrifuge tests and viscous effects can be defined much better

from field tests.

8.4.2 Field twitch test results

A series of in situ T-bar and cone twitch tests were carried out and the results reported

by Schneider et al (2004; see also Randolph, 2004). The reference penetration resistance

corresponds to that measured at a penetration rate of 20 mm/s. The reported data are

presented in Figure 8.13 for the field T-bar twitch tests and in Figure 8.14 for the field

cone twitch tests. As may be seen, there is much more scatter in the cone test data than

in the T-bar test data, but it is evident that both sets of data illustrate significant

reduction in normalised resistance from unity to about 0.8, as the rate was reduced by a

factor of ~50. The normalised resistance increased back to unity as the rate was reduced

by a further factor of ~40.

The data have been fitted with backbone curves computed using Equation 8.3 and the

best-fit curves deduced following the least squares method. In Figure 8.13 (a), the

backbone curve from Watson & Suemasa (2000) is adopted, and a combination of

Vo = 1080, λ = 0.15 and cv = 1.6 m2/year is found to give a very good fit to the field

T-bar test data. For the modified backbone curve of Randolph & Hope (2004), the best-

fit results are obtained for Vo = 620, λ = 0.15 and cv = 3.3 m2/year, as shown in Figure

8.13 (b). Note that, although the same value of λ was found to fit both the field and

centrifuge test data, the values of Vo obtained for the field data are considerably higher

than those for the centrifuge test data. Furthermore, the values of cv deduced are much

higher (particularly where the backbone curve from Randolph & Hope (2004) was

adopted) than the values measured from the oedometer tests performed on the in situ

test samples (between 1 and 1.25 m2/year at yield stress) and from Rowe cell tests

(between 0.35 and 0.54 m2/year).


8-13

The fitted results for the field cone tests are presented in Figure 8.14. The best-fit curve

is obtained using Vo = 1280, λ = 0.18 and cv = 1.9 m2/year. However, it was not

possible to fit the curve through the scatter observed at rates within a factor of 10 less

than the reference point.

8.5 Summary for effects of penetration rate

This chapter has discussed the effects of penetration rate both in the partially drained

and fully undrained conditions. When the rate (for undrained penetration) was reduced

by about two orders, penetration resistances for the various shaped model penetrometers

could increase up to 100 % higher than their corresponding undrained resistances due to

partial consolidation effects.

In the centrifuge, ‘twitch’ tests for the model T-bar and ball penetrometers have

demonstrated extremely similar resistance profiles, regardless of the penetration rate.

The reason may be attributed to the relative size and flow pattern of the two

penetrometers. The model ball, with a larger diameter, is expected to experience a delay

of partial consolidation effects compared to the T-bar, but this is compensated by more

rapid pore pressure dissipation for axisymmetric flow than for plane strain flow. It was

later found that for a diameter ratio (dball/dT-bar) of 2 to 3, consolidation rates for the two

penetrometers would be similar.

It has been shown that the consolidation coefficient (cv) of the clay can be estimated by

fitting the data obtained from twitch tests to a ‘backbone’ curve that gives the relation of

normalised resistance (q/qref) against non-dimensional velocity (V). The values of cv

deduced from curve fitting, based on the T-bar backbone curve from Watson &

Suemasa (2000), were similar to the cv values measured (at very high stress level) in the

oedometer tests and in Rowe cell tests, in addition to the cv values interpreted using data

from consolidation of the sample in centrifuge. Nevertheless, when the T-bar backbone

curve derived after Randolph & Hope (2004) was adopted, the results were more similar

to the cv values measured at yield stress in the oedometer tests.

The profile of normalised resistance against non-dimensional velocity for the model ball

penetrometer was found to plot horizontally to the right of the profile for the model

T-bar, by a factor of about 2.5. This suggests that the normalised consolidation rate for

the ball is relatively higher than for the T-bar, by a ratio of about 2.5. In an attempt to

assess the relative rates of consolidation for the T-bar and ball, the rates of consolidation


8-14

for a rectangular footing and a circular footing were evaluated, as an analogy to the

penetration problems. Theoretical solutions presented by Davis & Poulos (1972)

showed that, for an infinitely deep layer, the rate of consolidation was determined by the

area of the footing and was virtually independent of the shape. These were in agreement

with the results of model T-bar and ball twitch tests performed in the first centrifuge

sample, but in contrast with the results of model T-bar twitch tests with different

projected areas, conducted in the second centrifuge sample. Further investigation is

required to verify the rate of consolidation for the penetration problem.

The model plate penetrometer has shown similar normalised resistances to the model

ball, despite the ‘actual’ resistance profiles for the plate being lower than for the ball in

both the twitch tests and the normal (undrained) tests. This suggests that the two

penetrometers have similar rates of consolidation, regardless of their difference in shape

(in agreement with the theoretical findings shown by Davis & Poulos, 1972). However,

the reason for the model plate demonstrating lower ‘actual’ resistance profiles than the

model ball and T-bar was not identified.

In the undrained region, the model twitch tests have demonstrated relatively

insignificant viscous effects compared to the field twitch tests reported by Schneider et

al (2004). Perhaps, this is partly because the reference velocity (Vref) of the centrifuge

tests was fairly close to the transition point from undrained to partially drained

conditions. However, the field tests have 2 orders of magnitude greater ‘undrained’

range than the centrifuge tests, since (vd)field for a field T-bar test is 160 times greater

than (vd)centrifuge for a model T-bar test. More apparent reduction in resistance due to

reduced viscous effects relative to the reference resistance was noted for the field tests

than for the centrifuge tests. Therefore, the field tests can define viscous effects much

better than the centrifuge tests.

A new correction factor for viscous effects was introduced and applied to the ‘original’

backbone curves derived by Watson & Suemasa (2000) and by Randolph & Hope

(2004). The resulting backbone curves can fit the field twitch test data reasonably well.

However, the values of cv deduced from the curve fitting methods were somewhat

higher than the values measured at yield stress in the oedometer tests carried out on in

situ test samples. In addition, the correction factor for viscous effects would inevitably

pollute the partially drained region of the backbone curves and cause them to shift

slightly to the left from their original positions. One option to avoid this would be to


8-15

derive the constants of the backbone curves after the correction for viscous effects has

been applied.

CONCLUSIONS

9-1

9 CONCLUSIONS

This research has studied the penetration and extraction resistance profiles of different

types of penetrometers in soft clay. The penetrometers of interest include the cone,

T-bar, ball and plate. The primary focus has been on correlating the penetration

resistances of these penetrometers with the undrained shear strengths measured from the

vane shear and laboratory tests. Since the shear strength, and thus the penetration

resistance, are strain rate dependent, model variable rate (‘twitch’) penetration tests

were performed in the centrifuge, in order to study the strain rate effects under

undrained and partially drained conditions. Additionally, the data obtained from the

model twitch tests were used to deduce the coefficients of consolidation of the

reconstituted clay samples. Finally, the potential of cyclic penetration and extraction

tests on the full-flow penetrometers has also been explored, with a view to measuring

the remoulded shear strength.

This final chapter summarises the main findings arising from the research, and then

provides recommendations for future research.

9.1 Findings of research

In situ testing and sampling for the research were carried out at Burswood, a local site

near Perth, in Western Australia.

Results of the in situ penetrometer tests showed that the cone resistance is very sensitive

to the corrections for pore pressure and overburden pressure effects. Typically, the

correction for pore pressure effects caused an increase in resistance of about 17 % from

the measured to total cone resistance. The correction for overburden pressure effects

subsequently resulted in a reduction in resistance of about 31 % from the total to net

value, giving a net reduction of about 19 % from the measured to net cone resistance.

Hence, errors in estimating the pore pressure and overburden pressure effects may lead

to quite large inaccuracy in the net cone resistance, and thus the derived shear strength,

particularly for soft sediments.

In contrary to the cone, a full-flow penetrometer has shown to be relatively insensitive

to such corrections, which resulted in a decrease in resistance of only about 4 % from

the measured to net value. Therefore, in practice, the measured tip resistance of a full-

flow penetrometer is often used directly to derive the undrained shear strength profile.

CONCLUSIONS

9-2

Also, the full-flow penetrometers were found to demonstrate very consistent resistance

profiles both during penetration and extraction, with a difference of around 15 %

between the highest and lowest profiles, and a standard deviation of 15 %. However, the

cone penetrometer gave similar penetration resistance at shallow depths, but

increasingly higher resistance at greater depths − a phenomenon that is also common in

offshore results. During extraction, the cone penetrometer gave generally higher

resistance than the full-flow penetrometers.

The effect of surface roughness of the T-bar penetrometer on its resistance was

examined by performing T-bar tests, one with a machined smooth surface and the other

with a lightly sand-blasted surface. During penetration, the smooth T-bar seemed to

show slightly lower penetration resistance than the rough T-bar, in agreement with the

variation predicted by the theoretical T-bar resistance (Randolph & Houlsby, 1984).

However, during extraction, the resistance for the smooth T-bar was higher than for the

rough T-bar, which is contrary to the theoretical prediction. This may be due to the

smooth T-bar causing less remoulding or softening of the soil locally during

penetration, and thus, leading to higher extraction resistance for the smooth T-bar than

for the rough T-bar.

The standard field T-bar (250 mm x 40 mm) exhibited slightly higher penetration

resistance profile (~6 % on average) than the smaller field T-bar (160 mm x 40 mm).

This difference may be attributed to slight changes in stratigraphy. Such small

difference in resistance between the two field T-bars is believed to be statistically

insignificant. Besides, model T-bar tests conducted in the centrifuge indicated that the

aspect ratio (L/d) does not have any obvious effect on the T-bar resistance, at least for

L/d of 4 to 10. Hence, it can be concluded that any aspect ratio in the range 4 to 8 for

the T-bar would be appropriate.

The empirical bearing factors, N, back-calculated from the vane shear strength and

laboratory strengths were found to vary much more widely than the theoretical

solutions, both for the cone and full-flow penetrometers (Lu et al, 2004; Randolph &

Houlsby, 1984; Randolph et al, 2000). In addition to the effects of sample disturbance,

the main reason for the contradictory findings is due to the fact that the theoretical

solutions ignore the effects of strain rate, strain softening and strength anisotropy, each

of which factors can significantly influence the shear strength, and hence the N values

(Randolph, 2000; Einav & Randolph, 2005).

CONCLUSIONS

9-3

For comparison with Burswood, the T-bar and cone factors for two other clay sites (one

located onshore in Norway: Onsøy; and the other located offshore in Australia:

Laminaria) were also obtained from the report of NGI-COFS (2004a). All the three sites

are underlain with lightly overconsolidated clay with plasticity index, IP, mainly in the

range 30 to 50 %. The T-bar factors appear to vary to a lesser degree from one site to

another, compared to the cone factors. Particularly, the mean value of the cone factor

derived from simple shear strength (Nkt, SS) for Onsøy is ~21 % higher than for

Burswood (a, see Table 7.2), whereas that for Laminaria is ~6 % lower than for

Burswood. On the other hand, the mean values of NT-bar, SS for Onsøy and Laminaria are

~8 and 7 % respectively, higher than for Burswood (a, see Table 7.3). Also, the average

value of Nkt derived from vane shear strength for Onsøy is ~30 % higher, whilst that for

Laminaria is ~2 % lower than for Burswood, but the NT-bar, vane values for Onsøy and

Laminaria are on average ~9 and 7 % respectively, higher than for Burswood. Similar

findings were also noted by Long & Gudjonsson (2004), where the total range of Nkt for

three sites (one containing laminated clay and the other two containing lightly

overconsolidated clays, both with IP of about 22 %) is greater than the total range of

NT-bar, with triaxial compression strength being the reference strength.

It should be pointed out that, although the cone factors vary more widely from one site

to another compared to the T-bar factors (and most probably to the ball and plate factors

as well), the cone penetrometer is in fact superior to the full-flow penetrometers for

stratification purposes, since the cone resistance is more sensitive to stratigraphic

changes than the tip resistances of the full-flow penetrometers. Additionally, the pore

pressure measured in the cone penetration test also helps identify the stratigraphic

sequence. In contrast, pore pressures measured at the shoulder position (u2 position;

Lunne et al, 1997b) for the full-flow penetrometers, although not presented in the thesis,

were found to be very close to, or even slightly lower (for some depths) than the static

water pressure. This may be due to soil flowing around the full-flow penetrometers

impeding the measurement of pore pressure at the u2 position.

Recommendations for the choice of N values to be adopted for the various shaped

penetrometers have been provided in Table 7.4. However, one should bear in mind that

these recommendations serve merely as a general guide for preliminary assessment of

the undrained shear strength profiles, and may be limited to normally consolidated to

lightly overconsolidated clays with moderate plasticity index (typically in the range 30

CONCLUSIONS

9-4

to 50 %). The N values for other clays of much different characteristics may differ

significantly from the suggested N values given in the table, as have been noted by

DeJong et al (2004) in varved clay and by Long & Gudjonsson (2004) in organic

materials.

