UNIAXIAL BEHAVIOUR OF SUCTION CAISSONS IN · UNIAXIAL BEHAVIOUR OF SUCTION CAISSONS IN SOFT...

UNIAXIAL BEHAVIOUR OF SUCTION CAISSONS IN

SOFT DEPOSITS IN DEEPWATER

by

Wen CHEN B. Eng. (in Civil), M.Sc. (in Geotech.)

A thesis submitted for the degree of

Doctor of Philosophy

at

The University of Western Australia

Centre for Offshore Foundation Systems

School of Civil and Resource Engineering

July 2005

i Abstract

Centre for Offshore Foundation Systems The University of Western Australia

ABSTRACT

Suction caissons are a cost-effective alternative to traditional piles in deep to

ultradeep waters. No design rule has been available on the axial capacity of suction

caissons as part of the mooring system in soft sediments. In this research, a series of

centrifuge tests were performed using instrumented model caissons, to investigate the

axial capacity and radial stress changes around caissons during installation,

consolidation and vertical pullout in normally consolidated, lightly overconsolidated

and sensitive clays. Total pressure transducers instrumented on the caisson wall were

calibrated for various conditions. The radial total stress acting on the external wall

varied almost linearly during penetration and extraction of the caisson, with smaller

gradients observed during post-consolidation pullout. Minimum difference was found

in the penetration resistance and the radial total stress for caissons installed by jacking

or by suction, suggesting that the mode of soil flow at the caisson tip is similar under

these two types of installation. Observed soil heave showed that the soil particles at the

caisson tip flow about evenly outside and inside the caisson during suction installation.

Comparison was made between measurements and various theoretical predictions, on

both the radial stress changes during caisson installation, and the radial effective stress

after consolidation. Significant under-predictions on excess pore pressure changes,

consolidation times and external shaft friction ratios were found for the NGI Method,

based on the assumption that the caisson wall is accommodated entirely by inward

motion of the clay during suction installation. Obvious over-predictions by the MTD

approach were found in both stress changes and shaft capacity of the caissons. A

simple form of cavity expansion method was found to give reasonable estimations of

stress changes and post-consolidation external shaft friction. A model for predicting the

penetration resistance of suction caissons in clay was evaluated. Upper and lower

bound values of external shaft friction ratio during uplift loading after consolidation

were derived. Uplift capacity of caissons under sustained loading and cyclic loading

were investigated, revealing approximately 15 to 30% reduction of the capacity

compared to that under monotonic loading. External shaft friction ratios and reverse

end-bearing capacity factors were both found to be significantly lower than those under

monotonic loading.

ii Acknowledgement


ACKNOWLEDGEMENTS

I would like to express my sincere gratitude to my supervisor, Professor Mark

Randolph, who provided me with the opportunity to undertake this research. His ability

to sight through the ‘fog’ helped me so much when I was perplexed in dilemma; his

encouragement gave me power to face the frustration; his guidance forms an invaluable

fortune in my life.

Special thanks go to Mr. Don Herley, who helped me prepare the sample, ramp up the

centrifuge and arrange each instrument throughout my research. I would also like to

thank Dr. Andrew House, former Ph.D. student of COFS, for his generous suggestions

on the caisson study based on his experience. The model caisson was gauged first by

Mrs. Simone Fedoriczuk and later by Mr.Tuarn Brown; their work was essential in

achieving the accurate measurements and is greatly appreciated. Expert support from

electrical technicians, Mr. John Breen and Mr. Shane De Catania, is especially

appreciated, their patience and kindness was critical in the problem solving process. I

would like to thank Mr. Gary Davies and his colleagues in the Civil Engineering

Workshop for their high efficiency and accuracy in the fabrication work; their

endeavour ensures every urgent work to be finished on time. Assistances from Mr.

Binaya Bhattarai and Mrs. Clare Bearman, during both calibration tests and ring shear

tests, are acknowledged. Many thanks go to Mr. Wayne Galbraith, whose programming

work turned the high-precision control system into reality in the centrifuge. I am also

grateful to the IT support from Mr. Wenge Liu, and technical help from Mr. Bart

Thompson.

Many thanks go to Dr. Yuxia Hu, who gave me valuable assistance on FEM analysis on

caissons although this was not included in this thesis. I would like to thank A. Prof.

Barry Lehane for the open discussion.

I would like to express my sincere gratitude to the following persons for language

checking on my thesis: Dr. George Vlahos (Chapters 1, 3 - 5); Dr. Susan Gourvenec

(Chapter 6); Dr. Chris Martin (Chapter 7) and Mr. Mark Richardson (Chapters 8 - 9 and

proof reading). Without Mr. Mark Richardson, the sensitive clay sample could not have

been achieved in the centrifuge tests; his help is greatly acknowledged. Many thanks go

to Dr. George Vlahos for his friendship and encouragements on the research. I must

iii Acknowledgement


also thank Dr. Mostafa Ismail, Dr. Qin Lu, Dr. Shambhu Sharma, Ms. Sarah Elkhatib,

Mr. Manh Tran, Mr. Edgard Barbosa-Cruz, Mr. Nobutaka Yamamoro, Dr. Dong Fang

Liang, Dr. Conleth O'Loughlin and Dr. Christophe Gaudin for their help during

different stages of my research. Open discussions with Professor Martin Fahey,

Dr. Christ Martin and A. Professor David Airey are well appreciated.

Help from Mrs. Monica Mackman and other staffs in COFS on general issues are

specially acknowledged.

I received financial support from the University Postgraduate Scholarship for

International Students (UPA-IS), Fee-waiver Scholarship and AD HOC Scholarship

from UWA, and the TOP-UP scholarship from the School of Civil & Resource

Engineering, these supports are sincerely acknowledged. I would also like to appreciate

the International Society of Offshore and Polar Engineers (ISOPE) for awarding me the

‘2004-2005 ISOPE Offshore Mechanics Scholarship for Outstanding Students’.

I would like to acknowledge the help from Professor J.Y. Shi, Professor W.B. Zhao,

Professor Z.Z. Yin (Hohai University, China) and Professor F.H. Lee (formerly

National University of Singapore) on my way to geotechnical research.

Finally, I would like to express my sincere gratitude to my parents for their persistent

love, support and encouragement. Without their understanding and sacrifice, I could

not have achieved anything.

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.

Wen Chen

July, 2005

“God always takes the simplest way.”

-- Albert Einstein

iv Contents


TABLE OF CONTENTS

ABSTRACT......................................................................................................................i

ACKNOWLEDGEMENTS............................................................................................ii

TABLE OF CONTENTS ..............................................................................................iv

NOTATIONS..................................................................................................................xi

1 INTRODUCTION................................................................................................. 1-1

1.1 SUCTION CAISSONS IN DEEP AND ULTRADEEP WATERS ............. 1-1

1.2 PROBLEMS AND AIMS............................................................................. 1-3

1.3 TEST PROGRAM ........................................................................................ 1-5

1.4 THESIS STRUCTURE................................................................................. 1-6

2 LITERATURE REVIEW..................................................................................... 2-1

2.1 GENERAL INTRODUCTION..................................................................... 2-1

2.2 INSTALLATION BEHAVIOUR................................................................. 2-1

2.2.1 Penetration Resistance .......................................................................... 2-1

2.2.2 Necessary Underpressure...................................................................... 2-5

2.2.3 Allowable Underpressure...................................................................... 2-5

2.2.4 Factor of Safety..................................................................................... 2-5

2.2.5 Soil Heave inside Caisson..................................................................... 2-6

2.2.6 Field Example: Suction Anchor for Na Kika FDS ............................... 2-6

2.2.6.1 Penetration analysis........................................................................... 2-7

2.2.6.2 Actual self-weight penetration .......................................................... 2-7

2.2.6.3 Applied underpressure and flow rate ................................................ 2-8

2.2.6.4 Monitored soil heave inside caisson ................................................. 2-8

2.2.6.5 Summary ........................................................................................... 2-9

2.2.7 Uncertainties for Installation............................................................... 2-10

2.2.7.1 Soil flow after passing the first stiffener......................................... 2-10

2.2.7.2 Mode of soil flow under suction ..................................................... 2-10

2.2.8 Radial Stress Changes during Installation .......................................... 2-13

2.2.8.1 Measurements of radial stresses...................................................... 2-13

v Contents


2.2.8.2 NGI method..................................................................................... 2-15

2.2.8.3 Cavity expansion method................................................................ 2-16

2.2.8.4 Strain path method .......................................................................... 2-17

2.2.8.5 MTD method................................................................................... 2-18

2.3 RELAXATION DURING CONSOLIDATION......................................... 2-20

2.4 VERTICAL PULLOUT CAPACITY......................................................... 2-22

2.4.1 Failure Modes ..................................................................................... 2-22

2.4.2 End-bearing Capacity.......................................................................... 2-23

2.4.2.1 Unsealed pullout ............................................................................. 2-23

2.4.2.2 Sealed pullout.................................................................................. 2-23

2.4.2.3 Sealed (base-vented) pullout........................................................... 2-24

2.4.3 Shaft Friction during Vertical Pullout................................................. 2-25

2.4.3.1 Measurements ................................................................................. 2-25

2.4.3.2 Current design method.................................................................... 2-26

2.4.3.3 NGI method..................................................................................... 2-26

2.4.3.4 MTD and CEM method .................................................................. 2-27

2.4.3.5 Discussion ....................................................................................... 2-28

2.5 CONCLUSIONS......................................................................................... 2-28

3 EXPERIMENTAL APPARATUS AND SOIL PROPERTIES........................ 3-1

3.1 INTRODUCTION ........................................................................................ 3-1

3.2 MINIATURE TOTAL PRESSURE TRANSDUCER.................................. 3-1

3.3 INSTRUMENTED MODEL CAISSONS.................................................... 3-1

3.4 CALIBRATION CHAMBER FOR PRESSURE CELLS............................ 3-3

3.5 RING SHEAR APPARATUS ...................................................................... 3-4

3.6 CENTRIFUGE MODELLING: SCALING LAWS ..................................... 3-4

3.7 FIXED BEAM CENTRIFUGE FACILITIES .............................................. 3-6

3.7.1 Strong-box............................................................................................. 3-7

3.7.2 Actuators ............................................................................................... 3-7

3.7.3 Slip Rings.............................................................................................. 3-7

3.7.4 Syringe Pump........................................................................................ 3-8

vi Contents


3.7.5 Load Cells ............................................................................................. 3-8

3.8 T-BAR PENETROMETER .......................................................................... 3-8

3.9 PORE PRESSURE TRANSDUCERS.......................................................... 3-9

3.10 SOIL SAMPLES........................................................................................... 3-9

4 PEFRORMANCE OF MINIATURE TOTAL PRESSURE TRANSDUCERS

ON CAISSONS IN CLAY.................................................................................... 4-1

4.1 INTRODUCTION ........................................................................................ 4-1

4.2 FACTORS AFFECTING STRESS MEASUREMENTS............................. 4-1

4.2.1 Stress Cell Geometry and Properties .................................................... 4-1

4.2.2 Soil Properties ....................................................................................... 4-2

4.2.3 Environmental Conditions .................................................................... 4-3

4.3 SCHEME OF CALIBRATION TESTS ....................................................... 4-3

4.4 TEST RESULTS........................................................................................... 4-4

4.4.1 Calibration Tests in Water .................................................................... 4-4

4.4.2 Calibration tests in Kaolin Clay ............................................................ 4-6

4.4.2.1 Undrained calibration tests in kaolin clay......................................... 4-6

4.4.2.2 Drained calibration tests in kaolin clay............................................. 4-9

4.4.3 Variation of Initial Values of TPTs in Different Media...................... 4-10

4.4.4 Cross-sensitivity to Axial Loading on Caisson................................... 4-11

4.4.5 Calibration Tests in the Centrifuge ..................................................... 4-12

4.4.5.1 Static movement in water................................................................ 4-12

4.4.5.2 Sustained loading in water .............................................................. 4-12

4.4.5.3 Application to caisson penetration in clay in the centrifuge........... 4-13

4.5 CONCLUSION........................................................................................... 4-14

5 STUDYING THE INTERFACE CHARACTERISTICS BETWEEN SUCTION

CAISSON AND CLAY......................................................................................... 5-1

5.1 INTRODUCTION ........................................................................................ 5-1

5.2 RING SHEAR APPARATUS ...................................................................... 5-2

5.3 DESCRIPTION AND GENERAL PRINCIPLES........................................ 5-2

5.4 SOIL SAMPLE PREPARATION ................................................................ 5-3

vii Contents


5.4.1 Fabrication of Top Platen...................................................................... 5-3

5.4.2 Sample Filling ....................................................................................... 5-3

5.4.3 Sample Consolidation ........................................................................... 5-4

5.4.4 Forming the ‘shear plane’ ..................................................................... 5-4

5.4.5 Residual Strength Measurement ........................................................... 5-4

5.5 TEST RESULTS........................................................................................... 5-5

5.5.1 Sample 1: Smooth Ring Platen, NC clay .............................................. 5-5

5.5.2 Sample 2: NC clay, Sand-blasted Ring Platen...................................... 5-6

5.5.3 Sample 3: LOC clay, Sandblasted Ring Platen..................................... 5-7

5.5.4 Sample 4: Sensitive clay, Sandblasted Ring Platen .............................. 5-7

5.6 CONCLUSIONS........................................................................................... 5-8

6 AXIAL CAPACITY OF CAISSONS INSTALLED IN CLAY BY JACKING

AND BY SUCTION .............................................................................................. 6-1

6.1 INTRODUCTION ........................................................................................ 6-1

6.2 CYCLIC T-BAR TESTS FOR SENSITIVITIES OF CLAY....................... 6-3

6.3 FORMULAE FOR CALCULATING AXIAL CAPACITY........................ 6-7

6.4 PENETRATION RESISTANCE.................................................................. 6-8

6.4.1 Installation in NC Clay ......................................................................... 6-8

6.4.1.1 Jacked installation............................................................................. 6-8

6.4.1.2 Suction installation.......................................................................... 6-10

6.4.1.3 Re-installation in disturbed sites ..................................................... 6-15

6.4.1.4 Summary ......................................................................................... 6-17

6.4.2 Installation in LOC Clay ..................................................................... 6-17

6.4.3 Installation in Sensitive Clay .............................................................. 6-21

6.5 AXIAL CAPACITY DURING PULLOUT ............................................... 6-24

6.5.1 Unsealed Pullout in NC Clay .............................................................. 6-24

6.5.2 Sealed Pullout in NC Clay .................................................................. 6-28

6.5.2.1 Pullout after consolidation .............................................................. 6-29

6.5.2.2 Equivalent solid pile test ................................................................. 6-33

6.5.2.3 Immediate pullout ........................................................................... 6-34

6.5.3 Sealed Pullout in LOC Clay after Consolidation ................................ 6-36

6.5.4 Sealed Pullout in Sensitive Clay after Consolidation ......................... 6-37

viii Contents


6.6 CONCLUSIONS......................................................................................... 6-38

7 RADIAL STRESS CHANGES AROUND CAISSONS IN CLAY................... 7-1

7.1 INTRODUCTION ........................................................................................ 7-1

7.2 EXPERIMENTS IN NC CLAY ................................................................... 7-1

7.2.1 Analysis of Radial Stresses during Installation..................................... 7-1

7.2.1.1 Measured σri , derived σ ri and ∆ui during installation...................... 7-1

7.2.1.2 NGI method....................................................................................... 7-3

7.2.1.3 Cavity expansion method.................................................................. 7-4

7.2.1.4 Strain path method ............................................................................ 7-4

7.2.1.5 MTD method..................................................................................... 7-5

7.2.1.6 Comparison between predictions and measurements ....................... 7-6

7.2.2 Relaxation of Radial Stress during Consolidation ................................ 7-7

7.2.2.1 t50 and t90 ........................................................................................... 7-8

7.2.2.2 Post-consolidation radial effective stress........................................ 7-10

7.2.3 Radial Stress Changes and Shaft Friction during Pullout ................... 7-11

7.2.3.1 Pullout after consolidation .............................................................. 7-11

7.2.3.2 Immediate pullout ........................................................................... 7-15

7.3 EXPERIMENTS IN LOC CLAY............................................................... 7-15

7.3.1 Analysis of Radial Stresses during Installation................................... 7-16

7.3.1.1 σri , σ ri and ∆ui during suction installation: test B13SCC.............. 7-16

7.3.1.2 σri , σ ri and ∆ui during suction installation: test B13sus ................ 7-21

7.3.1.3 σri , σ ri and ∆ui during suction installation: test B13cyc................ 7-23

7.3.1.4 σri , σ ri and ∆ui during jacked installation: test B13JCC................ 7-25

7.3.1.5 Summary ......................................................................................... 7-27

7.3.2 Relaxation of Radial Stresses during Consolidation........................... 7-27

7.3.2.1 t50 and t90 ......................................................................................... 7-27


7.3.3 Radial Stress Changes and Shaft Friction during Pullout ................... 7-31

7.4 EXPERIMENTS IN SENSITIVE CLAY................................................... 7-33

7.4.1 Analysis of Radial Stresses during Installation................................... 7-33

7.4.1.1 σri , σ ri and ∆ui during installation: test B14cyc ............................ 7-33

7.4.1.2 σri , σ ri and ∆ui during suction installation: test B14susa............... 7-37

ix Contents


7.4.1.3 σri , σ ri and ∆ui during suction installation: test B14SCC.............. 7-40

7.4.1.4 σri , σ ri and ∆ui during suction installation: test B14sus ................ 7-41

7.4.1.5 Summary ......................................................................................... 7-42

7.4.2 Relaxation of Radial Stresses during Consolidation........................... 7-43

7.4.2.1 t50 and t90 ......................................................................................... 7-43


7.4.3 Radial Stresses Changes and Shaft Friction during Pullout................ 7-46

7.5 CONCLUSIONS......................................................................................... 7-48

8 SUCTION CAISSONS UNDER SUSTAINED LOADING AND CYCLIC

LOADING IN CLAY............................................................................................ 8-1

8.1 SUSTAINED LOADING ............................................................................. 8-1

8.1.1 Sustained Loading in NC clay .............................................................. 8-1

8.1.2 Sustained Loading in LOC clay ............................................................ 8-5

8.1.3 Sustained Loading in Sensitive Clay .................................................... 8-7

8.1.3.1 Pure sustained loading ...................................................................... 8-7

8.1.3.2 Sustained loading after monotonic loading..................................... 8-10

8.1.4 Summary ............................................................................................. 8-11

8.2 CYCLIC LOADING................................................................................... 8-12

8.2.1 Cyclic Loading in NC clay.................................................................. 8-12

8.2.2 Cyclic Loading in LOC clay ............................................................... 8-15

8.2.3 Tests in Sensitive Clay........................................................................ 8-17

8.2.4 Summary ............................................................................................. 8-19

8.3 CONCLUSIONS......................................................................................... 8-19

9 CONCLUSIONS AND FUTURE WORK .......................................................... 9-1

9.1 MAIN FINDINGS ........................................................................................ 9-1

9.1.1 Interface Normal Stress Measurements in Clay in the Centrifuge........ 9-1

9.1.2 Interface Friction Angle between Caisson and Clay............................. 9-2

9.1.3 Installation and Axial Pullout of Caisson ............................................. 9-2

9.1.3.1 Caisson installation ........................................................................... 9-2

9.1.3.2 Relaxation during consolidation ....................................................... 9-4

9.1.3.3 Caisson pullout.................................................................................. 9-5

9.1.4 Behaviour under Sustained Loading and Cyclic Loading .................... 9-6

x Contents


9.2 FUTURE WORK.......................................................................................... 9-6

REFERENCES.........................................................................................................Ref-1

xi Notations


NOTATIONS

Roman

Aext external surface area Aint internal surface area Abase gross cross-sectional area of caisson Aplug cross-sectional area of soil plug Atip cross-sectional area of caisson annulus ch coefficient of horizontal consolidation cv coefficient of vertical consolidation d caisson diameter d shearing distance deq equivalent diameter f caisson shaft friction Fs factor of safety Fs,plug factor of safety when soil plug contacts the caisson lid g gravitational acceleration at earth G shear modulus Gs specific gravity h height h distance between the pile tip and the point of interest hs soil heave hs,act actual soil heave hs,pre predicted soil heave Hi normalised initial total radial stress after installation J jacked installation k gradient of shear strength with depth K0 lateral earth pressure coefficient at rest Kc final radial effective stress after consolidation L caisson length Lnominal nominal depth of caisson during installation Lmax maximum embedment of caisson m mass N scaling ratio for centrifuge model, i.e. test acceleration level Nc end-bearing capacity factor NT-bar bearing capacity factor for T-bar penetrometer p pressure p'0 original in situ mean effective stress in the soil p'i mean effective stress in the soil just after caisson installation P vertical force q bearing pressure r radius rmon nominal radius of the centrifuge R radius of the concentric ring on the ring shear apparatus R overconsolidation ratio used by the CEM analysis S suction installation St clay sensitivity su undrained shear strength

xii Notations


us average undrained shear strength along the caisson embedment t caisson wall thickness t time t50 50% consolidation time t90 90% consolidation time ts thickness of suction-affected area T torque T non-dimensional consolidation time u0 hydrostatic pressure uo hydrostatic pressure outside the caisson lid ui hydrostatic pressure inside the caisson lid v velocity va velocity after installation decreasing vb velocity before installation decreasing V vertical load V non-dimensional velocity Wcais caisson weight (submerged) Wplug soil plug weight (submerged) w water content z embedment of caisson (below mudline) zfinal final depth of instllation for caissons zj embedment of caisson when TPT leaving jacking-affected area zs embedment of caisson when TPT entering suction-affected area zplug embedment of caisson when soil plug contacts the lid of caisson zTPT embedment of TPT zTPT,c embedment of TPT after consolidation zTPT,i embedment of TPT at the end of installation zTPT,j embedment of TPT when leaving jacking-affected area zTPT,s embedment of TPT when entering suction-affected area

Greek

α shaft friction ratio αext shaft friction ratio on the external wall of caisson αint shaft friction ratio on the internal wall of caisson δp peak interface friction angle δr residual interface friction angle ∆ (as prefix) used to denote change from initial or reference value ∆p net pressure during installation or pullout of the caisson ∆ua allowable underpressure

∆uapp applied underpressure ∆ui excess pore pressure generated during installation ∆un necessary underpressure φ friction angle γsat saturated unit weight of soil γ effective unit weight of soil θa chain angle (above horizontal) λ coefficient used by CEM

xiii Notations


µ coefficient used by CEM ρ density ρ area ratio σ stress σri radial total stress during installation σ ri radial effective stress during installation σ v0 vertical effective stress at rest τ frictional resistance in cohesive soil ω angular rotation (of centrifuge)

Principal Subscripts / Superscripts

ave average e external f failure i internal i installation m model max maximum min minimum nom nominal o external p prototype u undrained r radial sub submerged v vertical ' effective 0 original

Principal Abbreviations

API American Petroleum Institute CAF Cell Action Factor CEM Cavity Expansion Method CC Closed-ended pullout after Consolidation CI Closed-ended pullout immediately after Installation CLA Centre Line Average COFS Centre for Offshore Foundation Systems DAQ Data Acquisition DPA Deep Penetration Anchor FDS Floating Developing System FEM Finite Element Method FPSO Floating Production Storage and Offloading FPU Floating Production Unit GDS Geotechnical Digital Systems ISOPE International Society of Offshore and Polar Engineering LB Lower Bound LL Liquid Limit LOC Lightly Overconsolidated Clay

xiv Notations


MIT Massachusetts Institute of Technology MTD Marine Technology Dictorate NC Normally Consolidated NGI Norwegian Geotechnical Institute OC Over Consolidated OC Open-ended pullout after Consolidation OCR Overconsolidated Ratio OI Open-ended pullout immediately after Installation PI Plasticity Index PL Plastic Limit PLS Piezo-Lateral Stress cell PPT Pore Pressure Transducer SPM Strain Path Method TLP Tension Leg Platform TPT Total Pressure Transducer UB Upper Bound UWA University of Western Australia VIV Vortex-induced Vibrations VLA Vertical Loaded Anchor YSR Yield Stress Ratio

Chapter 1 1-1 Introduction


1 INTRODUCTION

1.1 SUCTION CAISSONS IN DEEP AND ULTRADEEP WATERS

The offshore oil and gas industry is moving rapidly towards deep and ultradeep waters

(Figure 1.1), with some locations reaching depths between 1500 m and 3000 m

(El-Gharbawy et al., 1999; Aubeny et al., 2001; Colliat & Dendani, 2002; Sparrevik,

2002). According to the Infield Worldwide Offshore Energy Database

(www.infield.com), a threefold growth is anticipated in deepwater platform prospects

during the next five years. Twenty-four production platforms exist in water depths

greater than 500 m, but during the next 5 years, 66 potential deepwater platform

installations are expected in the Gulf of Mexico, West Africa, North Sea, Offshore

Brazil, Timor Sea and South China Sea. In Western Australia, several major gas

discoveries, including Jansz, Io and Geryon, have been made in 1999 - 2001 on the

Exmouth Plateau, Carnarvon Basin, under water depths of 1000 - 2000 m (see

Figure 1.2).

Various types of deepwater facilities (Figure 1.3) such as Tension Leg Platforms (TLP),

SPARs, Floating Production Storage and Offloadings Vessels (FPSOs) and

Semi-submersibles, have been developed to face the offshore challenges (Andersen et

al., 1993; Clukey et al., 1995; Colliat J-L., 2002; Loez, 2002; Huang et al., 2003). For

the ultradeep water, it seems the future will see continued use of FPSOs. According to

Mustang Engineering Research (2004), there are 97 FPSOs already in operation, 19

under construction and 16 more planned up to the end of 2004 (Figure 1.4). FPSOs

have flourished in Australia with 8 already in service, 2 under construction and another

1 planned. Extensive use of FPSOs can also be seen in the offshore oil and gas industry

around other continents.

Deepwater mooring technology is critical in securing offshore drilling and production

vessels under various hostile conditions, with loads arising from waves, wind, loop

currents and even tsunami. Methods for anchoring these offshore facilities have

evolved from the traditional catenary mooring systems to taut-leg mooring systems,

where the angle between the mooring line and the mudline may be as high as 40 to 50

(Ehlers et al., 2004). Several types of offshore foundations, such as vertically loaded

(drag embedment plate) anchors (VLAs), suction caissons, deep penetration anchors

(DPAs) and suction embedded plate anchors (SEPLAs) have been brought into service



for the anchoring system (Dendani & Colliat, 2002; Audibert et al., 2003; O’Loughlin

et al., 2004; Ehlers et al., 2004). Among these choices, suction caissons are considered

as both economical and effective due to their ability to resist combined vertical and

horizontal loading, and the relatively simple installation procedures (Solhjell et al.,

1998; Tjelta, 2001). Suction caissons are a practical alternative to driven piles,

especially in ultradeep waters (> 2000 m), because of the relatively high cost of the

latter (Colliat, 2002).

A suction anchor is generally a large diameter stiffened cylindrical shell, opened at the

bottom and closed at the top (see Figure 1.5), with large diameters of about 2 to 8 m and

length to diameter ratios of 6 or less (Andersen et al., 2005). Suction caissons are

installed initially by their self-weight, with further penetration achieved by pumping the

water out through an opening in the top lid of the caisson, thus developing

underpressure (or suction) sufficient to force the caisson downwards (Ehlers et al.,

2004). After installation, the lid is sealed and ‘passive suction’ is developed to provide

resistance to transient vertical and inclined load (Morrison et al., 1994; Andersen &

Jostad, 2002). The mooring line is then attached to the caisson through a pad-eye

located at approximately 2/3 of the embedment depth, which is close to the centroid of

the lateral soil resistance for the typical normally to lightly overconsolidated clay in

deepwater (Ehlers et al., 2004).

One of the earliest uses of suction caissons was in the form of short concrete ‘tricells’

(Andersen et al., 1993; Støve et al., 1992) to anchor the tension leg platform (TLP) for

the Snorre field in 310 m water depth. By 1999, water depths where the caissons had

been used had increased to 1400 m in the Gulf of Mexico, for anchoring the Hoover-

Diana SPAR (Colliat, 2002). At the same time, suction caissons were also installed

successfully in 400 m water depth at Laminaria in the Timor Sea by Woodside Energy

Ltd, as the mooring anchor for the Northern Endeavour FPSO. The soil there was

calcareous mud, although the soil exhibits ‘clay like’ properties (Erbrich & Hefer,

2002). In Africa, suction caissons were installed as mooring anchors for the Girassol

FPSO in 2001, in water depths of 1400 m (Dendani & Colliat, 2002). In 2002, suction

caissons were successfully installed in water depths of 2000 m, as anchors for the Na

Kika Floating Developing System (FDS) which was jointly developed by Shell & BP,

in the Mississippi Canyon Area of the Gulf of Mexico (Newlin, 2003a). Caissons in

this project were installed with good positioning, orientation and verticality (Newlin,

2003b). The successful application proved that suction caissons are a reliable anchoring



system in ultradeep water. To date more than 485 suction caissons have been installed

successfully at more than 50 locations in water depths to nearly 2000 m until 2004. A

summary of the available information on suction caissons is given in Table 1.1

(Andersen et al., 2005; Ehlers, et al., 2004).

Table 1.1 Summary of installed suction caissons for deepwater floaters

Year Field Floater Water depth

(m) Size (d×L)

(m×m)

No. Operator

1991 Snorre1) TLP 335 30×13 4 Saga

1996 Norne1) FPSO 375 5×10 12 Statoil

1997 Marlim P19-P263) Semi FPU 770-1000 4.7×13 32 Petrobras

1997 Schiehallion4) FPSO 400 6.5×12 14 BP

1997 Aquila5) FPSO 850 4.5-5×16 8 Agip

1998 Laminaria6) FPSO 400 5.5×13 12 Woodside

1998 Marlim P183) FPSO 900 4.7×20 2 Petrobras

1999 Marlim P353) FPSO 810-910 4.8×17 6 Petrobras

1999 Troll C1) Semi FPU 350 5×15 12 Norsk Hydro

2000 Hoover-Diana7) SPAR 1500 6.4×32 12 Exxonmobil

2001 Girassol2) FPSO 1350 4.5×17 16 TFE

2002 Horn Mountain7) SPAR 1650 5.5×27.4-29 9 BP

2002 Na Kika7) FDS 1920 4.3×23.8 16 Shell/BP

2003 Devils Tower7) SPAR 1700 5.8×34.8 9 Dominion

2003 Holstein7) SPAR 1280 5.5×36.3-38.4 16 BP

2004 Thunder Horse7) Semi FPU 1830 5.5×27.5 16 BP

2004 Mad Dog7) SPAR 1600 7.6×14.6 11 BP

2004 Atlantis7) Semi-PQ 2130 NA 12 BP

1) North Sea 2) West Africa 3) Offshore Brazil 4) West of Shetlands 5) Adriatic Sea 6) Timor Sea 7) Gulf of Mexico 8) Offshore Nigeria

1.2 PROBLEMS AND AIMS

The early focus on suction caisson capacity was on quasi-horizontal and moment

loading imposed by the catenary mooring line (Andersen et al., 1993; Allersma et al.,



1999; McCarron & Sukumaran, 2000), as part of anchoring systems in shallow water

areas. Caissons used in deep to ultradeep water areas generally have embedment ratios,

L/d, in the range 2.5 to 5; for such geometries, the capacity under purely horizontal

motion is typically double the vertical capacity. The latter will therefore govern design

once the loading angle exceeds 40 to the horizontal, as will be the case for taut or semi-

taut mooring systems (Clukey et al., 2004), such as for deepwater anchoring systems for

FPSOs (Figure 1.6). This has led to increased attention on estimating the external shaft

capacity of suction caissons, which typically represents 40 to 60% of the total vertical

capacity (Huang et al., 2003), and the magnitude and time scale of consolidation effects

following installation. Various types of sediments exist in the deepwater areas, for

instance, normally consolidated to lightly overconsolidated clay in the Gulf of Mexico,

ultra-high plasticity soft clays in West Africa, and fine-grained carbonate soil with

clay-like properties in Australia (Randolph et al., 2005). However, currently there are

no established design guidelines for the shaft capacity of suction caissons in soft marine

clay, the calculation of which is mainly based on conventional design methods for open-

ended driven piles (e.g. API, 1993). Andersen & Jostad (2002) have argued that the

method of installation leads to external friction that may be 30 to 40% lower than that

for driven piles.

Compared to driven tubular piles, suction caissons have lower wall thickness and thus a

much larger ratio of diameter (d) to wall thickness (t), with d/t values in the range of 60

to 200 rather than a typical range of 30 to 50 for piles. Even if the full volume of steel

were accommodated by outward soil movement, it would be expected that the resulting

radial stress and pore pressure changes for suction caissons would be lower than during

pile installation. Additionally, Andersen & Jostad (2002) have argued that during

suction installation the wall thickness of the caisson is accommodated by purely inward

motion of the soil, thus reducing the increase in radial total stress outside the caisson,

compared to that during jacked installation. Their conceptual model leads to very

localized excess pore pressures (arising from shearing only) and thus rapid

consolidation with times for 50% consolidation generally less than one day. The shaft

friction during installation is taken as the remoulded shear strength, while the long-term

value was estimated as 58 to 65% of the intact shear strength.

The pattern of soil flow at the caisson tip, and the proportion of the caisson wall that is

accommodated by inward or outward displacement of the soil, has important

consequences for quantifying the shaft friction around suction caissons and the rate of



consolidation. The axial capacity of suction caissons is associated closely with the

variation of the radial stress during installation, consolidation and pullout of the caisson.

Exactly how the variation of radial stress is affected by the mode of installation (jacking

or suction) for thin-walled suction caissons has not been studied. It is important to find

a theoretical approach for predicting the radial stress changes around suction caissons

and hence the external shaft friction. Therefore, the present research on the axial

behaviour of caissons in clay is aimed at solving the following issues:

• Differences between caissons installed by jacking and by suction, in terms of :

o penetration resistance; o changes in radial total stresses, excess pore pressures and radial effective

stresses during installation; o radial effective stresses after consolidation; o vertical pullout capacity after consolidation.