At this stage, it is necessary to develop empirical correlations for the full-flow

penetrometers for each site, as for the cone penetrometer. Ultimately, as more empirical

N values from different sites are accumulated in the database, soil characteristics that

cause the variations of N may be clearly identified and thus quantified. For example, by

conducting parallel cone and T-bar penetration tests, there is the potential to interpret

differences in the cone and T-bar factors, perhaps to indicate differences in

overconsolidation ratio or in rigidity index (Randolph, 2004). Also, by performing

parallel T-bar and ball penetration tests, relative values of the T-bar and ball factors may

help quantify the effects of strain softening or infer the anisotropic strength ratio.

However, in order to facilitate these, it is of utmost importance that the shear strength

data used for the empirical correlations are of high quality, in order to minimise

complications due to sample disturbance when evaluating laboratory strengths.

The effect of penetration rate on the tip resistance was investigated, in the light of

variable rate model penetration tests (referred to as ‘twitch’ tests) carried out in the

centrifuge. The twitch test involves pushing the penetrometer into the soil with the

initial (undrained) penetration rate being successively halved over several steps, with

the penetrometer being advanced by a set distance (1 to 2 diameters) at each stage.

It was found that penetration resistances for the various shaped model penetrometers

increased up to 100 % higher than their undrained resistances due to partial

consolidation, as the penetration rate was reduced by about two orders of magnitude.

It has been shown that the consolidation coefficient, cv, of a particular soil can be

estimated by fitting the twitch test data to a ‘backbone’ curve that gives the relationship

of normalised tip resistance against non-dimensional velocity, V = vd/cv, where v is the

penetration rate and d the diameter of the penetrometer. The values of cv derived based

on the T-bar backbone curve from Watson & Suemasa (2000) were similar to the cv

values measured (at very high stress level) in the oedometer tests and in Rowe cell tests,

in addition to the cv values interpreted using data from consolidation of the sample in

centrifuge. Nevertheless, when the T-bar backbone curve derived after Randolph &

Hope (2004) was adopted, the results were more similar to the cv values measured at

CONCLUSIONS

9-5

yield stress in the oedometer tests.

The curve of normalised resistance against non-dimensional velocity for the ball

penetrometer seemed to plot horizontally to the right of the T-bar curve, by a factor of

about 2.5, suggesting that the relative rate of consolidation for the ball to the T-bar is in

a ratio of about 2.5. This may also imply that if the ball diameter is about 2 to 3 times

the T-bar diameter, the actual consolidation rates for the two penetrometers may be

similar. Nevertheless, according to the theoretical analysis presented by Davis & Poulos

(1972), the consolidation rate for a surface footing was primarily dependent on the area

of the footing and was virtually independent of the shape, at least for length to breadth

ratio up to 5 for the rectangular footing. Using the rates of consolidation under the

surface footings as an analogy to the rates of consolidation around the penetrometers,

the similar (‘actual’) consolidation rates observed in the model T-bar and ball twitch test

results were due to the fact that the projected areas of the two model penetrometers were

rather similar (100 mm2 for the model T-bar compared to 111.2 mm2 for the model ball).

However, there was no apparent evidence for the consolidation rate varying with

projected area in the results of model T-bar twitch tests, with areas ranging from 100 to

250 mm2. Further investigation is required to verify the rate of consolidation for the

penetration problems.

A correction factor for viscous effects, expressed as a ratio of hyperbolic functions, was

applied to the ‘original’ backbone curves derived by Watson & Suemasa (2000) and

Randolph & Hope (2004), for which the undrained regions of the curves have been

ignored. The resulting backbone curves can fit the field twitch test data reported by

Schneider et al (2004), reasonably well, but the curves were shifted slightly to the left

from their original positions due to the correction factor. Therefore, it is suggested that

the constants of the backbone curves should be derived after the correction factor for

viscous effects has been applied. Also, the viscous effects can be defined better from the

field tests than from the centrifuge tests, because (vd)field is much greater than

(vd)centrifuge, typically by a factor of 160 for a T-bar penetrometer, hence giving an

‘undrained’ range for the field tests two orders of magnitudes greater than for the

centrifuge tests.

Finally, field cyclic penetration and extraction tests were performed at specific depths

for each full-flow penetrometer. These tests comprised displacement cycles of ±0.5 m

about the relevant depth, recording the penetration and extraction resistances over five

CONCLUSIONS

9-6

full cycles. The results showed that both the penetration and extraction resistances

continued to degrade through the 5 cycles, but at a reducing rate, with the resistances

stabilising at a fully remoulded value. The sensitivity of the soil derived from the cyclic

tests is consistent with that observed in the vane shear tests. The cyclic T-bar test

presented by Long & Gudjonsson (2004), for which the test was performed in a uniform

lightly overconsolidated clay, also indicated that the T-bar resistance stabilised after 5 to

6 cycles. Nevertheless, in varved clay, the results of cyclic tests for the T-bar and ball

penetrometers presented by DeJong et al (2004) showed that the tip resistances

continued to degrade over more than 10 cycles before stabilising at the fully remoulded

values. Additionally, the remoulded value for the T-bar was higher than for the ball and

plate (which were similar) in the varved clay. However, the reasons for the

discrepancies were not identified (DeJong et al, 2004).

9.2 Recommendations for future work

Clearly, further studies are required to understand the characteristics of soft clays more

thoroughly, in order to allow assessments of additional soil properties from the results

of the various shaped penetrometer tests.

In the short term, the following studies are suggested:

• Validate the N values derived from the laboratory tests performed on the tube

samples collected from the Burswood site, because the overall quality of the tube

samples was moderate to poor by NGI’s criterion (Lunne et al, 1997a), as already

noted in Chapter 5. High quality Sherbrooke block samples were obtained recently

from the same site. Therefore, the effects of sample disturbance on the tube samples

may be assessed by comparing the results obtained for the tube samples with that

for the block samples.

• Investigate the locations for pore pressure measurement for the full-flow

penetrometers. Although not presented in the thesis, the pore pressure measured at

the shoulder position (u2 position; Lunne et al, 1997b) for the full-flow

penetrometers was found to be very close to the static water pressure, which is not

particularly useful. Alternative locations such as at the face of the penetrometer (u1

position) and behind the friction sleeve (u3 position; Lunne et al, 1997b) are

suggested for measuring the pore pressure.

• Confirm the position of the T-bar ‘backbone’ curve. Although there is no

CONCLUSIONS

9-7

theoretical basis at this stage, it is believed that the backbone curve that gives the

relationship between normalised resistance against non-dimensional velocity is

unique for each penetrometer. However, as has been observed, the T-bar backbone

curve derived by Watson & Suemasa (2000) was found to differ from that derived

by Randolph & Hope (2004). It is necessary to understand the reason causing such

discrepancy in the experimental results from the two independent studies.

Furthermore, it may be required to examine if the backbone curve derived from the

centrifuge model tests is the same as that derived from the field tests.

• Confirm the trends of the T-bar and cone backbone curves in the viscous region by

means of field penetration tests, preferably undertaken at other clay sites.

• Establish a backbone curve for the ball penetrometer, so that drainage conditions

for the ball penetration test at a given rate can be assessed. Additionally, the

relative rate of consolidation for the ball to the T-bar needs to be verified.

In the long term, the following studies are suggested:

• Continue collecting N data for the various shaped penetrometers, from different

sites, particularly for sites containing varved clays, moderate to heavily

overconsolidated clays, extremely high plasticity (IP > 100 %) clays etc, in order to

provide a thorough guide for the choice of N values to be adopted for the different

penetrometers for different types of sediments.

• Incorporate soil characteristics such as strain rate, strain softening and strength

anisotropy in the theoretical analysis, so that their combined effects can be

assessed. The theoretical solutions must be improved through continuing

verification and calibration with empirical data. Ultimately, the aim is to compute

appropriate N values to be adopted for the various penetrometers, based on sound

theoretical analysis, to evaluate the shear strength profiles required for design

purposes, with the effects of strain rate, strain softening and strength anisotropy

clearly addressed. Furthermore, only when each of the soil characteristics is clearly

understood may it be possible to quantify confidently the effects of these soil

characteristics and interpret additional soil properties such as rigidity index or

overconsolidation ratio, from parallel in situ penetrometer tests conducted using

different types of penetrometers.

• Collaborate with the industry to investigate and carry out in situ twitch tests. It has

CONCLUSIONS

9-8

been shown that the twitch test data, in conjunction with a backbone curve, can be

used to deduce a value for the in situ consolidation coefficient. However, there are

some challenges in the operation of the twitch test that require thorough

consideration, before it can be deployed, especially offshore. For example, one of

the main practical challenges is to modify commercial equipment to allow varying

penetration rate by 2 or 3 orders of magnitude.

REFERENCES

R1

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TABLES

Parameters Type of test and source

a b c m

Cone: Randolph & Hope (2004) 1 2.65 0.90 0.83

Cone: After Randolph & Hope (2004)* 1 2.65 1.5 0.83

T-bar: Randolph & Hope (2004) 1 2.77** 0.84 1.23

T-bar: After Randolph & Hope (2004)* 1 2.1 1.1 1.6

T-bar: House et al (2001) 1 2.77** 2.47 1.30

T-bar: Watson & Suemasa (2000) 1 2.77 0.57 1.45

Note:

* Constants derived based on data from Randolph & Hope (2004), without correction of viscous effects

** Drained resistance of 3.77 times undrained resistance assumed, based on Watson & Suemasa (2000)

Table 2.1 Summary of constants derived for the backbone curves

TABLES

Item Calibration factor

Load cell 0.9914 MPa/bit

Pore pressure transducer 1.0336 kPa/bit

Unequal area ratio, α 0.699

Table 4.1 Calibration details for field penetrometer test apparatus

TABLES

Borehole Tube type External diameter, dE*

Tube thickness, t* dE/t *

Outside cutting edge

angle** (mm) (mm) (°)

BH1 Stainless steel 76 1.5 50.7 15

BH1 Plastic 76 2.5 30.4 15

BH2 Stainless steel 102 2 51.0 7

BH2 Plastic 111 3.5 31.7 9

Note:

1. All sampling tubes have a length of 750 mm and zero inside clearance

* Ladd & DeGroot (2003) recommended a minimum external diameter of 76 mm and tube thickness such that dE/t > 45 be used

** Hight & Leroueil (2003) suggested outside cutting edge angle of 5° be used

Table 5.1 Dimensions of sampling tubes

TABLES

Index properties Water content

Overconsolidation ratio Coefficient of consolidation Compression index

Depth Unit weight ωi ωf σ'v σ'yield OCR cvy

cv at 5xσ'yield

Cc Cc/(1+eo) ei Δe/ei Sample quality*

Type of test

(m) (kN/m3) (%) (%) (kPa) (kPa) (m2/yr) (m2/yr) CRSC 4 18.35 14.6 23 70 95.4 164.7 1.727 4.00 0.652 0.430 0.599 0.135 3

CRSC 5 11.40 15.7 58 36 59.4 91.0 1.530 5.91 0.50 0.689 0.284 1.531 0.069 2

CRSC 6 5.97 14.2 81 47 33.9 51.6 1.522 1.11 0.50 1.030 0.337 2.156 0.044 2

CRSC 7 6.33 14.3 98 57 35.3 68.5 1.938 4.23 1.15 1.339 0.374 2.619 0.013 1

CRSC 9 12.11 16.0 49 57 63.4 91.0 1.435 1.22 0.73 0.536 0.242 1.315 0.080 3

CRSC 10 10.00 14.9 68 16 52.1 109.9 2.111 1.25 0.59 0.982 0.365 1.804 0.063 3

CRSC 11 10.00 14.9 69 36 52.1 103.3 1.983 1.17 0.67 0.964 0.361 1.824 0.084 3

CRSC 12 8.22 15.1 68 44 43.5 85.1 1.958 0.96 0.49 0.825 0.309 1.802 0.073 3

CRSC 13 8.22 15.0 69 48 43.5 70.2 1.614 1.00 0.74 0.822 0.302 1.824 0.057 2

CRSC 14 13.75 15.3 57 40 73.0 130.0 1.781 2.80 0.827 0.341 1.505 0.052 2

CRSC 15 14.94 15.4 64 46 79.7 106.3 1.333 2.48 0.859 0.337 1.699 0.091 3

CRSC 16 4.11 14.0 98 69 26.4 54.6 2.064 2.82 <1.2 1.392 0.398 2.601 0.039 2

CRSC 17 2.54 14.6 103 53 20.1 63.5 3.152 1.80 1.19 1.203 0.331 2.740 0.040 2

CRSC 18 11.40 16.2 56 36 59.4 95.6 1.608 1.73 0.87 0.715 0.295 1.494 0.050 2

Note: * Sample quality rated based on NGI’s criterion (refer to Table 5.3)

Table 5.2 Summary of CRSC tests

TABLES

Δe/ei** OCR

1* 2 3 4

1 - 2 <0.04 0.04 - 0.07 0.07 - 0.14 >0.14

2 - 4 <0.03 0.03 - 0.05 0.05 - 0.10 >0.10

4 - 6 <0.02 0.02 - 0.035 0.035 - 0.07 >0.07

Note: * Sample quality rated based on NGI’s criterion (Lunne et al, 1997a)

1 – Very good to excellent;

2 – Good to fair;

3 – Poor;

4 – Very poor.