• Accuracy of theoretical predictions of the measured radial stress changes during

installation of the caisson and after consolidation.

• Shaft friction and reverse end-bearing capacity during pullout, allowing for:

o effect of sustained loading; o effect of cyclic loading.

1.3 TEST PROGRAM

To resolve the issues put forward above, a series of physical model tests on suction

caissons were undertaken in the beam centrifuge at the University of Western Australia.

Tests were performed in reconstituted kaolin clay, but with variation in

overconsolidation ratio and sensitivity, in order to simulate soil properties encountered

in deepwater sediments. The research focused on the radial stress changes and the axial

capacity of caissons during installation, consolidation and axial pullout. Experiments

started with a series of calibration tests on the reliability of the pressure cells for

measuring the radial stress changes on caissons in clay, with tests performed in both 1 g

and high g conditions on the centrifuge. Then the interface friction characteristics

between the caisson and the clay were studied by ring shear tests. Finally, centrifuge

model tests were undertaken in normally consolidated (NC) clay, lightly

oveconsolidated (LOC) clay and sensitive clay. Radial stress changes were measured at

different elevations of the caisson during installation, consolidation and vertical pullout,

for caissons installed either by jacking or by self-weight penetration followed by



suction. Measurements were compared with various theoretical predictions in order to

find an effective approach for design purposes. Radial stress variations around caissons

under both sustained loading and cyclic loading were then measured and analysed.

1.4 THESIS STRUCTURE

This thesis consists of 9 chapters, the outline for which is described below:

Chapter 2 provides a literature review of previous studies on suction caissons in clay.

The current design rule is discussed.

Chapter 3 summarises the physical modelling equipment and experimental details

relevant for the testing program. The fixed beam geotechnical centrifuge is described.

Instruments such as the T-bar penetrometer, pore pressure transducers and total pressure

transducers are introduced. Design and instrumentation of the model caissons used in

this research are described. Soil properties of the kaolin clay used in this research are

introduced.

Chapter 4 addresses the calibration tests on the total pressure transducers in clay. A

series of calibration tests are carried out both at 1 g in a modified triaxial apparatus, and

under high g conditions on the centrifuge.

Chapter 5 presents the results of the ring shear tests. The residual interface friction

angles between the caisson shaft and NC, LOC and sensitive clays are obtained. An

effective step for improving the quality of ring shear tests is put forward.

Chapter 6 focuses on the axial capacity of suction caissons in soft clay, during both

installation and vertical pullout. The penetration resistance is compared for caissons

installed either by jacking or by self-weight penetration followed by suction. Soil heave

and factor of safety during suction installation are analysed. A model for prediction of

the shaft friction and end-bearing capacity during caisson installation is assessed.

Cyclic T-bar tests are performed to investigate the sensitivity of the clay. Sealed

pullout capacity of caissons installed by the two methods is also compared. Normalised

axial capacity during sealed pullout is obtained. Lower bound external shaft friction

ratios are derived from the vertical sealed pullout capacity after consolidation. The

axial capacity of a small solid pile with an equivalent diameter to the model caisson is

also tested.

Chapter 7 discusses the external radial stress changes measured during installation,



consolidation and vertical pullout of the caisson. Measurements are compared for

caissons installed by jacking and by suction. Comparison is made between the

measured radial stresses and various theoretical predictions. Data are presented for the

time needed for 90% consolidation and the final radial effective stress after

consolidation, and comparisons are made between such measurements and theoretical

predictions. Finally, upper bound external shaft friction ratios of caissons during

vertical pullout after consolidation are derived from the measured radial stress at failure.

In Chapter 8, the capacity ratios of caissons subjected to sustained loading and cyclic

loading are discussed. The measured radial stress changes around the caisson during

these two loadings are analysed. External shaft friction ratios are derived from the

measured radial stress when the caisson is loaded to failure for these two types of

loadings.

Chapter 9 summarises the major findings and recommendations for further work on this

topic.

Chapter 2 2-1 Literature Review


2 LITERATURE REVIEW

2.1 GENERAL INTRODUCTION

The axial behaviour of suction caissons during installation and vertical pullout in clay

has been studied by field tests (Hogervorst, 1980; Andersen et al., 1993; Newlin, 2003a,

b), 1 g laboratory tests (El-Gharbawy, 1998; Whittle et al., 1998; Luke, 2002; Andersen

& Jostad, 2004), centrifuge tests (Steensen-Bach, 1992; Watson, 1999; House et al.,

1999; Cao et al., 2002a, b; Randolph & House, 2002) and numerical analysis (Hu et al.,

1999; Deng & Carter, 2000; Zdravkovic et al., 2001; AG, 2002; Andersen & Jostad,

2002; Andersen et al., 2004; Cao et al., 2002c; Templeton, 2002; Supachawarote et al.,

2004).

A detailed list of the experimental work carried out in clay up to date is summarised in

Table 2.1 (Andersen et al., 2005). Two aspects including 1) installation and 2) holding

capacity of the caisson behaviour were generally investigated during these experiments.

The design information for predicting installation performance includes: 1) self-weight

penetration depth, 2) required underpressure with depth, 3) allowable underpressure as a

function of depth, 4) soil heave as a function of depth, and 5) maximum recommended

penetration depth including computed factor of safety against plug failure (Andersen &

Jostad, 1999; Offshore Technology Research Center, 2001; Ehlers et al., 2004).

2.2 INSTALLATION BEHAVIOUR

2.2.1 Penetration Resistance

According to the existing methods for predicting caisson installation (Andersen &

Jostad, 1999), the penetration resistance, Qtot, of suction caisson anchors is calculated as

the sum of the integrated interface shear strength along the external and internal skirt

walls, the end-bearing resistance and resistance from external protrusions (see Figure

2.1):

extrastipsidetot QQQQ ++= (2.1)

where

Qside = interface shear strength (sum of internal and external)

Qtip = tip resistance from skirt tip



Qextras = resistance from internal stiffeners, external protrusions such as

pad-eye

External and internal shaft resistance is estimated by integrating the remoulded strength

over the external and internal embedded surface of the caisson. The remoulded strength

may be estimated by either an effective stress or a total stress approach

(Andersen et al., 2005).

Table 2.1 Experimental studies on suction anchors (after Andersen et al., 2005)

Year Test site Test type Test Description Reference

1985 Gullfaks

North Sea Full scale field test Installation and extraction of 2 large diam.

(6.5×22 M) concrete cylinders Tjelta et al. (1986)

1989 NGI/ Lysaker Large scale field model tests

Monotonic and cyclic TLP loads 10° from vertical.

Dyvik et al. (1993) Andersen et al. (1993)

1991 Focomorto Field model test Installation of concrete skirt pile. O’Neill et al. (1991)

1991 ISMES Centrifuge Installation and monotonic & cyclic vertical load.

Renzi et al. (1991)

1991 DGI Centrifuge Installation and uplift tests, two uniform shear strength profiles.

Fuglsang et al. (1991)

Steensen-Bach (1992)

1991 NGI/Lysaker Large scale field model tests

Monotonic and cyclic lateral loads 10o from horizontal.

Keaveny et al. (1994)

1990-93 LCPC Centrifuge Monotonic and cyclic uplift tests on two different size caissons. One lateral test.

Clukey et al. (1993/95)

1996 Tordis Field test Installation and removal of skirted anchor (5x8 m). Incl. 3 months testing of fibre rope.

Offshore Engr. (1996a)

1996 Marlin Field test Installation, testing and removal of 3.6 m diameter, 18 m long skirted anchor.

Offshore Engr. (1996b)

1998 MIT 1g lab model test Installation & capacity. Miniature caisson. Clay Whittle et al. (1998)

1998 UWA Centrifuge Monotonic and cyclic lateral loads Randolph et al. (1998)

1996-99 UWA Centrifuge Installation, monotonic & cyclic capacity, combined loading

Watson et al., (2000)

1997-99 GeoDelft Centrifuge Installation, monotonic & cyclic capacity Andersen et al. (2003)

2000 C-Core Centrifuge Installation and undrained uplift Cao et al. (2002a)

1998-04 Univ. of Texas, Austin

1 g lab models Installation and monotonic capacity. Kaolin Olson et al. (2003) Rauch et al. (2004)

2001 UWA Centrifuge Installation and undrained uplift House & Randolph (2001)

2002 C-Core Centrifuge Uplift capacity for installation with or without suction. Kaolin

Clukey & Phillips (2002)

In the effective stress approach, the external friction, fs, is determined as:



tanδσ ris ⋅′=f (2.2)

where σ'ri is the radial effective stress during installation and δ is the interface friction

angle between the caisson and soil, which can be determined from ring shear tests.

Particularly low values of δ may be relevant where the caisson surface is painted

(Dendani & Colliat, 2002).

In the total stress approach, the remoulded shear strength f is determined either from 1)

direct measurements of the strength of remoulded samples, or 2) the intact undrained

shear strength, su, divided by the sensitivity, St, expressed as:

uut

αssS1

==f (2.3)

where α is referred to as the shaft friction ratio. This calculation is taken as the same as

that for piles in API RP2A(1993), which is in fact based on the ‘α ‘ method put forward

by Randolph & Murphy (1985), although the wall thickness of caissons is generally

much less than that of open-ended piles.

The tip resistance (Qtip,i) of the ith bearing surface is computed by multiplying the ith

steel tip area (Atip,i) with the corresponding undrained shear strength (su.i) using the

end-bearing capacity factor (Nc,i), as shown below:

( ) itip,iuiciitip, AzγsNQ ⋅⋅′+⋅= (2.4)

where

Nci = bearing capacity factor for ith surface

sui = local undrained shear strength for ith surface

γ' = effective unit weight of soil

zi = embedded depth of ith surface (but limited to the height of internal stiffeners)

Atip,i = area of ith bearing surface (caisson tip or stiffener)

Therefore, if the total stress approach is used, the total penetration resistance of caissons

can be expressed as follows (Chen & Randolph, 2004a):

( ) ( ) aiAaiτbiAextAusαn

1iitip,AγizuisciNA∆ptotQ base −−+−+∑

=+′+=⋅= (2.5)

where

∆p = net penetration pressure



Abase = cross-sectional area of the sealed caisson base

us = average undrained shear strength along the caisson embedment depth

Aext = external area of caisson shaft in contact with soil

Ai-b = internal area of the caisson shaft in contact with soil

Ai-a = area of internal shaft above upper edge of pad-eye stiffener

τi-a = nominal friction for internal shaft above the first stiffener

This formula is a simplified version of that presented by House & Randolph (2001),

based on the assumption of equal internal and external friction ratio, α.

Generally some allowance for the effects of remoulding should be made on the bearing

resistance on internal stiffeners within the caisson (Andersen et al., 2005). Also, the

internal shaft resistance above the first ring stiffener is generally reduced, allowing for

remoulding of the soil and incomplete flow of the soil back against the internal caisson

wall (Erbrich & Hefer, 2002). The internal shaft friction of the caisson is determined by

the flow mechanism of the soil after passing the first internal stiffener. Three modes of

flowing are considered to be possible: 1) full attachment, 2) no attachment and 3) partial

attachment, with some water entrapped between the soil and the wall. The above three

modes are shown in Figures 2.2a - c.

Full attachment is possible for very soft clay with low strength ratio, su/σ′v0, and high

sensitivity, for example normally consolidated high plasticity clays. The full

detachment case (Figure 2.2b), giving rise to a free-standing soil plug inside the caisson

and thus zero shaft friction, is thought to be possible for stiff clay, with high strength

ratio, su/σ′v0, and low sensitivity, for example, heavily overconsolidated clay. The

intermediate case (see Figure 2.2c), will occur for NC clay or LOC clay with medium

strength ratio and sensitivity. The soil plug will be free-standing for a short distance

before collapsing, entrapping some water between the soil and the internal wall. For

that case, the internal friction would be lower than that below the first stiffener,

although it should be larger than zero (Chen & Randolph, 2005).

Measurements of the penetration resistance from model tests or field installations allow

compatible sets of Nc and α values to be derived by Equation 2.5. Randolph et al.

(1999) suggested using Nc = 7.5, and α = 0.3 to 0.5. Cao et al. (2002a) obtained α

values of 0.25 to 0.3 by assuming Nc = 9.5, according to centrifuge tests on model

caissons. House (2002) gave a slightly higher value of α = 0.35 for stiffened caissons,



also by conducting centrifuge tests. Newlin (2003b) presented a range of 0.28 to 0.43

for α. The range of Nc and α values suggests that further study is needed, particularly

spanning clays of differing sensitivity.

2.2.2 Necessary Underpressure

According to Andersen & Jostad (1999), the necessary underpressure (see Figure 2.3a),

∆un, needed to install the caisson is calculated by:

( ) plugtotn AWQ∆u ′−= (2.6)

where

W′ = submerged weight during installation

Aplug = cross-sectional area of the soil plug

2.2.3 Allowable Underpressure

The allowable underpressure (see Figure 2.3b), ∆ua, or the capacity of the soil plug to

resist uplift failure is the sum of the reverse end-bearing capacity plus the internal shaft

friction of the anchor (Andersen & Jostad, 1999):

pluguinsidetipu,ca AsαAsN∆u ⋅⋅+⋅= (2.7)

where su,tip is the undrained shear strength at the caisson tip and Ainside is the internal

surface area of the caisson. The Nc value is generally taken as 6.2 to 9.0, depending on

the depth/diameter ratio during penetration (Andersen & Jostad, 1999).

2.2.4 Factor of Safety

The factor of safety (Fs) with respect to large plug heave can be calculated by two

different ways. A simple approach is to define it as the ratio of the computed allowable

underpressure and the predicted necessary underpressure (Ehlers et al., 2004):

nas ∆u∆uF = (2.8)

A more logical approach is based on the material coefficient for the strength used to

calculate the uplift capacity of the clay plug at caisson tip level, assuming a material

coefficient of unity on the remaining components of internal plug resistance (Andersen

& Jostad, 1999); this can be expressed as follows:



insideuplugapplied

plugtipu,cs AsαAp

AsNF

⋅−⋅

⋅⋅= (2.9)

where papplied is the actual suction pressure applied during installation. The logic for the

second approach is that the internal plug resistance contributes equally to the required

underpressure and the plug resistance, and so uncertainty in the internal resistance

should not affect the estimation of safety (Andersen et al., 2005). A minimum value of

1.5 is generally recommended for Fs. The maximum recommended depth of penetration

in soft clays may be 7 to 15 times the diameter, depending on the shaft friction, unit

weight of soil and safety factor (Ehlers et al., 2004).

2.2.5 Soil Heave inside Caisson

The soil heave inside the caisson during installation is estimated by assuming all the

clay replaced by the caisson wall and stiffener goes into the caisson during installation

by underpressure (suction) (Andersen & Jostad, 1999). The caisson length must be

increased by the soil heave height in order to achieve the target embedment of the

caisson.

It is useful to review a field project of caissons installed by suction in clay, including

both design and installation procedures.

2.2.6 Field Example: Suction Anchor for Na Kika FDS

A well documented case report of suction caisson installation in ultradeep water is

provided by Newlin (2003a, b). In August of 2002, suction caissons were successfully

installed in a water depth of 2000 m, as mooring anchors for the Na Kika Floating

Developing System (FDS) in the Mississippi Canyon Area of the Gulf of Mexico

(GOM), while hook-up with the FDS occurred in July of 2003. This latest, and deepest,

application of suction caissons will be reviewed in detail.

The Na Kika FDS is designed for handling oil and gas production. This FDS is a

semi-submersible hull with a displacement of approximately 60,000 tonnes moored

with a 16-leg, semi-taut mooring system in a 4 × 4 cluster configuration. An elevation

view of the suction anchor as part of the mooring system for the FDS is shown in Figure

2.4.

The seabed soils comprise normally consolidated clay with plasticity index, Ip, between

35 and 60. Variations of typical soil properties, effective unit weight (γ′) and undrained



shear strength (su), with the depth below the seabed, are shown in Figure 2.5. The

caissons (see Figure 2.6) had a diameter (d) of 4.3 m, a length of 25 m and a thickness

(t) of 22 mm (increased to 51 mm at the padeye). The caisson was designed to

penetrate to a depth (L) of 23.8 m, giving an aspect ratio, L/d, of 5.5. A thick-walled

shoe with external chamfer was used at the caisson tip.

2.2.6.1 Penetration analysis

Calculations for penetration analyses included soil resistance, necessary underpressure

to install the caisson to designed penetration, allowable underpressure with respect to

large soil heave and expected soil heave inside the caisson, using the method described

by Andersen & Jostad (2002).

The shaft friction along the caisson is taken as a function of the undrained shear

strength and shaft friction ratio, α, which is defined in Equation 2.3. Since the

sensitivity index, St, varies between 2.35 to 3.55, the derived α values thus vary

between 0.28 and 0.43. The St values are consistent with typical range of 2 to 4 for

normally consolidated clays in the GOM.

Based on the α values above, the expected range of necessary underpressure (∆un) to

install the pile to full penetration of 25 m is calculated by Equation 2.6, and the result

with respect to penetration depth is plotted in Figure 2.7a for the NE anchor group, as

an example. From the expected self-weight penetration of 12.2 m, ∆un increases

steadily with depth, with local variations attributed to the slightly higher strength. At

the designed penetration depth of 23.8 m, the predicted range of ∆un is 200 to 300 kPa

for all anchor groups. It should be noted that, in terms of the analysis, the installation is

assumed to be continuous without significant delays.

The allowable underpressure (∆ua) for the NE anchor group is computed from Equation

2.7, and shown in Figure 2.7b. It also increases gradually with penetration depth. At

the designed penetration depth of 23.8 m, ∆ua ranges between 515 and 620 kPa. The

factor of safety is calculated by Equation 2.8, and is approximately five at the expected

self-weight penetration decreasing to two at full penetration. This means suction

installation is safe without causing significant soil heave inside the caisson.

2.2.6.2 Actual self-weight penetration

Actual self-weight penetrations for each of the anchor groups are presented in Table



2.2. Self-weight penetrations averaged 13 m, slightly over the expected 12.2 m from

penetration analyses, with no significant trends observed when comparing anchor

groups. The average self-weight penetration depth is around 50% of the final

embedment of the caisson.

Table 2.2 Observed self-weight penetration of Na Kika FDS (after Newlin 2003b)

Anchor group Average self-weight penetration

NE 13.1

SE 13.0

SW 13.4

NW 12.7

2.2.6.3 Applied underpressure and flow rate

The applied underpressure and flow rate during suction operations are plotted versus

caisson penetration in Figure 2.8. As expected, the actual underpressure falls between

the lower bound and upper bound necessary underpressures, showing that the suction

installation was performed well, with no plug failure occurring during installation. For

each caisson, the flow rate typically reached a maximum of around 3000 lpm (or 0.05

m3/s) at the start of suction operations, and then gradually decreased to around 2000

lpm (or 0.03 m3/s) at grade.

In general, the pump operations took about an hour for each caisson to install the

caisson to the target depth. Time taken for completion of each caisson installation,

including any downtime, is shown in Figure 2.9. The minimum time was 62 minutes,

while the maximum time was 244 minutes (due to hose collapse from an obstruction).

2.2.6.4 Monitored soil heave inside caisson

Soil plug elevation was measured relative to the nominal seabed in select caissons using

a simple “dipstick” tool inserted in the pump interface. The tool had paint marks every

15 cm, from which a soil elevation was read accurate to the nearest paint mark.

Table 2.3 gives the six soil plug measurements performed at Na Kika. The

measurements indicate that soil heave inside the caisson was insignificant, less than the



expected 0.66 m (which assumed that the displaced soil goes fully outside the caisson

during self-weight penetration, and then goes fully inside the caisson during the suction

operation). The lack of soil heave inside the caisson may be due to the external bevel at

the bottom of the caisson, perhaps causing the majority of displaced soil to be directed

outside the caisson.

Table 2.3 Observed soil heave inside caisson during suction installation of Na

Kika FDS (after Newlin 2003b)

Anchor group Mooring leg Average soil heave

(mm)

NE L1-P6

L2-P9

+150

+150

SE L5-P8

L7-P13

+150

0

SW None measured

NW L15-P1

L16-P3

+300

-150

2.2.6.5 Summary

The Na Kika suction caissons were installed successfully, with the required

underpressures falling within the expected range. The main conclusions from Na Kika

caisson installation are as follows:

• For penetration analysis, the most likely undrained shear strength profile was

used with α = 0.28 to 0.43. Actual underpressure concurred with the predicted

range.

• Actual self-weight penetrations of all piles averaged 13 m versus the expected

12.2 m.

• The measured soil heave inside the caisson averaged 0.1 m, much less than the

expected 0.66 m, possibly due to the bevel installed at the tip of the caisson.

Therefore, technology is mature for installing suction caissons in ultradeep waters.

Predictions of the penetration resistance agree well with the measurements.



2.2.7 Uncertainties for Installation

2.2.7.1 Soil flow after passing the first stiffener

A review of previous studies shows general agreement that the frictional resistance on

the inner and outer walls may be taken as equal to the remoulded shear strength, which

can be predicted with a relatively high degree of accuracy (Andersen et al., 2005).

However, there seems to be uncertainty regarding the effect of internal ring stiffeners,

which could cause much lower friction above the first stiffener. Internal ring stiffeners

may also create a gap between the soil and the internal wall, after the soil has passed the

first internal stiffener; some water may be trapped in such a gap and make the

estimation of the shaft friction in that region difficult (Chen & Randolph, 2005).

2.2.7.2 Mode of soil flow under suction

Uncertainty also exists on the mode of soil flow at the caisson tip during suction

installation.

Based on the finite element analysis, Andersen & Jostad (2002) proposed that for a flat

tip caisson, the soil displaced by the caisson wall will move 50% outside and 50%

inside the caisson during jacked installation, and 100% inside during installation by

underpressure (see Figure 2.10). This assumption is widely adopted in current designs

(Clukey & Phillips, 2002; Huang et al., 2003; Ehlers et al., 2004; Andersen et al., 2005).

It should be noted that in the study by Andersen & Jostad (2002), although an attempt

was made to model large penetration of the caisson wall within the finite element

analyses, this was accomplished by incrementally transforming soil elements beneath

the advancing caisson wall from soil to steel, rather than by a true large penetration

analysis.

The long term capacity of the caisson may be affected by the soil flow during

installation, since inward flow will lead to lower external radial stresses, and thus lower

shaft friction, compared to caissons installed by jacking (Andersen & Jostad, 1999).

There are some research results supporting such an assumption. By conducting 1 g

laboratory tests on a miniature caisson in the resedimented Boston Blue Clay, Whittle et

al. (1998) stated that almost the entire volume of soil displaced by the wall moves

inside the soil plug. It should be pointed out that the model caisson they used has a

length of 330 mm, an outside diameter of 50.8 mm and a wall thickness of 1.45 mm; the

influence of soil self-weight, which is a very important issue for the mode of soil flow



during deep penetration, was not reflected in their tests, therefore the results may not be

applicable to real caissons. Renzi et al. (1991) mentioned a reduction of ~50 kPa in the

penetration pressure for a certain depth during suction installation, compared to that

under jacking. However, no detailed data was presented on the comparison and the

conclusion is thus not convincing.

Centrifuge tests reported by Andersen et al. (2003) show that, when the caisson tip

passed a piezometer which was embedded at 9.5 m depth in the clay, there was some

reduction in the measured excess pore pressure at 0.71 m away from the external wall

(see Figure 2.11); this reduction is considered to be the result of underpressure.

However, it can be seen clearly that the reduction was rather small, being less than 5

kPa for two tests (although in the third test it was much larger, which did not agree with

the other two and was thus considered unreliable); it cannot therefore be taken as a

proof of the assumed soil flow mode. After that reduction, the excess pore pressure

continued to increase with further penetration depth of the caisson, showing that the

effect of suction was at most transient. Unfortunately no data from jacked installations

in the same set-up was reported in their paper, and comparison cannot be made directly

between the variations of external excess pore pressures for the two types of

installation. In addition, measurement of plug heave from these tests suggested that

the soil displaced by the caisson wall moves inward during suction installation. At a

penetration depth of 7 diameters, about half the maximum penetration depths, the

volume of the soil heave inside the caisson actually increased more than the volume of

the displaced clay. In fact, the caisson they tested had a very large acpect ratio, L/d, of

~14.5, which caused the applied underpressure to exceed the allowable underpressure at

around half of the penetration depth, and appears to have resulted in plug failure inside

the caisson during the following penetration (see Figure 2.12). The factor of safety (Fs)

during suction installation was lower than unity, and was obviously lower than the

minimal value of 1.5 suggested by Andersen et al. (2005). Therefore, the results may

not prove a very reliable guide to soil flow for more typical aspect ratios of L/d ~ 5 or 6,

and higher factors of safety against plug failure.

Further finite element analysis (Andersen et al., 2004) using PLAXIS v. 8.1 led to

similar results as that of Andersen & Jostad (2002). Their analysis shows that

penetration by self-weight (jacking) gives a significant increase in the mean total stress

outside the caisson, while that by underpressure (suction) only results in a modest

increase, or no increase at all.



Clukey (2005) reported field installation of suction caissons, with a diameter of 6.50 m

and a design embedment of 24 m, in normally consolidated clay in the Gulf of Mexico.

The wall thickness at the caisson tip was 51 mm, of which only 22 was flat, and the

remainder (29 mm) had an external 4:1 taper. The measurements suggest that during

suction installation all the soil displaced by the 22 mm flat portion of the tip moved

inside the caisson. It should be noted that the final penetration depth, i.e., the depth

where soil contacted the top of the caisson, was indicated by ‘mud’ being pumped from

inside of the caisson. Uncertainty may result from such indirect measurements.

At the same time, there also exist some opposite observations. By conducting

centrifuge tests on suction caissons in clay, House (2002) observed very similar pullout

capacity for caissons installed either by jacking or by suction. Measurements of pore

pressure dissipation from instrumented centrifuge model tests (Cao et al., 2002b), or

from extraction of caissons at different periods after installation (Dendani & Colliat,

2002), both suggest times for 50% consolidation of the order of weeks to months rather

than 1 day as suggested by Andersen & Jostad (2002). This indicates that the excess

pore pressure generated outside the caisson during suction installation is larger than if

all the soil particles displaced by the caisson tip were drawn inside the caisson.

Movement of the soil particles at the caisson tip can be indicated by the internal heave

of the soil plug.

Field tests reported by Newlin (2003b) used a simple ‘dipstick’ tool inserted in the

pump interface to measure directly the elevation of soil plug inside the caisson. The

measured values of plug heave were less than a quarter of the value calculated assuming

accommodation of the caisson wall by inward movement alone during the suction phase

of penetration. In that case the caisson tip was chamfered to encourage outward

displacement of the soil, which might have influenced the pattern of soil flow. It is

interesting to find that although conducted in similar soil conditions (both in NC clay in

the Gulf of Mexico) using caissons with similar tips (both with an outward chamfered

part), installation tests by Newlin (2003b) lead to a totally different conclusion

compared to that by Clukey (2005).

According to the above analysis, whether or not all the soil displaced by the caisson

wall goes inside the caisson during suction installation is still unknown, and needs

further research. This problem, however, can be studied by measuring the radial stress

changes on the external wall and pullout capacity after consolidation, for caissons

installed either by jacking or by suction (Chen & Randolph, 2004a, b).



2.2.8 Radial Stress Changes during Installation

The radial stress changes, and resulting excess pore pressures, outside the caisson are

expected to be significantly affected by the proportion of the soil displaced by the

caisson wall that flows outward or inward. Excess pore pressures (∆ui) and thus the

radial total stress (σri) generated on the external wall of the caisson during installation

are thus important in assessing the mode of soil flow at the caisson tip.

For a thin-walled caisson, there are several theoretical methods for predicting the

external ∆ui and σri. These methods include the NGI method (Andersen & Jostad,

2002) shown above, a simplified cavity expansion method (CEM) put forward by

Randolph (2003), the strain path method (SPM) developed by Baligh (1985, 1986), and

the MTD method proposed by Jardine & Chow (1996) for driven piles. It should be

noted that, except for the NGI method which assumes a totally inward flow of soil at the

caisson tip during suction installation (Figure 2.13a), the CEM and the MTD method

are based on the assumption that all the soil moves outside the caisson during

installation (Figure 2.13b), while the SPM assumes a mode between totally inside and

totally outside. Comparison between measurements and these theoretical predictions

could help identify the mode of soil flow at the caisson tip during caisson installation.

In the following section, previous experimental studies are reviewed, followed by the

various predictive approaches.

2.2.8.1 Measurements of radial stresses

There have been various experimental studies aimed at measuring the variation of

excess pore pressures and radial total stresses during the deep penetration of solid piles.

The Piezo-Lateral Stress (PLS) cell was introduced by the Massachusetts Institute of

Technology (MIT) in 1978 on an instrumented model pile with a diameter of 38.4 mm

(Azzouz & Morrison, 1988). The PLS is capable of providing simultaneous

measurements of the total horizontal stress (σh), the excess pore pressure (∆u) and the

shear stress (fs) acting on the cylindrical pile shafts (Whittle, 1992). Total pressure

transducers and Druck (PDCR-81) transducers were applied by Coop & Wroth (1989)

in Oxford for an in situ model pile (IMP) with a diameter of 80 mm. Instrumented solid

piles were developed by Bond & Jardine (1991) at Imperial College to measure the

radial total stress, shear stress and the pore pressure acting on the pile shaft. These tests

were able to measure the total radial stress σri and the excess pore pressure ∆ui



simultaneously, and the derived radial effective stress was analysed with respect to the

shaft friction measured directly. Such tests with instrumented piles pioneered the age of

radial stress measurements on pile installation, and generated important results with

high accuracy. However, all these tests used the solid piles, with transducers having a

thickness equivalent to or larger than 10 to 20 mm. These instruments are obviously

unsuitable for installation on model caissons where the wall thickness is only 0.5 mm

(see Chapter 3).

Cao et al. (2002b) used Druck (PDCR-81) miniature pore pressure transducers (PPTs)

to measure the excess pore pressure generated in the soil during caisson installation. In

their tests, the PPTs were located to be about 20 mm away from the outside wall of the

caisson, at different elevations. However, the Druck PPT has a diameter of 6.4 mm and

a length of 11.4 mm, which is also very large, compared to the wall thickness of the

model caisson; the existence of such a massive object near the model caisson could

affect the flow mode in the soil and cause unpredictable errors to the measurements. In

addition, the variation of the radial total stress cannot be provided by the PPTs, and to

date no direct measurement of radial stresses on open pipe piles or caissons have been

reported.

Previous research shows that the accuracy of measuring normal stress in clay has been

rather low (Dewoolkar et al., 1998; Egan & Merrifield, 1998). Lee et al. (2004) made

improvements on the measurements by using an Entran EPL-D12 pressure cell

supported by a solid plate on the bottom, and reported an accuracy of 70 - 80%.

However, the thickness of the Entran cell was two times the wall thickness of the model

caisson in this research, and was thus obviously unsuitable.

For thin-walled model caissons, such as those used in the centrifuge model tests

presented later, diaphragm type total pressure transducers (TPTs) appear the only

option. There is, however, uncertainty regarding the performance of such TPTs in

centrifuge modelling, particularly as they are penetrated through the soil in a high g

environment, and the caisson is then subjected to cyclic and sustained loading. Careful

calibration exercises will be presented later, to validate the choice of TPTs and their

performance.

It will also be decided that there was insufficient space within the model caisson to

accommodate miniature pore pressure transducers, such as those used by Cao et al.

(2002b). Instead, excess pore pressures will be deduced from the combined



measurements of radial total stress and shaft friction, with radial effective stresses

deduced from the latter.

2.2.8.2 NGI method

In what will be referred to as the NGI method (Andersen & Jostad, 2002), the excess

pore pressures generated around the external shaft of the caisson are limited to those

arising from shearing and remoulding of the soil at constant mean total stress, assuming

that the volume of the caisson wall is accommodated entirely by soil flowing inwards

into the caisson during suction installation. Therefore, there is only shear-induced

excess pore pressure outside the caisson (Figure 2.13a), which is much lower compared

with that of jacked installation or driven open-ended piles.

Following the assumptions in the NGI method for suction installation (Andersen &

Jostad, 2002), the excess pore pressure generated in the remoulded zone during

penetration for caissons installed by suction can be expressed as:

rt

uv0

0i tanδS

sσ32K1∆u

⋅−′+

= (2.10)

where K0 is the in situ earth pressure coefficient, St is the sensitivity of the clay, and σ′v0

is the original vertical effective stress in the soil. The excess pore pressure immediately

after suction was removed, however, is supposed to be calculated as the initial effective

octahedral stress, which is the first part of Equation 2.10. It should be noted that such a

model would result in a larger excess pore pressure acting on the external wall of the

caisson immediately after installation, compared to that during installation.

When calculating the radial effective stress, the same equation as that in the API RP2A

(1993) was adopted by the NGI method, and can be expressed as follows:

r

u

tr

uri tanδ

sS1

tanδαs

σ ==′ (2.11)

where δr is the residual interface friction angle (Chow, 1997) between caisson and clay.

Consequently, the external radial total stress relative to u0 during suction installation of

caissons can be expressed as follows for the NGI method:

0v0

0ri σ3K21uσ ′+

=− (2.12)



2.2.8.3 Cavity expansion method

In the cavity expansion method (CEM), the installation of piles is simulated by the

expansion of a cylindrical cavity in the soil mass by an amount equal to the volume

displaced by the pile. The soil is considered to behave as a rigid-plastic, incompressible

solid surrounding the cavity, and as a linearly deformable solid beyond the plastic

region. Vesic (1972) considered a cylindrical cavity with initial radius of Ri expanded

by a uniformly distributed internal pressure, p (Figure 2.14). If this pressure is

increased, a cylindrical zone around the cavity will pass into a state of plastic

equilibrium. This plastic zone will expand until the pressure reaches an ultimate value,

pu. At this moment the cavity will have a radius, Ru, and the plastic zone around the

cavity will extend to a radius, Rp. Beyond that radius, the rest of the mass remains in a

state of elastic equilibrium. Closed-form solutions for radial total stress and excess pore

pressure around the cavity in terms of the shear strength, su, and shear modulus, G, were

obtained by using the simple Tresca criterion as the constitutive model (Gibson &

Anderson, 1961; Vesic, 1972). This simple version of CEM was used by Randolph &

Wroth (1979) as the basis for assessing the time scale of (radial) consolidation around

driven piles, assuming plane strain condition within a horizontal slice.