** Δe/ei, Δe = ei – eo or εv (1 + ei) and ei = Gsωi

Table 5.3 NGI’s criterion for sample quality

TABLES

Index properties Atterberg limits Loading Water content Depth Unit

weight ωi ωf ωL ωP Plasticity

index εaf su Type of test

(m) (kN/m3) (%) (%) (%) (%) (%) (%) (kPa) UU 1 3.59 - 69 75 59 28 31 15.0 7.9 UU 2 2.38 13.4 119 112 81 38 43 6.2 18.6 UU 3 3.77 15.4 71 68 52 24 28 8.2 15.1 UU 4 8.09 14.3 74 89 74 30 44 15.0 9.7 UU 5 12.32 15.8 52 53 63 25 38 10.5 19.2 UU 6 13.52 15.3 61 59 63 27 36 6.1 31.2 UU 7 11.27 15.1 61 59 68 28 40 12.8 20.4 UU 8 9.83 14.7 72 72 72 32 40 7.4 23.0 UU 9 18.20 14.7 62 75 72 31 41 3.4 48.2 UU 10 14.52 15.4 55 55 63 31 32 3.9 46.1 UU 11 16.74 14.7 63 68 66 32 34 4.2 54.6 UU 12 4.60 14.0 95 96 75 34 41 8.8 15.2 UU 13 6.79 13.9 89 88 75 37 38 7.8 18.0

Check Test 1* 6.33 93 86 37 49

Note: * Additional test using fresh specimen to justify values of initial water content and Atterberg limits

Table 5.4 Summary of UU tests

TABLES

Index properties Consolidation Loading Water content Depth Unit

weight ωi ωf σ'v σ'h εv εa εaf Δuf su Type of

test

(m) (kN/m3) (%) (%) (kPa) (kPa) (%) (%) (%) (kPa) (kPa)

Δe/ei* Sample quality**

CAU TXC 1 5.60 13.9 91 81 32.5 26.0 2.6 1.7 6.0 18.2 17.8 0.037 2 CAU TXC 2 6.06 13.9 94 79 34.4 27.5 3.6 2.6 6.2 21.2 21.0 0.050 2 CAU TXC 4 15.90 13.9 84 86 83.1 66.5 7.1 2.2 3.3 44.2 43.8 0.103 3 CAU TXC 5 17.11 14.7 69 66 90.1 72.1 4.1 2.2 3.8 50.1 49.8 0.063 2

CAU TXE 1 5.43 14.1 89 75 31.8 25.4 7.8 3.6 -12.4 -11.5 26.0 0.111 4 CAU TXE 2 15.71 14.5 73 77 82.0 65.6 3.5 1.9 -8.2 -12.7 29.3 0.053 2 CAU TXE 3 17.31 14.8 71 72 91.3 73.0 2.9 1.5 -6.3 -8.8 31.0 0.044 2 CAU TXE 4 5.26 13.8 87 89 31.1 24.9 4.3 2.4 -12.2 -9.4 10.0 0.062 3

Note: * Δe/ei computed at the end of consolidation

** Sample quality rated based on NGI’s criterion (refer to Table 5.3)

Table 5.5 Summary of CAU triaxial tests

TABLES

Index properties Consolidation Loading Water content Depth Unit

weight ωi ωf σ'v σ'h εv εa γf Δuf su Type of

test

(m) (kN/m3) (%) (%) (kPa) (kPa) (%) (%) (%) (kPa) (kPa)


CAU SS 1 5.89 14.1 97 82 33.7 26.9 3.5 2.6 28.5 12.5 13.5 0.049 2 CAU SS 3 5.81 14.2 84 86 33.4 26.7 10.9 5.8 16.7 9.1 7.0 0.158 4 CAU SS 4 14.94 14.5 63 67 77.7 62.2 10.6 8.7 16.5 20.1 28.0 0.170 4 CAU SS 5 14.94 15.1 61 62 77.7 62.2 9.4 8.1 11.6 21.5 28.6 0.152 4 CAU SS 6 4.77 14.5 109 29.1 23.3 4.0 3.0 24.0 7.2 20.7 0.054 3 CAU SS 9 10.51 14.5 72 64 54.6 43.7 12.7 10.9 22.5 15.6 24.1 0.193 4

CAU SS 10 12.66 15.6 59 58 65.5 52.4 5.8 4.3 11.1 17.8 28.5 0.095 3 CAU SS 11 9.42 14.3 81 77 49.4 39.5 8.1 6.6 21.8 13.4 20.8 0.119 3

Note: * Δe/ei computed at the end of consolidation

** Sample quality rated based on NGI’s criterion (refer to Table 5.3)

Table 5.6 Summary of CAU simple shear tests

TABLES

Index properties Consolidation Penetration Water content Depth Unit

weight ωi ωf σ'v (=σ'h) εv εa qin qout |qout/qin| su Test

No. (m) (kN/m3) (%) (%) (kPa) (%) (%) (MPa) (MPa) (kPa)


T-bar 1 15.00 15.2 62 55 68 5.5 2.6 0.37 35.2 0.089 3 T-bar 2 13.50 15.7 52 49 61 4.8 2.0 0.55 -0.11 0.20 52.4 0.083 3 T-bar 3 17.29 14.6 78 62 79 7.6 3.6 0.78 -0.19 0.24 74.3 0.113 3 T-bar 4 12.50 15.2 56 52 56 4.9 2.4 0.36 -0.26 0.72 34.3 0.082 3 T-bar 5 9.75 14.4 75 62 45 6.1 2.2 0.42 -0.11 0.26 40.0 0.092 3 T-bar 6 4.25 14.3 77 68 23 5.3 3.6 0.29 -0.09 0.31 27.6 0.079 3

Note: Isotropic consolidation pressure (σ'v = σ'h) is equal to the mean effective stress (p') at the sample depth

* Δe/ei computed at the end of consolidation

** Sample quality based on NGI’s criterion (refer to Table 5.3)

Table 5.7 Summary of T-bar in triaxial tests

TABLES

Index properties Consolidation Loading Type of Depth Unit In situ Yield Water content Stage 1 Stage 2 OCR

test weight σ'v σ'v ωi ωf σ'v1 σ'h1 σ'v2 σ'h2 σ'v1/σ'v2 εaf or γf* su 1** su 2

**

(m) (kN/m3) (kPa) (kPa) (%) (%) (kPa) (kPa) (kPa) (kPa) (%) (kPa) (kPa) SHANSEP

CAU TXC 6 6.60 14.1 36.5 68.9 99 75 78.7 44.7 44.4 36.4 1.773 5.8 32.3 27.9 CAU TXC 7 10.86 14.9 56.7 94.1 69 57 103.2 63.0 65.2 52.8 1.583 2.8 37.0 33.4

CAU TXE 5 7.79 14.1 41.6 73.7 95 74 82.2 48.5 53.2 42.3 1.545 -21.1 23.2 20.2 CAU TXE 7 10.68 14.7 55.6 92.7 70 60 107.8 63.5 67.8 53.8 1.590 -12.1 28.8 24.5

CAU SS 7 7.45 13.7 40.0 72.2 77 66 78.0 45.6 49.3 39.6 1.582 19.3 29.5 26.6 CAU SS 8 12.99 15.6 68.5 113.1 53 53 120.5 74.3 82.7 66.0 1.457 16.1 43.0 39.4

Note: * εaf is the axial strain at failure for triaxial sample; γf is the shear strain at failure for simple shear sample

** su 1 is the measured value after SHANSEP procedure is followed

su, 2 is the value adjusted to correspond to in situ stress level

Table 5.8 Summary of testing undergone SHANSEP procedure

TABLES

Value Soil property

Box 1 Box 2

Effective unit weight, γ' (kN/m3) 4.4 5.1

Reconstituted friction angle, φ'centrifuge (°) 32 29

Moisture content, ω (%) 66~86 55~67

Liquid limit, ωL (%) 98 98

Plastic limit, ωP (%) 39 39

Plasticity index, IP (%) 59 59

Coefficient of consolidation, cv (m2/yr) ∼0.5 ∼0.75

Compression index, Cc ~0.82 0.66~0.74

Particle size < 0.02 mm (%) 75 75

Particle size < 0.006 mm (%) 40 40

Clay content (< 0.002 mm) (%) 10 10

Table 6.1 Soil properties for reconstituted Burswood clay

TABLES

Parameter Scaling factor (model/prototype)

Length 1/n

Acceleration n

Density 1

Stress 1

Mass 1/n3

Force 1/n2

Time (consolidation) 1/n2

Table 6.2 Scaling relationships for centrifuge models

TABLES

Penetrometer type Projected area Offset per 100 kPa change in uo

(mm2) (kPa)

T-bar (20 mm x 5 mm) 100 19.63

T-bar (30 mm x 5 mm) 150 13.09

T-bar (40 mm x 5 mm) 200 9.82

T-bar (50 mm x 5 mm) 250 7.85

Ball (diameter = 11.9 mm) 111.2 17.65

Plate (diameter = 11.2 mm) 98.5 19.92

Table 6.3 Offset of tip resistance due to error generated from changes in normal

stress on the load cell

TABLES

Index properties Water content Overconsolidation ratio Coefficient of

consolidation Compression index Prototype depth range

Unit weight ωi ωf σ'v σ'yield OCR cvy

cv at 5xσ'yield

Cc Cc/(1+eo) Type of test

(m) (kN/m3) (%) (%) (kPa) (kPa) (m2/yr) (m2/yr) Box 1

CF 1 CRSC 1 3.3~5.8 14.8 79 48 14.5~25.5 42.2 2.91~1.65 1.5 0.6 0.817 0.395

CF 1 CRSC 2 11.3~13.8 15.1 71 47 49.7~60.72 73.9 1.49~1.22 1.1 0.5 0.826 0.305

Box 2

CF 2 CRSC 3 1.3~3.8 15.6 65 39 6.6~19.4 58.0 8.8~2.99 1.1 0.5 0.691 0.257

CF 2 CRSC 4 11.7~14.2 15.8 59 34 59.7~72.4 85.9 1.44~1.19 1.2 0.7 0.676 0.275

CF 2 CRSC 5 6.6~9.1 15.7 66 39 33.7~46.4 63.5 1.88~1.37 1.0 0.5 0.741 0.278

CF 2 CRSC 6 16.1~18.6 16.1 55 36 82.1~94.9 116.4 1.42~1.23 0.7 0.5 0.659 0.279

Table 6.4 Summary of CRSC tests on samples from centrifuge testing

TABLES

Index properties Consolidation Loading Water content Prototype

depth range Unit weight ωi ωf

σ'v* σ'h εv εa εa or γf Δuf su Type of test

(m) (kN/m3) (%) (%) (kPa) (kPa) (%) (%) (%) (kPa) (kPa) CF Box 1 CF 1 TXC 1 0~18.9 14.7 75 60 83.3 66.7 8.6 5.9 2.7 36.6 31.5 CF 1 TXE 1 1.4~19 14.7 73 65 83.6 66.9 7.8 4.6 -7.6 -8.0 18.5 CF 1 SS 1 6.8~10.8 14.6 83 75 47.5 38.0 6.2 4.6 3.2 7.6 22.2 CF 1 SS 2 14.9~18.9 14.9 66 66 83.1 66.5 8.0 6.3 4.4 15.0 34.5 CF Box 2 CF 2 TXC 1 1.0~20 15.4 63 53 102.0 81.6 7.3 4.8 2.5 39.6 34.3 CF 2 TXE 1 2.7~20 15.7 64 51 102.0 81.6 7.9 4.8 -12.6 -4.1 27.8 CF 2 SS 1 4.3~8.3 15.3 64 63 42.3 33.9 3.8 2.9 9.7 14.3 17.0 CF 2 SS 2 14.7~18.7 15.6 56 55 95.4 76.3 7.2 6.1 11.0 29.0 34.5 CF 2 SS 3 1.4~5.4 15.1 63 67 27.5 22.0 3.2 2.2 4.9 7.8 13.8 CF 2 SS 4 10.5~14.5 15.4 62 59 74.0 59.2 6.8 4.8 6.7 20.8 25.2

Note: * Test samples were consolidated to the stress levels experienced in the centrifuge at the sample base

Table 6.5 Summary of triaxial and simple shear tests on samples from centrifuge testing

TABLES

CAU triaxial compression CAU simple shear Average*

Depth (m)

su, TXC (kPa)

Depth (m)

su, SS (kPa)

Depth (m)

su, av* (kPa)

5.6 17.8 5.9 13.5 5.6 14.3

6.1 21 9.4 20.8 6.1 15.3

15.9 43.8 10.5 24.1 15.9 35.5

17.1 49.8 12.7 28.5 17.1 37.9

14.9 28

14.9 28.6

Note: * Average undrained shear strength, su, av = (su, TXC + su, TXE + su, SS)/3, where su, TXC , su, TXE and su, SS are the shear strengths from triaxial compression, triaxial extension and simple shear tests respectively

Table 7.1 Laboratory shear strengths adopted for calculating N values for the field

penetrometers

TABLES

Nkt, TXC Nkt, av Nkt, SS Nkt, vane

Range Mean ± SD3 Range Mean ± SD Range Mean ± SD Range Mean ± SD

Theoretical Nkt

4

(a)1 9.5 − 10.9 10.3 ± 0.8 12.6 − 13.6 13.2 ± 0.4 13.2 − 16.5 14.6 ± 1.5 10.0 − 14.6 12.5 ±1.5 12.5 − 13.0 Burswood

(b)1 7.6 − 10.0 8.5 ± 1.1 10.0 − 12.4 10.9 ± 1.1 10.1 − 12.8 11.7 ± 1.1 8.6 − 11.8 10.3 ± 1.0 12.5 − 13.0

Onsøy2 10.4 − 13.8 12.0 ± 1.7 15.9 −17.5 16.5 ± 0.9 16.8 − 18.7 17.6 ± 1.0 13.7 − 19.3 16.3 ± 1.6 13.6

Laminaria2 9.8 − 12.9 11.2 ± 1.2 N/A N/A 12.9 − 15.3 13.7 ± 1.0 9.8 − 13.9 12.3 ± 1.3 13.4 − 13.7

Centrifuge (Burswood clay) N/A N/A N/A N/A N/A N/A 5.5 − 8.6 7.1 ± 1.1 N/A

Note: 1. Nkt values are derived using cone resistance profiles (a) presented in Figure 4.17; (b) reported by Schneider et al (2004)

2. Results for Onsøy and Laminaria clays are quoted from the report of NGI-COFS (2004a)

3. Standard deviation, SD, is calculated using Equation 7.1

4. Theoretical Nkt values are calculated using Equation 7.2 (Lu et al, 2004)

Table 7.2 Summary and comparison of cone factors (Nkt) for various clay sites

TABLES

NT-bar, TXC NT-bar, av NT-bar, SS NT-bar, vane

Range Mean ± SD5 Range Mean ± SD Range Mean ± SD Range Mean ± SD

(a)1 7.4 − 11.5 8.9 ± 1.5 9.7 − 14.3 11.5 ± 1.8 9.8 − 15.1 11.6 ± 1.5 9.1 − 11.9 10.6 ± 0.7 Burswood

(b)2 7.8 − 8.6* 8.2 ± 0.6* 10.7* 10.7* 9.2 − 12.1 10.2 ± 1.3 7.8 − 10.2 9.2 ± 0.8

Onsøy3 7.5 − 9.9 8.5 ± 1.2 11.1 − 13.1 11.9 ± 1.1 11.4 − 13.6 12.5 ± 1.1 10.2 − 15.0 11.6 ± 1.5

Laminaria3 8.6 − 12.1 10.2 ± 1.6 N/A N/A 11.4 − 14.3 12.4 ± 1.3 9.0 − 12.8 11.3 ± 1.5

Centrifuge4 (Burswood clay) N/A N/A N/A N/A 10.2 − 12.0 11.0 ± 0.5 8.7 − 11.1 9.8 ± 0.7

Note: 1. NT-bar ranges and mean values are the overall results for T-bars (250 mm x 40 mm) and (160 mm x 40 mm)

2. NT-bar values are derived using T-bar (250 mm x 40 mm) results reported by Schneider et al (2004)

3. Results for Onsøy and Laminaria clays are quoted from the report of NGI-COFS (2004a)

4. These are the overall results for all model T-bars tested in the centrifuge

5. Standard deviation, SD, is calculated using Equation 7.1

* Only two data points from depth range 5.6 to 6.1 m are available

Table 7.3 Summary and comparison of T-bar factors (NT-bar) for various clay sites

TABLES

NTXC Nav NSS Nvane4

Empirical range

Recommended value ± 2SD3

Empirical range

Recommended value ± 2SD

Empirical range


Empirical range


Cone1 7.6 − 13.8 10.5 ± 3.1 10.0 − 17.5 13.6 ± 5.7 10.1 − 18.7 14.4 ± 5.0 8.6 − 19.3 12.9 ± 5.1

T-bar1 7.4 − 12.1 9.2 ± 1.8 9.7 − 14.3 11.7 ± 0.6 9.2 − 15.1 11.9 ± 1.4 7.8 − 15.0 10.7 ± 2.1

Ball2 7.0 − 10.0 8.5 ± 1.7 9.2 − 12.5 10.9 ± 2.2 9.7 − 13.7 11.2 ± 2.2 8.8 − 11.0 10.1 ± 2.0

Plate2 7.7 − 11.8 9.4 ± 1.9 9.9 − 14.8 12.1 ± 2.4 10.8 − 15.7 12.5 ± 2.5 10.1 − 12.3 11.4 ± 2.3

Note: 1. The empirical ranges are the total ranges of N collected for Burswood, Onsøy and Laminaria clays

2. Only data from Burswood are available, so the standard deviation (SD) is calculated by assuming 10 % coefficient of variance (COV)

3. SD is calculated using Equation 7.1

4. Nvane data for the centrifuge model penetrometers are not included in the empirical ranges and recommended values

Table 7.4 Summary of bearing factors (N) for the various shaped penetrometers

FIGURES

Burswood site

Burswood site

Note: maps are extracted from www.whereis.com.au

Figure 3.1 Map of test location at Burswood

FIGURES

TB2

See Figure 4.1 (b) for details

TB1

C2

P1

P2

V4

V5STB1

STB2

BH1

BH2

TB3

TB4

B1

B2

C3

C4

ElectricalPole

ReferencePoint

16.6 m

MAIN TESTING AREA

25 m

C1

62m to Graham Farmer Freeway

~40m to Swan River

N

Figure 4.1 (a) Test location layout: Location of main testing area

FIGURES

0

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34

Distance from Reference Point (m)

Dis

tanc

e fr

om R

efer

ence

Poi

nt (m

)

TB1,TB2:Smooth T-bar TB3,TB4:Rough T-bar STB1,STB2:Small T-bar P1,P2:Plate

B1,B2:Ball C2,C3,C4:Cone V4,V5:Vane BH1,BH2:Borehole

B2

B1

TB4

TB3

P2

P1

STB1

STB2

TB1 V4

C2

BH2

BH1

TB2

N

~40m to Swan River

C3

C4

V5

Figure 4.1 (b) Test location layout: Details of main testing area

FIGURES

Figure 4.2 Diagram of field cone penetrometer

~36 mm

35.7 mm (AT)

Water seal

15.80 mm

Screw thread

25.10 mm (AN)

Small hole

~30 mm

9.5 mm

Connection rod

u2 position

FIGURES

Figure 4.3 Diagram of field T-bar penetrometer

250 mm

15.80 mm

25.10 mm

Connection rod

Screw thread

Diameter: 38.9 mm

Sand blasted surface cylinder bar Machine

smoothed for both ends

Water seal

Small holeu2 position

FIGURES

Cone

Plate

T-bar (250x40)

Ball

Note: The porous elements have not been assembled on the cone penetrometer

Figure 4.4 Photograph of field cone, T-bar, ball and plate penetrometers

FIGURES

Figure 4.5 Diagram of shear vane

60 mm

Vane blade

2.1 mm

130 mm

60 mm

Connection to slip

coupling

29 mm

17.3 mm

45°

45°

59 mm

25.4 mm

FIGURES

Figure 4.6 Photographs of truck and saturating a penetrometer

FIGURES

Figure 4.7 Photograph of frame for jacking the shear vane

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8Tip resistance, q (MPa)

Dep

th (

m)

Cone 3 (meas'd)Cone 3 (total)Cone 3 (net)

T-bar 3 (meas'd)T-bar 3 (net)

σvo

Figure 4.8 Corrections of field cone and T-bar data

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.1 0.2 0.3 0.4 0.5 0.6Tip resistance, qin (MPa)

Dep

th (

m)

T-bar 1 (smooth)

T-bar 2 (smooth)

T-bar 3 (rough)

T-bar 4 (rough)

T-bar* (rough)

Note: * T-bar resistance profile reported by Schneider et al (2004)

Figure 4.9 (a) Comparison of smooth and rough T-bar penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0Tip resistance, qout (MPa)

Dep

th (

m)

T-bar 1 (smooth)

T-bar 2 (smooth)

T-bar 3 (rough)

T-bar 4 (rough)

* Abrupt reductions where cyclic tests were performed

Note: * Cyclic penetration and extraction test results were extracted and will be presented later

Figure 4.9 (b) Comparison of smooth and rough T-bar extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0.6 0.7 0.8 0.9 1 1.1 1.2 1.3Ratio smooth/rough T-bar resistance

Dep

th (

m)

PenetrationExtraction

Figure 4.10 Ratios of average smooth to rough T-bar resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Smaller T-bar 1

Smaller T-bar 2

T-bar 3

T-bar 4

Figure 4.11 (a) Comparison of smaller and standard T-bar penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Smaller T-bar 1

Smaller T-bar 2

T-bar 3

T-bar 4

Depths where cyclic tests were

conducted

Figure 4.11 (b) Comparison of smaller and standard T-bar extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Cone 1Cone 2Cone 3Cone 4SCone*T-bar 3T-bar 4

Note: * Cone profile obtained from a seismic cone reported by Schneider et al (2004)

Figure 4.12 (a) Comparison of cone and T-bar penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Cone 3

Cone 4

T-bar 3

T-bar 4


conducted

Figure 4.12 (b) Comparison of cone and T-bar extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-2 -1 0 1 2 3 4 5 6Friction ratio (%)

Dep

th (

m)

Cone 1

Cone 2

Cone 3

Cone 4

Penetration

Extraction

Figure 4.13 Comparison of friction ratios from cone tests

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.4 -0.2 0 0.2 0.4 0.6 0.8 1Bq values

Dep

th (

m)

Cone 1Cone 2Cone 3Cone 4SCone*

Penetration

Extraction

Note: * Bq profile obtained from a seismic cone reported by Schneider et al (2004)

Figure 4.14 Comparison of Bq values from cone tests

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Ball 1

Ball 2

Ball*

T-bar 3

T-bar 4

Note: * Ball resistance profile reported by Schneider et al (2004)

Figure 4.15 (a) Comparison of ball and T-bar penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Ball 1

Ball 2

T-bar 3

T-bar 4


conducted

Figure 4.15 (b) Comparison of ball and T-bar extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Plate 1

Plate 2

T-bar 3

T-bar 4

Figure 4.16 (a) Comparison of plate and T-bar penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Plate 1

Plate 2

T-bar 3

T-bar 4


conducted

Figure 4.16 (b) Comparison of plate and T-bar extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.4 -0.2 0 0.2 0.4 0.6Tip resistance, qnet (MPa)

Dep

th (

m)

Average ConeAverage T-barAverage ST-barAverage BallAverage Plate

Figure 4.17 Summary of tip resistance profiles for all various penetrometers

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1 1.2Ratio qout/qin

Dep

th (

m)

Average ConeAverage T-barAverage ST-barAverage BallAverage Plate


conducted

Figure 4.18 Comparison of ratios of extraction to penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 10 20 30 40 50Undrained shear strength, su (kPa)

Dep

th (

m)

Vane 4: Peak

Vane 4: Remoulded

Vane 5: Peak

Vane 5: Remoulded

Figure 4.19 Peak and remoulded shear strengths from field vane tests

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 2 4 6 8 10Sensitivity, St

Dep

th (

m)

Vane 4Vane 5Average

Figure 4.20 Sensitivity of clay from field vane tests

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

ConeT-barSmaller T-barBallPlateVane: PeakVane: Remoulded

Note: Bearing factor of N = 10.5

used for all penetrometers

Figure 4.21 Comparison of undrained shear strength profiles

FIGURES

8.89

9.29.49.69.810

10.2

-0.3 -0.2 -0.1 0 0.1 0.2 0.3Tip resistance, q (MPa)

Dep

th (m

)

0

0.2

0.4

0.6

0.8

1

1 2 3 4 5 6Cycle number

Deg

rada

tion

fact

or


FIGURES

3.53.73.94.14.34.54.74.9


Dep

th (m

)

8.58.78.99.19.39.59.79.9


Dep

th (m

)