The process of installing a displacement pile is more complex than that encountered by

the pressuremeter, due to different strain conditions in the vicinity of both ends of the

pile along the major length (Figure 2.15). For the purpose of simplification, as the

CEM was applied to simulate pile penetration, the local details such as soil heave and

precise details around the pile tip were neglected (Randolph & Wroth, 1979). In fact,

these areas are quite limited and have little effect over the major length of the pile

(Steenfelt et al., 1981), and such a simulation was deemed valid. A slightly more

sophisticated version of this approach was adopted by Randolph et al. (1979), taking

account of the change in the mean effective stress around piles as the soil is sheared and

remoulded.

A generalised form of the CEM for open-ended piles was discussed recently by

Randolph (2003), with the post-installation radial stresses expressed as

( ) pppuu 0iriiri0ri ∆+′+′−σ′=∆+σ′=−σ (2.13)

where p'0 and p'i are respectively the original in situ mean effective stress and that just

after caisson installation, and ∆p is the increase in mean total stress due to caisson

installation. The difference between p'0 and p'i represents shear-induced excess pore



pressures, while the increase in mean total stress may be estimated from the mean total

stress according to the traditional CEM (Randolph & Wroth, 1979). Some uncertainty

exists in the magnitude of the bracketed term in Equation 2.13, although its contribution

will be small (limited to less than the remoulded shear strength).

An alternative is to estimate the excess pore pressure, ∆ui, directly from the two

components of decrease in mean effective stress and increase in mean total stress, using

the solution for cylindrical cavity expansion (Gibson & Anderson, 1960) for the latter,

to give

⎟⎟⎠

⎞⎜⎜⎝

⎛ρ+σ′+

⎟⎟⎠

⎞⎜⎜⎝

⎛−≈∆

uuvo

0

ti s

Glns3K21

S11u (2.14)

in which ρ is the area ratio of the caisson (~4t/d, shown in Figure 2.16), and G/su is the

rigidity index, Ir. The first component of excess pore pressure is identical to the ∆ui

from the NGI method (see Equation 2.10).

In the CEM, the excess pore pressure generated outside the caisson during installation is

composed of two parts: shear induced and expansion induced excess pore pressures

(see Figure 2.13b). The external excess pore pressure predicted by CEM is thus larger

than that predicted by the NGI method for caissons during suction installation.

The radial effective stress, σ ri, however, can be estimated by the identical expression

adopted by API RP2A and also the NGI method, as shown in Equation 2.11.

According to Equations 2.13, 2.14 and 2.11, the radial total stress relative to u0, can

therefore be expressed more directly by the following equation:

⎟⎟⎠

⎞⎜⎜⎝

⎛+′+

⎟⎟⎠

⎞⎜⎜⎝

⎛−+=−

uuvo

0

tr

u

t0ri s

Gρlnsσ3

K21S11

tanδs

S1uσ (2.15)

2.2.8.4 Strain path method

Simulation of pile installation, initially for full-displacement piles, was provided by the

strain path method (SPM) developed by Baligh (1985, 1986). In the SPM, the

undrained deep penetration of a single pile in clay can be modelled by a stationary point

source within a ‘flowing’ soil mass. The soil is modelled as an ideal, incompressible,

inviscid, irrotational fluid, which flows around the point source. The strain field around

the advancing pile is obtained by ignoring the shear strength of the soil and any friction

between the object and the soil (Clayton et al., 1998). By applying potential flow



theory, typical strain paths for soil elements at different radial distances from the

centreline of a simple pile were presented by Baligh (1985) (see Figure 2.17). Stress

changes are then deduced by integrating the equations of equilibrium, taking account of

the stress-strain response of the soil.

The SPM was extended by Chin (1986) and Baligh et al. (1987) to investigate the strain

induced due to the penetration of a simple sampling tube (or open-ended pile) in

saturated clay, by simulating the tip of the open-ended pile by an annular ring source.

The deformations, strain and octahedral shear strain (γoct) contours around such an

open-ended sampler with d/t = 40 were determined and are shown in Figure 2.18.

Further development was achieved by implementing the MIT-E3 model into the above

approach, and the stress field around the open-ended pile (d/t = 40) was presented by

Whittle (1992). In the SPM, it can be seen that the soil displaced by the pile volume is

pushed both inside and outside the pile, rather than completely into the soil plug.

The SPM provides an effective approach for assessing the stress field around piles

quantitatively during installation, under the assumption of zero shaft friction at the

pile-soil interface, which is needed to make the calculation feasible. However, the SPM

is somewhat complex for general applications, since stress changes cannot easily be

expressed in terms of underlying properties, and most published results are restricted to

particular soil models and parameters. Predictions here are based on the study of an

open-ended pile reported by Whittle & Baligh (1988), using the MIT-E3 model. The

normally consolidated Empire clay on which their calculations were based is considered

to be the closest in soil properties to NC kaolin clay, among those cases published using

the SPM. In fact, the range of results for different soil types is small for the SPM, since

the result is dominated by the pile geometry (Clayton et al., 1998). The prediction by

SPM (Whittle & Baligh, 1988) gives ∆ui = 1.05σ'v0, and σ ri = 0.23σ'v0 for an open-

ended pile with d/t = 40. These results can be adjusted for the caisson with d/t = 60,

which is close to the pile with d/t = 40.

2.2.8.5 MTD method

Extensive field research was carried out by researchers at Imperial College, jacking a

closed-ended pile with a diameter of 100 mm into different natural clays (Figure 2.19).

The pile was instrumented to measure radial total stress, pore pressure and shear stress

along the pile shaft (Lehane et al., 1994). A pile design method, generally referred to as



the MTD method, was subsequently proposed by Jardine & Chow (1996), calibrated

through an extensive database of pile load tests. Since the MTD prediction of the stress

field around open-ended piles was extrapolated from the measurements in Lehane’s

work using the small diameter closed-ended pile, his work will be discussed below in

this section.

The MTD method includes an ‘h/d’ effect, first proposed by Bond (1989), where h is

the distance between the point of interest and the tip of the pile, and d is the diameter of

the pile. The ‘h/d’ effect on the radial total stress during pile penetration in various clay

sites is illustrated in Figure 2.20 (in which R = d/2), taken from Lehane & Jardine

(1994). Another point considered by the above approach is the effect of yield stress

ratio (YSR), or overconsolidation ratio (OCR), on the radial total stress around the pile.

The variation of the normalised radial total stress during pile installation, Hi (defined as

(σri – u0)/σ v0), with YSR, is shown in Figure 2.21. The MTD method made some

improvement on the pile studies by considering stress history and sensitivity of the soil.

Although their tests focused on closed-ended piles, Jardine & Chow (1996) stated that

the excess pore pressure generated during the penetration of open-ended piles can be

estimated by using the same expression as for the solid piles, but replacing the diameter,

d, by an equivalent solid pile of diameter, deq, with the same volume of steel (thus deq =

d√ρ where ρ is the area ratio).

The expression proposed by Lehane (1992) for radial total stress σr relative to

hydrostatic pressure u0 can be written as:

( ) 0v0.2

eq10.4

0ri σdhYSR43.uσ ′=− − (2.16)

while the excess pore pressure ∆ui generated adjacent to the pile during installation is

( ) v00.2

eq0.5

i σdhRYS7.2∆u ′= − (2.17)

where YSR is the overconsolidation (or yield stress) ratio and σ'v0 is the in situ vertical

effective stress.

The radial effective stress can be derived from the above two equations and expressed

as:

( ) ( ) v00.2

eq0.5

v00.2

eq0.42

ri σdh2.7YSRσdh3.4YSRσ ′−′=′ −− (2.18)

Development of the radial stresses acting on the external wall of the caisson after

installation will be discussed below.



2.3 RELAXATION DURING CONSOLIDATION

As consolidation proceeds, some relaxation of radial total stresses is to be expected

around a suction caisson, just as for displacement piles in clay as described by Lehane

& Jardine (1994).

The magnitude of stress relaxation during consolidation is important in determining the

final radial effective stress at the caisson wall, and hence the long-term shaft friction.

For displacement piles, Lehane & Jardine (1994) quantified the initial total radial stress

(after installation) in terms of the coefficient Hi = (σri – u0)/σ'v0 and the final radial

effective stress after consolidation as Kc = σ'rc/σ'v0. For low overconsolidation ratios,

typical values of these ratios for full-displacement piles, were ~2 (see Figure 2.21) and

0.8 to 1 respectively (see Figure 2.22), implying a stress relaxation of over 50% of the

‘potential’ radial effective stress at the pile wall.

The expression for the effective stress ratio after consolidation proposed by Lehane et al.

(1994) and Jardine & Chow (1996) is

( )[ ] ( ) 0.20eq

0.42t10c h/dYSR S0.870log0.016YSR2.287.0K −−+= (2.19)

The low embedment ratio, L/d, and thin-walled nature of suction caissons take them

well outside the database assembled by Lehane (1992) and Chow (1997) for

displacement piles in clay, which forms the basis of the MTD pile design approach

(Jardine & Chow, 1996). However, it is of interest to compare the radial effective stress

measured after consolidation with that predicted using the MTD approach.

For full-displacement piles, Randolph (2003) postulated an expression for the final

radial effective stress ratio of

⎟⎟⎠

⎞⎜⎜⎝

⎛σ′∆

⋅µ⋅λ

+µ

+σ′σ′

=σ′σ′

=0v

i

0v

ri

0v

rcc

uR

1nRK l (2.20)

where R is the overconsolidation (or yield stress) ratio, and λ and µ were two

parameters taken as 1 and 5 respectively in order to fit measured data from field pile

tests assembled by Lehane (1992) and Chow (1997). Values of σ′ri and ∆ui can be

obtained from Equations 2.11 and 2.14 in the CEM.

The above expression was viewed as somewhat speculative by Randolph (2003), but it

is interesting to assess later how it performs for the thin-walled suction caissons.



In the NGI method, Andersen & Jostad (2002) do not give an explicit expression for the

radial effective stress after consolidation, but instead refer to the post-consolidation

mean effective stress, relative to the in situ value. That ratio is estimated to lie in the

range 0.53 to 0.61 for the various clays considered, with a value of 0.55 for kaolin. The

post-consolidation radial effective stress ratio, Kc, will depend on the relative

magnitudes of the radial, vertical and circumferential stresses around the caisson.

Times needed for 50% and 90% consolidation after the caisson is installed by suction

were predicted by the NGI method (Andersen & Jostad, 2002), and are shown in Table

2.4. It can be seen in the table that the NGI method predicts t50 as less than 2 days, and

t90 as less than 40 days for the NC or LOC clays in the Gulf of Mexico. For NC kaolin

clay, t50 is less than 1 day, while t90 is 6 days. They also stated that the dissipation time

for excess pore pressures is significantly higher for caissons installed by jacking

(self-weight), compared to those installed by suction.

Comparison of the final radial effective stress after consolidation measured for caissons

installed by jacking and by suction can reveal the difference of these two types of

installation, and the effect on axial behaviour of the caisson. Further comparison

between measurements and theoretical predictions can identify the mode of soil flow

during suction installation.

Table 2.4 Predicted 50% and 90% consolidation times after caissons installed by

suction in clay (after Andersen & Jostad, 2002)

Clay type Ip t50 t90

(%) (days) (days)

Offshore Africa 80 2.1 55

Gulf of Mexico 55-60 1.5 37

Gulf of Mexico 35-40 1 21

Drammen 25-30 0.5 15

Kaolin clay 30 0.2 6



2.4 VERTICAL PULLOUT CAPACITY

2.4.1 Failure Modes

For caissons serving as taut-wire mooring anchors for deepwater platforms, the chain

loading angle relative to the horizon is generally larger than 40 . The axial capacity

thus governs design for the holding load (Mello et al., 1998; Huang, et al., 2003). The

axial capacity of suction caissons in clay has been studied by Renzi et al. (1991),

Morrison et al. (1994), EI-Gharbawy & Olson (1999), Deng & Carter (2000),

House & Randolph (2001), Zdravkovic et al. (2001) and Andersen & Jostad (2002).

The uplift capacity depends heavily on assumptions regarding the degree of ‘passive

suction’ that can be relied upon under different types (and time-scales) of loading, and

this introduces considerable complexity to the problem. There are essentially three

different modes of failure to be considered (Fuglsang & Steensen-Bach, 1991;

Randolph & House, 2002), as shown in Figure 2.23. All three modes involve shearing

between the caisson shaft and the external soil, with limiting shear stress most

conveniently expressed in terms of an interface friction ratio, α, times the average (or

simple shear) shear strength.

The other component of resistance for suction caissons is determined by the failure

mode (Randolph & House, 2002), and includes (in addition to the submerged weight of

the caisson):

1. shearing resistance between the soil plug and the internal caisson surface, plus

the reverse end-bearing resistance of the annular caisson tip;

2. weight of the internal soil plug plus any (long term) tensile capacity available at

the base of the soil plug;

3. reverse end-bearing resistance of the full caisson area.

For convenience, these three failure modes will be referred to as ‘unsealed’, ‘sealed

(base-vented)’ and ‘sealed’ respectively. The terms ‘unsealed’ and ‘sealed’ refer to the

condition assumed for the caisson lid, while ‘base-vented’ refers to where a hydraulic

short-circuit precludes development of suction at the caisson base. The reverse end-

bearing capacity factor, Nc, is affected most extensively by the mode of failure, and will

be discussed in detail below.



2.4.2 End-bearing Capacity

2.4.2.1 Unsealed pullout

For caissons pulled out with an unsealed lid (see Figure 2.23a), failure will occur by

sliding along the internal wall of the caisson or, where internal ring stiffeners are used,

along a cylinder defined by the internal diameter of the stiffeners.

The reverse end-bearing resistance of the annular caisson tip, qu, can be estimated using

a standard bearing capacity approach as

v0ucu σsNq ′−⋅= (2.22)

but with Nc ~ 7 to 8 (Randolph & House, 2002), reflecting the near plane strain

geometry of the annular tip.

2.4.2.2 Sealed pullout

For the sealed caisson during pullout (see Figure 2.23c), the base resistance may be

estimated using Equation 2.22, but with the more usual Nc value appropriate for a deep

circular foundation (such as a pile). Several researchers have reported results from

physical model tests, either on the laboratory floor (at 1 g) or conducted under enhanced

gravity in a geotechnical centrifuge.

Fuglsang and Steensen Bach (1991) reported centrifuge and laboratory model tests on

the sealed and unsealed uplift capacity of suction caissons in overconsolidated kaolin

clay. Their centrifuge data suggested an Nc value between 6.5 and 8.5.

Clukey & Morrison (1993) performed centrifuge tests on suction caissons with L/d ~ 2

in normally consolidated soil, and obtained an Nc value of around 11. Based on results

from 1 g model tests in kaolin clay, El-Gharbawy & Olson (1999) recommend a reverse

end-bearing capacity factor of 9, irrespective of embedment depth. Test results of

Watson et al. (2000) suggest that the bearing resistance in tension is similar in

magnitude to that in compression. The value of Nc is customarily taken as 9, an

appropriately conservative value given the strain-softening nature of the response as the

caisson is extracted. Luke (2002) also suggested that the tension bearing resistance is

similar to the compression bearing resistance in terms of the magnitude (thus Nc ~ 9),

although the caisson they tested has a rather small aspect ratio (L/d ~ 1). Randolph &

House (2002) derived Nc values on the order of 14 - 15, based on centrifuge tests of

model caissons in normally consolidated kaolin clay. This value is extremely high, and



is similar to end-bearing factors achievable in deep bearing for a cone penetrometer.

Interestingly, this high bearing factor was achieved with only small displacements. The

small distance (less than 1 diameter of the caisson) between the caisson tip and the

bottom of the sample in their tests is considered to account for, at least partially, the

high Nc values. It should also be noted that if verticality is not well maintained during

the pullout, lateral resistance could exaggerate the capacity.

At present, there is no agreement on which Nc value should be used in designing the

uplift capacity of caissons. Further experimental research is needed in order to resolve

such uncertainty.

2.4.2.3 Sealed (base-vented) pullout

In addition to the static load (e.g. hull buoyancy) exerted by the mooring line, suction

anchors can also be subjected to environmental loads with low frequency (e.g. mean

wind and loop current) and high frequency (e.g. hurricane and storm wave load) (Huang

et al., 2003). Loop current loading may last days to weeks, and therefore may be

considered as sustained loading (Clukey et al., 2004). Hurricane loads are usually

applied quickly (seconds to minutes), and are thus cyclic loading. Behaviour of suction

caissons under long-term sustained loading and cyclic loading requires investigation

(House, 2002; Clukey et al., 2004).

For sealed caissons under long-term or cyclic axial loading, the intermediate mode of

failure (see Figure 2.23b) is suggested as appropriate, since it is difficult to guarantee a

good hydraulic seal at the bottom of caissons subjected to sustained or cyclic loading

(Randolph & House, 2002). With the dissipation of pore pressures, the reverse

end-bearing (i.e. passive suction) could be significantly reduced, as revealed by

centrifuge tests reported by Clukey & Phillips (2002) and Randolph & House (2002).

Clukey & Phillips (2002) showed that the axial capacity of caissons reduced by less

than 13% for a sustained loading lasted around 2 months, compared to that during

monotonic pullout; the Nc value was obtained as 9.4. It should be noted that the applied

load in Clukey & Phillips (2002) was inclined 40° to the horizontal direction, making

the end-bearing different from that for monotonic uplift. Randolph & House (2002)

reported a reduction of 20% in the vertical pullout capacity during sustained loading,

compared to that during monotonic tensile loading, and obtained Nc as ~ 9 in sustained

loading. However, this value was derived by using a shaft friction ratio deduced from

the axial capacity of an unsealed pullout test in a disturbed site, with uncertainty



existing in such an extrapolation. Measurements of the radial stress changes of the

caisson during sustained loading may be helpful in providing a reliable shaft friction

ratio, from which the Nc value may be derived.

Cyclic environmental loading has two potentially compensating effects: cyclic

degradation of soil strength and soil resistance increase due to loading rate effects (Bea

& Audibert, 1979). The former appears to prevail over the latter since centrifuge tests

undertaken by Clukey et al. (1995) show that, after 100 cycles, the capacity under

cyclic loading reduced to 48% of that during static loading, although the loading was

inclined and may have caused a gap to develop between the caisson and the soil.

Recent centrifuge tests reported by Randolph & House (2002) showed that the capacity

under cyclic loading was reduced to around 80% of the monotonic pullout capacity.

Appropriate Nc values and α values need to be determined for caissons under cyclic

loading. Measurement of radial stress changes around the caissons could provide an

explanation of the changes in axial capacity during cyclic loading.

2.4.3 Shaft Friction during Vertical Pullout

In previous research, the shaft friction ratio, α, for caissons during vertical pullout was

determined either from the measured uplift capacity of unsealed caissons (House,

2002), or from theoretical prediction, such as the API design rule, NGI method, CEM

and MTD method. Previous work is reviewed in the following section.

2.4.3.1 Measurements

There has been limited research reporting the shaft friction ratio for suction caissons

from physical modelling tests. Centrifuge tests undertaken by Clukey & Morrison

(1993) show an α value of 0.8 for caissons pulled out vertically in normally

consolidated clay, by adopting Nc as 9.4. Based on unsealed pullout tests of model

caissons in the centrifuge, House (2002) obtained α as 0.45 - 0.5 in normally

consolidated clay, and 0.45 - 0.9 for overconsolidated clay. In his analysis, equal α

values were assumed inside and outside the caisson, and the deduced average α value

was then applied to sealed caissons leading to Nc = 15. The high Nc value suggests that

the α value for sealed caissons may have been higher.



2.4.3.2 Current design method

Prediction of the shaft friction of suction caissons during pullout in soft marine clay has

tended to be based on conventional design methods used for open-ended driven piles

(API RP2A, 1993), expressed as

usα ⋅=f (2.23)

in which the α value can be calculated by:

0.5v0u )σs0.5(α −′= for 1.0σs v0u ≤′ (2.24)

0.25v0u )σs0.5(α −′= for 1.0σs v0u >′ (2.25)

For normally consolidated clay with moderate sensitivities (2 - 3), the predicted α value

during caisson pullout is generally unity. However, suction caissons are different from

driven piles in two aspects: the large ratio of caisson diameter to wall thickness (d/t

=100 - 200) and the different installation method. These differences may cause the

results from driven piles (d/t = 30 - 60) to not be applicable to large diameter,

thin-walled suction caissons (Huang et al., 2003). The differences may lead to lower

external excess pore pressures generated during installation, and lower external radial

effective stresses after consolidation, resulting in lower α values for suction caissons

compared to driven piles. Therefore, direct extrapolation of the design rule for driven

piles to suction caissons may result in over-prediction of the shaft capacity (or α value).

2.4.3.3 NGI method

For the NGI method, it was suggested by Andersen & Jostad (2002) that the different

installation methods of jacking (including driving) or the use of suction will lead to

different axial capacities. In their work, small strain finite element analysis suggested

that during jacking or self-weight penetration the clay displaced by the caisson wall at

the skirt tip would flow 50% inside and 50% outside the caisson; by contrast during

suction installation, all the clay displaced by the wall was found to move inside the

caisson, with no tendency for outward movement of the clay below the skirt tip. As a

result, the interface friction along the external shaft of the caisson would be different for

self-weight penetration (taken as similar in mechanism to jacked or driven installation)

and suction installation. Any clay pushed outwards will lead to increased external

pressure and thus higher excess pore pressures during installation. After consolidation,

the effective stress level and local shear strength of the clay should be higher than if no



outward movement occurs, as hypothesised for suction installation. Thus the long-term

interface friction mobilised during pullout should be lower for caissons installed by

suction than for when they are jacked. Andersen & Jostad (2002) gave predictions on

the α values for caissons installed by suction in NC clay, as shown in Table 2.5.

It can be seen in the table that for suction caissons in NC clay with moderate sensitivity

(St < 3.0) and plasticity index (IP = 25 - 50%), the NGI method gave an α value of 0.65,

which is obviously lower than that (unity) for jacked installation. The time to reach

90% pore pressure dissipation for most situations is taken as less than 2 months.

Table 2.5 α values along outer skirt wall 2 months after installation by

underpressure in normally consolidated clays

IP <25% 25 - 50% >50%

St >3.0 0.58 0.65 0.65

St<3.0 0.58 0.65 1.95/St

2.4.3.4 MTD and CEM method

The radial effective stress after consolidation for MTD and CEM can be obtained by

Equations 2.19 and 2.20, respectively. According to Chow (1997), the radial effective

stress when the pile is loaded to failure can be obtained by applying a reduction factor,

K, to the measured σ′rc, in order to consider the influence of change of loading direction.

The external shaft friction ratio, α, during pullout of the caisson after consolidation can

thus be estimated by:

u

rrcs

tanδσKα

⋅′⋅= (2.26)

K can be taken as 0.8, according to field tests on piles by Chow (1997), although detailed identification of this value for thin-walled caissons is necessary.

It should be emphasised that the CEM and MTD method will both predict lower shaft capacity for suction caissons than for solid piles, or open-ended piles. The high d/t ratio for suction caissons leads to lower excess pore pressures during installation, and hence lower post-consolidation radial effective stresses, even assuming that all the soil displaced by the caisson moves outside the caisson.



2.4.3.5 Discussion

It can be seen in the above analysis that the NGI prediction of shaft friction is different

from the result of the API design method, while the effectiveness of the CEM and MTD

for analysing thin-walled suction caissons is unknown.

In fact, as for analysis on piles, the external α value of caissons during vertical pullout

can be derived from the measured radial effective stress, σ′rf, as the caisson is loaded to

failure, by the basic expression:

u

rfs

tanδσα

⋅′= (2.27)

where δ is the interface friction angle between the external wall and soil.

The α value which is derived from the measured radial effective stress can thus be

compared to theoretical predictions, in order to reveal a reasonable design scenario.

2.5 CONCLUSIONS

Research on suction caissons in soft clay has been carried out by field tests, 1 g

laboratory tests, centrifuge tests and numerical modelling. Field practice shows that

suction caissons can be installed successfully in deep and ultradeep waters. Review of

previous work shows that the following problems need to be resolved:

1. The pattern of internal soil flow after passing the first ring stiffener,

during caisson penetration.

2. The mode of soil flow at the caisson tip during suction installation, and

the difference with that during jacked installation.

3. Effectiveness of existing theoretical methods for predicting the excess

pore pressure generated outside the caisson during suction installation.

4. Variations of external radial stresses around the caisson during

consolidation, and the reliability of theoretical predictions of the final

radial effective stress acting on the external wall of the caisson after

consolidation.

5. Time scale for consolidation in the soil after caisson installation.

6. Values of α and Nc during vertical uplift of sealed caissons in soft clay

after consolidation.



7. Modified values of α and Nc during sustained loading and cyclic loading,

and the corresponding variations in radial stresses.

In this thesis, results of centrifuge tests on suction caissons are reported, according to

the test program described in Chapter 1, using model caissons that are instrumented

with miniature total pressure transducers, in order to find answers to the problems

identified above. Details of the instrumented model caissons will be described in

Chapter 3.

Chapter 3 3-1 Experimental Apparatus and Soil Properties


3 EXPERIMENTAL APPARATUS AND SOIL PROPERTIES

3.1 INTRODUCTION

This chapter provides a general overview of the instrumented model caisson, the total

pressure transducer (TPT) used on the model caisson, the calibration system for the total

pressure transducer, the ring shear apparatus, the UWA centrifuge facilities and the

corresponding scaling laws used in the centrifuge modelling. Details of the apparatus

used in the test program are described, including equipment developed specifically for

individual tests. In addition, the engineering properties of the kaolin clay used in this

research are presented, as well as the T-bar device employed for in situ measurement of

the soil strength.

3.2 MINIATURE TOTAL PRESSURE TRANSDUCER

Radial stress changes around the suction caissons during installation and pullout in the

clay were measured by miniature total pressure transducers (TPTs). All TPTs used in

this research are Kyowa PS type (manufactured by the Kyowa Electronic Instruments

Co., Ltd), having a diameter of 6 mm and a thickness of 0.6 mm. The capacity is 1000

kPa (corresponding to PS-10 KA type) (Figure 3.1) for the cells on the shaft of the

caisson, and 500 kPa (corresponding to PS-5 KA type) for that recording the external

hydrostatic pressure on the lid of the caisson. The Kyowa PS type pressure transducer

is a diaphragm type pressure cell, which has a foil strain gauge and Wheatstone bridge

in a small thin membrane. Some working parameters of the Kyowa PS-10 kA

transducer are as follows:

• capacity: 1000 kPa;

• rated output: 0.892 mV/V;

• safe excitation: 3 V;

• bridge resistance: 120 Ω.

3.3 INSTRUMENTED MODEL CAISSONS

Model suction caissons were designed according to the geometry of the caisson used in

the field. Two model caissons, namely caisson 1 and caisson 2, were fabricated for the



tests in the centrifuge. The overall geometries are very similar between these two

caissons; both have a length (L) of 120 mm, an external diameter (d) of 30 mm and a

wall thickness (t) of 0.5 mm, representing a length of 14.4 m, a diameter of 3.6 m and a

wall thickness of 0.06 m at prototype scale for centrifuge tests undertaken at 120 g. A

pad-eye is located at 0.4L (48.5 mm, model scale) from the caisson tip.

For caisson 1 (see Figure 3.2), there are two stages of internal stiffener, with the wall

thickness increasing to 1 mm between 41 mm and 56.5 mm from the tip (opposite the

pad-eye), and then to 1.5 mm for the next 7 mm (housing total pressure transducers)

before reverting to the 0.5 mm wall. The variation of wall thickness (t) with the

distance from the tip (h) of caisson 1 is shown in Table 3.1. Above the stiffener, two

holes were drilled diametrically opposite each other, with a diameter of 6.5 mm and a

depth of 0.7 mm. These can accommodate the TPTs allowing for the thickness of the

adhesive on which they are seated. Then, the two transducers were glued on the shaft of

the caisson with the measuring surface facing outside. The distance from the TPTs to

the caisson tip is 60 mm at model scale (representing 7.2 m at 120 g). Both caissons

were made of 6061 T6 aluminium, the surface of the caissons was anodised after

sand-blasting to resist corrosion, leaving a slightly roughened surface with a CLA

roughness of 2.5 µm. Details of the design of model caisson 1 are shown in Figures

3.2 - 3.3, while photographs of the caisson after instrumentation are shown in Figures

3.4. The leads were covered with epoxy and exited the top of the caisson inside a

groove machined along the internal shaft (see Figure 3.5). The surface of the transducer

was made flush with the caisson shaft (see Figure 3.6).

Table 3.1 Wall thickness (t) in terms of the distance from the tip (h) of the caisson

(in model scale)

Caisson 1 Caisson 2

h

(mm)

t

(mm)

h

(mm)

t

(mm)

0.0 - 41.0 0.5 0.0 - 35.0 0.5

41.0 - 56.5 1.0 35.0 - 45.0 1.5

56.5 - 63.5 1.5 45.0 - 56.5 1.0

63.5 - 120 0.5 56.5 - 120 0.5



Details of caisson 2 are shown in Figure 3.7. Caisson 2 has the same length, diameter,

wall thickness and degree of sand-blasting as caisson 1 (i.e. L × d × t = 120 × 30 × 0.5

mm), except that the total pressure transducers are situated lower in the caisson, and the

distance between the TPTs and the caisson tip is 40 mm (or 4.8 m prototype) here. The

1.5 mm thick internal stiffener started at 35 mm away from the tip, and the wall

thickness then decreased in stages to 1 mm at 45 mm from the tip, and reverting to 0.5

mm at 56.5 mm from the tip (see Table 3.1). The pad-eye has the same geometry as

that of model caisson 1, although its orientation was perpendicular to the two TPTs, to

avoid possible influence on the pressure measurements during loading of the caisson.

Other instrumentation included a load cell to measure axial force applied to the caisson,

a miniature pore pressure transducer (PPT) in the caisson lid to measure the internal

water pressure, and a TPT on top of the caisson lid to monitor the external water

pressure. A pneumatic valve was also built into the caisson lid to allow venting of the

caisson during self-weight or jacked installation, and sealing of the caisson during

suction installation or (sealed) pullout. Details of the connection between the caisson

and the load cell are shown in Figure 3.8, while the arrangement of the instrumented

model caisson on the centrifuge is shown in Figure 3.9.

3.4 CALIBRATION CHAMBER FOR PRESSURE CELLS

Calibration tests were used to investigate the accuracy of the pressure cells. Limited

calibration tests on the total pressure transducer were carried out in a triaxial system.

The basic triaxial apparatus (see Figure 3.10) consists of a 50 kN compression machine

with a stainless steel cell, which can operate up to a maximum pressure of 3 MPa. It

was adapted specifically for calibration tests under both undrained and drained

conditions. Details of the wiring connections between the caisson and the top cap are

shown in Figure 3.10. This calibration chamber has the advantage of controlling the

applied pressure to a very accurate degree through the Geotechnical Digital Systems

(GDS) controller. The GDS controller (see Figure 3.11) is a microprocessor controlled

hydraulic actuator for the precise regulation and measurement of liquid pressure and

liquid volume change. It was programmed to ramp and cycle the pressure and volume

change linearly with respect to time. The GDS controller has a pressure and volume

capacity of 2 MPa and 2000 cm3, respectively. Feedback control on the calibration tests

was achieved via a computer equipped with digital-to-analogue and analogue-to-digital



cards, and data was collected via a 12-bit data acquisition card. The calibration system

for TPTs in water, including the pressure generation system and the data acquisition

system, is expressed in Figure 3.12. Test results are presented in Chapter 4.

3.5 RING SHEAR APPARATUS

The interface friction angle between the clay and the aluminum alloy used for the model

caissons was measured in a Bromhead-type ring shear apparatus (WF25850) (see

Figure 3.13), which was purchased from Wykeham Farrance International Ltd. This

ring shear apparatus shears at a rate between 0.018 and 45 mm/min. The top platen

simulates the external caisson wall, and therefore was manufactured from the same

material (namely 6061 T6 aluminium). The top platen was subsequently sandblasted to

a corresponding centre line average (CLA) roughness of 2.5 µm, before it was anodised

to resist corrosion (see Figure 3.14). The soil sample was taken from the strong-box

immediately after the caisson tests finished, and was then filled into the concentric ring,

to a depth of 5 mm thick, and an inner and outer diameter of 70 and 100 mm,

respectively (see Figure 3.14). A view showing how the vertical load is applied through

the top platen and how the shear is transferred by the torque arms is shown in Figure

3.15. Vertical load was applied by a porous bronze loading platen through a

counter-balanced 10:1 ratio lever. The settlement of the sample during consolidation

and the shearing was monitored with the dial gauge bearing on top of the load hanger.

Torque force transmitted through the sample was measured by the two proving rings.

Results of the ring shear tests are presented in Chapter 5.

3.6 CENTRIFUGE MODELLING: SCALING LAWS

Centrifuge modelling is used widely in the investigation of geotechnical problems

(Schofield, 1980). Examples of the application of centrifuge tests for offshore

geotechnical problems were presented by Murff (1996).

The self-weight stresses in the model are enhanced by the centrifugal acceleration, in

order to give stresses (and shear strengths) that are homologous in model and prototype.

Essentially, all stresses and strains model as 1:1 between model and prototype. In

centrifuge modelling, all linear dimensions of the model are scaled down N times, and a

centrifugal acceleration of N times earth’s gravity (g) is applied during the test, where N

is called the scaling ratio. Therefore, the vertical stress, σm, at model depth, hm, can be



obtained as

mm hgNρσ ⋅⋅⋅= (3.1)

where ρ is the density of the model object, and the corresponding prototype vertical

stress is

pp hgρσ ⋅⋅= (3.2)

where hp is the prototype depth. Since hp/hm = N, the same vertical stress is achieved

between the model and the prototype object. The basic principles of centrifuge

modelling for geotechnical problems have been elaborated by Schofield (1980) and

Taylor (1995). The scaling laws relevant to this research are detailed in Table 3.2.

Table 3.2 Centrifuge scaling laws

Parameter Scaling relationship

(model/prototype)

Gravity N

Length 1/N

Density 1

Mass 1/N3

Force 1/N2

Area 1/N2

Stress 1

Strain 1

Time (consolidation) 1/N2

The inertial acceleration field at a radius, r, generated by the angular rotation, ω , of the

centrifuge results in a normal component of acceleration given by

rωgN 2 ⋅=⋅ (3.3)

During the centrifuge tests, a nominal radius, rnom, is adopted, and the angular velocity

of the centrifuge will achieve the desired gravity level at that point. The optimal value

of this nominal radius was recommended by Schofield (1980) as



s0nom h32rr −=

(3.4)

where r0 is the distance between the axis of the centrifuge and the base of the sample,

while hs is the height of the sample.

It should be noted that because of the variation of acceleration with radius, there is a

discrepancy between the stress experienced by the model in the centrifuge and the ideal

(equivalent) prototype. The actual vertical stress produced by the centrifuge at a certain

depth, z, from the surface of the object can be calculated as:

rdrρωσ zr

r 2

zmin

min∫

+= ( ) ⎥⎦

⎤⎢⎣

⎡+⋅⋅=

min

2

nom

minnom

2

2rzz

rr

rρω

( ) ⎥⎦

⎤⎢⎣

⎡+⋅⋅⋅⋅=

min

2

nom

min

2rzz

rr

gρN (3.5)

where rmin is the distance between the centrifuge axis and the surface of the model

object (Figure 3.16). The discrepancy between Equation 3.5 and Equation 3.1 should

be accounted for when calculating the vertical stress of the soil sample during

centrifuge tests. In this research, all vertical pressures including the hydrostatic

pressure and the vertical effective stress have been calculated using Equation 3.5.

3.7 FIXED BEAM CENTRIFUGE FACILITIES

All tests in this research were performed on the fixed beam centrifuge at the University

of Western Australia. With a radius of 1.8 m, the Acutronic Model 661 geotechnical

centrifuge is rated at 40 g-tonnes, which enables a package weight of 200 kg to be

accelerated to a maximum 200 g. Under the acceleration of 120 g used in this research,

the package can reach a maximum weight of 333 kg. The fixed beam geotechnical

centrifuge at UWA is depicted in Figure 3.17. A full description of the equipment and

associated facilities can be found in Randolph et al. (1991).

All instruments were monitored and recorded by the remote computer in the centrifuge

control room. During this research program, the control panel of the Data Acquisition

(DAQ) system was upgraded from a QBASIC to a LABVIEW system. In addition, the

A/O card was upgraded from 12 bit in the old system to 16 bit in the new system. This

means the resolution was increased approximately 15 times (0.000152 mV/bit instead of

0.0024 mV/bit).



3.7.1 Strong-box

Clay samples were consolidated in aluminium strong-boxes (Figure 3.18) where all

caisson tests in this research were performed. With internal dimensions of 390 mm by

650 mm by 325 mm high, the sample can represent a clay bed of up to 47 m by 78 m by

39 m at 120 g, although certain distances from the edges should be allowed to avoid

boundary effects.

3.7.2 Actuators

All caissons and T-bar tests were performed ‘in flight’ by the electronic actuator (see

Figure 3.18), either in displacement or load control mode. It should be noted that the

actuator can apply both monotonic loading and cyclic loading.

With the old DAQ system used for the first half of this research, the actuator was

controlled by a 30 Volt DC variable speed servo-motor, which was capable of

delivering displacement rates from 0 to 3.156 mm/s with a maximum axial thrust of

6.5 kN. Two Penny and Giles rectilinear potentiometers monitor displacements to a

maximum vertical stroke of 250 mm, and a horizontal stroke of 180 mm. The actuators

were upgraded half-way through by using high resolution HEDS-5640 Optical Encoders

(Figure 3.19). The resolution was thus substantially improved up to 1024 counts per

revolution. This proved to be especially useful when a stress-hold was required during

the consolidation stage immediately after caisson installation.

3.7.3 Slip Rings

Both dual hydraulic/pneumatic slip rings and single phase 250 Volt 10 Amp mains

power slip rings were used in this research. The dual slip ring (Figure 3.20) can pass

any combination of air or water through to the centrifuge simultaneously. The air was

generally used to control the open and closed states of the valve in the suction caisson

lid, and the water was used to compensate evaporation during spinning of the

centrifuge. Single slip ring and dual slip ring were used respectively during

consolidation of the sample and while the caisson was being tested. The units also have

the capacity to carry DC volts via two auxiliary electrical slip rings, where the data

collected from various instrumentations on the centrifuge arm (Figure 3.21) were

digitised (A/D conversion) and then transferred to the control room. In-flight motion

was recorded by the high speed digital cameras mounted on the centrifuge package,



then displayed and recorded in the control room. Developed and fabricated by the Civil

Engineering Workshop at UWA, the slip rings are situated on the central axis of the

centrifuge.

3.7.4 Syringe Pump

Suction installation of the caisson in this research was realised by a motor-driven

syringe pump, which was powered by a 70 Watt RE036-072 Maxon motor combined

with a GP032A planetary gear head capable of delivering torque up to 4.5 Nm. Details

of the syringe pump were described by House (2002). A resolution of 500 encoder

counts per revolution was provided by a Hewlett-Packard photoelectric optical digital

encoder. The 50 mm diameter aluminium piston has a maximum stroke of 190 mm,

and the maximum volume of water it can accommodate is 370×103 mm3. The

maximum drive rate of the motor shaft is 3 mm/s. A pore pressure transducer (±1400

kPa capacity) is located within the syringe pump to record pressures developed in

response to suction or purging of the fluid within the stainless cylinder. The syringe

pump is housed within the centrifuge platform (Figure 3.22) and was designed to

sustain a maximum pressure of 700 kPa.

3.7.5 Load Cells

A ±3 kN capacity axial load cell was used in the original system and has been described

by House (2002). However, its resolution was not suitable for the encoder used in the

new system. Therefore, a ±2 kN axial load cell (Figure 3.23) was used instead in the

centrifuge tests with the new system for measuring the axial force during installation

and axial pullout of the suction caissons.

3.8 T-BAR PENETROMETER

The in situ undrained strength of the soil was investigated using a T-bar penetrometer,

which was developed and introduced to the centrifuge by Stewart & Randolph (1991).

Recently, it has been applied to the field both onshore (Stewart & Randolph, 1994) and

offshore (Randolph et al., 1998; Randolph, 2004; Lunne, et al., 2005). Fabricated by

the Civil Engineering Workshop at UWA, the T-bar penetrometer (Figure 3.24) used

here comprises a 5 mm diameter by 20 mm long bar attached at right angles to the end

of a vertical shaft. The shaft was instrumented with a ±370 N capacity load cell suitable



for determining penetration resistances up to 3.7 MPa (for 100 mm2 bar area). During

the tests, the bearing resistance, q, was recorded continuously and transferred to the

DAQ system. The major advantage of the T-bar penetrometer over other penetrometers

such as the cone, is that the soil is allowed to flow around and over the T-bar during

penetration, therefore, the soil overburden pressure is equilibrated above and below the

bar, and the corrections that are necessary for the cone are avoided. Correlation

between the net bearing pressure and the undrained shear strength can be expressed by a

simple equation:

ubar-T sNq ⋅= (3.6)

where NT-bar is the bearing capacity factor, and su is the undrained shear strength of the

soil. The value of NT-bar has been determined through plasticity analysis, with a lower

limit of 9.2 for a fully smooth interface, and an upper limit of 11.92 for a fully rough

interface (Randolph & Houlsby, 1984). In these tests, an intermediate value of 10.5 was

adopted for NT-bar according to the recommendations by Randolph & Houlsby (1984)

and Stewart & Randolph (1991).

3.9 PORE PRESSURE TRANSDUCERS

Miniature pore pressure transducers (PPTs) were used to monitor both the excess pore

pressure in the clay during consolidation of the sample, and the internal pore pressure

inside the caisson during installation and pullout stages. All PPTs used in this research

were Druck PDCR 81 type (see Figure 3.25) with an external diameter of 6.4 mm and a

length of 11.4 mm. The capacity of the PPTs is 700 kPa, working at a nominal

excitation voltage of 5 volts. Sensitivity of the transducers is 0.023 mV/V/kPa. Pore

water pressures as opposed to total stresses, were ensured by the use of a ceramic filter

located at the tip of each PPT. Reaction of the cells to the input pressures is very linear,

with a standard non-linearity & hysteresis of ±0.2%. External fittings were fabricated

for the PPTs used on top of the caisson.

3.10 SOIL SAMPLES

In this research, caisson tests were performed in reconstituted kaolin clay, with various

overconsolidation ratio and sensitivity. The tested samples include normally

consolidated (NC) clay, lightly overconsolidated (LOC) clay and sensitive clay. The



NC clay was consolidated and tested both at 120 g. Firstly, commercially available dry

kaolin powder was mixed to slurry with a water content of 120% (twice the liquid limit)

of the particular clay soil (Stewart, 1992; House, 2002). During mechanical mixing, the

slurry was de-aired with a vacuum pump to ensure a saturation ratio close to unity.

Then, the slurry was manually placed within the strong-box over a 10 mm deep sand

drain. Internal standpipes were placed in the corners of the sample to facilitate

communication between the free water surface and the external standpipe, thereby

avoiding an increase in pore pressures beneath the low permeability clay layer. Finally,

the slurry was then consolidated at 120 g in the centrifuge, targeting a final depth of 150

mm, which was designed to accommodate the full length of the caisson after installation.

During consolidation, fluid was added to the package in-flight through the hydraulic

slip-ring to compensate for evaporation losses. The external standpipe was set with an

overflow that maintained a constant water level within the sample, and therefore a

constant mass of the package. Three pore pressure transducers were generally installed

at different depths within the sample to monitor consolidation progress through the

dissipation of excess pore pressures. Once consolidated, T-bar penetration tests were

performed to assess the in situ strength of the sample before commencement of the

foundation tests. Other properties of the NC kaolin clay are shown in Table 3.3, with

some properties taken from Stewart (1992), but with some additional tests performed.

The strength gradient of the NC clay is ~1.2 kPa/m, and the strength ratio su/σ′v0 is

around 0.18.

The method for preparing the LOC sample was basically similar to that of the NC

sample, except that it was consolidated at 180 g while tested at 120 g, resulting in an

overconsolidation ratio, OCR, of 1.5. It was shown by later T-bar tests that the major

difference between such a sample and the NC sample is the magnitude of the strength

profile, with much larger gradients (~1.7 kPa/m) for the LOC sample. Besides, a slight

curvature exists in the strength profile of the LOC sample within the top 10 mm, due to

‘topping up’ the slurry during consolidation. Such a curvature disappeared with further

consolidation as the whole test program generally lasted for 1 week. Key properties of

the LOC were measured and are shown in Table 3.3.

A sensitive clay sample was successfully created, with the help of Mr. Mark Richardson

(current PhD student in COFS) and Dr. Susan Burns (Academic visitor from University

of Virginia), since no ‘recipe’ for the constitution of such a soil sample was available

when this research commenced. The sensitive sample was created by dissolving



sodium hexametaphosphate (Na6O18P6) in water for 24 hours prior to mixing; the

concentration ratio of such a dispersant relative to water was 15 g/L. It should be noted

that the water content of the slurry was chosen as 70%, which is much lower than that

(120%) for the NC and LOC samples, although this gave the same slurry consistency.

The slurry was then consolidated at 120 g using the same method as for NC clay, and

the time needed for 90% consolidation was found to be around 2 days, which is longer

than that of the samples with lower sensitivity. Back-figured time for T90 suggested a cv

value of ~2 m2/year (0.063 mm2/s) for the sensitive kaolin clay. Such a value is a bit

smaller than that of normal samples, suggesting that the dispersant has an effect in

reducing the void ratio of the soil.

Table 3.3 Soil properties for NC, LOC and sensitive kaolin clay

Property NC clay LOC clay Sensitive clay

Specific Gravity, Gs 2.60 2.60 2.60

Liquid limit, LL (%) 61 NA NA

Plastic limit, PL (%) 27 NA NA

Average water content, w (%) 47 45 46

Consolidation coefficient, cv (m2/year) 2.6 2.4 2.0

Undrained strength ratio, su/σ'vo 0.18 0.24 0.19

Effective density, γ' (kN/m3) 6.7 7.2 7.3

Coefficient of earth pressure at rest, K0 0.65 0.70 0.55

Sensitivity factor, St 2 - 2.8 2 - 2.5 4 - 5

Values of the lateral stress ratio at rest, K0, were not measured directly. K0 for the NC

kaolin clay was taken from Andersen & Jostad (2002). For the overconsolidated clay,

Mayne & Kulhawy (1982) suggested that K0 may be predicted using (K0)NC×OCRsin φ′,

where (K0)NC is the value of Κ0 for the NC clay, and φ′ is the effective stress friction

angle of the soil. By taking φ′ as 21 , then K0 for the LOC clay used in this research can

be estimated as 0.74. K0 for the LOC kaolin clay (OCR = 1.5) can also be interpolated

between the values of NC kaolin clay (OCR = 1.0, K0 = 0.63) and Drammen clay (OCR

= 4, K0 = 1.0) reported by Andersen & Jostad (2004), and the result is 0.70, which is



very close to the prediction by Mayne & Kulhawy (1982). Therefore, K0 for the LOC

clay was adopted as 0.70, while for the sensitive clay, K0 of 0.55 was estimated from

values for sensitive clays in Offshore Africa reported by Andersen & Jostad (2002).

Chapter 4 4-1 Performance of Total Pressure Transducers


4 PEFRORMANCE OF MINIATURE TOTAL PRESSURE

TRANSDUCERS ON CAISSONS IN CLAY

4.1 INTRODUCTION

As stated in Chapter 2, a type of total pressure transducer with high accuracy and

suitable geometry is needed, in order to measure the radial total stress on the thin-

walled model caisson used in this research. Previous researchers have shown that the

‘piston’ type pressure cells worked well for solid piles (Coop & Wroth, 1989; Bond &

Jardine, 1991), but such cells are generally large and therefore not suitable for open-

ended piles. For the open-ended piles or thin-walled suction caissons, diaphragm type

cells with smaller thickness should be used, built into the caisson wall (of 0.5 mm

thickness). The diaphragm type pressure cells generally have foil strain gauges

connected in a Wheatstone bridge (Figure 4.1) on a small thin membrane, making the

thickness of the cell rather small. Diaphragm type total pressure transducer (Kyowa

PS-10kA) (see Figure 3.1) was chosen for this study.

In this chapter, the accuracy of the Kyowa PS 10kA type miniature total pressure

transducer (TPT), for measuring radial stress changes around caissons in clay under

various loading conditions, was evaluated through a series of calibration tests. Such

tests were performed at 1 g in a modified triaxial apparatus and at high g in the

centrifuge. Monotonic loading, unloading, cyclic loading and sustained loading were

applied on the pressure cell and the corresponding accuracy was assessed. Initial

changes of the cell reading when shifting the pressure cell among air, water and the

kaolin slurry, and cross-sensitivity of the TPTs, were also evaluated.

4.2 FACTORS AFFECTING STRESS MEASUREMENTS

There are mainly three categories of factors to consider during the pressure cell

measurements in a soil sample (Weiler & Kulhawy, 1982): 1) stress cell properties and

geometry, 2) properties of the soil in which the cell is placed, and 3) environmental

conditions.

4.2.1 Stress Cell Geometry and Properties

The stress cell geometry and properties determine the influence of the shape and



stiffness on the stress measurements. Research reported by Weiler & Kulhawy (1982)

shows that the ratio of diameter to thickness (d/t) of the cell needed to be no less than 5,

to minimize the error caused by the wall thickness of the cell itself. The Kyowa PS

10kA type TPT has a diameter of 6 mm, and a thickness of 0.6 mm, giving a d/t ratio of

10, which is larger than the required lower limit. The thickness of such a pressure cell

is larger than the wall thickness (0.5 mm) of the model caisson. However, caissons in

the field generally have a stiffener to reinforce the pad-eye (Newlin, 2003a). Similarly,

the model caissons used here also have a thickened wall thickness of 1.5 mm at that part

(see Figures 3.2 and 3.7). Such a design provides enough space for the transducers,

allowing for the thickness of the adhesives applied during instrumentation.

Another concern about the cell property is the ‘arching effect’, which plays an

important role in the accuracy of stress measurements using the diaphragm type

pressure cells (Weiler & Kulhawy, 1982; Labuz & Theroux, 2005). ‘Arching’ is used

here to refer to the ability of a particular medium to support itself when an external

support is removed. When a pressure is applied to the diaphragm, it tends to deflect

from the original place; this movement will result in reduced pressure as the soil

‘arches’ over the deflected diaphragm (see Figure 4.2). The degree to which this

arching occurs is critical in the stress measurements. The accuracy of measurements of

the pressure cells may be evaluated by a cell action factor (CAF), or so-called

registration ratio (Hvorslev, 1976), defined as follows (Weiler & Kulhawy, 1982):

applied

measured

pp

CAF = (4.1)

where measuredp is the pressure measured by the cell, and appliedp is the pressure applied to

the cell. The closer the value of CAF to unity, the more accurate is the measurement.

Measurements of the diaphragm type pressure cell can be affected by the in-plane

(lateral) compression, since the strain gauges inside the cells are used to measure the

radial and tangential bending strains rather than directly measuring vertical diaphragm

deflection (Brown & Pell, 1967). How the pressure cells react to the axial load applied

on the caisson during centrifuge tests needs to be investigated.

4.2.2 Soil Properties

The boundary condition of the chamber where the pressure cells are calibrated is very

important to the stress-strain response of the soil. A K0 condition of the chamber



generally causes obvious hysteresis loop on the cell measurements during cyclic loading

(see Figure 4.3), since the ratio of σ′h to σ′v will vary during unloading-reloading loops.

This can be avoided by the use of direct stress control in a triaxial apparatus (Weiler &

Kulhawy, 1982). Therefore, a modified triaxial apparatus (see Figure 3.10) was chosen

for carrying out the calibration tests in this research, with cyclic loading applied, to test

the performance of the pressure cells.

4.2.3 Environmental Conditions

Stress cells will give the best results in a dry, controlled environment measuring

relatively static stresses (Weiler & Kulhawy, 1982). However, when measuring the

radial stress changes on suction caissons, TPTs are subjected to a more challenging

environment. For instance, TPTs are required to respond accurately under long-term

loading during consolidation, and sustained loading or cyclic loading during pullout. In

addition, precise responses to the thermal environment changes when TPTs travel from

water to kaolin clay during installation and clay to water during extraction are

necessary. In order to investigate the influence of such factors, the reliability of the

TPTs was tested under sustained loading and cyclic loading, and the changes in the

readings of the TPTs were also evaluated when the TPTs were shifted among air, water

and kaolin slurry.

4.3 SCHEME OF CALIBRATION TESTS

The overall test scheme of calibration tests on TPTs in clay was divided into two steps.

The pressure cell was first calibrated in a triaxial apparatus in the laboratory, then it was

tested in the centrifuge.

The triaxial apparatus was adapted specifically for this project (see section 3.4). The

test started with calibration in water, and then progressed to clay. The readings of the

pressure cells under both loading and unloading phases were recorded. For tests in clay,

the TPTs were tested under both undrained (under either static or cyclic loading) and

drained (under sustained loading) conditions. Changes in the initial values of the

transducers within different media were investigated by shifting the transducer among

air, water and clay slurry. In the centrifuge, the cells were tested in both water and clay.

Tests in water included suspending the caisson to measure the response under sustained

loading, then the caisson was pushed downwards and pulled upwards, to obtain the



response of the pressure cells with respect to changing hydrostatic pressure. Details of

the above test scheme are summarised step by step as follows:

1. The TPTs were calibrated in water, with the applied pressure increased from

zero to 300 kPa, and then decreased to zero.

2. The TPTs were tested in clay. For the undrained conditions, monotonic loading

(from 0 - 300 kPa) and unloading (from 300 - 0 kPa), cyclic loading between 0

and 300 kPa and sustained loading under 150 kPa for 13.3 hours were applied

on the TPTs. Then, under drained conditions, the TPTs were tested under

sustained loading for a maximum of 98 hours.

3. Changes in the initial values of the TPTs in different media, including air, water

and kaolin slurry were assessed.

4. Cross-sensitivity of the TPTs relative to axial loading applied to the caisson was

evaluated.

5. TPTs were calibrated in water at 120 g in the centrifuge, when the caisson was

moved up and down slowly.

6. Finally, the stability of the readings of TPTs in the centrifuge was assessed

under sustained loading for 24 hours at 120 g.

Test results are presented and analysed below.

4.4 TEST RESULTS

Calibration tests on the pressure cells started in water. The overall purpose of the

calibration tests on TPTs in water is to obtain a calibration factor, which is the ratio of

the applied pressure to the electronic output of the TPT, to be useed to factor the

electronic output of tests in clay.

4.4.1 Calibration Tests in Water

The input and output system for the calibration test for TPTs in water was shown

previously in Figure 3.12. Two modifications were made to the triaxial system. The

first step was to replace the clay sample with a plastic sleeve, where water can flow in

and out of the annulus between the caisson and the sleeve. The second step was to

modify the top-cap of the triaxial sample to allow wires to pass through (Figure 4.4).



Then the caisson with the two pressure cells was placed on the bottom cap of the triaxial

apparatus (Figure 4.5), and a plastic sleeve with a diameter of 75 mm and a length of

150 mm was used to cover it. The top-cap was supported by the plastic tube (Figure

4.5b). A black lead connecting the TPTs on the caisson exited the top-cap and then

connected to the signal channel on the bottom plate of the triaxial apparatus (Figure

4.5b). After placing the caisson and connecting all the wires for transferring the signal,

the stainless steel cylinder of the triaxial apparatus was closed and gradually filled with

water. The water flow was stopped when the cylinder was full. Then different levels of

water pressure were applied using the GDS pressure controller. At the same time, the

responses of the TPTs were recorded automatically. The data acquisition system for the

triaxial calibration tests in water with respect to the pressure generation was shown

previously in Figure 3.12.

The TPT readings were recorded in bits, thus a relationship between the measured

pressure in bits and the applied pressure (in kPa) could be obtained for each loading

phase. Since the TPTs were located at the mid-height of the caisson and the largest

pressure to be encountered at that point in the centrifuge tests was around 200 kPa, the

maximum pressure applied in the calibration tests was 300 kPa. The applied pressure

was increased in stages of 10 kPa until a final pressure of 300 kPa was reached. This

was then reduced to zero in 10 kPa increments.

It should be noted that, in each loading stage, time was allowed for the pressure

controller to become stable, as indicated by the static reading on the water volume in the

GDS pressure controller. Altogether, seven loading tests and seven unloading tests

were performed in two individual series. The response of the TPTs varied linearly with

the applied pressure during both the loading and unloading phases (Figure 4.6). When

the applied stress was below 300 kPa, the gradients of the loading and unloading curves

were quite close for both transducers. In Figure 4.6 the average calibration factors

during loading tests were 0.560 kPa/bit for TPT1 and 0.524 kPa/bit for TPT2, while

during unloading tests they were 0.563 kPa/bit and 0.530 kPa/bit respectively. After 5

series of loading and unloading tests, the average calibration factors were 0.560 kPa/bit

and 0.537 kPa/bit for TPT1 and TPT2, respectively. The difference between the

calibration factors during loading and unloading was less than 1%. Judging from the

response of the TPTs, the results of the calibration tests in water were quite

encouraging.



4.4.2 Calibration tests in Kaolin Clay

In clay, arching effects together with several other factors mentioned in section 4.2

could affect the accuracy of pressure cells. Therefore, calibration tests were carried out

in the kaolin clay in both undrained and drained conditions. The former case

corresponds to the instantaneous reaction of the cells to the external pressure, while the

latter can reveal whether or not such a transducer is reliable during the longer term,

when the effective stress (and stiffness) of the surrounding clay varies as it consolidates

or swells. Tests under drained conditions are considered to be practically relevant for

measuring the radial stress changes around caissons during consolidation.

4.4.2.1 Undrained calibration tests in kaolin clay

Firstly, a kaolin clay sample was taken from clay consolidated at a vacuum pressure of

120 kPa, with an effective density (γ ) of 6.8 kN/m3, and a moisture content (w) of

51%. The average undrained shear strength was 12 kPa along the depth according to

the 1 g T-bar tests, with homogeneous conditions over the whole depth. The diameter

of the sample was 75 mm with a height of 150 mm, which accommodated the caisson

completely (see Figure 4.7a). Subsequently, the caisson was inserted manually and

statically into the centre of the clay sample, to ensure full contact between the pressure

cells and the surrounding soil. Then the sample was placed onto the bottom cap of the

triaxial apparatus before being covered by a membrane. The wiring connection was the

same as that during the calibration tests in water. The caisson and the clay sample after

being coated with the membrane is shown in Figure 4.7b. In the undrained test, the

drainage valve on the bottom cap was sealed throughout the calibration process, so that

water was prevented from flowing out from within the membrane and the volume of the

sample remained unchanged (∆e = 0) during the whole process.

The undrained calibration was carried out in loading, unloading, cyclic loading and

sustained loading conditions. Load was increased from 0 to 100 kPa in 20 kPa

increments and from 100 to 300 kPa in 50 kPa increments, and then reduced to zero in

50 kPa increments. Cyclic loading was achieved by applying the load in a sequence of

0, 50, 100, 40, 80, 100, 200, 120, 160, 180, 200 kPa, thus cycling the load between 0

and 100 kPa, and between 100 and 200 kPa. Sustained loading was applied by

maintaining the applied pressure at 150 kPa for 800 minutes (13.3 hours) and 100 kPa

for 260 minutes (4.3 hours) while recording the change in the TPT readings. The output

was transformed from bits into pressure (in kPa) using the calibration factors obtained



from the calibration tests in water.

The measured and applied pressure are plotted in Figures 4.8a - d for the loading,

unloading, cyclic loading and sustained loading tests under undrained conditions. It can

be seen in the graphs that the measured results are close to the theoretical predictions

for all four tests. During unloading, no hysteresis loop was observed in the plots since

the response was almost identical to that under loading. This was strong evidence that

any ‘arching effect’ was reduced to a very small extent by adopting such a TPT design

on the caisson. CAFs (see Equation 4.1) during monotonic loading and unloading in

kaolin clay are summarised in Table 4.1. It can be concluded from Table 4.1 that

during the monotonic loading process, the CAFs are 0.988 (error = –1.2%) and 0.997

(error = –0.3%) for TPT1 and TPT2, respectively. During unloading the CAFs are

0.984 (error = –1.6%) and 0.962 (error = –3.8%) respectively for these two cells. These

errors are slightly larger than during loading, although the overall accuracy is still very

high.

Table 4.1 CAFs for TPTs calibrated in kaolin during loading and unloading

Load Unload papp

(kPa) CAFTPT1 CAFTPT2 papp

(kPa) CAFTPT1 CAFTPT2

40 0.933 0.947 300 0.984 0.961

100 0.999 1.041 250 0.992 0.964

150 0.997 1.005 200 0.987 0.962

200 0.996 0.999 150 0.989 0.969

250 1.002 1.000 100 0.985 0.966

300 1.001 0.993 50 0.966 0.951

Average 0.988 0.997 Average 0.984 0.962

The CAFs of the TPTs under cyclic loading (see Figure 4.8c) are listed in Table 4.2. It

can be seen that under cyclic loading, the measurement is of high precision, with CAFs

being 0.998 (error = –0.2%) and 0.992 (error = –0.8%) for TPT1 and TPT2,

respectively. In addition, the hysteresis loop shown in Figure 4.3 (Weiler & Kulhawy,

1982) was not found in the results. Although it was mentioned in Weiler & Kulhawy



(1982) that the hysteresis loop is small and often disappears upon cyclic loading for the

triaxial calibration, the inaccuracy introduced in the first loop during installation of the

caissons could cause unreliable stress measurements. In Figure 4.8c, the linear

relationship between the applied pressure and the TPT response under cyclic loading

eliminates this concern.

Table 4.2 CAFs for TPTs under cyclic loading in kaolin clay

papp CAFTPT1 CAFTPT2 papp CAFTPT1 CAFTPT2

(kPa) (kPa)

50 0.971 0.999 300 0.998 0.979

100 0.990 1.010 220 1.006 0.978

40 1.007 1.022 260 1.003 0.981

80 0.992 1.009 280 1.000 0.978

100 0.995 1.009 300 1.000 0.979

200 0.999 1.002 250 1.003 0.978

120 0.991 0.989 200 1.003 0.976

160 0.991 0.990 150 1.005 0.971

180 0.991 0.987 100 1.018 0.963

200 0.988 0.999 50 1.021 0.907

Average 0.998 0.992

Table 4.3 CAFs for TPTs under sustained loading in kaolin clay

Time papp CAFTPT1 CAFTPT2

(minute) (kPa)

0 150 0.988 1.011

800 150 0.997 1.008

800 100 1.012 1.030

1060 100 0.986 1.025

Average 0.996 1.018

Another concern on the accuracy of the pressure cells under sustained loading (see



Figure 4.8d) was removed by the test results shown in Table 4.3. It can be seen in the

table that, under a continuous pressure of 150 kPa for 13.3 hours (800 minutes), the

CAFs of the pressure cells drifted only –0.9% (from 0.988 to 0.997) in TPT1 and 0.3%

(from 1.011 to 1.008) in TPT2; under a sustained pressure of 100 kPa for 4.3 hours,

they drifted –2.6% (from 1.012 to 0.986) and –0.5% (from 1.030 to 1.025) respectively.

Such a consistent response again confirms the reliability of the pressure cells used here.

4.4.2.2 Drained calibration tests in kaolin clay

The above tests were all performed under undrained conditions, for which no

consolidation within the clay samples was allowed during the whole testing progress.

Therefore, there was no volume change in the soil. However, under drained conditions

the effective stress and stiffness of the clay is changing, and this may affect any

tendency to arch, and cause the cell to under-register or over-register. Hence, TPTs

were also calibrated under drained conditions in clay.

Methods for preparing the saturated kaolin clay samples, and the lead connections on

the triaxial apparatus were both the same as those used previously in the undrained

calibration tests. The caisson was first installed into the clay sample before being

installed in the triaxial apparatus. After closing the stainless steel cylinder, water

pressure was applied through the GDS digital pressure controller until reaching a real

value of papp = 158 kPa (original target was 150 kPa), which was maintained throughout

the process. At the same time, the drainage valve was opened to allow water to flow

and excess pore pressure to dissipate. Variations of the TPT response during

installation in the clay, filling the cylindrical chamber with water and applying external

pressure are shown in Figure 4.9a. Unfortunately during consolidation, TPT2 stopped

working as indicated by the output suddenly exceeding the upper limit. Therefore, only

the response of TPT1 was measured.

It took 65 minutes to fill the cylinder with water and for the pressure controller to

stabilise. However, for safety a delay of 17.36 hours (62507 seconds) was allowed

before starting the consolidation. The volume change of the clay sample was measured

by the volume of water passing into the triaxial cell during consolidation; the value

increased gradually with time during consolidation (Figure 4.9b). The reaction of TPT1

increased to 164 kPa immediately, probably due to the sudden opening of the drainage

valve, and then decreased gradually when the applied pressure remained unchanged



during the drainage phase (Figure 4.9a). During the whole consolidation process, the

readings of TPT1 varied around 158 ± 5 kPa, the discrepancy between the measured

pressure and the applied pressure was small and decreased as consolidation progressed.

At the end of consolidation, the TPT readings were almost constant, which is in

accordance with the observation that the change in the volume of the sample stopped

(see Figure 4.9b). At this stage, the consolidation of the soil came to an end, and the

measured pressure approached the applied pressure. The average CAF over the

consolidation period was 0.992 (Figure 4.9c), which is only –0.8% smaller than the

theoretical value of unity. The maximum error was only 5% within the whole

consolidation process.

Another calibration test in clay was performed on a second sample. Immediately after

the drainage valve was opened, the reaction of TPT1 fluctuated between 144 - 162 kPa,

probably due to the ‘shock’ from valve opening. Under the applied pressure of 154

kPa, the soil sample consolidated with time and approached full consolidation after

around 98 hours (Figure 4.10). During the consolidation process, the measured pressure

was relatively stable, varying between 142.45 - 155.60 kPa, which was very close to the

applied pressure, although some discrepancy existed during the early stages of

consolidation (Figure 4.10a). The average CAF during the drained period was 0.969,

which was also very close to unity (under-registered by only –3.1%), and the minimum

recorded pressure was 142.45 kPa at 125267 s (see Figure 4.10a), corresponding to a

CAF of 0.925.

As a result, it can be concluded from the above two calibration tests in clay that the

Kyowa PS-10KA pressure cell can present very reliable measurements (with the CAF

larger than 0.925; the error is less than 7.5%) during consolidation under sustained

loading in clay at 1 g.

4.4.3 Variation of Initial Values of TPTs in Different Media

The influence of changing media for the pressure cells was evaluated by shifting the

TPTs among air, water and clay. Such an evaluation is important for the radial stress

measurements on caissons since the TPTs would move from air into water, and then

into the kaolin clay in the centrifuge tests. The variations of the initial values of the

TPTs when shifting from air to water and from air to kaolin slurry are depicted in

Figure 4.11 for four typical tests. By setting the reading of the TPTs in air as zero, the

actual pressure readings of the TPTs in air and water are as listed in Table 4.4. It is



shown that the TPTs tend to give a lower value in water than in air under the same

condition; the average change in initial values from two such tests was –4.2 kPa. When

the cells shifted from air into kaolin slurry, the average change was –3.3 kPa. These

results suggest a change of 0.9 kPa in the initial value when TPTs enter from water into

clay. This change is rather small compared to the surrounding pressure of 100 - 200

kPa that could be encountered in the caisson tests.