Figure 4.23 (a) Cyclic penetration response for T-bar 3 (4 m and 9 m)

FIGURES

13.513.713.914.114.314.514.714.9

-0.4 -0.2 0 0.2 0.4Tip resistance, q (MPa)

Dep

th (m

)

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

4.25 m9.25 m14.25 m

Figure 4.23 (b) Cyclic penetration response for T-bar 3 (14 m and summary)

FIGURES

3.53.73.94.14.34.54.74.9


Dep

th (m

)

13.513.713.914.114.314.514.714.9


Dep

th (m

)

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

4.30 m14.30 m


FIGURES

3.53.73.94.14.34.54.74.9


Dep

th (m

)

13.513.713.914.114.314.514.714.9


Dep

th (m

)

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

4.25 m14.25 m


FIGURES

3.73.94.14.34.54.74.95.1


Dep

th (m

)

13.713.914.114.314.514.714.915.1


Dep

th (m

)

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

4.35 m14.35 m


FIGURES

3.63.8

44.24.44.64.8

5


Dep

th (m

)

13.613.8

1414.214.414.614.8

15


Dep

th (m

)

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

4.30 m14.30 m


FIGURES

3.63.8

44.2

4.44.64.8

5


Dep

th (m

)

13.613.8

1414.214.414.614.8

15


Dep

th (m

)

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

4.25 m14.25 m


FIGURES

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or

T-bar 3 T-bar 4ST-bar 1 ST-bar 2Ball 1 Ball 2Plate 1 Plate 2

Depth: 4.3 m

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or


Depth: 14.3 m

Figure 4.31 Summary of degradation parameters for cyclic penetrometer tests

FIGURES

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or


Depth: 4.3 m

0

0.2

0.4

0.6

0.8

1


Deg

rada

tion

fact

or


Depth: 14.3 m

Figure 4.32 Smoothed degradation curves for cyclic penetrometer tests

FIGURES

(a) Sample from depth 8.40 - 9.15 m

(b) Sample from depth 12.90 – 13.65 m

Figure 5.1 X-ray of tube samples collected from the field

Cluster of large shell pieces

Visible cracks

Tiny shell, no cluster of shells

Potential cracks

FIGURES

Figure 5.2 Schematic diagram of CRSC apparatus

SAMPLE

Pore pressure transducer

Cutting ring Porous discs

Burette

Valve closed during testing

O-rings

Piston

Load cell

Top drainage

Bottom drainage

FIGURES

Figure 5.3 Schematic diagram of triaxial apparatus

Top drainage

Bottom drainage

SAMPLE

Membrane

Loading ram

Top cell

External LVDT

Top cap

Top O-rings

Cell shroud

Triaxial pedestal

Bottom O-rings

To air-water interface cylinder

Loading frame

Top porous disc

Bottom porous disc

Internal load cell

FIGURES

Figure 5.4 Schematic diagram of simple shear apparatus

SAMPLE Membrane

Base carriage Roller bearing Internal load cells

Horizontal motor

Vertical motor

Horizontal LVDT

Roller bearing

Vertical LVDT

Roller bearing

Sleeve

FIGURES

Figure 5.5 Schematic diagram of T-bar in triaxial apparatus

6 mm

25.4 mm (1 inch)

Triaxial cell

SAMPLE

External LVDT

Submersible LVDTs

Base Pedestal

Porous discs

Top cap

T-bar penetrometer (180 mm long)

Top drainage

Plug sleeve

Penetrometer sleeve (200 mm long)

Electrical connection for penetrometer

Electrical connection for LVDT

O-rings

Membrane

FIGURES

(a) Laboratory T-bar penetrometer

(b) T-bar tip

Figure 5.6 Photographs of T-bar penetrometer for T-bar triaxial test

FIGURES

10 12 14 16 18Unit weight, γ (kN/m3)

0

2

4

6

8

10

12

14

16

18

20

0 20 40 60 80 100 120Water content (%)

Dep

th (m

)Initial moisture

ωP ωL

Weathered crust,

clay layer in between

Soft clay,large shell fragments dominant

Soft clay,less and

smaller shell fragments

Silty clay andsand layer

Figure 5.7 Natural water content, Atterberg limits and unit weight profiles

FIGURES

75 μ

m

150

212

300

425

600

1.18

mm

2.36

4.75

0

10

20

30

40

50

60

70

80

90

100

0.0001 0.001 0.01 0.1 1 10Particle size (mm)

% P

assi

ng0

10

20

30

40

50

60

70

80

90

100

% R

etai

ned

CLAY FRACTION SILT FRACTION SAND FRACTIONFINE COARSE FINE MEDIUM COARSE

0.002 0.02 0.06 0.2 0.6 2

GRAVEL FRACTION

MEDIUM

0.006 Note: Classification of particle size is based on Australian Standard (AS 1289, 2000)

Figure 5.8 Grading curve for Burswood clay material

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 50 100 150 200 250 300 350Stress (kPa)

Dep

th (m

)

Total stress

Pore pressure

Effective stress

For lab testing

Figure 5.9 In situ vertical stress profiles

FIGURES

0 50 100 150 200

Effective stress, σ'v (kPa)

Vertical stress profile

Yield stress profile

Measured yield stress

0

2

4

6

8

10

12

14

16

18

20

0 1 2 3 4

Overconsolidation ratio, OCR

Dep

th (m

)

OCR = 4e-0.4Depth+1.65

Figure 5.10 OCR and yield stress profiles

FIGURES

0 0.2 0.4 0.6

Compression ratio, Cc/(1+eo)

0

2

4

6

8

10

12

14

16

18

20

0 0.5 1 1.5

Compression index, Cc

Dep

th (m

)

Figure 5.11 Compression index and compression ratio profiles

FIGURES

0 1 2 3 4cv (m

2/year)

At yield stress (Cvy)Cv at 5 yield stress

0

2

4

6

8

10

12

14

16

18

20

0 1 2 3

Void ratio, e

Dep

th (m

)Initial void ratio

In situ void ratio

Figure 5.12 Void ratio and consolidation coefficient profiles

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1

T-bar resistance, qin (MPa)

Dep

th (m

)

Field T-barField smaller T-barLaboratory T-bar

Figure 5.13 (a) Comparison of laboratory and field T-bar penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0

T-bar resistance, qout (MPa)

Dep

th (m

)


Figure 5.13 (b) Comparison of laboratory and field T-bar extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1

Ratio qout/qin

Dep

th (m

)


Figure 5.13 (c) Extraction to penetration ratios for laboratory and field T-bars

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 10 20 30 40 50 60 70 80

Undrained shear strength, su (kPa)

Dep

th (m

)

Field T-bar

Field cone

CAU TXC

CAU TXE

CAU SS

UU compression

Laboratory T-bar

Figure 5.14 (a) Undrained shear strength profiles from laboratory testing

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 10 20 30 40 50 60 70 80


Dep

th (m

)

Field T-bar

Field cone

CAU TXC

TXC (SHANSEP)

CAU TXE

TXE (SHANSEP)

CAU SS

SS (SHANSEP)

Figure 5.14 (b) Undrained shear strength profiles including SHANSEP results

FIGURES

Figure 6.1 Photograph of centrifuge and actuator

Actuator

Penetrometer

FIGURES

Figure 6.2 (a) Photograph of model cone penetrometer

FIGURES

Figure 6.2 (b) Penetrometer rod attached with a model T-bar tip

FIGURES

(a) Hand vane

(b) Vane calibration scale

Figure 6.3 Photographs of hand vane apparatus

FIGURES

0

50

100

150

200

250

300

350

0 50 100 150 200 250 300 350 400 450 500 550 600 650

Length of strongbox (mm)

Wid

th o

f str

ongb

ox (m

m)

TB1-TB6: T-bar C1, C3-C4: Cone B1-B3: BallP1-P3: Plate HV11-HV13: Vane

Tube sample

C1

C3

TB1

TB2

TB3

TB4

TB6

B1

B3

P1

P3

HV11 HV12HV13

Row 1Row 2 Row 3 Row 4Row 5

TB5(twt)B2

(twt)

P2(twt)

C4(twt)


FIGURES

0

50

100

150

200

250

300

350

0 50 100 150 200 250 300 350 400 450 500 550 600 650

Length of strongbox (mm)

Wid

th o

f str

ongb

ox (m

m)

TB11-13, 51-53: (20x5) TB21-23: (30x5)TB31-33: (40x5) TB41-43: (50x5)HV21-HV26: Vane

Tube sample

TB11

TB13 TB51

TB53TB21

TB23 TB31

TB33

TB41

TB43

HV21

HV22HV23

HV24

HV25

HV26

Row 1Row 2 Row 3 Row 4 Row 5

TB12(twt)

TB52(twt)

TB22(twt)

TB42(twt)

TB32(twt)


FIGURES

0

5

10

15

20

25

0 20 40 60 80 100 120Root time (minutes)1/2

Sam

ple

settl

emen

t (m

m)..

90% degree of consolidation


FIGURES

0

2

4

6

8

10

12

14

16

18

-20 0 20 40 60 80 100 120Root time (minutes)1/2

Sam

ple

settl

emen

t (m

m)..

90% degree of consolidation


FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.1 0.2 0.3 0.4 0.5Tip resistance, qin (MPa)

Dep

th (

m)

T-bar 1

T-bar 2

T-bar 3

T-bar 4

T-bar 6


FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 4

T-bar 6

T-bar 5 (twitch)

Reverted to 1 mm/s

End Step 1

End Step 2End Step 3

End Step 4

End Step 5

End Step 6End Step 7

End Step 8

Figure 6.8 (b) Box 1: Comparison of normal and twitch tests for T-bars (20x5)

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.5 -0.4 -0.3 -0.2 -0.1 0Tip resistance, qout (MPa)

Dep

th (

m)

T-bar 1

T-bar 2

T-bar 3

T-bar 4

T-bar 6

T-bar 5 (twitch)

Extracted at 1 mm/s as standard tests

Figure 6.8 (c) Box 1: Comparison of extraction resistances for T-bars (20x5)

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Cone 1: Meas'd

Cone 1: Bq=0.45

Cone 1: Bq=1

Cone 3: Meas'd

Cone 3: Bq=0.45Cone 3: Bq=1

Cone 4: Twitch

T-bar 4

T-bar 6

Figure 6.9 (a) Box 1: Comparison of cone and T-bar (20x5) penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.4 -0.3 -0.2 -0.1 0 0.1Tip resistance, qout (MPa)

Dep

th (

m)

Cone 1: Meas'd

Cone 1: Bq=0.17

Cone 1: Bq=1

Cone 3: Meas'd

Cone 3: Bq=0.17Cone 3: Bq=1

Cone 4: Twitch

T-bar 4

T-bar 6

Figure 6.9 (b) Box 1: Comparison of cone and T-bar (20x5) extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Ball 1

Ball 3

Ball 2 (twitch)

T-bar 4

T-bar 6

Figure 6.10 (a) Box 1: Comparison of ball and T-bar (20x5) penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Ball 1

Ball 3

Ball 2 (twitch)

T-bar 4

T-bar 6

Figure 6.10 (b) Box 1: Comparison of ball and T-bar (20x5) extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Plate 1

Plate 3

Plate 2 (twitch)

T-bar 4

T-bar 6

Figure 6.11 (a) Box 1: Comparison of plate and T-bar (20x5) penetration resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Plate 1

Plate 3

Plate 2 (twitch)

T-bar 4

T-bar 6

Figure 6.11 (b) Box 1: Comparison of plate and T-bar (20x5) extraction resistances

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.3 -0.2 -0.1 0 0.1 0.2 0.3 0.4Tip resistance, q (MPa)

Dep

th (

m)

Field T-barAverage T-barAverage ConeAverage BallAverage Plate

Figure 6.12 Box 1: Summary of tip resistances for all model penetrometers

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1Ratio qout/qin

Dep

th (

m)

T-bar 1

T-bar 2

T-bar 3

T-bar 4

T-bar 6

Figure 6.13 Box 1: Ratios of extraction to penetration resistances for T-bars (20x5)

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1 1.2Ratio qout/qin

Dep

th (

m)

Average T-bar

Average Cone

Average Ball

Average Plate

Figure 6.14 Box 1: Summary of resistance ratios for all model penetrometers

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)

T-bar 51 (20 x 5)

T-bar 53 (20 x 5)


FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)

T-bar 51 (20 x 5)

T-bar 53 (20 x 5)

Figure 6.15 (b) Box 2: Comparison of extraction resistances for T-bars (20x5)

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 21 (30 x 5)

T-bar 23 (30 x 5)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)


FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 21 (30 x 5)

T-bar 23 (30 x 5)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)

Figure 6.16 (b) Box 2: Extraction resistances of T-bars 30x5 and 20x5

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 31 (40 x 5)

T-bar 33 (40 x 5)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)


FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

T-bar 41 (50 x 5)