Table 4.4 Initial value of TPTs in different media

TPT from air to water TPT from air to kaolin clay

Test Media TPT1

(kPa)

TPT2

(kPa)

Test Media TPT1

(kPa)

TPT2

(kPa)

Air 0 0 Air 0 0 A1

Water –6.4 –1.0 K1

Kaolin –4.0 –1.5

Air 0 0 Air 0 0 A2

Water –7.8 –1.5 K2

Kaolin –5.7 –2.1

Average change from air to water: – 4.2 kPa

Average change from air to kaolin clay: – 3.3 kPa

4.4.4 Cross-sensitivity to Axial Loading on Caisson

During the penetration and pullout of the caisson in the centrifuge, the stress along the

axis of the caisson may have some effect on the measurement of the TPTs. A test was

therefore designed to evaluate the cross-sensitivity (Figure 4.12). To avoid causing

damage to the caisson itself in the centrifuge, it was placed vertically on the ground at 1

g, but with an axial load applied, with a magnitude no less than that applied in the

centrifuges tests. From previous projects on similar caisson tests undertaken by House

(2002), the maximum axial load experienced by the caisson was approximately 300 N.

Yet in the calibration tests here the maximum axial force applied was 1,270 N, to cover

some unpredictable issues during cyclic loading tests. The strength of the aluminium

and the slenderness of the caisson itself were both issues of consideration for choosing

such a maximum load. The response of the TPTs under the axial load on the caisson

was recorded.

The horizontal stress measurements of the TPTs were not affected by the axial load on



the caisson (Figure 4.13). Under an axial load of 1,270 N, the output of TPT1 only

increased by 5.5 kPa, and in TPT2 decreased by 2.7 kPa, compared to the readings at

rest. It can be inferred from the quasi-linear relationship that under the normal loading

limit of 300 N on the caisson, the reading in the TPTs would only change by ~1.3 kPa,

which is small compared to the corresponding horizontal pressure of ~200 kPa during

centrifuge tests. These tests indicate a negligible cross-sensitivity of the TPTs to axial

loading.

4.4.5 Calibration Tests in the Centrifuge

The Kyowa PS-10kA type TPTs were calibrated in the centrifuge, in order to get

confirmation of the performance of the pressure cells during typical centrifuge tests on

caissons.

4.4.5.1 Static movement in water

The caisson bearing the TPTs was attached to the actuator (Figure 4.14). The

calibration test was undertaken at 120 g in water by installing and pulling out the

caisson at a very slow speed (around 0.1 mm/s) in-flight. The slow speed was adopted

so as to limit the influence from any water currents caused by this movement. Since

TPT2 ceased working during the previous test, it was replaced and re-calibrated before

the beginning of the centrifuge test. The response of the TPTs and the variation of the

CAFs at 120 g versus the depth under the water level are shown in Figures 4.15a & b.

Both transducers worked well at 120 g in the centrifuge with almost identical results to

the theoretical hydrostatic line (see Figure 4.15a). The variation of the CAF stays close

to unity, with 0.960 being the lower bound and 1.011 the upper bound (see Figure

4.15b). The average CAFs along the penetration depth of 7.4 m were 0.992 (error =

–0.8%) and 0.989 (error = –1.1%) for TPT1 and TPT2 respectively. As a result, it can

be concluded that the Kyowa PS-10KA can give very reliable measurements during

static movement of caissons in water in the centrifuge.

4.4.5.2 Sustained loading in water

To investigate the performance of the TPTs under sustained loading at high g

conditions, the TPTs on the caisson were suspended at a certain depth below the water

surface. A constant rate of water flow, which was equivalent to the evaporation of the

water in flight, was introduced into the strongbox during the test. The centrifuge started



from 1 g and was ramped up to 120 g for the test. The response increased to

approximately 87 to 88 kPa after the acceleration became stable. From t = 2608 s (or

0.72 hours) to t = 87901 s (or 24.4 hours), the average readings of the TPT1 and TPT2

were 87.4 kPa and 88.29 kPa, respectively, with very stable readings for both cells (see

Figure 4.16).

4.4.5.3 Application to caisson penetration in clay in the centrifuge

This investigation was continued by calibrating the TPTs in kaolin clay in the

centrifuge. In such tests, the caisson was installed vertically by self-weight penetration

followed by suction installation into the kaolin clay. Variations of the measured radial

stress (σr) changes with the depth of tip of the caisson during installation in two typical

centrifuge tests are shown in Figures 4.17 and Figure 4.18 respectively. The first one

was tested in NC clay using caisson 1 (see Figure 3.2), and the second was tested in

LOC clay using caisson 2 (see Figure 3.7). Both tests were carried out at 120 g with the

caissons hung just over the mudline and submerged in water. For the test in NC clay

using caisson 1, the distance from the TPTs to the caisson tip is 7.2 m at 120 g. It can

be seen in Figure 4.17 that during installation, σr increases with depth above 7.2 m, and

the measured values and gradient agree well with those of the hydrostatic line. This is

reasonable since the TPTs are travelling in the water in this region. As soon as the

TPTs pass 7.2 m and enter the clay, σr increases at an obviously larger gradient. The

gradient of the radial total stress relative to the hydrostatic pressure, d(σr − u0)/dz, is

6.61 kPa/m.

For the test in LOC clay using caisson 2 (see Figure 3.7), similar results are obtained,

except that the turning point in the measured σr occur at 4.8 m depth (see Figure 4.18),

which is the distance between the TPTs and the tip of this caisson. After the TPTs leave

the water and enter the soil at 4.8 m depth, the measured radial total stress also increases

at a larger gradient. The gradient of σr − u0 here is 7.88 kPa/m, which is obviously

larger than that in NC clay. This shows that the measured radial total stress changes

respond reasonably to the stiffness of the ambient clay, since the LOC clay has a larger

gradient of undrained shear strength ratio (su/σ′v0) than the NC clay (see Table 3.3). At

the end of the test in LOC clay, there was some decrease in the measured radial total

stress, which is discussed in detail in Chapter 7.

Based on the above analysis of TPT readings in two typical caisson tests in the



centrifuge, it can be seen that the TPTs present reasonable measurements during caisson

installation, with rapid response to the change in applied pressures, and are therefore

suitable for use in this research.

4.5 CONCLUSION

In this chapter, miniature total pressure transducers (TPTs) were applied to measure the

horizontal pressure acting on the shaft of suction caissons. The performance of the

Kyowa PS-10KA pressure cell in clay was evaluated through a series of calibration

tests. Calibration tests on the pressure cells were performed both at 1 g using a

modified triaxial apparatus, and in high g conditions using the centrifuge. Under 1 g

conditions, the accuracy of the TPTs was assessed under loading, unloading, cyclic

loading and sustained loading. The accuracy of the cells in clay was evaluated under

both undrained and drained conditions. The change in initial values of the TPTs in

different media and the cross-sensitivity to axial loading were also studied. Finally, the

TPTs were calibrated in water in the centrifuge, during both static movement and

sustained conditions; the measurements of the TPTs during caisson installation in NC

and LOC clays were also assessed. The following conclusions can be presented on the

Kyowa PS-10KA pressure transducer and the stress measuring system on the suction

caissons:

1. The response of the TPTs tested varies linearly with the applied pressure both in

water and clay.

2. The pressure cell adopted in this research has an appropriate geometry that can be

installed on the model suction caisson. The design limits any arching effect to a

minimal amount.

3. The pressure cell adopted in this study has high accuracy for radial stress

measurements on caissons in clay under both 1 g and high g conditions in the

centrifuge, during both static conditions and during movements.

4. The accuracy of the miniature TPTs tested in a triaxial cell under loading,

unloading and cyclic loading, sustained loading in kaolin clay under undrained

conditions is better than 98%.

5. The cells performed well under sustained loading. The average accuracy of the

TPTs under drained conditions in clay is 97%, although this varies with

consolidation time with a lowest value of 92.5% (representing a maximum error of



7.5%).

6. The change in the initial value of the TPTs in different media is small. The

measurement will only change less than 1.5 kPa when entering from water into

kaolin clay, as will be encountered during caisson installation.

7. Calibration tests under water at 120 g gave an average accuracy of the TPTs of

99%, with the lowest recorded value being 96%; the pressure cells also gave

constant readings under sustained loading in the centrifuge.

8. During static penetration of the instrumented caisson in clay at 120 g, the TPTs

gave reasonable responses to the applied pressure.

In conclusion, the miniature TPTs chosen in this work performed well on the suction

caisson both in water and in clay. They appear to yield reliable measurements in the

centrifuge under various loading conditions. The measured radial total stress during a

series of caisson tests in the centrifuge will be presented and analysed in detail in

Chapter 7.

Chapter 5 5-1 Interface Characteristics by Ring Shear Tests


5 STUDYING THE INTERFACE CHARACTERISTICS

BETWEEN SUCTION CAISSON AND CLAY

5.1 INTRODUCTION

Shaft friction of the suction caisson is not measured directly in this research, yet it can

be determined by the interface friction angle and the radial effective stress at the point

of interest, using the fundamental expression:

tanδσ r ⋅′=sf (5.1)

where

sf = shear stress;

rσ′ = radial effective stress acting on the pile shaft at failure;

δ = interface friction angle between pile and soil.

Once the radial total stress is measured, the radial effective stress (σ′r) can be inferred,

by assuming that the excess pore pressure (∆u) is small relative to the hydrostatic

pressure. In fact, this can be controlled in experiments, by setting the shearing speed to

a very low value, thus preventing the excess pore pressure from building up. The value

of δ has been found to be linked best to residual conditions, owing to the large relative

displacement during pile (or caisson here) installation (Lehane, 1992; Chow, 1997).

The interface friction angles can be determined from several types of interface tests,

such as shear box tests (Kulhawy & Peterson, 1979; Desai et al., 1985), simple shear

tests (Kjellman, 1951; Roscoe, 1970; Budhu, 1983), rod shear tests (Brummund &

Leonards, 1973; Jewell & Randolph, 1988), constant stiffness interface shear tests

(Ooi & Carter, 1987) and ring shear tests (Bishop et al., 1971; Bromhead, 1979;

Yoshimi & Kishida, 1981; Kelly, 2001). The main advantage of ring shear tests is that

the distance the sample can be sheared is unlimited, so that residual conditions can be

ensured. Therefore, ring shear tests were adopted here to measure the residual friction

angle between the caisson and the soil.



5.2 RING SHEAR APPARATUS

There are two main categories of ring shear apparatus: the Bishop (1971) model, and

the Bromhead (1979) model. Compared to the ‘Bishop’ equipment, the ‘Bromhead’

equipment is much simpler, and thus is more practical to use. Based on the ring shear

tests performed at Imperial College with the ‘Bishop’ apparatus and at Fugro Limited

with the ‘Bromhead’ equipment, Ramsey et al. (1998) stated that in most cases δr is

relatively insensitive to the type of ring-shear apparatus. For the present tests, a

Bromhead Ring Shear apparatus (WF25850) (see Figure 3.15) was used. An

introduction to this apparatus was given in Chapter 3.

5.3 DESCRIPTION AND GENERAL PRINCIPLES

The sample was prepared by first kneading clay, at a water content of ~ 50%, into the

concentric ring, to form an annular sample with 5 mm thickness, 70 mm inner diameter

and 100 mm outer diameter. Details of the top platen and concentric rings were shown

in Figure 3.14. Then, the sample was compressed vertically between porous bronze

loading platens (see Figure 3.14) by means of a counter-balanced loading system with a

10 : 1 ratio lever. Rotation was imparted to the base platen and lower platen by means

of a variable speed motor and gearbox through a worm drive. This caused the sample to

shear, forming a shear surface between the soil and upper platen.

The torque, T, is measured continuously and may be converted to an average shear

stress, τ, acting on the interface, given by

( )31

32 RRπ

32

Tτ−⋅⋅

= (5.2)

where R1 and R2 are respectively the inner and outer radii of the annular sample.

The tests are conducted under fully drained conditions, with the sample immersed in

water, so that the average effective stress σ′v is known. Equation 5.1 then allows the

interface friction angle to be determined. The value of δ usually decreases as the test

progresses, reaching a residual value after around 10 mm displacement (Chow, 1997).

The settlement of the upper platen during consolidation was monitored by means of a

sensitive dial gauge bearing on the top of the load hanger. Torque transmitted through

the sample was measured by two proving rings.



5.4 SOIL SAMPLE PREPARATION

5.4.1 Fabrication of Top Platen

Two different top platens were fabricated in the workshop, to simulate different

roughness of the caisson surface. One platen was smooth (anodised without

sandblasting); the other was sandblasted to a centre line average (CLA) roughness of

2.5 µm, being the same roughness as that of the model caisson, then it was anodised to

resist corrosion. Both platens were made from the same material, namely 6061 T6

aluminium, as for the model caissons.

5.4.2 Sample Filling

The lower porous platen and sample container assembly were removed from the

machine by undoing the two knurled retaining nuts and lifting this assembly clear of the

two locating studs. This was facilitated by swinging the proving ring turrets and load

hanger clear of the water bath. The remoulded clay was taken directly from the

consolidated kaolin samples (NC, LOC or sensitive clays) used for tests on suction

caissons in the centrifuge. No special technique was needed to avoid disturbance since

the sample was to be remoulded in the ring shear tests, except that the samples were

stored carefully under airtight conditions in a storage room where constant temperature

and moisture content were maintained before the ring shear tests commenced. The

remoulded kaolin sample was then kneaded evenly into the annular cavity using a small

spatula. The top of the sample was then scraped level with the top of the confining

rings, and the assembly placed in position, on the locating studs.

The upper platen was then situated; it was located on the centring pin, on which a light

smear of grease was applied. The water bath need not be removed during this

operation. At this point the adjusting rods should be located at the appropriate radii on

the torque arm, before the load hanger was swung into position. The settlement dial

gauge was then mounted to bear on top of the loading yoke screw adjuster. Time was

allowed for it to come to equilibrium under the load of the top platen. If too much load

was applied too quickly, excessive ‘squeeze’ or loss of soil through the clearance

between the upper platen and confining rings occurred. During the tests in this

research, soil ‘squeezing’ was found to be a very disturbing problem during the early



stages of the experiments, resulting in several tests being spoiled due to the additional

friction thus caused. Such problems are discussed later.

5.4.3 Sample Consolidation

The sample was then consolidated under vertical load, providing the desired normal

effective stress on the horizontal planes. In this research, since the caisson has a

prototype length of 14.4 m and the effective unit weight of the kaolin clay is 6.7 kN/m3,

considering a K0 of 0.65 and a passive earth pressure factor of 2, the lower bound and

upper bound of the average lateral effective stress acting on the caisson shaft were

31 kPa and 97 kPa, respectively. While some exploratory tests were performed under a

vertical effective pressure, σ v0, of 100 - 125 kPa, all other tests were undertaken under

a σ v0 of 50 kPa. During both the consolidation and shear processes, the sample was

flooded to prevent it from drying out.

5.4.4 Forming the ‘shear plane’

After consolidation, the samples were subjected to a relative displacement in excess of

1.2 m by a series of shearing pulses performed manually at a rate of 500 mm/min. Each

pulse was followed by a pause period of 3 minutes. Such a shearing stage was imposed

in order to form a ‘shear plane’, to simulate the displacement history experienced by the

soil element adjacent to the caisson shaft during installation (Ramsey et al., 1998). This

was sufficient to ensure that residual conditions were achieved in the shear zone.

5.4.5 Residual Strength Measurement

The proving rings were aligned to bear at right angles to the torque arms, which transfer

the torque force to the proving rings. Then the samples were allowed to reconsolidate

for 24 hours before the shearing commenced. The ring shear apparatus allows a built

in velocity of shearing between 0.018 - 45 mm/min. The slowest rate of 0.018 mm/min

was selected so as to ensure at least 95% dissipation of any excess pore pressures over 1

mm displacement. If higher shear rates were chosen, rheological or viscous shear

strength effects could be induced.

The vertical settlement of soil samples was carefully monitored to ensure no abrupt

change in the structure of the soil occurred. Readings from the two proving rings were

recorded every minute during the first 30 minutes of shearing and every 5 minutes

afterwards, until the two readings became stable for a time interval of 4 hours, when the



residual strength has been reached for most tests. The shearing distance observed in the

tests varied between 5 and 10 mm.

Since the torque T is given by the mean load on the proving rings, F multiplied by the

distance between them (L), i.e.

( ) 2L/FFT 21 ⋅+= (5.3)

Considering Equation 5.2, the shear stress τ can be calculated as

( )( )3

132

21

RRπ4LFF3

τ−⋅

⋅+⋅= (5.4)

The normal effective stress σ′v is given by

( )21

22

v RRπPσ−⋅

=′ (5.5)

where P is the total vertical load on the sample, i.e. 10 times the load on the hanger plus

the weight of the top platen. Hence,

( ) ( )( ) PRR4

LRRFF3tanδστ

31

32

21

2221

rv ⋅−⋅

⋅−⋅+⋅==

′ (5.6)

5.5 TEST RESULTS

5.5.1 Sample 1: Smooth Ring Platen, NC clay

Three tests were performed on the NC sample taken from centrifuge test Box 2, with a

smooth platen. By using Equations 5.2 - 5.6, the measured values of torque force from

the proving rings versus shearing distance (F - d), vertical settlement versus shearing

distance (s - d), shear stress versus shearing distance (τ - d), and interface friction angle

versus shearing distance (δ - d) are shown in Figures 5.1 - 5.3 for tests S1-1 (σ v0 = 100

kPa), S1-2 (σ v0 = 100 kPa) and S1-3 (σ v0 = 125 kPa), respectively. These plots show

that the residual interface angle was 15 , 26 and 24 , respectively, for the above three

tests. The results varied significantly between each individual test.

Figure 5.4 expresses the relationship between the interface friction angle and the

plasticity index of clay soils (Lemos, 1986; Tika, 1989), based on a large variety of ring

shear tests. For PI = 30% for the NC kaolin clay used in this research, the residual

angle should be around 11 to 18 , depending on whether the interface is rough or

smooth. However, the measured values (average 22 ) were much larger than that

estimated from Figure 5.4 for a relatively smooth aluminium plate in clay. In addition



to the strange values obtained for the residual friction angle, Figures 5.1 - 5.3 also show

that the shearing distance needed for developing the residual state was generally more

than 15 mm, which was much larger than the result of less than 10 mm found by other

researchers (Chow, 1997).

Further inspection of the test procedures revealed a common phenomenon for the above

three tests, i.e. kaolin clay was ‘squeezed’ into the gap between the upper platen and

confining rings, when the ‘shearing surface’ was created by pulses of fast shearing.

Existence of the extra clay would introduce extra friction during shearing; therefore, the

measured friction angle was larger than the real value. Upon realizing this, a step was

added after the test procedure shown in section 5.4.4. After the shear plane was formed,

the platen was taken out and clay in the gap was carefully removed, then the fast

shearing pulses were imposed again under the vertical pressure, until no clay or only

very trivial clay was visible in the gap. Then the sample was reconsolidated for 24

hours before the formal shearing started. In subsequent tests, the vertical stress was

decreased to 50 kPa to assist in preventing the soil from squeezing out.

5.5.2 Sample 2: NC clay, Sand-blasted Ring Platen

In this test, the extra step of removing the clay in the gap was included. Such a test was

carried out with the NC clay taken from the centrifuge test Box 6. The platen used

(Figure 3.14) was sand-blasted to a CLA roughness of 2.5 µm, the same roughness as

that of the model caissons tested in the centrifuge. After shearing, a layer of clay was

found to adhere to the platen (Figure 5.5), indicating that the shearing occurred inside

the clay, rather than right at the interface of the plate and the soil.

Under a vertical effective stress of 50 kPa, the measured results for (F - d), (s - d),

(τ - d) and (δ - d) are shown in Figure 5.6 and Figure 5.7 respectively for tests S2-1 and

S2-2. A comparison of the variation of the interface friction angle with the shearing

distance between these two tests is shown in Figure 5.8. The peak values of the

interface friction angle were respectively 19.4 and 19.1 for these two tests, while the

residual values of the interface friction angle were 17.7 and 17.5 , respectively, with an

average value of δr = 17.6 . The overall decrease in the friction angle was less than 2

degrees within a shearing distance of 10 mm (see Figure 5.8). This indicates that the

residual state has been reached after such a distance of shearing. This also shows that

the influence of the clay trapped in the gap of the top platen and concentric rings had

been eliminated, and the additional step seemed to be effective.



5.5.3 Sample 3: LOC clay, Sandblasted Ring Platen

A lightly overconsolidated clay (LOC) sample was taken from centrifuge test Box 13.

The clay has an overconsolidation ratio of 1.5, an effective density of 7.2 kN/m3, and a

sensitivity of 2 - 2.5. The platen used was sandblasted to a CLA roughness of 2.5 µm.

Under a vertical effective stress of 50 kPa, the measured results for (F - d), (s - d),

(τ - d) and (δ - d) are shown in Figure 5.9 and Figure 5.10 respectively for test S3-1 and

S3-2. The peak values of the interface friction angle from these two tests were

19.4 and 18.3°, respectively, while the residual values of the interface friction angle

were 18.3 and 17.9 , with an average δr of 18.1 . It took 3 mm and 7 mm of shearing

distance respectively in these two tests to reach the residual state (Figures 5.9 and 5.10).

The difference between the peak friction angle and the residual value was less than 1.5°.

Such a small difference also indicated that the soil entrapped in the gap between the top

platen and concentric rings was quite trivial, and the result was reliable.

5.5.4 Sample 4: Sensitive clay, Sandblasted Ring Platen

A more sensitive clay sample was taken from centrifuge test Box 14; methods for

preparing such a sample have been shown in Chapter 3. The sensitivity of such a

sample was 4 - 5, which is higher than 2 - 2.8 for the standard kaolin clay, according to

the cyclic T-bar tests. The effective density was 7.3 kN/m3. The sand-blasted platen

was used in the ring shear tests, using the modified test procedure, i.e. the squeezed soil

was removed before the formal shearing. The test results under a vertical effective

stress of 50 kPa are depicted in Figure 5.11 and Figure 5.12. These graphs show that

the peak friction angle was 14.8 and 14.3 respectively for test S4-1 and S4-2, after a

shearing distance of 6 - 7 mm, with an average value of 14.6 . The residual friction

angle was 11.3 and 12.1 , respectively for the above tests, with an average δr of 11.7°,

which was markedly smaller than those obtained in the NC and LOC clays. The

difference between the peak and residual values was around 3 , which was slightly

larger than those measured in NC and LOC clays. As discussed later in Chapter 6,

during the centrifuge tests on caissons in sensitive clay, soil samples were impossible to

retrieve using the tubular sampler either before or after the caisson tests, since the

remoulded soil was too soft to be retained by the sampler. This is consistent with the

measurement of a rather lower δr in sensitive clay.



5.6 CONCLUSIONS

Ring shear tests were carried out in a Bromhead-type ring shear apparatus to investigate

the interface friction angle for the model caisson used in this study. It can be concluded

from the results that:

1. Removal of the soil entrapped between the concentric ring and the top platen

after creating the shearing surface was found to be essential in obtaining the

correct residual friction angle for the Bromhead-type ring shear apparatus. It is

suggested that such a step should be added into the operation manual.

2. For the sand-blasted surface with the same roughness as the model caisson,

shearing occurred in the soil, rather than at the interface between soil and plate.

3. A peak interface friction angle δp was reached shortly (generally 0.1 - 0.2 mm)

after the beginning of the shearing; then it dropped gradually towards the

residual value.

4. The peak interface friction angle δp and the residual interface friction angle δr,

for caissons in NC, LOC and sensitive clay samples in this research are shown

in Table 5.1.

Table 5.1 Interface friction angles between caisson and clay

Clay δp (°)

δr (°)

NC clay (OCR = 1, St = 2 - 2.8) 19.4 17.6

LOC clay (OCR = 1.5, St = 2 - 2.5) 18.9 18.1

Sensitive clay (St = 4 - 5) 14.6 11.7

Chapter 6 6-1 Axial Capacity of Caissons in Clay


6 AXIAL CAPACITY OF CAISSONS INSTALLED IN CLAY BY

JACKING AND BY SUCTION

6.1 INTRODUCTION

In this chapter, the penetration resistance and vertical pullout capacity of suction

caissons were studied by centrifuge tests, in normally consolidated (NC), lightly

overconsolidated (LOC) and sensitive clays. The caisson was installed either by

jacking, or by self-weight penetration followed by suction installation, in order to

investigate the influence of the method of installation on the behaviour of caissons.

During uplift, the caisson was pulled out either with an open lid or a sealed lid.

Variation of the axial capacity of caissons with time were investigated by extracting the

caisson either immediately after installation, or after a period of consolidation. A

description of the centrifuge facility, the geometry of the model caissons, and the

geotechnical properties of the kaolin clay samples used here have been presented in

Chapter 3.

The caisson was (quasi-rigidly) connected to an actuator that prevented horizontal

movement or rotation during installation and loading (see Figures 3.8 - 3.9), but allowed

vertical movement under either displacement or load control. Each caisson test was

performed along the centre line of the strong-box to avoid any bending effects due to

the horizontal component of the acceleration field (see Figure 6.1). In each sample box,

tests were undertaken according to the following alternative modes:

Test ‘OC’: the caisson was pulled out with the drainage valve open (O) after

consolidation (C) for 1 hour (model time, representing 1.7 years at prototype scale)

following installation.

Test ‘OI’ and ‘OC*’: the caisson was first pulled out with the drainage valve open (O)

immediately (I) after installation (Test OI). It was then re-installed at the same site, 1

hour of consolidation (model time) was allowed before it was finally pulled out

vertically (Test OC*).

Test ‘CC’: the caisson was pulled out with the drainage valve closed (C) after

consolidation (C) for 1 hour model time following installation.

Test ‘CI’: The caisson was pulled out with the drainage valve closed (C) immediately

(I) after installation.



Test ‘sus’: sustained loading was applied to the caisson after consolidation for 1 hour

(model time) at 120 g; then it was pulled out vertically with a sealed lid.

Test ‘cyc’: cyclic loading was applied to the caisson after consolidation for 1 hour

(model time) at 120 g; then it was pulled out vertically with a sealed lid.

The strength profile of the clay at the time of each caisson test was assessed from T-bar

penetration tests conducted at the beginning (A), middle (B), and end (C) of the testing

series in each box, as shown in Figure 6.1.

The pullout capacity of caissons was measured both immediately after installation and

after consolidation, and also for closed and open conditions. Unfortunately, in some of

the tests misrouting of the drainage lines led to cavitation, preventing a proper ‘closed’

condition being achieved, and reducing CC and CI capacities accordingly. The problem

did not occur in the earlier tests, and the routing of the drainage lines was later adjusted

to avoid this problem, with further tests on closed caissons undertaken.

All tests started with the caisson suspended just above the surface of the kaolin clay, but

entirely submerged in water. Installation of the caisson was then carried out either by

jacking (indicated by ‘J’ in the name) or by suction (indicated by ‘S’ in the name).

Jacked installation was undertaken by driving the actuator at a constant rate of 2 mm/s

with the force recorded by a 2 kN (or 3 kN) axial load cell and penetration of the

caisson monitored by a displacement transducer located on the actuator. Suction

installation was also performed using the actuator, however it first penetrated the

caisson until the jacking load reached the nominated caisson weight (in this case 16 N

in NC clay, representing a prototype submerged weight of 230 kN). Once this load was

reached, the actuator was programmed to maintain the load, and further installation was

achieved by extracting water using a computer-controlled syringe pump. Typically,

self-weight penetration occurred down to a penetration of 55 to 65 mm or

approximately half the final embedment.

The penetration rate of 2 mm/s was chosen in order to achieve undrained conditions at

the caisson tip, with non-dimensional velocity of V = vt/cv (where v is the actual rate, t

the wall thickness and cv the consolidation coefficient of around 0.1 mm2/s (House

et al., 2001)) of about 10. According to Randolph (2004), T-bar tests suggested that

drained and undrained limits for V are 0.1 and 10 respectively. Therefore, the caisson

installation was largely undrained. In early tests, there was a slight delay (some 100

seconds model time, representing over 2 weeks at prototype scale) in initiating the



suction installation, but later tests eliminated this delay, such that the whole installation

was virtually uninterrupted at a rate of ~2 mm/s. This gave a total installation time of

about 1 minute, which represents 1 week at prototype scale. Pullout tests were

conducted at a rate of 0.3 mm/s, thus V ~ 1.5 for open-ended tests, and is partially

drained. During sealed pullout, non-dimensionalising the velocity using the diameter,

d, instead of the wall thickness leads to V ~ 90, so undrained.

6.2 CYCLIC T-BAR TESTS FOR SENSITIVITIES OF CLAY

According to the α method in API RP2A (1993), the shaft friction during penetration of

open-ended piles in clay can be expressed as:

us sα ⋅=f (6.1)

where ‘fs’ is the shaft friction, α is the local shaft friction ratio, and su is the undrained

shear strength at the point of interest. Since us can be determined by in situ T-bar tests,

the shaft friction can be obtained once an α value is determined. Generally for piles in

clay, the friction during installation is taken as the remoulded shear strength of the clay

(API RP2A, 1993). Thus α may be obtained from the sensitivity of the clay, St, by:

tS1α = (6.2)

The sensitivity of the clay can be measured by cyclic T-bar tests as proposed by

Watson et al. (2000). Details of the T-bar penetrometer (see Figure 3.24) have been

described in Chapter 3. The T-bar has a projected area of 100 mm2, (5 mm diameter by

20 mm long) and was penetrated at 3 mm/s and extracted at 1 mm/s. Generally the

strength gradient of the NC sample increased gradually through the period of testing,

from ~1.0 kPa/m initially to ~1.3 kPa/m (using prototype depth) by the end of testing

(Figure 6.2). It can be seen in the graph that the undrained shear strength (su) varies

almost linearly with the depth of penetration, which indicates that the sample is

normally consolidated (NC clay). The average gradient of su with depth is 1.2 kPa/m.

During extraction of the T-bar, the deduced shear strength (using the same T-bar factor

of 10.5) is typically about 80% of the strength measured during penetration.

According to the cyclic T-bar tests carried out by Watson et al. (2000), the fully

remoulded T-bar resistance was just under 50% of the initial penetration resistance after

9 cycles of tests. The corresponding sensitivity of the NC kaolin clay they measured



was thus ~2, although their results suggested a trend of further strength reduction with

more cycling. Similar tests were performed in this research in order to investigate the

sensitivity of the soil samples used in these tests. The strength profiles of continuous

cycles of T-bar tests in the same site are shown in Figure 6.3a. The extraction speed

during cyclic T-bar tests was intentionally set to 3 mm/s after the first cycle, in both NC

and LOC clays, to avoid possible influence from consolidation.

There are several factors, such as the fast velocity of movement during extraction, and

difference in temperature of soil and water, that may result in an asymmetrical cyclic

T-bar profile in the final cycle. This was subsequently adjusted, by taking zero strength

as the mid-point between penetration and extraction loops on the final cycle of the test,

to avoid offsets in the T-bar readings. Taking test B12TB1 (see Figure 6.3a) for

example, the original undrained shear strength gradient is 1.16 kPa/m, while the

gradient after adjustment is 1.19 kPa/m, which is 3% larger than the original one.

Therefore, the original strength gradient of the T-bar tests, rather than the gradient

shown in Figure 6.3, is used later for calculating the axial capacity and radial stress

changes.

The resistance ratio, defined as the ratio of the penetration resistance for a given cycle

to the original resistance at certain depths, was used for evaluating the remoulded state

of the soil. Depths between 10 and 12 m below the mudline were chosen for analysis,

since this is the depth range where the caisson achieves the major part of its shaft

capacity.

Table 6.1 Variation of the resistance ratio with the number of cycles of T-bar test

in NC kaolin clay

Number of cycle B12TB1

(NC) 1 2 3 4 5 6 7 8 9

10 m 1.00 0.66 0.50 0.45 0.42 0.40 0.39 0.37 0.36

11 m 1.00 0.61 0.53 0.48 0.45 0.40 0.39 0.37 0.36

12 m 1.00 0.65 0.56 0.52 0.48 0.46 0.44 0.42 0.42

Variations of the resistance ratio with the number of cycles within the above depth

range of T-bar test B12TB1 are listed in Table 6.1, and profiles at various depths are



shown Figure 6.3b. It can be seen in this table and graph that during the last three

cycles the values of the resistance ratios are already very close, ranging from 0.36 to

0.44, with an average value of 0.39; the ultimate remoulded state was achieved after

8 - 9 cycles of penetration and extraction. The sensitivity index, St, defined as the ratio

of the undisturbed and remoulded undrained shear strength of the soil, is therefore

2.3 - 2.8. Considering the independent tests results of Watson et al. (2000) on similar

samples, St can be adopted as 2 - 2.8 for the caisson analysis in the NC kaolin clay.

One of the criticisms of using reconstituted kaolin for model testing is that the

sensitivity is considered lower than for typical natural clays (Andersen & Jostad, 2002).

However, the sensitivity measured here is close to typical values of 3 to 4 for

ultradeep water sediments in the Gulf of Mexico and West Africa shown by Andersen

& Jostad (2004). A possible explanation lies in the method used to measure the

sensitivity, since sample disturbance will tend to reduce the peak strength measured for

the intact soil, leading to an underestimate of the sensitivity. The in situ T-bar test will

also cause stress-softening during insertion (Einav & Randolph, 2005), so that the initial

penetration resistance does not reflect the full strength. However, while the sensitivity

may be underestimated by comparing initial and final penetration resistances, the

remoulded strength may be estimated directly from the final cycle resistance.

Application of the in situ T-bar test could be useful in deepwaters, where undisturbed

soil samples are extremely difficult to obtain (Lunne, 2001).