T-bar 43 (50 x 5)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)


FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.4 -0.2 0 0.2 0.4 0.6Tip resistance, q (MPa)

Dep

th (

m)

Field T-barAvg T-bar (20x5)Avg T-bar (30x5)Avg T-bar (40x5)Avg T-bar (50x5)

Figure 6.19 Box 2: Summary of tip resistances for all various model T-bars

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1Ratio qout/qin

Dep

th (

m)

T-bar 11 (20 x 5)

T-bar 13 (20 x 5)

T-bar 51 (20 x 5)

T-bar 53 (20 x 5)

Figure 6.20 Box 2: Ratios of extraction to penetration resistances for T-bar (20x5)

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1Ratio qout/qin

Dep

th (

m)

Field T-bar

Field ST-bar

Box 1: T-bar (20x5)

Avg T-bar (20 x 5)

Avg T-bar (30 x 5)

Avg T-bar (40 x 5)

Avg T-bar (50 x 5)

Figure 6.21 Box 2: Ratios of extraction to penetration resistances for all T-bars

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

HV 11: Peak

HV 12: Peak

HV 13: Peak


FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

HV 21: Peak HV 22: Peak



HV 21: Remoulded HV 22: Remoulded




FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Field vane: Peak

Field vane: Remoulded

Box 1: HV (peak)

Box 2: HV (peak)

Box 2: HV (remoulded)

Figure 6.24 Average results of hand vane tests for Boxes 1 and 2

FIGURES

0

2

4

6

8

10

12

14

16

18

20

0 2 4 6 8 10Sensitivity, St

Dep

th (

m)

Field vane

Box 2: HV

Figure 6.25 Sensitivity of reconstituted clay sample (Box 2)

FIGURES

0

1

2

3

4

5

6

7

8

0 50 100 150Time (hours)

Stra

in (

%)

ε v

ε a

(a) Re-consolidation in triaxial cell

0

10

20

30

40

50

60

70

80

90

100

-0.4 -0.2 0 0.2 0.4

Tip resistance, qT-bar (MPa)

Pene

tratio

n (m

m)

qT-bar(out) qT-bar(in)

(b) Model T-bar resistances

0

10

20

30

40

50

60

70

80

90

100

0 0.05 0.1 0.15 0.2

Axial strain, εa (%)

Pene

tratio

n (m

m)

(c) Axial strain during testing

0

10

20

30

40

50

60

70

80

90

100

-0.04 -0.02 0 0.02 0.04

Volumetric strain, εv (%)

Pene

tratio

n (m

m)

(d) Volumetric strain during testing

Figure 6.26 T-bar test in triaxial sample recovered from Box 1

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Field T-barBox 1: T-bar (20x5)Box 1: ConeBox 1: BallBox 1: PlateBox 1: Vane (peak)Box 1: TXCBox 1: TXEBox 1: SSBox 1: Lab T-bar

Suspect


Note: Bearing factor of N = 10.5 used for

all penetrometers

FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Field T-bar

Box 2: T-bar (20x5)

Box 2: T-bar (30x5)

Box 2: T-bar (40x5)

Box 2: T-bar (50x5)

Box 2: Vane (peak)

Box 2: TXC

Box 2: TXE

Box 2: SS


FIGURES

0

2

4

6

8

10

12

14

16

18

20


Dep

th (

m)

Field T-bar

Box 1: T-bar (20x5)

Box 2: T-bar (20x5)

Box 1: Vane (peak)

Box 2: Vane (peak)

Figure 6.29 Comparison of undrained shear strength profiles for Boxes 1 and 2

FIGURES

y = 0.3836x - 1.5297y = 0.5603x - 0.3708

y = 0.5634x - 2.0815

0

2

4

6

8

10

12

14

16

18

20

0 10 20 30 40 50 60 70 80


Dep

th (m

)

CAU TXC

CAU TXE

CAU SS

su, av

TXC best-fit

TXE best-fit

SS best-fit

Figure 7.1 Best-fit trends for laboratory su data to evaluate the average shear strength, su, av = (su, TXC + su, TXE + su, SS)/3

FIGURES

0

2

4

6

8

10

12

14

16

18

20

4 6 8 10 12 14 16 18

Nkt, TXC , Nkt, SS or Nkt, av = qcnet/(su, TXC , su, SS or su, av)

Dep

th (m

)CAU TXCCAU SSLab averageLu et al (2004)

Open symbols: Cone-(a)

Solid symbols: Cone-(b) (Schneider et al, 2004)

Figure 7.2 Cone factors, Nkt, calculated using different laboratory su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

4 6 8 10 12 14 16 18

Nkt, vane = qcnet/su, vane

Dep

th (m

)Cone-(a)Cone-(b) (Schneider et al, 2004)Model cone (centrifuge)Lu et al (2004)

Figure 7.3 Cone factors, Nkt, calculated using vane su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

NT-bar = qT-bar/su

Dep

th (m

)

CAU TXCCAU SSLab average

Black solid: T-bar-(a1) (250 mm x 40 mm)Red open: T-bar-(a2) (160 mm x 40 mm)Blue solid: T-bar-(b) (250 mm x 40 mm) by Schneider et al (2004)

Figure 7.4 T-bar factors, NT-bar, calculated using different laboratory su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

NT-bar, vane = qT-bar/su, vane

Dep

th (m

)

T-bar-(a1) (250x40)

T-bar-(a2) (160x40)

T-bar-(b) (250x40)(Schneider et al, 2004)

Figure 7.5 T-bar factors, NT-bar, calculated using vane su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

NT-bar, SS = qT-bar/su, SS

Dep

th (m

)

T-bar (20 x 5)

T-bar (30 x 5)

T-bar (40 x 5)

T-bar (50 x 5)

Figure 7.6 T-bar factors, NT-bar, calculated using simple shear strengths for model T-bars tested in the centrifuge

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

NT-bar, vane = qT-bar/su, vane

Dep

th (m

)

Box 1: T-bar (20 x 5)

Box 2: T-bar (20 x 5)

Box 2: T-bar (30 x 5)

Box 2: T-bar (40 x 5)

Box 2: T-bar (50 x 5)

Figure 7.7 T-bar factors, NT-bar, calculated using hand vane strengths for model T-bars tested in the centrifuge

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

Nball = qball/su

Dep

th (m

)CAU TXCCAU SSLab average

Open symbols: Ball-(a)

Solid symbols: Ball-(b) (Schneider et al, 2004)

Figure 7.8 Ball factors, Nball, calculated using different laboratory su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

Nball, vane = qball/su, vane

Dep

th (m

)

Ball-(a)

Ball-(b) (Schneider etal, 2004)Model ball (centrifuge)

Figure 7.9 Ball factors, Nball, calculated using vane su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

Nplate = qplate/su

Dep

th (m

)

CAU TXCCAU SSLab average

Figure 7.10 Plate factors, Nplate, calculated using different laboratory su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

2 4 6 8 10 12 14 16

Nplate, vane = qplate/su, vane

Dep

th (m

)

Field plate

Model plate (centrifuge)

Figure 7.11 Plate factors, Nplate, calculated using vane su values

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.4 -0.3 -0.2 -0.1 0 0.1 0.2 0.3 0.4 0.5Tip resistance, qnet (MPa)

Dep

th (

m)

Normal T-bar

T-bar 5

Cone 4

Ball 2

Plate 2

Twitch tests for sample Box 1

Extracted at 1 mm/s after twitch tests

Start of step 5 for test T-bar 5

Figure 8.1 Resistance profiles of twitch tests for various shaped model penetrometers

FIGURES

0

2

4

6

8

10

12

14

16

18

20

-0.6 -0.4 -0.2 0 0.2 0.4 0.6 0.8Tip resistance, qnet (MPa)

Dep

th (

m)

Normal T-bar (20x5)

T-bar 12 (20x5)

T-bar 22 (30x5)

T-bar 32 (40x5)

T-bar 42 (50x5)

T-bar 52 (20x5)

Twitch tests for sample Box 2

Extracted at 1 mm/s after twitch tests

Figure 8.2 Resistance profiles of twitch tests for model T-bars with various aspect ratios

FIGURES

0

0.5

1

1.5

2

2.5

3

3.5

4

0.1 1 10 100 1000V = vd / cv

q / q

ref

Watson & Suemasa (2000)

House et al (2001)

After Randolph & Hope (2004)

Box 1: cv (m2/yr) = 0.45

Box 2: cv = 0.4Box 1: cv = 1

Box 2: cv = 0.9

Box 1: cv = 1.8

Box 2: cv = 1.6

Model T-bar (20 x 5)

Figure 8.3 Evaluation of consolidation coefficient using data from T-bar twitch tests based on various backbone curves

FIGURES

0

0.5

1

1.5

2

2.5

3

3.5

4

0.1 1 10 100 1000V = vd / cv

q / q

ref

After Randolph & Hope (2004)

Box 1: cv (m2/yr) = 0.6

Model cone

Figure 8.4 Evaluation of consolidation coefficient using data from cone twitch test

FIGURES

0

0.5

1

1.5

2

2.5

3

3.5

4

0.1 1 10 100 1000V = vd / cv

q / q

ref

Box 1: T-bar 5

Box 1: Ball 2

Box 1: Plate 2

Box 1: Cone 4


cv = 0.45 m2/year

Figure 8.5 Non-dimensional plot for various shaped model penetrometers

FIGURES

0

0.5

1

1.5

2

2.5

3

3.5

4

0.1 1 10 100 1000V = vd / cv

q / q

ref

Box 2: T-bar 12 (20x5)

Box 2: T-bar 22 (30x5)

Box 2: T-bar 32 (40x5)

Box 2: T-bar 42 (50x5)

Box 2: T-bar 52 (20x5)


cv = 0.4 m2/year

Figure 8.6 Non-dimensional plot for different model T-bar penetrometers

FIGURES

(a) Consolidation curves for square and rectangular footings

(b) Consolidation curves for square and circular footings

Figure 8.7 Degree of consolidation (settlement), Us, versus time factor, Tv, for surface footings − reproduced from Davis & Poulos (1972)

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

Box 1: T-bar 5: cv (m2/yr) = 0.45

Box 1: T-bar 5: cv (m2/yr) = 1

W&S (2000): no viscous effects

W&S (2000): with viscous effects

After R&H (2004): no viscous effects

After R&H (2004): with viscous effects

Figure 8.8 (a) Fitting results of model T-bar twitch test in Box 1: focus in viscous region

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

Box 1: T-bar 5: cv (m2/yr) = 0.6

Box 1: T-bar 5: cv (m2/yr) = 1.3





Figure 8.8 (b) Fitting results of model T-bar twitch test in Box 1: overall best-fit

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000 10000V = vd / cv

q / q

ref

Box 1: Cone 4: cv (m2/yr) = 0.6



Figure 8.9 (a) Fitting results of model cone twitch test in Box 1: focus in viscous region

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000 10000V = vd / cv

q / q

ref

Box 1: Cone 4: cv (m2/yr) = 0.8



Figure 8.9 (b) Fitting results of model cone twitch test in Box 1: overall best-fit

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

Box 2: T-bar 12 (20x5)

Box 2: T-bar 22 (30x5)

Box 2: T-bar 32 (40x5)

Box 2: T-bar 42 (50x5)

Box 2: T-bar 52 (20x5)

W&S (2000): no viscouseffectsW&S (2000): withviscous effects

cv = 0.4 m2/year

Figure 8.10 (a) Fitting results of various model T-bar twitch tests in Box 2, using backbone curve from Watson & Suemasa (2000): focus in viscous region

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

Box 2: T-bar 12 (20x5)

Box 2: T-bar 22 (30x5)

Box 2: T-bar 32 (40x5)

Box 2: T-bar 42 (50x5)

Box 2: T-bar 52 (20x5)

W&S (2000): no viscouseffectsW&S (2000): with viscouseffects

cv = 0.5 m2/year

Figure 8.10 (b) Fitting results of various model T-bar twitch tests in Box 2, using backbone curve from Watson & Suemasa (2000): overall best-fit

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

Box 2: T-bar 12 (20x5)

Box 2: T-bar 22 (30x5)

Box 2: T-bar 32 (40x5)

Box 2: T-bar 42 (50x5)

Box 2: T-bar 52 (20x5)

After R&H (2004): noviscous effectsAfter R&H (2004): withviscous effects

cv = 0.9 m2/year

Figure 8.11 (a) Fitting results of various model T-bar twitch tests in Box 2, using backbone curve after Randolph & Hope (2004): focus in viscous region

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

Box 2: T-bar 12 (20x5)

Box 2: T-bar 22 (30x5)

Box 2: T-bar 32 (40x5)

Box 2: T-bar 42 (50x5)

Box 2: T-bar 52 (20x5)

After R&H (2004): noviscous effectsAfter R&H (2004): withviscous effects

cv = 1 m2/year

Figure 8.11 (b) Fitting results of various model T-bar twitch tests in Box 2, using backbone curve after Randolph & Hope (2004): overall best-fit