Cyclic T-bar tests were also performed in the LOC sample (OCR = 1.5). Variations of

the resistance ratio with cycles of T-bar penetration between 10 - 12 m for parallel tests

B13TB1 and B13TB2 are reported in Table 6.2. Profiles of these tests are presented in

Figure 6.4 and Figure 6.5. These plots were also adjusted from the original

measurements, using the same technique as that adopted in NC clay, to reach

symmetrical strengths between installation and pullout during the final cycle. These

graphs show that the strength profile of the LOC sample also increased linearly with

depth, for both the undisturbed samples and the remoulded sample. The gradients of the

original (unadjusted) undrained shear strength, dsu/dz, were 1.77 kPa/m and 1.76 kPa/m

for tests B13TB1 and B13TB2. These gradients were around 1.5 times those measured

in NC clay, indicating that the LOC sample was stronger than the NC sample at the

same depth. Data in Table 6.2 show that in the last three cycles, individual resistance

ratios varied in a narrow range between 0.40 and 0.46. Such a range was slightly higher

than that of 0.36 - 0.44 for NC clay, indicating a slightly smaller sensitivity for the LOC



clay. According to Equation 6.2, St can be obtained as 2.2 - 2.5 for the LOC kaolin clay

with an OCR of 1.5. For convenience, St was adopted as 2 - 2.5 for the LOC clay in

this study.


in LOC kaolin clay


(LOC) 1 2 3 4 5 6 7 8 9 10 11

10 m 1.00 0.67 0.59 0.55 0.52 0.51 0.50 0.48 0.45 0.44 0.42

11 m 1.00 0.64 0.56 0.52 0.51 0.50 0.48 0.47 0.46 0.45 0.43

12 m 1.00 0.65 0.57 0.53 0.52 0.51 0.49 0.48 0.46 0.45 0.44


(LOC) 1 2 3 4 5 6 7 8 9 10 11

10 m 1.00 0.69 0.62 0.55 0.53 0.49 0.47 0.45 0.44 0.42 0.42

11 m 1.00 0.67 0.61 0.55 0.52 0.49 0.47 0.46 0.44 0.43 0.42

12 m 1.00 0.68 0.62 0.56 0.53 0.51 0.48 0.46 0.45 0.43 0.40

Two cyclic T-bar tests were performed in the sensitive sample after the caisson tests

finished, with the penetrometer cycling between 8 - 12 m, as shown in Figures 6.6 - 6.7.

Except for a slight curvature observed during the early stage of penetration in the first

cycle, the T-bar resistance varied almost linearly with penetration depth. These graphs

were obtained from the original measurements following the same technique used in

NC clay to eliminate asymmetry. Analysis of the initial penetration resistance from

T-bar tests undertaken during the program of caisson tests showed that the strength

gradient experienced a substantial increase, reaching ~1.6 kPa/m at the end of the

experiments, from a measured value of ~1.2 kPa/m at the beginning of the tests. This

increase is considered to be caused by secondary consolidation, which appears more

marked for this clay mixture. The variations of the resistance ratio with cycles of T-bar

penetration at the depth of 11 m for parallel tests B14TC1 (see Figure 6.6b) and

B14TC2 (see Figure 6.7b) are listed in Table 6.3. Since the cycling distance was rather

short, the same speed was adopted both during installation and extraction. The average

resistance ratio was 0.22 after 12 cycles of penetration, suggesting a sensitivity of ~4.6.



Subsequently a sensitivity of 4 - 5 was adopted for the sensitive clay used in this

research.


in sensitive kaolin clay

Number of cycle B14TC1

(Sensitive) 1 2 3 4 5 7 8 9 10 11 12

10 m 1.00 0.48 0.37 0.32 0.28 0.26 0.25 0.24 0.23 0.22 0.20

10.5 m 1.00 0.48 0.39 0.33 0.31 0.28 0.27 0.25 0.24 0.23 0.22

11 m 1.00 0.48 0.38 0.33 0.30 0.28 0.26 0.25 0.24 0.23 0.22

Number of cycle B14TC2

(Sensitive) 1 2 3 4 5 7 8 9 10 11 12

10 m 1.00 0.46 0.35 0.30 0.28 0.26 0.25 0.25 0.24 0.23 0.22

11 m 1.00 0.47 0.37 0.32 0.29 0.27 0.25 0.24 0.23 0.22 0.21

12 m 1.00 0.48 0.38 0.34 0.31 0.29 0.28 0.26 0.25 0.25 0.23

6.3 FORMULAE FOR CALCULATING AXIAL CAPACITY

The net installation pressure (∆p) applied during jacked installation, self-weight

penetration and pullout is expressed as the ratio between the axial force P measured by

the load cell (set to zero before installation), and the gross cross-sectional area of the

caisson, Abase, according to:

baseAP∆p = (6.3)

For the suction installation period, the penetration pressure is calculated as:

( )

baseAplugAuuholdstressP

∆pio ⋅−+−= (6.4)

where Pstress-hold is the (nominal) self-weight of the caisson, uo and ui are the external and

internal pore pressures measured at the caisson lid respectively, and Aplug is the

(maximum) internal cross-sectional area of the caisson.

For caissons during installation, the penetration resistance can be expressed by Equation



2.5 in Chapter 2.

It is difficult to estimate the internal friction above the first internal stiffener, although

back-analysis of the axial capacity measured in NC clay revealed a difference of only

10% on the deduced α values during installation by varying the shaft friction in that

region from zero to full strength. Therefore, a nominal value of 0.5 kPa has been taken

for the (average) internal shaft friction above the upper edge of the first stiffener, τi-a,

for caissons in NC clay (see Figure 6.8c). For the caisson tests in LOC clay (OCR =

1.5), considering the differences in both the strength gradient and the geometry of the

internal stiffener, an average nominal shaft friction of 1 kPa was adopted for τi-a in the

analysis. For the sensitive clay, the same α values were taken both above and below

the internal stiffener (see Figure 6.8a), since the α value is small, as indicated by high St

values derived from cyclic T-bar tests.

6.4 PENETRATION RESISTANCE

6.4.1 Installation in NC Clay

Comparison was made between the penetration resistance of caissons installed by

jacking and by suction in the NC sample. The final depths of penetration for the two

types of installation were also compared.

6.4.1.1 Jacked installation

Jacked installation tests were performed by choosing ‘displacement control’ in the servo

motor of the actuator, the caisson was installed at velocity of 2 mm/s. The

corresponding normalised velocity (V = vt/cv, cv is the consolidation coefficient, 0.082

mm2/s, or 2.6 m2/year for NC kaolin clay) was thus 12.2, which is larger than 10, and

thus is undrained penetration according to Randolph (2004). Therefore, jacked

installation tests were all performed in an undrained state. During jacked installation,

the drainage valve in the caisson lid was kept open to ensure no excess pore pressure

built up inside the caisson. It was found to be essential to ensure purely vertical

movement of the caisson during penetration. Otherwise, readings from the axial load

cell became unreliable due to horizontal thrust on the caisson and thus bending

moments exerted on the load cell. The degree of verticality of the caisson can be

assessed by the readings of the two total pressure transducers (TPTs) that are located at



the same height on the caisson shaft. Any significant difference in the readings of the

two pressure cells would indicate tilt of the caisson, and the potential for bending

moments to occur. Replacing the earlier pinned connection on top of the caisson with a

more rigid connection helped to ensure verticality of the caisson during installation and

greatly improved the accuracy of measurements.

Table 6.4 Penetration rate (v)*, nominal depth of installation (Lnominal), penetration

resistance (∆p) and shaft friction ratio (α) during installation by jacking

and by suction in NC clay

Test γ'

(kN/m3)

dsu/dz

(kPa/m) Lnominal

(m)

v

(mm/s)* ∆p

(kPa)

α

B2JOI 6.93 1.17 13.92 2.00 125.5 0.41

B2JOC 6.93 1.17 14.05 2.00 105.1 0.38

B4JCI 6.76 1.13 14.02 2.00 127.0 0.40

B6JOC 6.94 1.27 13.87 2.00 115.2 0.40

B6JCC 6.94 1.00 14.38 2.00 115.3 0.39

B8JOC 6.59 1.17 14.10 2.00 124.1 0.38

Average (by jacking) 6.85 1.15 14.06 2.00 118.7 0.39

B3SCI 6.91 1.08 14.02 1.23 112.3 0.36

B3SOI 6.80 1.34 13.86 1.09 121.1 0.41

B3SCC 6.91 1.08 14.02 0.70 125.6 0.41

B9SOI 6.76 1.20 14.06 2.05 117.9 0.39

B10SOC 6.90 1.13 14.18 1.80 102.5 0.37

B10SCI 6.90 1.30 13.78 1.89 111.2 0.40

B10SCC 6.90 1.20 13.83 1.97 110.8 0.38

B11SOC 6.80 1.28 14.11 1.77 105.1 0.36

B12SCC 6.80 1.17 14.12 1.85 110.8 0.38

Average (by suction) 6.85 1.20 14.00 1.59 113.0 0.38

Average (all) 6.85 1.18 14.02 1.76 115.3 0.39

Note: ‘J’ for jacked installation and ‘S’ for suction installation, ‘*’ is shown in model scale.

The net penetration pressure (resistance) of the caisson measured during jacked



installation was calculated by Equation 6.3. The variation of this pressure and the

internal pore pressure with penetration depth of the caisson for six tests in NC clay are

shown in Figures 6.9 - 6.14 for tests B2JOI, B2JOC, B4JCC, B6JOC, B6JCC and

B8JOC, and compared in Figure 6.15. It can be seen that the net penetration pressures

were very close in value at the same depth, and increased almost proportionally to the

squared depth.

Internal pore pressures for test B4JCI and B6JCC were not recorded since signals from

the internal PPT were not received during the test. Judging from results in other tests,

the internal pore pressure increased almost linearly with penetration depth, and agreed

well with the hydrostatic pressure except at the end of penetration. The close match

over most of the penetration depth is reasonable since the drainage valve was vented

during the whole process of jacked installation. However, once the soil plug contacted

the PPT, the recorded internal pore pressure surged, and the corresponding depth of

installation is defined as the nominal depth of installation (Lnominal), as shown in Figures

6.9 - 6.14, and summarised in Table 6.4. During jacked installation, the average

nominal depth of the caisson was 14.06 m in prototype scale. It should be noted that

Lnominal is an estimate of the actual depth of installation (including any overdriving),

rather than for evaluating the soil heave. It was estimated from the internal PPT

readings, and many factors such as scour at the top of the soil plug suggested by Clukey

(2005) may affect the accuracy of measurement. In fact, later tests installed by suction

installation (see next section) revealed that Lnominal is much larger than the depth where

the soil plug first contacted the caisson top.

6.4.1.2 Suction installation

Suction installation following self-weight penetration is a more complex process

compared to jacked installation. The caisson was penetrated by the differential pressure

developed using suction (underpressure) across the caisson lid, instead of the jacking

force. The suction pressure (shown as positive for the underpressure in this research) is

calculated as the difference between the hydrostatic pressure and the measured pore

pressure inside the caisson at the point of interest.

During suction installation, the net penetration resistance was defined as the ratio of the

net penetration force (Pstress-hold plus the force due to the underpressure) and the

cross-sectional area, as shown in Equation 6.4. Variations of the measured internal pore



pressure with penetration depth of the caisson in nine tests in NC clay are shown in

Figures 6.16 - 6.24 for tests B3SCI, B3SOI, B3SCC, B9SOI, B10SOC, B10SCI,

B10SCC, B11SOC and B12SCC. Also shown in the graphs are the variation of the load

cell readings (expressed as pressure in terms of the cross-sectional area of the caisson),

internal pore pressure recorded by the internal PPT, syringe pump pressure and

hydrostatic pressure. During self-weight penetration, the axial load increased with the

depth of penetration, the force applied by the syringe pump was zero and the internal

PPT recorded hydrostatic pressure, since the drainage valve was vented. Suction

installation was initiated once the caisson had stopped or nearly stopped moving under

the preset stress-hold (simulating self-weight of the caisson in the field), the drainage

valve was then shut by applying high pressure to the piston of the drainage valve, then

the syringe pump was activated and an underpressure was created inside the caisson.

After suction installation started, the water pressure inside the caisson decreased almost

linearly with penetration depth, the stress-hold was maintained during suction

installation, and the syringe pump pressure was observed to increase almost linearly

with penetration depth, until the end of penetration (in the last 1 m) when the pressure

increased rapidly due to the difficulty of further movement of the caisson. When the

readings in the internal PPT increased suddenly, the syringe pump was stopped and the

drainage valve was vented.

Tests named ‘B3’ were carried out using the old control system, where a short stop

occurred when the system was switched from self-weight penetration to the subsequent

suction installation. Other tests were performed with the new control system where the

transformation between self-weight penetration and suction installation was continuous,

although at a reduced speed. It should be noted that in the new control system, the

stress-hold was set as 16 N (equivalent to ~23 kPa) during installation. Variations of

net penetration pressure with penetration depth derived from these nine tests are shown

in Figure 6.25.

Two typical plots of the embedment (indicated by the prototype depth of tip) of the

caisson versus time (in model scale) are shown in Figures 6.26 and 6.27. The

penetration rate (v, in model scale) of suction installation in the old system in test

B3SCI (Figure 6.26) was 1.23 mm/s, corresponding to a normalised velocity (V = vt/cv)

of 7.5, which is partly drained. By contrast in test B11SOC (Figure 6.27) installed by

the new system, the velocity was raised to 1.77 mm/s, which corresponds to a V of 10.8,

and thus undrained. It can be seen that in test B11SOC no stop existed within the



transition from jacking to suction installation, although the penetration rate decreased to

0.12 mm/s. Penetration rates (v, in model scale) during suction installation for various

tests in NC clay are summarised in Table 6.4. The average v is 1.59 mm/s, and V is

9.2, thus suction installation is essentially undrained.

The nominal depth (Lnominal, in prototype scale) where the readings of internal PPTs

surged during various suction installation tests (see Figures 6.17 - 6.24) is shown in

Table 6.4. It can be seen that the average Lnominal during nine tests was 14.00 m, which

is similar to that of 14.06 m measured during the jacked installation.

The penetration resistance during suction installation also increased quadratically with

the depth of installation (Figure 6.25). The largest resistance is found to be 120 kPa -

130 kPa at the final installation depth. The resistance measured during suction

installation is close to that measured during jacked installation, indicating that similar

radial effective stresses developed on the interface during these two types of

installation.

Comparison of the penetration resistance (expressed as a net pressure according to

Equation 6.4) versus depth of the caisson tip from typical tests is shown in Figure 6.28.

A total of 15 installation results are presented, with 6 jacked caissons and 9 caissons

installed using suction, from 9 separate soil samples.

The resistance profiles versus penetration depth during jacked installation (labelled with

a ‘J’) and suction installation (labelled with an ‘S’) are very similar (Figure 6.28).

When the caisson is fully installed, the resistance for both types of installation is around

5 - 6 times the self-weight ‘pressure’. At around 6 to 7 m depth of penetration, a step

change in resistance occurs for some of the early tests where suction installation was

used (in Box 3). In those tests, the change from jacked installation to suction

installation (once the nominal self-weight had been reached) took around 100 seconds

(equivalent to 2 weeks at prototype scale), allowing some dissipation of excess pore

pressure and resulting in an increase in resistance. Of interest is that even after such a

break, the penetration resistance seems to return gradually to the trend from the tests

using jacked installation, or using suction installation in tests where this delay was

avoided.

A more detailed comparison of the penetration resistance of caissons installed by

suction using the old system and the new system is shown in Figure 6.29. In this figure,

tests labelled with ‘B3’ were installed by suction using the old system, including a



pause during installation, while those labelled with ‘B10’ and ‘B11’ were installed

almost continuously using the new system. In the old system, the increase in

penetration resistance due to the time delay was ~80%, while in the new system,

variation in the resistance was very small. It is interesting to find that the penetration

resistances in the two cases converge rapidly for penetration depths greater than 10 m.

Therefore, it can be inferred that increases in resistance due to consolidation may be

reduced by further penetration.

Predictions of the necessary underpressure (∆un), allowable underpressure (∆ua), actual

applied underpressure (∆uapp), soil heave (hs,pre) inside the caisson, and actual factor of

safety (Fs) during suction installation are considered to be important issues in design

(Andersen & Jostad, 1999; Ehlers et al., 2004). Methods for calculating ∆un, ∆ua are

shown in Equations 2.6 and 2.7, with the reverse end-bearing capacity factor Nc of the

soil plug adopted as 9 (Ehlers et al., 2004). The predicted soil heave is calculated by

assuming that the displaced soil at the caisson tip moves 50% inside and 50% outside

during self-weight penetration, and 100% inside the caisson during suction installation,

while the soil displaced by the internal stiffeners moves 100% inside the caisson

(Andersen & Jostad, 1999). The actual Fs is calculated by Equation 2.9, while the

actual soil heave hs,act (in prototype scale) is the distance between the overall length

(14.4 m) of the caisson, and the penetration depth where the applied underpressure

sharply increases, as the soil plug contacts the lid of the caisson.

The predicted ∆un, ∆ua, and hs,pre, and actual applied underpressure (∆uapp) and actual Fs

for a typical test (B11SOC) during suction installation in NC clay are shown in Figures

6.30a - d. It should be noted that the actual applied underpressure was supplied by the

syringe pump. Also shown in Figure 6.30b is ∆un, for the purpose of comparison. It

can be seen that at 4.95 m when the syringe pump was initiated, the applied

underpressure was 21.09 kPa, which is less than the allowable underpressure of 47.3

kPa, and plug failure did not occur at that moment. The applied underpressure lay

between the necessary and allowable underpressures for most of the subsequent suction

installation (see Figure 6.30b). This agrees with the field measurements of suction

installation (see Figure 2.8a) reported by Newlin (2003b). After the installation depth of

13.36 m, the applied underpressure increased abruptly (see Figure 6.30b), indicating

that the soil plug contacted the caisson lid. The corresponding actual Fs was 2.06 (see

Figure 6.30c), which is larger than the lower limit of 1.5 proposed by Andersen &



Jostad (1999), showing that the soil plug did not fail at that moment. However, contact

of the soil plug with the caisson lid blocked the outlet to the syringe pump and

subsequently caused the penetration rate to reduce to 0.36 mm/s at 13.90 m (see Figure

6.27), after which the caisson moved at a very slow speed under the exerted stress-hold.

The depth where the soil plug touched the caisson lid, 13.36 m, indicates an actual soil

heave length (hs,act) of 1.04 m, which is clearly lower than the predicted soil heave of

1.43 m (see Figure 6.30d). The corresponding factor of safety at that moment is defined

as Fs,plug.

A summary of the depth where suction installation started (zs), the actual depth of

penetration where the applied underpressure surged as the soil plug contacted the

caisson lid (zplug), the corresponding actual soil plug heave (hs,act) inside the caisson and

the factor of safety (Fs,plug), the final depth of installation (zfinal) where the syringe pump

was stopped, and the predicted soil heave (hs,pre) at the end of penetration for various

suction installation tests in NC clay are presented in Table 6.5. It should be noted that

only those tests for which the readings of syringe pump pressure are both available and

normal (see Figures 6.21 - 6.24) are presented.

Table 6.5 Values of zs, zplug, hs,act, Fs,plug, zfinal and hs,pre during suction installation

in NC clay (units in prototype scale)

Test zs

(m)

zplug

(m)

hs,act

(m)

Fs,plug zfinal

(m)

hs,pre

(m)

B10SCI 6.96 13.53 0.87 2.10 13.90 1.33

B10SCC 6.19 13.59 0.81 1.73 13.93 1.36

B11SOC 4.87 13.36 1.04 2.06 14.30 1.43

B12SCC 5.62 13.36 1.04 1.99 14.39 1.41

Average 5.91 13.46 0.94 1.97 14.13 1.38

It can be seen in Table 6.5 that the average actual soil heave of 0.94 m (in prototype

scale, or 4.25 mm in model scale) is obviously lower than the predicted value of 1.38 m,

which is based on the assumption that all the soil particles displaced by the caisson tip

move inward under suction. The corresponding volume (in model scale) of actual soil

heave is 5172 mm3. By subtracting the volume of the internal stiffeners (1310 mm3)



and that of epoxy and wires (1549 mm3) from the plug heave, the remaining volume of

soil heave (2313 mm3) amounts to only 45% of the volume displaced by the caisson

wall (5198 mm3) at that depth. This ratio is obviously smaller than that (100%)

assumed in the NGI method (Andersen & Jostad, 2002).

During caisson installation, the annular tip area is very small compared to the

cross-sectional area of the caisson. Randolph & House (2002) adopted a bearing

capacity factor, Nc, of 7.5, corresponding to deep bearing of a strip foundation

(Skempton, 1951). For the caisson used here, varying Nc between 7 and 12 for a fixed

shaft friction ratio (α) of 0.38 only leads to a difference of 10% in the total penetration

resistance (see Figure 6.31). Therefore, it is sufficient to adopt Nc = 7.5 to calculate the

tip resistance of the caisson during installation. Values of α back-figured from

Equation 2.5, using Nc = 7.5 for the tip resistance, during the various caisson installation

tests mentioned above are shown in Table 6.4.

It can be seen in Table 6.4 that when the caisson is installed in clay with similar

gradients of shear strength with depth, the values of α are very similar for caissons

installed by jacking and by suction. An average value of 0.39 is obtained for jacked

installation and 0.38 for suction installation. In general, the deduced values of α fall in

the range of 0.30 to 0.45 for NC clay, with an average value of 0.38. A typical

simulation of the measured penetration resistance of caissons is plotted in Figure 6.28,

adopting Nc = 7.5 and α = 0.39. It can be seen clearly that satisfactory agreement was

achieved for both the shape and magnitudes of the penetration resistance.

The average α value derived from the measurements during caisson installation,

however, is in reasonable agreement with the residual (or fully remoulded) strength

ratio measured by cyclic T-bar tests described in Table 6.1. The sensitivity (St) of the

NC kaolin clay was derived as 2 - 2.8 from the cyclic T-bar tests. α can be taken as 1/St

according to API RP2A (1993), giving an α of 0.36 - 0.5, which is only slightly larger

than 0.30 - 0.45 back-figured from the penetration resistance of the caisson as shown in

Table 6.4.

6.4.1.3 Re-installation in disturbed sites

In the field it is possible that a caisson may need to be re-installed in the same site to

rectify excessive inclination or other problems occurring during installation (Ehlers



et al., 2004). To investigate the disturbance from previous installation on the axial

capacity of the suction caissons, in some tests the caisson was re-installed at the same

site immediately after unsealed pullout; such tests were denoted by a ‘*’ in the name,

for example, test ‘B2JOC*’ was jacked in the same site as test ‘B2JOI’ which was

pulled out with a vented lid immediately after installation. Re-installation tests were

carried out either by jacking, such as tests B2JOC*, B5JOC* and B6JOC*; or by

suction, such as tests B3SOC* and B10SOC*. The nominal depth of installation and

the corresponding net penetration pressure are listed in detail in Table 6.6. Since

over-driving occurred in some early tests installed by jacking (i.e. caisson was installed

to a depth more than 14.4 m by applying excessive jacking force).

Table 6.6 Shaft friction ratio α during re-installation by jacking and by suction in

disturbed sites in NC clay

Test γ'

(kN/m3)

dsu/dz

(kPa/m) Lnominal

(m) ∆p

(kPa)

α

B2JOC* 6.93 1.17 14.27 128.2 0.38

B5JOC* 6.59 1.02 14.18 106.0 0.34

B6JOC* 6.59 1.07 13.87 118.6 0.38

Average (by jacking) 6.70 1.09 14.11 117.6 0.35

B3SOC* 6.80 1.36 14.20 120.8 0.38

B10SOC* 6.90 1.25 13.76 102.3 0.34

Average (by suction) 6.85 1.31 13.98 111.6 0.36

Average (all) 6.76 1.18 14.06 115.2 0.36

Note: ‘J’ for jacked installation and ‘S’ for suction installation, ‘*’ for re-installation in the same site.

A comparison of the resistance profiles for five re-installation tests is shown in Figure

6.32. The α values back-figured from these re-installation tests are shown in Table 6.6,

from which an average α value of 0.36 can be derived. The average measured α value

was around 10% smaller than the average α value measured for the undisturbed sites.

Such a decrease, however, is not very large and is assumed to result from further

gradual softening of the soil due to remoulding, partially compensated by some

dissipation of excess pore pressure and resulting strength recovery in the soil. A



comparison between the typical penetration resistances of caissons during the original

installation (signified by ‘OI’) and the reinstallation (signified by ‘OC*’) for both

jacked installation and suction installation is shown in Figure 6.33.

For re-installation tests in the disturbed sites, the predicted ∆un, ∆ua, hs,pre, actual applied

underpressure, ∆uapp, and actual Fs versus penetration depth of the caisson for a typical

test (B10SOC*) are shown in Figure 6.34. It can be seen that the soil plug contacted the

caisson lid at 11.73 m, which is smaller than that observed in undisturbed sites. This

reveals the danger of early plug heave for re-installation in the disturbed sites.

6.4.1.4 Summary

From the above analysis, it can be concluded that despite the difference in the

installation processes, there is no essential difference between the penetration resistance

of caissons installed by suction and by jacking in NC clay. The model used in this

thesis for predicting the penetration resistance during caisson installation gives

satisfactory prediction in NC clay, independent of installation methods. Adopting a

bearing capacity factor of Nc = 7.5, the interface friction factor is found to be in the

range of 0.30 to 0.45, with an average value of 0.38, for both jacked and suction

installation. Re-installation at the same site decreases the axial capacity by around 10%

for either installation process, and is likely to result in earlier plug heave compared to

tests in intact sites.

6.4.2 Installation in LOC Clay

Caisson tests undertaken in LOC clay all used model caisson 2 (see Figure 3.7), in

centrifuge test Box 13. Four tests labelled B13JCC, B13SCC, B13sus and B13cyc were

undertaken. Test B13JCC was installed by jacking, and the other three by suction.

Comparisons of the penetration resistance and penetration depth between these two

types of installation were subsequently made. In test B13JCC, jacked installation was

performed at a rate of 2 mm/s, and variations of the penetration resistance and internal

pore pressure with penetration depth are shown in Figure 6.35. The penetration

resistance also increased smoothly with depth, until reaching ~160 kPa at the end (~14

m) of penetration. At the end of penetration the internal pore pressure increased

suddenly as the PPT made contact with the soil surface. The nominal depth of

installation was thus 13.89 m, with a corresponding net penetration pressure of 157.8



kPa, as shown in Table 6.7. The caisson was installed to a final depth (zfinal) of 13.99 m

due to inertia (Figure 6.35).

For suction installation in LOC clay, a similar installation procedure was adopted as that

in NC clay, except that a larger stress-hold value of 30 N (representing a prototype

submerged weight of 431 kN) was used, in order to achieve sufficient embedment of the

caisson during self-weight penetration such that suction could be developed when the

syringe pump was started.

Table 6.7 Installation velocity (v), nominal depth of installation (Lnominal),

maximum installation pressure (∆p) and shaft friction ratio (α) during

installation by jacking and by suction in LOC clay (OCR = 1.5)

Test γ'

(kN/m3)

dsu/dz

(kPa/m) v

(mm/s)*

Lnominal

(m) ∆p

(kPa)

α

B13JCC (by jacking) 7.15 1.64 2.00 13.89 157.8 0.42

B13SCC (by suction) 7.15 1.64 1.80 13.87 165.0 0.43

B13sus (by suction) 7.18 1.76 1.85 13.64 154.8 0.38

B13cyc (by suction) 7.21 1.77 1.80 13.50 163.7 0.40

Average (by suction) 7.18 1.72 1.82 13.67 161.2 0.40

Average (all) 7.17 1.70 1.86 13.73 160.3 0.41

Note: ‘J’ for jacked installation and ‘S’ for suction installation, ‘sus’ for sustained loading, ‘cyc’ for cyclic loading, ‘*’ is in model scale.

Altogether three suction installation tests were performed in LOC clay. Variations of

the measured internal pore pressure (by the PPT), the load cell (expressed as pressure),

and the syringe pump pressure with penetration depth of the caisson are shown in

Figures 6.36 - 6.38 for tests B13SCC, B13sus and B13cyc. Variations of the caisson

embedment (expressed as the prototype depth of tip) versus model time during suction

installation for these three tests are shown in Figure 6.39 - 6.41. During self-weight

penetration, the caisson was installed to prototype depths of 6.97 m, 7.60 m and 7.69 m

respectively for tests B13SCC, B13sus and B13cyc. The average self-weight

penetration depth of 7.4 m is equivalent to around 50% of the overall length of the

caisson. After a short time delay, suction installation started at 7.31 m, 7.77 m and

7.85 m respectively for these three tests, at a penetration rate (in model scale) of 1.80,



1.90 and 1.89 mm/s, respectively. The average penetration rate of 1.86 mm/s

corresponds to a normalised velocity (V = vt/cv, cv = 0.076 mm2/s) of 12.3, which

means penetration was undrained in these suction installation tests.

Tests B13JCC and B13SCC were carried out in sites with similar strength gradients,

therefore comparison between these two tests is appropriate. The nominal depth of

installation for test B13SCC is 13.87 m, and the corresponding net penetration pressure

was 165.0 kPa. This is very close to the measured penetration resistance of 157.8 kPa

(at 13.89 m) of test B13JCC, which was installed by jacking in a site with similar soil

strength.

After turning around the strong-box, the nominal depths achieved in tests B13sus and

B13cyc were 13.64 m and 13.50 m; both these values were less than those in tests

B13SCC and B13JCC. Presumably there is more likelihood of a free-standing soil

column (at least for a given distance) in these tests, which could account for the lower

Lnominal values, since the strength gradient for the LOC clay was 40 - 50% larger than

that of the NC clay.

Variations of the necessary underpressure (∆un), allowable underpressure (∆ua), actual

applied underpressure (∆uapp), actual factor of safety (Fs) and predicted soil heave (hs)

versus depth of installation for tests B13SCC, B13sus and B13cyc are shown in Figures

6.42 - 6.44. It can be seen that when suction installation started, the applied

underpressure was below the allowable value, with the factor of safety larger than 2,

showing that no plug failure occurred at that time.

A summary of the depth where suction installation started (zs), the actual depth of

penetration where the applied underpressure surged as the soil plug contacted the

caisson lid (zplug), the corresponding actual soil plug heave (hs,act) inside the caisson and

the factor of safety (Fs,plug), the final depth of installation (zfinal), and the predicted soil

heave (hs,pre) at the end of penetration during each suction installation test (also see

Figures 6.36 - 6.38) in LOC clay are summarised in Table 6.8.

After suction installation started at 7.31 m in test B13SCC (see Figures 6.42, 6.36), the

applied underpressure was just above the necessary underpressure, and was well below

the allowable underpressure for most of the penetration depth, until after the depth of

13.75 m. The syringe pump pressure suddenly increased at 13.62 m as the soil plug

contacted the caisson lid. The corresponding actual Fs was 2.26, indicating that plug

failure did not occur then. The suction installation essentially stopped at 13.72 m,



where the penetration rate recuced to 0.18 mm/s (see Figure 6.39).

The depth where the soil plug made contact with the caisson lid in test B13sus was

13.52 m (see Figure 6.43b). While in test B13cyc this depth was 12.12 m (see Figure

6.44), which is smaller than in tests B13SCC and B13sus. This difference arises

because test B13cyc was carried out at the end of the test sequence in that box, and

installation became very difficult as the soil became rather stiff at that time; the

excessive applied underpressure caused an earlier plug failure in test B13cyc. Results

in this test are thus not considered in the average value of Table 6.8.

Table 6.8 Values of zs, zplug, hs,act, Fs,plug, zfinal, and hs,pre during suction installation

in LOC clay

Test zs

(m)

zplug

(m)

hs,act

(m)

Fs,plug zfinal

(m)

hs,pre

(m)

B13SCC 7.31 13.62 0.78 2.26 13.90 1.15

B13sus 7.77 13.52 0.88 3.33 13.72 1.12

*B13cyc 7.85 12.12 2.28 2.15 13.53 1.10

Average 7.54 13.57 0.83 2.80 13.81 1.12

Note: ‘*’ is abnormal, not considered in the average value.

Judging from the results of tests B13SCC and B13sus shown in Table 6.8, the average

depth where the soil plug touched the caisson lid was 13.57 m. The corresponding

average actual Fs,plug was ~2.80, thus no plug failure occurred then. The corresponding

actual soil heave can be derived as 0.83 m (in prototype scale, or 6.92 mm in model

scale), which is obviously less than the predicted value of 1.12 m. The volume (in

model scale) of actual soil heave was 4571 mm3. If the volume of the internal ring

stiffeners (1395 mm3) and the epoxy and wires (963 mm3) inside the caisson are

subtracted from the soil heave, the remaining volume (2213 mm3) is only 42% of the

volume that was occupied by the caisson wall (5240 mm3, ring stiffener not included).

This proportion is consistent with that of 45% for tests in NC clay.

A comparison of the penetration resistance for the above four tests in LOC clay is

shown in Figure 6.45, while the corresponding T-bar test results are shown in

Figure 6.46. The α values back-figured from each installation test by Equation 2.5 are



shown in Table 6.7. It can be seen in the table that the derived α during installation

varied between 0.38 - 0.43. The average α was 0.42 during jacked installation, and 0.40

during suction installation, with a small difference of 5% only between those two types

of installation. As a result, it can be inferred that there is no significant difference in

penetration resistance and nominal depth of installation between the caissons installed

by jacking and by suction in LOC clay. It should be noted that varying the shaft friction

above the first stiffener from zero to full strength leads to 12% change in the

back-figured α value.

The average α value of 0.41 measured during installation of the caissons in LOC clay,

however, is in good agreement with the sensitivity derived from the cyclic T-bar tests,

which gave a St of 2 - 2.5 for the clay with an OCR of 1.5 and thus indicated an α value

of 0.4 - 0.5, according to Equation 5.2. It is interesting to find that the α value in LOC

clay was just slightly higher than that for the NC clay (see Table 6.4), although the net

penetration resistance is much larger for the former. Thus, even though the α values for

the NC soil were slightly greater than 1/St, the trend of decreasing St and increasing α is

consistent between the NC and LOC tests.

6.4.3 Installation in Sensitive Clay

Caisson tests performed in sensitive clay all used model caisson 2 (see Figure 3.7), in

centrifuge box 14. Four tests labelled B14SCC, B14sus, B14cyc and B14susa were

conducted. Due to the sensitivity of the sample (as indicated by the failure to core the

sample after consolidation), a smaller stress-hold of 8 - 12 N (representing a prototype

submerged weight of 115 - 173 kN) was adopted in these tests, to prevent excessive

settlement of the caisson after installation. The net penetration resistance during these

four suction installation tests was calculated by Equation 6.4.

Variations of the measured net penetration resistance, the internal pore pressure

(measured by the PPT), the load cell pressure (expressed as pressure relative to the

cross-sectional area of the caisson), and the syringe pump pressure with penetration

depth of the caisson are shown in Figures 6.47 - 6.50 for tests B14SCC, B14sus,

B14cyc and B14susa. Variations of the caisson embedment (expressed as the prototype

depth of tip) versus model time during suction installation for these four tests are shown

in Figure 6.51 - 6.54. Under self-weight (jacked) penetration, the caissons were



installed to depths (in prototype scale) of 8.90 m, 9.27 m, 7.21 m and 6.34 m

respectively during the above four tests, with an average depth of 7.93 m. The average

vertical range of the suction-affected area after full installation was thus around 1.67 m,

thus giving enough depth for observing the radial stress changes in that region. The

penetration rates (in model scale) during suction installation for these four tests were

1.83, 1.79, 1.89 and 1.34 mm/s. The average installation velocity of 1.65 mm/s

corresponds to a normalised velocity (V = vt/cv, cv = 0.063 mm2/s) of 10.1, thus

undrained.

Variation of the necessary underpressure (∆un), allowable underpressure (∆ua), actual

applied underpressure (∆uapp), actual safety factor (Fs) and predicted soil heave (hs,pre)

with depth of installation for tests B14SCC, B14sus, B14cyc and B14susa are shown in

Figures 6.55 - 6.58. It can be seen in the graphs that the applied underpressure when

suction installation started was well below the allowable value; the corresponding Fs

was larger than 2, thus no plug failure occurred.

A summary of zs, zplug, hs,act, Fs,plug, zfinal, and hs,pre for various suction installation tests

(also see Figures 6.47 - 6.50) in sensitive clay are presented in Table 6.9. It should be

noted that due to insufficient experience on performing caisson installation tests in

sensitive clay, the caisson continued to settle obviously after the syringe pump stopped

working, under the applied axial force (stress-hold). The magnitude of the input

stress-hold was subsequently decreased to as low as 8 N in the last two tests (while the

actual value recorded by the load cell during penetration showed some variation from

this target value), to achieve a stable state of the caisson. The final depth of installation

(zfinal) of the caisson was thus larger than 14.4 m.

Table 6.9 Values of zs, zplug, hs,act, Fs,plug, zfinal and hs,pre during suction installation

in sensitive clay (units in prototype scale)

Test zs

(m)

zplug

(m)

hs,act

(m)

Fs,plug zfinal

(m)

hs,pre

(m)

B14SCC 9.12 13.62 0.78 2.69 15.30 1.31

B14sus 9.40 13.63 0.77 4.33 14.59 1.14

B14cyc 7.45 13.44 0.96 4.61 15.24 1.20

B14susa 6.74 13.45 0.95 3.17 14.77 1.25

Average 8.18 13.54 0.86 3.70 14.98 1.23



It can be seen in Table 6.9 that the average depth where the soil plug contacted the

caisson lid is 13.54 m, which corresponds to a soil heave of 0.86 m (or 7.17 mm in

model scale). This actual soil heave is far less than the average predicted value of 1.23

m. The corresponding volume of actual soil heave (in model scale) is 4736 mm3. By

subtracting the volume of the ring stiffener (1395 mm3) and that of the epoxy and wires

for that depth (959 mm3), the remaining volume amounts to 46% of the volume

displaced by the caisson wall (5227 mm3) during installation. This proportion is

consistent with those obtained from tests in NC and LOC clays.

The nominal depth of installation where readings of the internal PPTs surged, Lnominal,

for tests B14SCC, B14sus, B14cyc and B14susa are shown in Figures 6.47 - 6.50 and

Table 6.10. The average Lnominal value was 14.32 m for these four tests in sensitive clay.

Compared to the installation tests in the NC and LOC clays, the depths of nominal

installation in the sensitive sample were higher, indicating an easier installation process

in sensitive clay.

Profiles of the penetration resistance during these four tests in sensitive clay are shown

in Figure 6.59, while the profiles of undrained shear strength obtained from T-bar tests

corresponding to each test stated above are shown in Figure 6.60. The plot shows that

the strength gradient of the soil increased from 1.16 kPa/m when the first caisson test

was carried out, to 1.58 kPa/m when all tests were finished. The penetration resistance

profiles were similar in shape, although the values at a given depth were much larger for

later tests than for early tests. The penetration resistance of the soil was around 50 - 80

kPa at the depth of ~14 m (see Figure 6.59), lower than those measured in NC clay

(around 120 kPa) and in LOC clay (around 150 kPa), indicating that much smaller shaft

friction developed on the caisson in sensitive clay, compared to that in the soil with

lower sensitivity. The α values back-figured from each installation test by using

Equation 2.5 are shown in Table 6.10.

For the calculation in sensitive clay, an identical α value was adopted on both sides of

the caisson, according to the theoretical model suggested in this chapter. The table

shows that despite some differences in the penetration resistance at certain depths, the

derived α values were very close for the above four tests, varying between 0.15 and

0.18 with an average value of 0.16. This α value is just below the α value of 0.20 - 0.25

derived from the cyclic T-bar tests in sensitive clay (St = 4 - 5). This confirmed that the

model for calculating the α values during caisson installation (see section 6.3) is also



valid in sensitive clay.

Table 6.10 Installation velocity (v), nominal depth of installation (Lnominal),

maximum installation pressure (∆p) and shaft friction ratio (α)

during installation by jacking and by suction in sensitive clay

Test γ'

(kN/m3)

dsu/dz

(kPa/m)

v

(mm/s)

Lnominal (m)

∆p

(kPa)

α

B14SCC 7.30 1.16 1.67 14.40 60.5 0.16

B14sus 7.30 1.33 1.75 14.33 65.5 0.15

B14cyc 7.30 1.45 1.89 14.21 73.4 0.15

B14susa 7.30 1.58 1.30 14.28 67.9 0.18

Average 7.30 1.38 1.65 14.32 65.9 0.16

6.5 AXIAL CAPACITY DURING PULLOUT

In NC clay, suction caissons were pulled out vertically either vented or sealed following

installation by jacking or by suction, while in the LOC and sensitive clays all caissons

were pulled out with a sealed lid. The pullout tests were carried out either immediately

after installation (indicated by ‘I’ in the name of the test) or after consolidation

(indicated by the second ‘C’ in the name) for 1 hour at 120 g, which is equivalent to 1.7

years prototype time. The velocity during the pullout was chosen to be 0.3 mm/s which

corresponds to a non-dimensional velocity (V = vt/cv, t is the thickness of the caisson)

of 2.5 for the unsealed pullout, and 150 (V = vd/cv, d is the diameter of the caisson) for

the sealed pullout in NC clay. During sealed pullout tests in the LOC and sensitive

clay, values of V can be obtained as 118 and 143, respectively, thus undrained.

6.5.1 Unsealed Pullout in NC Clay

Variations of the unsealed uplift capacity and the internal pore pressure with the upward

movement of the caisson during immediate pullout in NC clay are shown in Figures

6.61 - 6.63, for tests B2JOI, B8JOI and B9SOI; the first two tests were installed by

jacking and the last by suction. Comparison of these three tests is shown in Figure 6.64.

The uplift capacity during unsealed pullout immediately after installation was found to

be very close for caissons installed by jacking and by suction, being around 100 kPa at



the full embedment of ~14 m for all three tests. It can be inferred that the underpressure

applied during suction installation had no obvious influence on the caisson capacity

during immediate pullout. In order to avoid influence from the soil strength and depth

of embedment when evaluating the capacity of the caisson, a normalised uplift capacity,

usp∆− , defined as the ratio of the uplift capacity (–∆p) to the average shear strength

( us ) over the maximum caisson embedment (Lmax), was used. It should be noted that

Lmax is generally larger than Lnominal used previously, after allowing for the settlement

during consolidation. The normalised uplift capacities of the above unsealed pullout

tests are listed in Table 6.11. The average normalised uplift capacity is 13.4.

Table 6.11 Uplift capacity (∆p) and shaft friction ratio (α) during unsealed pullout

of the caisson in NC clay (assuming αexternal = αinternal and Nc = 7.5)

Test Lmax

(m)

dsu/dz

(kPa/m)

su, tip

(kPa) ∆p

(kPa)

us∆p

− α

B2JOI 14.18 1.17 16.59 –101.9 12.3 0.38

B8JOI 13.97 1.10 15.37 –110.1 14.3 0.45

B9SOI 14.14 1.20 16.97 –116.0 13.7 0.40 OI

Average 14.10 1.16 16.31 –109.3 13.4 0.43

B2JOC 14.45 1.17 16.91 –171.2 20.3 0.65

B6JOC 14.05 1.07 15.03 –163.7 21.8 0.69

B8JOC 14.25 1.17 16.67 –167.9 20.1 0.73

B11SOC 14.32 1.28 18.41 –186.7 20.4 0.70

B12SOC 14.50 1.31 19.00 –192.8 20.3 0.68

OC

Average 14.33 1.20 17.20 –169.7 20.6 0.69

B2JOC* 14.54 1.23 17.88 –164.7 18.4 0.58

B6JOC* 13.95 1.17 16.32 –150.8 18.5 0.57

B3SOC* 14.24 1.34 19.08 –170.4 17.9 0.60

B10SOC* 13.80 1.25 17.25 –160.0 18.6 0.58

OC*

Average 14.13 1.21 17.07 –162.0 18.4 0.59

(Note: ‘*’ means re-installation in disturbed sites)



Uplift capacities from various unsealed pullout tests in NC clay after 1 hour of

consolidation at 120 g (equivalent to 1.7 years prototype time) are shown in Figures

6.65 - 6.69, for tests B2JOC, B6JOC, B8JOC, B11SOC and B12SOC. The first three

tests were installed by jacking, and the last two by suction. The unsealed uplift capacity

after consolidation was very close for both jacked installation and suction installation,

in addition to the similarity of the penetration resistance during installation (Figure

6.70). The undrained shear strength gradients (dsu/dz) of the NC clay in these unsealed

pullout tests ranged between 1.02 kPa/m and 1.30 kPa/m, with an average value of 1.17

kPa/m (see Figure 6.71). Some curvature appeared at the deeper part of the strength

profiles; therefore, nonlinear equations instead of linear equations were used in

analysing the corresponding capacity, so as to achieve the best simulation of the

strength. The normalised uplift capacity, usp∆− , during the unsealed pullout tests

after consolidation is shown in Table 6.11, and the average value is 20.6. Compared to

that during immediate pullout, the normalised capacity during consolidated tests

showed an increase of 54%.

Figure 6.72 shows four typical unsealed pullout tests, two of which were pulled out

after consolidation and two immediately after installation. For the two pullout tests

after consolidation, one was installed by suction (test B11SOC) and the other by jacking

(test B2JOC). The uplift capacity profiles were similar in both shape and magnitude.

At the very beginning of the two pullout tests, a ‘pseudo’ capacity developed, but it

suddenly dropped to a value which was 20 - 30% lower than the starting value. Such a

pseudo value, however, is considered to be the result of the suction developed between

the soil plug and the caisson lid, which prevented the caisson from moving away from

the soil plug, although it could not be maintained since the drainage valve was open. As

a result, the suction force was lost immediately after a small displacement. The

measured capacity after this abrupt decrease is considered to be the real pullout

capacity, being –186.7 kPa for test B11SOC and –171.2 kPa for test B2JOC.

The capacity following suction installation was around 7% larger than jacked

installation because the strength of the clay sample used for the suction installation test

was around 9% higher (with dsu/dz being 1.28 kPa/m for the suction installation site

versus 1.17 kPa/m for the jacked installation site). The normalised capacity was 20.4

and 20.3 for the above two tests, very similar for the two types of installation. Also

plotted in Figure 6.72 are two immediate pullout tests B9SOI (installed by suction) and

test B2JOI (installed by jacking), which have been reported previously, to investigate



the effect of ‘set-up’ on the pullout capacity of the caissons. The uplift capacities after

consolidation were larger than those without consolidation, which is attributed to the

increase in the radial effective stress around the caisson after dissipation of the excess

pore pressure, accompanied by recovery in the soil strength adjacent to the caisson body

after the severe disturbance.

Details of the uplift capacity in the early stage of these four unsealed pullout tests are

shown in Figure 6.73. It is interesting to notice from the graph that the failure

mechanisms are quite different for caissons undergoing immediate pullout and pullout

after consolidation. During immediate pullout, it took around 0.4 - 0.7 m (0.11 - 0.19d)

of displacement for the ultimate capacity to be developed, while for the pullout after

consolidation, ‘peak’ capacity was reached in an extremely short distance (around 0.1

m) of shearing, and then decreased gradually with further pullout. The difference in the

shearing modes between the immediate pullout and pullout after consolidation could be

attributed to the change in the structure of the soil surrounding the caisson after

consolidation: a type of ‘brittle’ failure occurred for pullout tests after consolidation,

compared to a ‘ductile’ type for immediate pullout.

House et al. (1999) suggested that different α values should be adopted for the external

wall and internal wall of the caisson during unsealed pullout, with α = 0.5 - 0.7 for the

external wall and α = 0.2 - 0.3 for the internal wall. However, their assumption was

based on a 1 g model test and may have been affected by details of their experimental

arrangement. The influence of the internal stiffener on the shaft friction ratio is difficult

to assess, especially after consolidation. Therefore, it is convenient to assume that the α

values are the same on both sides of the caisson during unsealed pullout. During

pullout, the axial capacity can be expressed as:

( ) ( )intextu

n

1itipiuicibase AAsα AγzsNA∆p ++′−=⋅− ∑

=

(6.5)

where the meaning of each symbol has been discussed in Equation 2.5, and Aint is the

area of the internal caisson shaft; ‘–’ is used before ∆p since the tensile pressure is

defined as negative. During unsealed pullout, a reverse end-bearing capacity factor, Nc,

of 7.5 has been adopted, since the tip area of the caisson is so small that the influence

from different Nc values would be minor (similar to the analysis during the caisson

penetration).



By assuming same α values on both sides of the caisson wall, the α values derived from

the measured capacity in NC clay for each pullout test after consolidation (named with

‘OC’) are summarised in Table 6.11, for three of the tests installed by suction and two

by jacking. The average α value during pullout was 0.69 for caissons installed by

jacking and 0.70 for caissons installed by suction. No discernible difference is apparent

in the shaft friction ratios during unsealed pullout between the two different types of

installation. The α value after consolidation increased 75% compared to the α value of

0.43 during immediate pullout of the caisson, due to the contribution from set-up effects

in the surrounding soil. It should be noted that the external α obtained in this way is a

lower bound, since the external α is likely to be larger than the internal value, due to the

more favourable conditions for consolidation outside the caisson compared to inside.

The uplift capacity of the caissons installed in the disturbed sites was also investigated.

As mentioned previously, these tests are indicated with a ‘*’ in the name. They were

pulled out in the disturbed sites after consolidation for 1 hour at 120 g. Profiles of

capacity during unsealed pullout for four typical tests: B2JOC* and B6JOC* installed

by jacking, and B3SOC* and B10SOC* installed by suction, are shown in Figure 6.74.

The graph reveals that the capacity in the disturbed sites was similar for both jacked and

suction installed caissons. Further comparison was made between the unsealed pullout

tests in the disturbed sites (labelled ‘OC*’) and the tests in the original sites (labelled

‘OC’). The uplift capacity in the disturbed site is smaller than in the undisturbed site

(Figure 6.75). According to Table 6.11, the normalised uplift capacities during

unsealed pullout after consolidation in the disturbed sites averaged 18.4, which was

around 10% smaller than observed in undisturbed sites. The α values back-figured

from tests in the disturbed sites are also listed in Table 6.11. The average α value for

tests ‘OC*’ was 0.59, which is 15% smaller than that measured (α = 0.70) in the

undisturbed site, and is larger than that during immediate pullout (α = 0.43). This

suggests that in the disturbed sites, the strength of the soil adjacent to the caisson is

partially regained with time, although it does not recover fully (i.e. equal to the strength

in the undisturbed site).

6.5.2 Sealed Pullout in NC Clay

During sealed pullout, the drainage valve was closed by applying a positive pressure up

to 400 kPa on the piston of the valve, to ensure the soil plug remained inside the caisson



during pullout. The uplift capacity during sealed pullout is composed of three parts: 1)

the reverse end-bearing capacity, 2) the external shaft friction, and 3) the submerged

weight of the soil plug and the caisson body. The uplift capacity ∆p is defined as the

axial force divided by the cross-sectional area of the caisson and can be expressed as

follows:

( ) plugextuext

n

1ibaseiuicibase WAsα AγzsNA∆p ++′−=⋅− ∑

=

(6.6)

where the meaning of each item has been described in Chapter 2.

It should be noted that the contribution from the pad-eye is also included in the above

formula. Sealed pullout tests were carried out for both suction-installed and jacked

caissons. Some of the successful sealed pullout tests installed by suction are reported

here. In order to verify the α values obtained here from caisson tests, two extra tests

were performed on a solid pile which has the same equivalent diameter to the caisson.

6.5.2.1 Pullout after consolidation

In test B11SCC where the caisson was installed by suction in NC clay, 1 hour of

consolidation time at 120 g was allowed, then the sealed caisson was pulled out

vertically at a velocity of 0.3 mm/s. Variation of the pullout capacity with the

embedment of the caisson for test B11SCC is shown in the left side of Figure 6.76a.

The capacity was –325 kPa and the maximum embedment was 13.86 m. The variation

of the pore pressure inside the caisson with the pullout depth is shown in Figure 6.76b.

The undrained shear strength gradient of test B11SCC is 1.26 kPa/m (Figure 6.77).

Also shown in Figure 6.76a is the uplift capacity of an unsealed pullout test, B11SOC,

performed in the same box. The uplift capacity of the sealed caisson was 1.6 times the

capacity of –187 kPa measured during the unsealed pullout. The difference may be

attributed to the reverse end-bearing capacity mobilised for the sealed caisson.

The normalised uplift capacities were 33.3 and 20.4 for the above two tests, giving a

value of the sealed pullout test 64% higher than in the unsealed pullout test. It should

be noted that the su at the tip of the caisson in Table 6.12 was calculated by the

nonlinear relationship between su and depth derived from the T-bar test, since the

strength profiles became increasingly non-linear at greater depths. The failure modes

between sealed pullout and unsealed pullout were totally different. It took around 0.9 m

(or 0.25d) of movement for the caisson to develop the ultimate sealed pullout capacity,



as compared to the very small displacement during unsealed pullout. The failure mode

was ductile for the sealed pullout while brittle for the unsealed pullout after

consolidation. Contribution from the reverse end-bearing capacity during sealed pullout

is much larger than that during unsealed pullout, and linked to the caisson diameter

rather than the wall thickness, which accounts for the more ductile response and longer

displacement to failure.

Table 6.12 Upper bound Nc during sealed pullout of the caisson in NC clay

Test dsu/dz Lmax

(m)

su, tip

(kPa) –∆p

(kPa) us∆p

− αext Nc

B11SCC 1.26 13.86 19.5 325 33.5 0.70 12.5

B12SCC 1.17 14.41 16.9 292 34.6 0.70 12.3

B3SCC 1.08 14.09 14.4 240 33.3 0.70 12.7

Average (test ‘CC’) 1.17 14.12 16.9 287 34.0 0.70 12.5

B12SCI 1.17 14.40 16.9 213 25.2 0.40 9.8

B3SCI 1.08 14.23 15.1 207 27.4 0.40 11.5

B4JCI 1.13 14.70 16.8 236 28.1 0.40 11.5

Average (test ‘CI’) 1.13 14.44 16.3 219 26.9 0.40 10.9

Note: ‘S’ for suction installation and ‘J’ for jacked installation

At this stage, it is convenient to adopt the external α during sealed pullout as the same

value measured in the unsealed pullout, since there is no logic for using different values

on the external wall. The α value for test B11SCC was taken as 0.7 obtained in the

unsealed pullout test B11SOC in the same box. As is generally applied in engineering

analysis, the uplift capacity (∆p) and the maximum embedment (Lmax) of the caisson are

used in the analysis, although in fact they do not occur at the same time. According to

the uplift capacity and the maximum embedment shown in Table 6.12, the reverse

end-bearing capacity factor Nc in test B11SCC is 12.5 by Equation 6.6. It should be

noted that this value of Nc is in fact an upper bound, since as already noted, the

estimated shaft friction ratio from the unsealed tests is a lower bound, and the thickness

of soil below the caisson tip at full embedment is only around one diameter of the

caisson. The upper bound α value and the corresponding lower bound Nc value are



discussed in the next chapter, based on measurements of the radial effective stresses

around caissons. It should be noted that Nc values obtained in Table 6.12 are subjected

to the influence of consolidation on the shaft friction, since full consolidation may not

have been achieved after 1.7 years of consolidation. The roughness of the caisson wall

may also affect the shaft friction and the bearing capacity. In addition, the caisson was

pulled out at a different rate compared to that used to obtain su in T-bar tests.

Therefore, there are a number of secondary effects that may influence the derived Nc

values.

Another sealed pullout test after consolidation, B12SCC, was carried out in Box 12.

The axial capacity during pullout is shown in the left side of Figure 6.78a, and the

variation of the internal pore pressure is shown in Figure 6.78b. It should be noted that

during pullout, suction was applied through the syringe pump as soon as an increase in

the internal pore pressure was noticed, in order to ensure the soil plug stayed within the

caisson. The suction force was applied by withdrawing the syringe pump, starting at a

velocity of 0.1 mm/s, and the velocity was adjusted according to the subsequent

readings of the internal PPT during pullout. The maximum embedment of the caisson

was 14.41 m, and the ultimate uplift capacity was –288 kPa after 1.08 m of movement

(see Figure 6.78). Using the undrained shear strength measured by T-bar test (see

Figure 6.79), the normalised uplift capacity for test B12SCC was 34.6, which is very

close to the value of 33.3 obtained in the parallel test B11SCC (see Table 6.12). By

adopting an external α value of 0.7 as used in test B11SCC (since no unsealed pullout

test was arranged in the same box of B12SCC), the Nc value can be calculated as 12.3,

which is quite close to that derived in test B11SCC (see Table 6.12).

Figure 6.80a shows the sealed pullout capacity of the caisson installed by suction in NC

clay (test B3SCC) in the old control system, where there was an interval of 100 seconds

between the self-weight penetration and the suction installation. The variation of the

underpressure is shown in Figure 6.80b, while the corresponding strength profile of the

soil is shown in Figure 6.81. The uplift capacity of the test B3SCC was –244 kPa, the

maximum embedment was 14.1 m, and the strength gradient was 1.08 kPa/m. The

normalised uplift capacity was found to be 33.3, which is also close to that for tests

B11SCC and B12SCC using the new control system. The average normalised uplift

capacity, us∆p− , for these sealed pullout tests is 34.0.



A sealed pullout test after consolidation was conducted on a caisson installed by

jacking. Due to incorrect arrangement of the drainage tubes, cavitation occurred after a

short distance of vertical displacement of the caisson. The capacity profile is shown in

Figure 6.82 and the strength profile is shown in Figure 6.83. The maximum embedment

of the caisson was 14.32 m, the ultimate pullout capacity of –251 kPa occurred after

0.62 m of vertical displacement. The normalised uplift capacity of test B4JCC was

31.3, smaller than the other sealed pullout test. The sudden drop in the uplift pressure

of test B4SCC was considered to be a result of cavitation during pullout. Data on the

internal pore pressure was unavailable since the PPT stopped working during the test.

Direct comparison of the sealed pullout capacity after consolidation therefore could not

be made between caissons installed by jacking and by suction in NC clay. Such a

comparison will be made through tests in LOC clay, as discussed later in section 6.5.3.

By assuming different Nc values during pullout, the external shaft friction ratio may be

estimated for the caissons extracted under sealed conditions. The results are

summarised in Table 6.13 for the suction-installed caissons, for three alternative

assumptions on the end-bearing capacity factor, Nc. The average α values are

surprisingly high, exceeding unity for Nc values less than 10.5. Even taking Nc as 12,

which is considered a likely upper limit, the average α value obtained was 0.77, which

is 10% greater than the average internal and external α values obtained from the

unsealed caisson tests.

Table 6.13 External α during sealed pullout after consolidation in NC clay by

assuming different Nc values

Test us∆p− Nc = 9 Nc = 10.5 Nc = 12

B11SCC 33.5 1.26 1.02 0.78

B12SCC 34.6 1.13 0.93 0.73

B3SCC 33.9 1.28 1.05 0.81

Average 34.0 1.22 1.00 0.77



6.5.2.2 Equivalent solid pile test

A small solid pile with an equivalent diameter to the model caisson ( dt2deq = ) was

fabricated, and two pullout tests after consolidation were performed in NC clay. These

tests were designed to provide a comparison of the shaft friction ratios derived from

tests on suction caissons. The small pile was fabricated from 6061 T6 aluminium, the

same material as the caisson, and was anodised after sandblasting to a CLA roughness

of 2.5 µm, the same roughness as that of the caisson (Figure 6.84). The solid pile has a

diameter of 7.68 mm at model scale, representing a diameter of 0.92 m at 120 g; the

cross-sectional area of the pile is equivalent to the caisson annulus used in this research.

Tests were performed in undisturbed sites in Box 12 following the caisson tests, in a site

with a soil strength gradient of 1.24 kPa/m. The solid pile was jacked at a speed of

2 mm/s at 120 g to the same embedment depth as the caisson, i.e. 120 mm at model

scale (or 14.4 m at prototype scale). After consolidation for 1 hour (1.7 years at

prototype time) in flight, the pile was pulled out vertically at a velocity of 0.3 mm/s.

The recorded axial force was divided by the cross-sectional area of the pile; the

resulting pressures during both installation and pullout are shown in Figures 6.85 and

6.86, for parallel tests B12pile1 and B12pile2. Details of the test results are listed in

Table 6.14.

Consistent results are found between these two tests, both during installation and

pullout after consolidation. The axial capacities expressed in pressure are different

from those measured in the caisson tests, due to different cross-sectional areas (0.67 m2

for the solid pile and 10.18 m2 for the prototype caisson, although the net sectional areas

are the same for both). Nc values are taken as 9 during both installation and pullout,

according to the recommendation of API RP2A (1993) for solid piles, since the depth of

soil below the pile at full embedment was 3.8deq, which is appropriate for developing

the Nc value adopted here.

Based on the data shown in the above plots and Table 6.14, the α value during

installation can be back-figured as 0.45 and 0.52 for the two tests, the average α value

of 0.49 is 29% larger than that of 0.38 derived from caisson tests in NC clay. During

pullout after consolidation, the derived α value was 0.90 for test B12pile1 and 0.96 for

test B12pile2, with an average α value of 0.93, which is very close to unity derived

from API RP2A (1993), and is 21% larger than that of 0.77 derived from pullout tests of

suction caissons. It can be seen that for a certain equivalent diameter, solid piles have



larger α values during both installation and pullout, compared to thin-walled caissons,

showing a trend of increased α with decreased d/t ratio for piles. The extrapolation

from the measured α value of solid piles to thin-walled caissons, by simply adopting an

equivalent diameter, will introduce significant over-predictions in both installation and

loading. Therefore, study (and especially measurements) of the real radial stress

changes and thus the development of shaft friction for thin-walled suction caissons is

very important in building the design rules; this will be discussed in detail in Chapter 7.

Table 6.14 Pullout tests on the equivalent solid pile after consolidation in NC clay

(assuming Nc = 9 during both installation and pullout)

Test dsu/dz Lmax

(m)

su, tip

(kPa) ∆pinstall

(kPa)

∆ppullout

(kPa)

αinstall αpullout

B12pile1 1.24 14.40 17.9 507.7 –546.4 0.45 0.90

B12pile2 1.24 14.17 17.6 519.2 –509.7 0.52 0.96

Average 1.24 14.29 17.8 513.5 –528.1 0.49 0.93

6.5.2.3 Immediate pullout

Sealed pullout tests in NC clay were also carried out immediately after installation. A

typical result for test B12SCI is shown in Figure 6.87, the corresponding undrained

strength profile is shown in Figure 6.88, with a gradient of 1.17 kPa/m. Also shown in

Figure 6.87a are the results of the immediate unsealed pullout test (B9SOI) and sealed

pullout test after consolidation (B12SCC). The graph shows that the ultimate pullout

capacity for B12SCI was –213 kPa, which occurred after 0.86 m (0.24d) of vertical

displacement. The largest embedment of the caisson was 14.4 m; the normalised uplift

capacity can thus be obtained as 25.2, which is 85% larger than that during unsealed

pullout in test B9SOI, and 27% less than that of test B12SCC during sealed pullout after

consolidation. In the former case, the difference is due to the reverse end-bearing

enforced during sealed pullout, while the latter was due to the increase in shaft friction

on the external wall after consolidation. During immediate pullout no consolidation

was allowed in the soil and the α value should be very similar to that during installation.

Therefore, the α value during immediate pullout in NC clay can be estimated as 0.4,



which is the same as during installation. The Nc value for immediate sealed pullout test

B12SCI can be back-figured as 9.8 by Equation 6.6, according to the measurements

shown in Table 6.15.

Table 6.15 Upper bound Nc during immediate sealed pullout of the caisson in NC

clay

Test dsu/dz Lmax

(m)

su, tip

(kPa) ∆p

(kPa) us∆p

− α Nc

B12SCI 1.17 14.40 16.9 –213 25.2 0.40 9.8

B3SCI 1.08 14.23 15.1 –207 27.4 0.40 11.5

B4JCI 1.13 14.70 16.8 –236 28.1 0.40 11.5

Average 1.13 14.44 16.3 –219 26.9 0.40 10.9

Results of another immediate sealed pullout test after suction installation, test B3SCI,

are shown in Figure 6.89, and the corresponding undrained shear strength profile is

shown in Figure 6.90. The ultimate uplift capacity and embedment were –207 kPa and

14.23 m, respectively, while the strength gradient of the soil was 1.08 kPa/m. The

normalised uplift capacity was thus 27.4, which is similar to the result of test B12SCI.

By adopting the α value during immediate pullout as 0.40, the Nc value derived from

Equation 6.6 was 11.5. The capacity profile of test B3SCI shows that during the later

stage the soil plug inside the caisson was pulled out due to some cavitation caused by

upward movement in the tubes connected to the drainage valve. During the early stage

of the pullout, the tubes are well below the water level and hence cavitation was not

developed and the end-bearing capacity factor observed at the start of pullout is reliable.

An immediate sealed pullout test B4JCI was performed on a caisson installed by

jacking. The variation of uplift capacity during pullout is shown in Figure 6.91. The

undrained shear strength gradient dsu/dz was 1.13 kPa/m (Figure 6.92). The internal

PPT broke during the test so no results for the pore pressure inside the caisson were

available. The ultimate uplift capacity was –232 kPa which occurred after 1.2 m of

upward displacement, and the maximum embedment was 14.7 m (see Figure 6.91). The

corresponding normalised uplift capacity was thus 28.1. By adopting the same α value

during immediate unsealed pullout, the upper bound Nc value is 11.5. It can be seen in



Table 6.12 that the average Nc value during immediate pullout after installation was

10.9, which is 10% lower than the value of 12 estimated from pullout tests after

consolidation. This difference may be attributed to the effect of consolidation in the

soil.

6.5.3 Sealed Pullout in LOC Clay after Consolidation

Sealed pullout tests were also performed in the lightly over-consolidated (LOC) clay,

for both jacked caissons and suction caissons, using model caisson 2. After

consolidation for 1 hour at 120 g (1.7 years at prototype scale), the caisson was pulled

out vertically at a velocity of 0.3 mm/s, with suction applied by the syringe pump at an

initial rate of 0.1 mm/s, to prevent the soil plug from falling out of the caisson.

The uplift capacity of the caisson installed by jacking (test B13JCC) is shown in the left

side of Figure 6.93a, while the variation of internal pore pressure during pullout is

shown in Figure 6.93b. The measured ultimate pullout capacity (∆p) was –379 kPa,

while the maximum embedment (Lmax) was 14.01 m. Considering a corresponding

dsu/dz of 1.64 kPa/m for the soil strength profile shown in Table 6.7, the normalised

capacity was 33.6, which matched well with the value of 34.0 obtained in NC clay.

Variation of the axial capacity with depth of the sealed pullout test (B13SCC) for the

caisson installed by suction in LOC clay is plotted in Figure 6.94. The ultimate pullout

capacity for test B13SCC was –389 kPa, and the maximum embedment was 13.92 m.

Thus the normalised capacity could be obtained as 34.2, which is only 2% different

from that of test B13JCC installed by jacking. A certain shearing distance was also

needed in developing the maximum uplift capacity in LOC clay. The capacity profiles

of the suction installation test B13SCC and jacked installation test B13JCC are

compared in Figure 6.95, from which close agreement can be found for both the uplift

capacity and penetration resistance. Therefore, it can be concluded that there is little

difference between the axial behaviour of the caissons installed by jacking and by

suction in LOC clay.