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

No viscous effects

Box 1: Vo = 300

Box 1: Vo = 200

Box 2: Vo = 200

Box 2: Vo = 150

Backbone curves from Watson & Suemasa (2000) with viscous effects

Figure 8.12 (a) Comparison of backbone curves from Watson & Suemasa (2000) with different Vo

FIGURES

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000V = vd / cv

q / q

ref

No viscous effects

Box 1: Vo = 150

Box 1: Vo = 100

Box 2: Vo = 100

Box 2: Vo = 70

Backbone curves from Randolph & Hope (2004) with viscous effects

Figure 8.12 (b) Comparison of backbone curves from Randolph & Hope (2004) with different Vo

FIGURES

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000 10000 100000V = vd / cv

q / q

ref

Field T-bar (Schneider et al, 2004)



cv = 1.6 m2/year

Figure 8.13 (a) Fitting results of field T-bar twitch tests reported by Schneider et al (2004), using backbone curve from Watson & Suemasa (2000)

FIGURES

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000 10000 100000V = vd / cv

q / q

ref

Field T-bar (Schneider et al, 2004)



cv = 3.3 m2/year

Figure 8.13 (b) Fitting results of field T-bar twitch tests reported by Schneider et al (2004), using backbone curve from Randolph & Hope (2004)

FIGURES

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000 10000 100000V = vd / cv

q / q

ref

Field cone (Schneider etal, 2004)

After R&H (2004): noviscous effects

After R&H (2004): withviscous effects

cv = 1.9 m2/year

Figure 8.14 Fitting results of field cone twitch tests reported by Schneider et al (2004), using the backbone curve from Randolph & Hope (2004)

APPENDIX A

CONSTANT RATE OF STRAIN CONSOLIDATION

(CRSC) TEST RESULTS

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

1 10 100 1000


Voi

d ra

tio, e

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.E+00 1.E-09 2.E-09 3.E-09

Permeability, k (m/s)

Voi

d ra

tio, e

0

1

2

3

4

5

6

7

8

9

10

0 100 200 300


cv (m

2 /yr)

0

5

10

15

20

25

30

0 100 200 300


% o

f Por

e pr

essu

re, u

(%)

Appendix A.1 CRSC 4: Depth = 18.35 m

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

0.E+00 1.E-10 2.E-10 3.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800 1000


cv (m

2 /yr)

0

5

10

15

20

25

0 200 400 600 800 1000


% o

f Por

e pr

essu

re, u

(%)


0

0.5

1

1.5

2

2.5

1 10 100 1000


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

0.E+00 1.E-10 2.E-10 3.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 100 200 300


cv (m

2 /yr)

0

2

4

6

8

10

12

14

16

18

20

0 100 200 300


% o

f Por

e pr

essu

re, u

(%)


0

0.5

1

1.5

2

2.5

3

1 10 100 1000


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

0.E+00 3.E-10 6.E-10 9.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 100 200 300 400 500


cv (m

2 /yr)

0

1

2

3

4

5

6

7

8

9

10

0 100 200 300 400 500


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0.E+00 1.E-10 2.E-10 3.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600


cv (m

2 /yr)

0

2

4

6

8

10

12

14

16

18

20

0 200 400 600


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.E+00 3.E-10 6.E-10 9.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600


cv (m

2 /yr)

0

2

4

6

8

10

12

14

16

18

0 200 400 600


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.0E+00 4.0E-10 8.0E-10 1.2E-09


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800


cv (m

2 /yr)

0

5

10

15

20

25

30

0 200 400 600 800


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.E+00 3.E-10 6.E-10 9.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800


cv (m

2 /yr)

0

2

4

6

8

10

12

14

16

18

0 200 400 600 800


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.E+00 3.E-10 6.E-10 9.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600


cv (m

2 /yr)

0

5

10

15

20

25

30

35

40

45

0 200 400 600


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1 10 100 1000 10000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0.0E+00 4.0E-10 8.0E-10 1.2E-09


Voi

d ra

tio, e

0

1

2

3

4

5

6

7

8

9

10

0 300 600 900 1200


cv (m

2 /yr)

0

0.5

1

1.5

2

2.5

3

3.5

0 300 600 900 1200


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

0.0E+00 5.0E-10 1.0E-09 1.5E-09


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 100 200 300 400


cv (m

2 /yr)

0

2

4

6

8

10

12

14

16

18

20

0 100 200 300 400


% o

f Por

e pr

essu

re, u

(%)


0

0.5

1

1.5

2

2.5

3

1 10 100 1000


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

0 1E-09 2E-09 3E-09 4E-09


Voi

d ra

tio, e

0

1

2

3

4

5

6

7

0 100 200 300


cv (m

2 /yr)

0

2

4

6

8

10

12

14

0 100 200 300


% o

f Por

e pr

essu

re, u

(%)

Appendix A.12 CRSC 16:Depth = 4.11 m

0

0.5

1

1.5

2

2.5

3

1 10 100 1000


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

0.E+00 1.E-09 2.E-09 3.E-09


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 100 200 300 400


cv (m

2 /yr)

0

5

10

15

20

25

30

0 100 200 300 400


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0.E+00 3.E-10 6.E-10 9.E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800


cv (m

2 /yr)

0

2

4

6

8

10

12

14

0 200 400 600 800


% o

f Por

e pr

essu

re, u

(%)


APPENDIX B

CAU TRIAXIAL TEST RESULTS

-10

-50

5

1015

2025

3035

40

0 20 40 60 80

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-5

0

5

10

15

20

25

30

35

0 20 40 60 80

Time (hours)

Stre

ss (k

Pa)

p'

q

0

0.5

1

1.5

2

2.5

3

0 20 40 60 80

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 20 40 60 80

Time (hours)

Ko

-10

-50

5

1015

2025

3035

40

0 5 10 15 20


Stre

ss (k

Pa) p'

q

∆u

0

5

10

15

20

25

30

35

40

0 10 20 30 40

Mean effective stress, p' (kPa)

Dev

iato

r str

ess,

q (k

Pa)

Figure B.1 CAU TXC 1: Depth = 5.60 m

σ'v = 32.5 kPa and σ'h = 26.0 kPa (Ko = 0.8)

-10

0

10

20

30

40

50

0 20 40 60

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

0

5

10

15

20

25

30

35

0 20 40 60

Time (hours)

Stre

ss (k

Pa)

p'

q

-0.5

0

0.5

1

1.5

2

2.5

3

3.5

4

0 20 40 60

Time (hours)

Stra

in (%

) ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 20 40 60

Time (hours)

Ko

-10

0

10

20

30

40

50

0 5 10 15 20


Stre

ss (k

Pa)

p'

q

∆u

05

10

1520

25

3035

40

4550

0 10 20 30 40


Dev

iato

r str

ess,

q (k

Pa)



-100

102030405060708090

100

0 20 40 60

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-10

0

10

20

30

40

50

60

70

80

0 10 20 30 40 50

Time (hours)

Stre

ss (k

Pa)

p'

q

0

1

2

3

4

5

6

7

8

0 20 40 60

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 10 20 30 40 50

Time (hours)

Ko

0

10

20

30

40

5060

70

80

90

100

0 5 10 15 20


Stre

ss (k

Pa)

p'

q

∆u

010

203040

5060

708090

100

0 20 40 60 80


Dev

iato

r str

ess,

q (k

Pa)



-20

0

20

40

60

80

100

120

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

60

70

80

90

0 20 40 60 80

Time (hours)

Stre

ss (k

Pa)

p'

q

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

0 20 40 60 80

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 20 40 60 80

Time (hours)

Ko

0

20

40

60

80

100

120

0 5 10 15 20


Stre

ss (k

Pa)

p'

q

∆u

0

20

40

60

80

100

120

0 20 40 60 80 100


Dev

iato

r str

ess,

q (k

Pa)



-80

-60

-40

-20

0

20

40

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-10

-5

0

5

10

15

20

25

30

35

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

1

2

3

4

5

6

7

8

9

0 50 100 150

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0 50 100 150

Time (hours)

Ko

-80

-60

-40

-20

0

20

40

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-70

-60

-50

-40

-30

-20

-10

0

10

0 10 20 30 40


Dev

iato

r str

ess,

q (k

Pa)

Figure B.5 CAU TXE 1: Depth = 5.43 m


-80

-60

-40

-20

0

20

40

60

80

100

0 20 40 60 80 100

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-10

0

10

20

30

40

50

60

70

80

0 20 40 60 80

Time (hours)

Stre

ss (k

Pa)

p'

q

0

0.5

1

1.5

2

2.5

3

3.5

4

0 20 40 60 80

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 20 40 60 80

Time (hours)

Ko

-80

-60

-40

-20

0

20

40

60

80

100

-20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-70

-60

-50

-40

-30

-20

-10

0

10

20

30

0 20 40 60 80


Dev

iato

r str

ess,

q (k

Pa)



-80

-60

-40

-20

0

20

40

60

80

100

0 20 40 60 80

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

60

70

80

90

0 10 20 30 40 50

Time (hours)

Stre

ss (k

Pa)

p'

q

0

0.5

1

1.5

2

2.5

3

3.5

0 10 20 30 40 50

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 10 20 30 40 50

Time (hours)

Ko

-80

-60

-40

-20

0

20

40

60

80

100

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-70

-60-50

-40-30

-20-10

010

2030

0 20 40 60 80 100


Dev

iato

r str

ess,

q (k

Pa)



-30

-20

-10

0

10

20

30

40

0 20 40 60 80 100

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-5

0

5

10

15

20

25

30

35

0 10 20 30 40 50 60 70

Time (hours)

Stre

ss (k

Pa)

p'

q

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

0 10 20 30 40 50 60 70

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 10 20 30 40 50 60 70

Time (hours)

Ko

-30

-20

-10

0

10

20

30

40

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-25

-20

-15

-10

-5

0

5

10

0 10 20 30 40


Dev

iato

r str

ess,

q (k

Pa)



APPENDIX C

CAU SIMPLE SHEAR TEST RESULTS

0

0.5

1

1.5

2

2.5

3

3.5

4

0 0.5 1 1.5 2 2.5 3Time (hours)

Stra

in (%

)

ε v

ε a

0

2

4

6

8

10

12

14

16

0 10 20 30 40 50Shear strain, γ (%)

Shea

r str

ess,

τxy

(kPa

)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss ra

tio

( τxy

/ σ'v )

0

5

10

15

20

25

30

35

40

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σv

∆u

0

5

10

15

20

25

30

35

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25 30s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)

Figure C.1 CAU SS 1: Depth = 5.89 m


0

2

4

6

8

10

12

0 5 10 15 20 25Time (hours)

Stra

in (%

)

ε v

ε a

0

1

2

3

4

5

6

7

8

9

0 10 20 30 40 50Shear strain, γ (%)

Shea

r str

ess,

τxy

(kPa

)

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

-5

0

5

10

15

20

25

30

35

40

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

5

10

15

20

25

30

35

40

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

1

2

3

4

5

6

7

8

9

10

0 5 10 15 20 25 30 35s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



0

2

4

6

8

10

12

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

35

0 10 20 30 40 50Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

70

80

90

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)σv

∆u

0

10

20

30

40

50

60

70

80

90

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

35

0 20 40 60 80s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)


σ'v = 77.7 kPa and σ'h =62.2 kPa (Ko = 0.8)

0

1

2

3

4

5

6

7

8

9

10

0 0.5 1 1.5 2 2.5 3Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

35

0 10 20 30 40 50Shear strain, γ (%)

Shea

r str

ess,

τxy

(kPa

)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

70

80

90

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

60

70

80

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

35

0 20 40 60 80s' = (σ1'+σ3')/2 (kPa)

t = (

σ1'- σ

3 ')/2

(kP

a)


σ'v = 77.7 kPa and σ'h =62.2 kPa (Ko = 0.8)

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

0 10 20 30 40 50Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

5

10

15

20

25

30

35

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

5

10

15

20

25

30

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

0 5 10 15 20 25 30s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



0

2

4

6

8

10

12

14

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

0 10 20 30 40 50Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

70

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

60

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

0 10 20 30 40 50s' = (σ1'+σ3')/2 (kPa)

t = (

σ1'- σ

3 ')/2

(kP

a)



0

1

2

3

4

5

6

7

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

0 10 20 30 40 50Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

-0.1

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

70

80

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

60

70

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

35

0 10 20 30 40 50 60 70s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



0

1

2

3

4

5

6

7

8

9

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

0 10 20 30 40 50Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

0 10 20 30 40 50s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



APPENDIX D

MODEL T-BAR IN TRIAXIAL TEST RESULTS

0

1

2

3

4

5

6

0 20 40 60 80

Time (hours)

Stra

in (

%)

ε v

ε a

0

10

20

30

40

50

60

70

80

90

100

0 0.2 0.4 0.6

Tip Resistance, qT-bar (MPa)

Pene

tratio

n (m

m)

qT-bar(in)

0

10

20

30

40

50

60

70

80

90

100

0 0.02 0.04 0.06

Axial Strain, εa (%)

Pene

trat

ion

(mm

)

0

10

20

30

40

50

60

70

80

90

100

280 290 300 310 320

Pore Pressure, u (kPa)

Pene

tratio

n (m

m)

Figure D.1 T-bar 1: Depth = 15.00 m

(p' = 68 kPa)

0

1

2

3

4

5

6

0 20 40 60 80

Time (hours)

Stra

in (

%)

ε v

ε a

0

10

20

30

40

50

60

70

80

90

100

-0.4 -0.2 0 0.2 0.4 0.6 0.8


Pene

tratio

n (m

m)

qT-bar(in)qT-bar(out)

0

10

20

30

40

50

60

70

80

90

100

0 0.05 0.1 0.15 0.2


Pene

tratio

n (m

m)

0

10

20

30

40

50

60

70

80

90

100

-0.06 -0.04 -0.02 0

Volumetric Strain, εv (%)

Pen

etra

tion

(m

m)


(p' = 61 kPa)

0

1

2

3

4

5

6

7

8

0 20 40 60 80

Time (hours)

Stra

in (

%)

ε v

ε a

0

10

20

30

40

50

60

70

80

90

100

-0.4 -0.2 0 0.2 0.4 0.6 0.8 1


Pene

tratio

n (m

m)


0

10

20

30

40

50

60

70

80

90

100

0 0.05 0.1 0.15


Pen

etra

tion

(m

m)

0

10

20

30

40

50

60

70

80

90

100

-0.04 -0.02 0 0.02


Pen

etra

tion

(m

m)


(p' = 79 kPa)

0

1

2

3

4

5

6

0 20 40 60 80

Time (hours)

Stra

in (

%)

ε v

ε a

0

10

20

30

40

50

60

70

80

90

100

-0.4 -0.2 0 0.2 0.4 0.6


Pene

tratio

n (m

m)


0

10

20

30

40

50

60

70

80

90

100

0 0.02 0.04 0.06 0.08


Pen

etra

tion

(m

m)

0

10

20

30

40

50

60

70

80

90

100

-0.06 -0.04 -0.02 0


Pen

etra

tion

(m

m)


(p' = 56 kPa)

0

1

2

3

4

5

6

7

0 20 40 60 80 100

Time (hours)

Stra

in (

%)

ε v

ε a

0

10

20

30

40

50

60

70

80

90

100

-0.2 0 0.2 0.4 0.6


Pene

tratio

n (m

m)


0

10

20

30

40

50

60

70

80

90

100

0 0.05 0.1 0.15


Pen

etra

tion

(m

m)

0

10

20

30

40

50

60

70

80

90

100

-0.06 -0.04 -0.02 0


Pen

etra

tion

(m

m)


(p' = 44 kPa)

0

1

2

3

4

5

6

0 20 40 60 80

Time (hours)

Stra

in (

%)

ε v

ε a

0

10

20

30

40

50

60

70

80

90

100

-0.2 0 0.2 0.4


Pene

tratio

n (m

m)


0

10

20

30

40

50

60

70

80

90

100

0 0.05 0.1 0.15


Pen

etra

tion

(m

m)

0

10

20

30

40

50

60

70

80

90

100

-0.15 -0.1 -0.05 0


Pen

etra

tion

(m

m)


(p' = 23 kPa)

APPENDIX E

TRIAXIAL AND SIMPLE SHEAR TESTS

FOLLOWING SHANSEP PROCEDURE

-10

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

60

0 20 40 60 80 100

Time (hours)

Stre

ss (k

Pa)

p'

q

0123456789

10

0 20 40 60 80 100

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 20 40 60 80 100

Time (hours)

Ko

0

10

20

30

40

50

60

70

0 5 10 15 20


Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

60

70

0 10 20 30 40


Dev

iato

r str

ess,

q (k

Pa)

Figure E.1 TXC 6 (SHANSEP): Depth = 6.60 m

Stage 1: σ'v1 = 78.7 kPa and σ'h1 = 44.7 kPa


-100

102030405060708090

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

6070

80

90

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

5

10

15

20

25

0 50 100 150

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

0 50 100 150

Time (hours)

Ko

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25 30


Stre

ss (k

Pa) p'

q

∆u

0

10

20

30

40

50

60

70

80

0 20 40 60


Dev

iato

r str

ess,

q (k

Pa)

Figure E.2 TXC 7 (SHANSEP): Depth = 10.86 m



-60

-40

-20

0

20

40

60

80

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-10

0

10

20

30

40

50

60

70

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

2

4

6

8

10

12

0 50 100 150

Time (hours)

Stra

in (%

) ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 50 100 150

Time (hours)

Ko

-60

-40

-20

0

20

40

60

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-50

-40

-30

-20

-10

0

10

20

0 20 40 60


Dev

iato

r str

ess,

q (k

Pa)

Figure E.3 TXE 5 (SHANSEP): Depth = 7.79 m



-80

-60-40

-20

0

20

40

6080

100

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-100

102030405060708090

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

2

4

68

10

12

1416

18

0 50 100 150

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 50 100 150

Time (hours)

Ko

-80

-60

-40

-20

0

20

40

60

80

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-70

-60

-50

-40-30

-20

-10

010

20

0 20 40 60 80


Dev

iato

r str

ess,

q (k

Pa)

Figure E.4 TXE 7 (SHANSEP): Depth = 10.68 m



024

68

101214

161820

0 5 10 15 20 25Time (hours)

Stra

in (%

)ε v

ε a

0

5

10

15

20

25

30

35

0 20 40 60Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9


Stre

ss r

atio

( τ

xy / σ

'v )

-10

0

10

20

30

40

50

60


Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

60


Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

35

0 10 20 30 40 50s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)

Figure E.5 SS 7 (SHANSEP): Depth = 7.45 m



0

2

4

6

8

10

12

0 5 10 15 20 25Time (hours)

Stra

in (%

)ε v

ε a

05

10

1520

253035

404550

0 10 20 30 40 50Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

-20-10

0102030405060708090

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

01020

30

40

506070

8090

100

0 10 20 30 40 50Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

05

10

1520

2530

35

404550

0 20 40 60 80s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)

Figure E.6 SS 8 (SHANSEP): Depth = 12.99 m



APPENDIX F

CRSC TEST RESULTS ON CENTRIFUGE

SAMPLES

0

0.5

1

1.5

2

2.5

1 10 100 1000


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

0 5E-10 1E-09 2E-09 2E-09


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

0 200 400 600


cv (m

2 /yr)

0

2

4

6

8

10

12

14

16

18

0 200 400 600


% o

f Por

e pr

essu

re, u

(%)

Appendix F.1 Box 1: CF 1 CRSC 1, depth at base = 5.8 m

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0 2E-10 4E-10 6E-10 8E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600


cv (m

2 /yr)

0

5

10

15

20

25

30

0 200 400 600


% o

f Por

e pr

essu

re, u

(%)

Appendix F.2 Box 1: CF 1 CRSC 2, depth at base = 13.8 m

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

0 2E-10 4E-10 6E-10 8E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800


cv (m

2 /yr)

0

5

10

15

20

25

0 200 400 600 800


% o

f Por

e pr

essu

re, u

(%)

Appendix F.3 Box 2: CF 2 CRS 3, depth at base = 3.8 m

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

1 10 100 1000 10000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

0 2E-10 4E-10 6E-10 8E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 500 1000 1500


cv (m

2 /yr)

0

5

10

15

20

25

30

35

0 500 1000 1500


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0 2E-10 4E-10 6E-10 8E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800


cv (m

2 /yr)

0

5

10

15

20

25

30

35

0 200 400 600 800


% o

f Por

e pr

essu

re, u

(%)


0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1 10 100 1000


Voi

d ra

tio, e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0 1E-10 2E-10 3E-10 4E-10


Voi

d ra

tio, e

0

0.5

1

1.5

2

2.5

3

3.5

4

0 200 400 600 800 1000


cv (m

2 /yr)

0

5

10

15

20

25

30

35

40

0 200 400 600 800 1000


% o

f Por

e pr

essu

re, u

(%)


APPENDIX G

CAU TRIAXIAL AND SIMPLE SHEAR TEST

RESULTS FOR CENTRIFUGE SAMPLES

-10

0

10

20

30

40

50

60

70

80

0 50 100 150 200

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-10

0

10

20

30

40

50

60

70

80

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

123

45

678

910

0 50 100 150

Time (hours)

Stra

in (%

) ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 50 100 150

Time (hours)

Ko

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25


Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

60

70

0 20 40 60 80


Dev

iato

r str

ess,

q (k

Pa)

Figure G.1 Box 1: CF 1 CAU TXC 1, depth at base = 18.9 m


-20

0

20

40

60

80

100

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-20

0

20

40

60

80

100

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

1

2

3

4

5

6

7

8

0 50 100 150

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 50 100 150

Time (hours)

Ko

0

10

20

30

40

50

60

70

80

90

0 5 10 15 20 25


Stre

ss (k

Pa)

p'

q

∆u

0

10

20

30

40

50

60

70

80

0 20 40 60 80 100


Dev

iato

r str

ess,

q (k

Pa)

Figure G.2 Box 2: CF 2 CAU TXC 1, depth at base = 20 m

σ'v = 102 kPa and σ'h = 81.6 kPa (Ko = 0.8)

-60

-40

-20

0

20

40

60

80

100

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

-10

0

10

20

30

40

50

60

70

80

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

1

2

3

4

5

6

7

8

9

0 50 100 150

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

0 50 100 150

Time (hours)

Ko

-60

-40

-20

0

20

40

60

80

100

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-50

-40

-30

-20

-10

0

10

20

0 20 40 60 80 100


Dev

iato

r str

ess,

q (k

Pa)

Figure G.3 Box 1: CF 1 CAU TXE 1, depth at base = 19 m


-60

-40

-20

0

20

40

60

80

100

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

∆u

0

20

40

60

80

100

0 50 100 150

Time (hours)

Stre

ss (k

Pa)

p'

q

0

1

2

3

4

5

6

7

8

9

0 50 100 150

Time (hours)

Stra

in (%

)

ε a

ε v

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 50 100 150

Time (hours)

Ko

-80

-60-40

-20

020

4060

80

100120

-25 -20 -15 -10 -5 0


Stre

ss (k

Pa)

p'

q

∆u

-60

-50

-40

-30

-20

-10

0

10

20

30

0 20 40 60 80 100


Dev

iato

r str

ess,

q (k

Pa)

Figure G.4 Box 2: CF 2 CAU TXE 1, depth at base = 20 m


0

1

2

3

4

5

6

7

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

0 5 10 15 20 25Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

-0.1

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

-10

0

10

20

30

40

50

60

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

510

1520

253035

404550

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

0 10 20 30 40 50s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)

Figure G.5 Box 1: CF 1 CAU SS 1, depth at base = 10.8 m

σ'v = 47.5 kPa and σ'h = 38 kPa (Ko = 0.8)

0

1

2

3

4

5

6

7

8

9

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

35

40

0 5 10 15 20 25Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

70

80

90

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

60

70

80

90

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

35

40

45

0 20 40 60 80s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

2

4

6

8

10

12

14

16

18

0 5 10 15 20 25Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

05

1015

2025

30

3540

4550

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

5

10

15

20

25

30

35

40

45

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

02

46

810

12

1416

1820

0 10 20 30 40s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



0

1

2

3

4

5

6

7

8

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

35

40

0 5 10 15 20 25Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

-0.1

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

20

40

60

80

100

120

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

01020

3040

506070

8090

100

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

35

40

0 20 40 60 80 100s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



0

0.5

1

1.5

2

2.5

3

3.5

0 5 10 15 20Time (hours)

Stra

in (%

)ε v

ε a

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.10.2

0.30.4

0.50.6

0.70.8

0.91

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

5

10

15

20

25

30

35

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

5

10

15

20

25

30

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)


σ'v = 27.5 kPa and σ'h = 22 kPa (Ko = 0.8)

0

1

2

3

4

5

6

7

8

0 5 10 15 20Time (hours)

Stra

in (%

)

ε v

ε a

0

5

10

15

20

25

30

0 5 10 15 20 25Shear strain, γ (%)

Shea

r st

ress

, τxy

(kP

a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss r

atio

( τ

xy / σ

'v )

0

10

20

30

40

50

60

70

80

90

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss

(kPa

)

σv

∆u

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25Shear strain, γ (%)

Stre

ss (

kPa)

σ'h

σ'v

0

5

10

15

20

25

30

0 10 20 30 40 50 60 70s' = (σ1'+σ3')/2 (kPa)

t = ( σ

1 '-σ3

')/2

(kPa

)



characterisation of soft soils for deep water

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