A comparison between the sealed pullout capacity of caissons after consolidation in

LOC clay and that in NC clay (test B12SCC) is also shown in Figure 6.95, which

reveals a much larger capacity in the LOC sample. The difference in capacity is

attributed to the larger strength gradient of the LOC clay, which is around 1.5 times that

of the NC clay. The external α values derived from the measured uplift capacity after



consolidation by assuming different Nc values are shown in Table 6.16. The external α

values would be well in excess of unity if Nc is set less than 10. By adopting a possible

upper bound Nc value of 12 as stated previously, external α values of 0.77 for suction

installation and 0.68 for jacked installation are obtained. The average external α of 0.73

is slightly smaller than the value of 0.77 obtained in NC clay. This is in agreement with

the trend in most pile design that the long term shaft friction would decrease with an

increase in the OCR (Kolk & van der Velde, 1996).

Table 6.16 External α during sealed pullout after consolidation in LOC clay by

assuming different Nc values

Test Lmax

(m) ∆p

(kPa) us∆p

− Nc = 9.5 Nc = 10 Nc = 10.5 Nc = 12

B13SCC 13.92 –389 34.2 1.14 1.07 0.99 0.77

B13JCC 14.01 –379 33.6 1.05 0.97 0.90 0.68

Average 13.97 –384 33.9 1.10 1.02 0.95 0.73

6.5.4 Sealed Pullout in Sensitive Clay after Consolidation

Sealed pullout tests were carried out in sensitive clay, using the model caisson 2. After

installation by suction, the caisson was held in place in the soil under a downward

stress-hold of 8 - 12 N for one hour at 120 g. The caisson was then pulled out vertically

at a velocity of 0.3 mm/s with the lid sealed. Variation of the uplift capacity versus the

depth of the caisson for test B14SCC is shown in the left side of Figure 6.96. It should

be noted that the syringe pump was activated at a velocity of 0.5 mm/s during the

pullout, to prevent the soil plug from falling out of the caisson. The uplift capacity for

test B14SCC was –296 kPa (see Figure 6.96); the normalised axial capacity can be

obtained as 33.3 at a strength gradient of 1.16 kPa/m (see Figure 6.60) for the soil.

Such a normalised axial capacity was consistent with that measured in NC clay

(averaged at 34.0) and the LOC clay (averaged at 33.9) with lower sensitivity. By

adopting an Nc value of 12 during pullout, the lower bound external α is obtained as

0.65.

Another monotonic sealed pullout test B14susa (Figure 6.97) was carried out at the end

of the experiment. It can be seen in the graph that the uplift capacity was –293 kPa.



However, since the corresponding strength gradient of the soil has increased to

1.59 kPa/m, the normalised axial capacity was 25.1, which is much lower than that of

B14SCC. The decrease can be attributed to the disturbance from two earlier sealed

pullout tests (sustained loading test B14sus and cyclic loading test B14cyc) adjacent to

this site. This also emphasises the importance of considering the normalised axial

capacity instead of the absolute capacity value in such comparisons.

6.6 CONCLUSIONS

The axial capacity of caissons during both installation and pullout in normally

consolidated (NC), lightly over-consolidated (LOC) and sensitive clays has been

investigated here through a series of centrifuge model tests. The tests involved caissons

installed in an undrained mode either by jacking, or using suction after initial self-

weight penetration. Using data from (a) penetration resistance, (b) deduced soil plug

heave, (c) unsealed and sealed uplift capacity after consolidation, it was found that there

was no consistent difference between the behaviour of caissons installed by either

method.

The caisson was installed in an undrained mode. For most of the penetration depth

(~14 m), the actual applied underpressure was equal to or just above the required

underpressure, showing that no plug failure occurred for most of the depth during

suction installation. Lengths of soil heave inside the caisson deduced during suction

installation are 0.94 m, 0.83 m and 0.86 m respectively in NC, LOC and sensitive clays

(see Table 6.17); they are all clearly smaller than the theoretical predictions which are

based on the assumption that all the soil displaced by the caisson wall moves inside the

caisson during suction installation. The percentages of inward soil flow at the caisson

tip with respect to the embedded caisson volume for tests in NC, LOC and sensitive

clays are 45%, 42% and 46% respectively. These observations suggest that the soils

flow about evenly at the caisson tip during suction installation in soft clay. This

observation is in contrast to the assumption (100% inward flow under suction) generally

accepted in industry. This is an important finding, and further research is needed in

order to quantify the flow mechanisms in the soil around the caisson wall during the

installation process, and consequent radial stress changes around the caisson during

installation, in order to establish design rules for the external shaft friction from first

principles.



For NC clay, the shaft friction ratio α ranged between 0.30 and 0.45 for both installation

processes, with an average of 0.38. Such a result is in accordance with the cyclic T-bar

results of St = 2.0 - 2.8, which suggests α = 0.36 - 0.5 during installation. For LOC clay

with OCR = 1.5, the shaft friction ratio α derived from the penetration resistance was

found to be 0.38 - 0.43, with an average value of 0.42 for jacked installation and 0.40

for suction installation. The α value measured during installation agrees well with the α

of 0.4 - 0.5 (corresponding to St = 2 - 2.5) derived from the cyclic T-bar tests in LOC

clay. For installation in sensitive clay (St = 4 - 5), the measured α value during

installation is 0.16, which is just below that derived from the cyclic T-bar tests (α = 1/

St = 0.20 - 0.25).

Table 6.17 Actual soil heave, predicted soil heave, percentages of inward and

outward soil flow at tip level for caissons installed by suction in NC,

LOC and sensitive clays

Soil Actual soil heave*

(m)

Predicted soil heave

(m)

Inward flow Outward flow

NC clay 0.94 1.38 45% 55%

LOC clay 0.83 1.12 42% 58%

Sensitive clay 0.86 1.23 46% 56%

Note: ‘*’ includes the soil heave caused by the ring stiffener and epoxy and wires inside caisson.

A model for predicting the penetration resistance of suction caissons was found to be

effective:

1. End-bearing capacity factor Nc = 7.5.

2. Shaft friction ratio on the external wall and internal wall below the first

stiffener: α = 1/St, with α = 0.30 - 0.45 (average value of 0.38) for NC clay, α

= 0.4 - 0.5 (averaged at 0.41) for LOC clay, and α ~ 0.16 for sensitive clay.

3. Average internal shaft friction above the first stiffener: τi-a ~ 0.5 kPa for NC

clay, τi-a ~ 1.0 kPa for the LOC clay, and α = 1/St for the sensitive clay.

Soil sensitivity St can be determined from in situ cyclic T-bar tests.



It should be noted that back-analysis shows that varying the shaft friction above the first

stiffener from zero to full strength leads to a difference of only 10% for tests in the NC

clay, 12% in the LOC clay and almost no change in the sensitive clay

During unsealed pullout in NC clay, a clear increase in the uplift capacity of the

caissons after consolidation was observed compared to those pulled out immediately

after installation. The reverse end-bearing capacity factor Nc can be taken as 7.5.

During immediate sealed pullout tests in NC clay, the average Nc value is 10.9; after

consolidation, the average normalised axial capacity, umin s∆p− , is 34.0, and the

average Nc value increased to 12.5, although some uncertainties exist in the estimate of

Nc values. It should be noted that the insufficient depth (one diameter of the caisson) of

soil below the caisson tip at full embedment may account for the high Nc value derived

from caisson tests, compared to Nc of 9 generally accepted in pile design. By adopting

an upper bound Nc of 12, the lower bound external α value after consolidation is 0.77 in

NC clay.

For sealed pullout tests after consolidation in LOC clay, the normalised uplift capacity,

umin s∆p− , is 33.9, which is close to that in NC clay. The normalised uplift capacity

was very close for caissons installed by jacking (34.2) and by suction (33.6). By

adopting an upper bound Nc value of 12, the lower bound external α value during

pullout is 0.73.

During sealed pullout after consolidation in sensitive clay, the normalised uplift

capacity agrees well with values measured in the NC and LOC clays, and varies in the

range 33 to 35. A lower bound αext of 0.65 can be obtained by using Nc = 12.

Re-installation in the disturbed sites would decrease the α values by ~10% during

installation and 15% during vertical pullout after consolidation.

Tests on a solid pile with the same equivalent diameter and surface roughness as the

model caisson show an α value of 0.48 during installation and 0.93 during vertical

pullout after consolidation in NC clay, by adopting Nc as 9. These α values are both

significantly higher than those derived from caisson tests. Therefore, a simple

extrapolation of the shaft friction ratios for solid piles to thin-walled caissons by taking

same equivalent diameters will introduce obvious over-predictions.



The upper bound external α value during pullout after consolidation can be analysed

from the radial effective stress around the caisson. This will be discussed with respect

to the measured radial stress changes around caissons in the next chapter.

Chapter 7 7-1 Radial Stress Changes around Caissons


7 RADIAL STRESS CHANGES AROUND CAISSONS IN CLAY

7.1 INTRODUCTION

The axial capacity of suction caissons during installation and monotonic uplift has been

discussed in the previous chapter. Lower bound external α values during sealed pullout

have been derived from the measured post-consolidation uplift capacity. In fact, the

shaft friction is linked directly to the radial effective stress, which in turn is related to

the radial total stress and excess pore pressures during installation, consolidation and

loading phases (Randolph, 2003). The radial stress changes around suction caissons in

clay, and their relationship with the axial capacity, as well as comparisons between

measurements and theoretical predictions, form the main focus of this chapter. During

installation, back-analysis of shaft friction allows deduction of excess pore pressures,

while after consolidation (once the excess pore pressures reduce to zero) the radial total

stress allows upper bound estimation of shaft friction, as will be explained later.

In the experiments reported in this chapter, the radial stress changes around suction

caissons in clay were measured in the centrifuge by miniature pressure cells embedded

in the external wall. These measurements were made simultaneously with those of the

axial capacity reported in Chapter 6, in NC, LOC and sensitive clays.

7.2 EXPERIMENTS IN NC CLAY

7.2.1 Analysis of Radial Stresses during Installation

7.2.1.1 Measured σri , derived σ ri and ∆ui during installation

Tests in NC clay were all carried out on model caisson 1 (see Figure 3.2). The radial

total stress (σri) acting on the external shaft of the caisson during installation was

measured by two total pressure transducers (TPTs) located 60 mm (7.2 m at prototype

scale) above the caisson tip. The radial stress changes (averaged from the

measurements of two TPTs) during five jacked installations are shown in Figures

7.1 - 7.5, for tests B4JCC, B4JCI, B5JOI, B6JOI and B8JOI, and a comparison is shown

in Figure 7.6. It can be seen from the graphs that during penetration up to 7.2 m, the

TPTs remained in water and the measured radial total stress was the hydrostatic

pressure, uo, which increased linearly with depth. At the moment when the TPTs



entered the clay, the gradient of the measured total stress suddenly increased. Similar

trends were found between the measured σri among these tests, and the maximum

difference between the gradients of the stress in the soil was 5%.

Radial stress changes versus penetration depth during eight suction installation tests are

presented in Figures 7.7 - 7.14, for tests B3SCC, B3SCI, B11SOC, B12SOC, B12SCI,

B12SCC, B12cyc and B12sus. Very similar trends to those observed during jacked

installation occurred in these tests installed by suction. The comparison shows a

maximum difference of 8% in the gradients of individual tests (see Figure 7.15). It

should be noted that for suction installation tests in NC clay presented here, the average

depth where suction installation started was around 5.91 m (see Table 6.5). The average

depth where the caisson was finally installed to was 14.0 m (see Table 6.4), the TPTs

thus entered the suction-affected area in soil by 0.9 m. This distance is not large, yet

there is no obvious difference between the gradients of the measured external σri of the

jacking-affected area and the suction-affected area (see Figure 7.15) Taking an

individual test B11SOC for example, suction installation started at 4.95 m, and the

depth that the TPTs penetrated in the suction-affected area was thus 2.15 m, as the

caisson was installed to 14.30 m (see Figure 6.23). No significant change occurred in

the gradient of the measured external σri of test B11SOC as the TPTs entered the

suction-affected area, at a penetration depth of the caisson of 12.15 m (which is the sum

of 4.95 m and 7.2 m) (see Figure 7.9). This suggests the similarity between the patterns

of soil flow at the caisson tip for installation by jacking and by suction.

Comparison of the radial stress changes between jacked and suction installation tests is

shown in Figure 7.16, where average values of radial stress were taken from the two

types of test. Values of the radial total stress at the same depth are very close for the

two types of installation (Figure 7.16). The average gradient of radial total stress σri,

relative to the hydrostatic pore pressure, was 6.3 kPa/m for the 5 tests installed by

jacking, and 6.7 kPa/m (6% higher) for the 8 tests installed by suction. The difference

is not considered significant, and the overall average gradient is 6.5 kPa/m, or ~5.5su.

The measured radial total stress, σri, comprises the sum of the hydrostatic pressure, u0,

the excess pore pressure, ∆ui, and the radial effective stress at the caisson wall, σ'ri, and

can be expressed as follows:

r

uiriiori tan

suuuδ

α+∆=σ′+∆=−σ (7.1)



where the radial effective stress is estimated using

r

uri tanδ

sασ

⋅=′ (7.2)

where α is the local shaft friction ratio, and δr is the residual interface friction angle

measured by the ring shear tests.

The excess pore pressure during installation, ∆ui, was not measured directly by pore

pressure transducers (PPTs) due to the excessive size of PPTs relative to the model

caisson, which would lead to disturbance of the test sites. Instead, ∆ui can be deduced

from Equations 7.1 and 7.2 by the following expression:

or

urioririi u

tanδsα

σuσσ∆u −⋅

−=−′−= (7.3)

where items σri, α, su, δr and u0 in the formula are as stated previously. From Table 6.4,

the average shaft friction gradient, αsu, is 0.46 kPa/m. By substituting the measured

interface friction angle of 17.6° (see Table 5.1) into Equation 7.2, the radial effective

stress derived from Equation 7.2 is 1.45 kPa/m, or 1.2su. According to Equation 7.3, the

excess pore pressure gradient is then calculated as 5.1 kPa/m, or ~4.3su.

A number of theoretical methods have been proposed for predicting the variation of

radial total stress σri, radial effective stresses σ'ri and excess pore pressures ∆ui, as

discussed previously in Chapter 2. Theoretical predictions from these methods of the

radial stress changes are then compared with the measured values presented above.

7.2.1.2 NGI method

The NGI method is based on the assumption that the volume of the caisson wall is

accommodated entirely by soil flowing inwards into the caisson during suction

installation (Andersen & Jostad, 2002). Therefore, only shear-induced excess pore

pressures are generated on the external wall of the caisson during suction installation,

and can be expressed by Equation 2.10 in Chapter 2.

For the NC kaolin clay tested here, the in situ earth pressure coefficient K0 is 0.65, and

the sensitivity St is 2 to 2.8 (see Table 3.3). This leads to an excess pore pressure (∆ui)

gradient of 3.6 ± 0.3 kPa/m, or (3.1 ± 0.2)su, which is clearly lower than the derived

gradient of ~5.1 kPa/m. It should be emphasised, however, that during the caisson tests

using caisson 1, the average distance that TPTs penetrated in the suction-affected area



was ~0.9 m, which is rather limited, and so it would be expected that the measured

excess gradient would, at most, only trend gradually towards the NGI value.

The radial effective stress, σ ri, and the radial total stress relative to u0, σri – u0, can be

calculated by Equations 2.11 and 2.12, respectively.

Using the parameters stated previously, the σ ri predicted by the NGI method is

1.6±0.3 kPa/m, or (1.4±0.3)su. The gradient of σri – u0 predicted by the NGI method

is 5.25 kPa/m, or 4.5su; this is lower than the measured value, due to the

under-prediction of ∆ui.

7.2.1.3 Cavity expansion method

A simple form of the CEM for open-ended piles was developed recently by Randolph

(2003), with the assumption that all the soil particles displaced by the caisson tip during

installation move outside the caisson. The excess pore pressure generated outside the

caisson during installation can be calculated by Equation 2.14.

The first component of excess pore pressure is identical to that in Equation 2.10 of the

NGI method. For the second component, adopting G/su = 100 to 150 and ρ = 0.066 for

the model caisson, the gradient of ∆ui may be estimated as 5.4 ± 0.5 kPa/m, or

(4.6 ± 0.4)su. This range lies just above the derived value of 5.1 kPa/m.

The radial effective stress, σ ri, is estimated by the identical expression adopted by API

RP2A and also the NGI method, as shown in Equation 2.11. Again, CEM would give a

gradient of 1.6±0.3 kPa/m, or (1.4±0.3)su in σ ri.

As shown in Equation 2.15, the radial total stress relative to u0, σri – u0, can be obtained

as 7.1±0.3 kPa/m, or (6.0±0.3)su , which is just above the measured value of 6.5

kPa/m.

7.2.1.4 Strain path method

The prediction by the SPM (Whittle & Baligh, 1988) gives ∆ui = 1.05σ'v0, and σ ri =

0.23σ'v0 for an open-ended pile with d/t = 40 in NC clay. The results for the actual

caisson d/t of 60 were not calculated directly, but were obtained approximately by

imposing a reduction factor of 93% on the results of the open-ended pile with a d/t ratio

of 40. This reduction factor was derived from the difference between the results of an

open-ended pile with an area ratio (ρ) of 0.1 (corresponding to d/t = 40) and those with



a ρ of 0.066 (corresponding to d/t = 60), by using Equation 2.14 in the cavity expansion

method. The results of SPM can then be expressed as ∆ui = 0.98σ′v0, σ ri = 0.21σ′v0,

and σri – u0 = 1.19σ′v0. Hence SPM predicts ∆ui of ~6.7 kPa/m (or ~5.7su), σ ri ~ 1.4

kPa/m (or 1.2su) and σri – u0 ~ 8.1 kPa/m (or 6.9su) for the caisson in this study, with an

over-prediction of 30% and 25% on the excess pore pressure and radial total stress

relative to the hydrostatic pressure, respectively. However, its prediction of the radial

effective stress matches well with the measurement.

7.2.1.5 MTD method

Chow (1997) stated that the excess pore pressure generated during the penetration of

open-ended piles can be estimated by using the same expression as for solid piles, but

replacing the diameter, d, by an equivalent solid pile of diameter, deq, with the same

volume of steel (thus deq = d√ρ where ρ is the area ratio).

Values of σri – u0 and ∆ui generated adjacent to the pile during installation can be

predicted by Equations 2.16 and 2.17, which were proposed by Lehane (1992) for solid

piles (or closed-ended piles). For the present tests in NC kaolin clay, YSR = 1, while

h = 7.2 m and deq is 0.92 m, giving h/deq = 7.8 for model caisson 1. The predicted

gradients for σri – u0 and ∆ui are therefore 15.5 kPa/m (~13.1su) and 12.3 kPa/m

(~10.4su) respectively. It can be seen that the predicted σri – u0 and ∆ui are both more

than double the measured values. The radial effective stress can be derived from

Equation 2.18, and the gradient of σ ri so obtained is 3.2 kPa/m (~2.7su), which is also

more than double the measured value shown previously. It should be noted that in

Equation 2.18 σ ri was deduced from the difference between two quantities, with one

proportional to YSR0.42 and another to YSR0.5; this is not consistent and may introduce

unexpected errors.

These discrepancies between the measurements (derivations) and MTD predictions

confirm the difficulty in extrapolating the stress field from results obtained mainly from

solid (or closed-ended) piles to the very thin-walled caissons considered here. This

agrees with observations of a solid pile (with an equivalent diameter to the caisson)

tested in NC clay (see Chapter 6), which shows significant over-prediction when

applying α values of solid piles directly to thin-walled caissons, using an equivalent

diameter.



It should be noted that the strength ratio, su/σ′v0, for the NC kaolin clay is much lower

than that of most natural clays. This difference also contributes to the over-prediction

of the MTD approach, and the SPM as well.

7.2.1.6 Comparison between predictions and measurements

A summary of the theoretical predictions and the measured σri – u0 and derived ∆ui and

σ ri during caisson installation, expressed in terms of su, are given in Table 7.1.

Table 7.1 Predictions and measurements (or derivations) of external σri – u0, ∆ui

and σ ri in terms of su during caisson installation in NC clay

Method (σri – u0)/ su ∆ui/su σ ri/su

Measured (derivations *) 5.5 4.3* 1.2*

Lower bound 4.5

(–24%)

2.9

(–67%)

1.1

(-) NGI

Upper bound 4.5

(–24%)

3.3

(–25%)

1.6

(-)

Lower bound 5.7

(+4%)

4.2

(–2%)

1.1

(-) CEM

Upper bound 6.3

(+15%)

5.0

(+16%)

1.7

(-)

SPM 6.9

(+26%)

5.7

(+33%)

1.2

(0%)

MTD 13.1

(+138%)

10.4

(+152%)

2.7

(+125%)

Note: values with ‘*’ are derived from measurements, values in the brackets are the difference between

predictions and the corresponding measurements (derivations), ‘-‘ means no comparison made

since values are originated from the same formula as for the measured value.

A comparison of the analytical predictions of the radial total stress relative to the

hydrostatic pore pressure, σri – u0, in NC clay for all four methods is shown in Figure

7.17, and compared with average values deduced from the series of tests on (a) jacked

caissons and (b) suction-installed caissons. It should be noted that the NGI predictions

are shown for the depth where the TPTs penetrated into the suction-affected area. As



noted previously, the measured data for the two methods of installation lie close

together and stay just above the lower bound CEM prediction; the NGI method is

around 80% of the measurement; the SPM over-predicts by about 30%; the MTD

prediction is more than double the measurement.

The derived excess pore pressure, ∆ui, during caisson installation is compared with

various theoretical predictions in Figure 7.18. The SPM (assuming full outward motion

of the soil) over-predicts the measured data by around 30%; the MTD over-predicts the

measurements by a factor of 2, while the NGI method (based on full inward motion of

the soil) under-predicts the data by around 30%.

It can be concluded from Table 7.1 that the NGI method, which assumes all soil

particles are drawn inside the caisson during suction installation, obviously

under-predicts the measurements; the MTD prediction extrapolated from closed-ended

piles significantly over-predicts the values; the SPM over-predicts moderately; the

modified CEM developed by Randolph (2003) can give a good prediction of the

stresses around the caisson during installation.

7.2.2 Relaxation of Radial Stress during Consolidation

After the caisson was installed to the target depth, for jacked installation the axial force

was reduced to the nominal self-weight of 16 N, while for suction installation the

syringe pump was stopped immediately whilst the self-weight load was maintained. It

is important to maintain a constant self-weight force during consolidation, rather than

fix the displacement of the caisson, since the latter led to unacceptably high tension

forces developing as the clay consolidated. Three tests including B11SOC, B12SOC

and B12SCC were chosen for analysis. It should be noted that the depths where suction

installation initiated were 4.87 m (see Figure 6.23), 5.40 m and 5.62 m (see Figure 6.24)

respectively for these tests. With the caisson penetrated to 14.30 m, 14.40 m and

14.39 m during these three tests, the depths that TPTs penetrated into the

suction-affected area were 2.23 m, 1.80 m and 1.52 m. Variations of the external radial

total stress of TPTs inside the suction-affected zone during consolidation will be

discussed below.



7.2.2.1 t50 and t90

Throughout the consolidation period, the vertical displacement of the caisson and the

axial force were recorded by the displacement transducer and load cell, respectively.

These variations are shown in Figure 7.19 (in model scale) for test B11SOC, from

which it can be seen that the self-weight was reasonably tightly controlled (given the 2

kN range of the load cell), while downward movement of the caisson amounted to

0.12 mm (14.4 mm at prototype scale). The movements continued slowly over the one

hour (3600 seconds) of the consolidation period, although the rate of settlement

decreased gradually. By assuming 100% consolidation after one hour consolidation

(although some error may be introduced, it should be limited judging from the very

slow settlement of the caisson at the end of consolidation), times for 50% and 90%

consolidation (t50 and t90) are estimated to be about 1326 and 3230 seconds respectively,

corresponding to 7 months and 18 months at prototype scale. The measured

embedment of the caisson during consolidation for test B12SOC is shown in Figure

7.20 (in model scale), from which a settlement of 0.13 mm (15.6 mm at prototype scale)

can be identified. Times for t50 and t90 are 1079 and 2818 seconds respectively,

corresponding to 6 months and 16 months at prototype scale. Measurements in another

parallel test B12SCC (Figure 7.21) show that the settlement during consolidation is 0.07

mm (or 8.4 mm at prototype), times for t50 and t90 are 1163 and 2596 seconds

respectively, corresponding to 7 months and 14 months at prototype scale. The average

t50 and t90 derived from the variation of the embedment were respectively 7 months and

16 months prototype time. Embedment of the caisson corresponding to t50 and t90 of the

above three tests is shown in Table 7.2.

Table 7.2 Variations of embedment and average σr – u0 during consolidation in

NC clay

Embedment (model scale)

(mm) σr – u0

(kPa) Test

at t0 at t50 at t90 at t100 at t0 at t50 at t90 at t100

B11SOC 119.22 119.28 119.33 119.34 43.23 35.63 29.54 28.02

B12SOC 120.71 120.78 120.83 120.84 43.47 39.60 31.10 29.72

B12SCC 119.99 120.03 120.05 120.06 47.18 36.68 27.98 25.80

Average 119.97 120.03 120.07 120.08 44.75 37.30 29.54 27.85



As consolidation proceeds, there is some relaxation of radial total stresses around the

caisson, just as for displacement piles in clay as described by Lehane & Jardine (1994).

In test B11SOC (Figure 7.22), the two TPTs gave reductions of 16.87 and 13.56 kPa

over the consolidation period, giving an average reduction of 15.2 kPa. This represents

about one third of the initial value of σri – u0 immediately after installation. Times for

50 and 90% consolidation deduced from the variation in radial total stress were about

146 and 1733 seconds respectively, corresponding to 1 month and 10 months

respectively at prototype scale.

The variation of σr – u0 for test B12SOC during consolidation is shown in Figure 7.23.

The reductions in the two pressure cells are 13.95 and 13.54 kPa, with an average

reduction of 13.75 kPa. Times for 50 and 90% consolidation deduced from the

variation in radial total stress are about 521 and 2187 seconds, corresponding to 3

months and 12 months at prototype scale. In test B12SCC (Figure 7.24), the average

relaxation in σr – u0 from the two TPTs is 21.78 kPa, while t50 and t90 are 762 s and

2068 s, corresponding to 4 months and 12 months at prototype scale. Values of

σr – u0 corresponding to t50 and t90 during consolidation of the above three tests are

summarised in Table 7.2.

The 50 and 90% consolidation times derived from measured embedment of the caisson

and σr – u0 in the above three tests are summarised in Table 7.3. It can be seen in the

table that the average 50% and 90% consolidation times are 3 and 11 months,

respectively. These values are somewhat lower than those derived from the embedment

of the caisson in Figures 7.19 to 7.21, possibly reflecting on-going secondary

consolidation of the clay which may have affected the settlement response.

Randolph (2003) presented dimensionless consolidation times of T50 and T90 of ~1 and

~10, respectively, where T is defined as cht/deq2, and deq is the equivalent diameter.

Taking cv = 2.6 m2/yr and ch ~ 3cv (allowing for partial swelling during the

consolidation process - Fahey & Lee Goh, 1995), and deq as the equivalent diameter of

the caisson, 0.92 m, this would give prototype consolidation times of 1.3 months and 13

months respectively. These times fall either side of those for 50% and 90%

consolidation derived from the measured radial stress changes shown in Table 7.3. This

supports the notion that significant outward displacement of the soil occurs, even during

suction installation, and certainly the measured consolidation times are much greater



than the typical times of ~1 day (50%) and 6 days (90%) suggested by Andersen &

Jostad (2002). This has repercussions for the design of suction caissons, since in most

developments there are only short delays between installation and the attachment of

mooring lines.

Table 7.3 Derived 50% and 90% consolidation time in NC clay

by embedment z by σr – u0

t50 t90 t50 t90 Test

Model

(s)

Proto.

(month)

Model

(s)

Proto.

(month)

Model

(s)

Proto.

(month)

Model

(s)

Proto.

(month)

B11SOC 1326 7 3230 18 146 1 1733 10

B12SOC 1079 6 2818 16 521 3 2187 12

B12SCC 1163 7 2596 14 762 4 2068 12

Average 1203 7 3024 16 333 3 1960 11

7.2.2.2 Post-consolidation radial effective stress

The magnitude of stress relaxation during consolidation is important in determining the

final radial effective stress at the caisson wall, and hence the long-term shaft friction.

For displacement piles, Lehane & Jardine (1994) quantified the initial total radial stress

(after installation) in terms of the coefficient Hi = (σri – u0)/σ'v0 and the final radial

effective stress after consolidation as Kc = σ'rc/σ'v0. For low overconsolidation ratios,

typical values of these ratios for full-displacement piles, were ~2 and 0.8 to 1

respectively (see Figure 2.22), implying a stress relaxation of over 50% of the

‘potential’ radial effective stress at the pile wall.

Table 7.4 summarises typical values of the key stress ratios measured from the three

suction-installation tests mentioned in the last section, namely B11SOC, B12SOC and

B12SCC. Here for the thin-walled caisson, the coefficient, Hi, is around 0.9, while the

final radial effective stress ratio, Kc, averages 0.57. The latter value is similar to,

although just below, the estimated in situ earth pressure coefficient, K0, for normally

consolidated kaolin clay. The average stress relaxation from the two tests is 37%. This



is lower than for full-displacement piles, as suggested by Randolph (2003) who

postulated an expression for the final radial effective stress ratio of Equation 2.20.

Table 7.4 Measured radial stress changes during consolidation in NC clay

Test i,TPTz

(m)

0ri uσ −

(kPa) v0

0ri

σuσ

′− c,TPTz

(m)

0rcrc uσσ −=′

(kPa) v0

rc

σσ

′′

0ri

rc

uσσ−′

B11SOC 7.10 43.23 0.90 7.12 28.02 0.58 0.65

B12SOC 7.28 43.47 0.88 7.30 29.72 0.60 0.68

B12SCC 7.19 47.18 0.93 7.20 25.80 0.52 0.55

Average 7.19 44.63 0.90 7.21 27.85 0.57 0.63

From the measured shaft friction during installation, and the interface friction angle of

17.6º, the quantity σ'ri/σ'v0 may be estimated as 0.21, while ∆ui/σ'vo is 0.79. Taking R

(name of yield stress ratio in CEM) as unity leads to a post-consolidation radial

effective stress ratio of 0.53, which is some 10% lower than the observed value of 0.63

(see Table 7.4).

According to Jardine & Chow (1996), the MTD prediction of Kc can be obtained by

Equation 2.19, which resulted in a Kc value of 1.08 to 1.12. This is more than 80%

greater than the measured ratio, underlining the need for caution when applying such

methods to pile or caisson geometries well outside the database on which the methods

are based.

The NGI method (Andersen & Jostad, 2002) does not give an explicit expression for the

radial effective stress after consolidation, and will not be discussed here.

7.2.3 Radial Stress Changes and Shaft Friction during Pullout

7.2.3.1 Pullout after consolidation

The radial total stresses measured during the pullout tests for the three high quality

suction installation tests B11SOC, B12SOC and B12SCC are shown in Figures 7.25 to

7.27, with the measured σr in each test averaged from the simultaneous measurements



of two transducers. Comparison of the average σr measured in these three tests is

depicted in Figure 7.28. At the start of extraction, the radial total stress dropped 7.3 kPa

for test B11SOC, 6.0 kPa for test B12SOC and 3.8 kPa for test B12SCC, in an

extremely short pullout distance. Part of this reduction in σri may be due to slight

cross-sensitivity of the TPTs to the axial force acting through the caisson wall, although

the reduction is consistent with similar reductions measured during tension loading of

piles, where a similar 20% reduction in radial effective stress was reported by Lehane &

Jardine (1994). It may also be seen that there is a small recovery in the measured radial

stress immediately after the local minimum, after which the stresses decrease almost

linearly with the pullout distance. The gradient of σr during pullout is smaller than that

during installation, with the difference due to the relaxation in total stress during

consolidation, and the absence of any mechanism to generate significant excess pore

pressure during pullout. Once the TPTs leave the soil and re-enter the water, the

recorded σr matches closely the hydrostatic pressure line, confirming the reliability of

the total stress measurements.

An upper bound to the external shaft friction ratio, α, may be estimated from the

measured radial stresses during pullout, assuming that excess pore pressures are zero at

the very beginning of failure. The variations of the uplift pressure (∆p = P/Abase) and

radial total stress relative to u0 with the movement of the caisson are shown in Figures

7.29 to 7.31, for tests B11SOC, B12SOC and B12SCC. It can be seen in the graphs that

at the very beginning of pullout, the uplift pressure increased to a peak value in a very

short distance of shearing for the unsealed pullout tests B11SOC and B12SOC, being

0.06 m in prototype scale for both tests, and dropped suddenly after passing that point.

One possible reason has been described as the breakage of the suction beneath the

caisson lid during the unsealed pullout; another reason is that the structure of the soil

regained during consolidation is suddenly destroyed. The peak point is thought to be

the failure moment. At that stage, the excess pore pressures are likely to be low since

very small movement has taken place. Taking zero excess pore pressure at such a

failure point would result in the radial effective stress at failure, σ rf, calculated as the

radial total stress relative to u0, being 20.72 kPa and 23.77 kPa respectively for tests

B11SOC and B12SOC.

The maximum uplift capacity for the sealed pullout test B12SCC, however, is

developed later than that in unsealed pullout. The caisson reaches its peak capacity

after a shearing distance of 0.89 m (or 0.25d) in prototype (Figure 7.31). This ductile



response is considered to be due to the development of the reverse end-bearing capacity

during pullout. The failure state was reached at a much higher load than in the unsealed

pullout, where the contribution from the caisson tip resistance is rather small. For the

sealed pullout, the failure point of the radial effective stress is considered to be the

minimum value measured by the TPTs during the early stage of pullout. The σ rf

measured in test B12SCC is 22.05 kPa, and after a short distance of oscillation, the

radial stress decreases almost linearly with further movement of the caisson (Figure

7.31b). A summary of the σ rf measured in the three tests B11SOC, B12SOC and

B12SCC is given in Table 7.5.

Table 7.5 Measured σ rf and external shaft friction ratio α when the caisson is

loaded to failure during pullout in NC clay

Test Embedment after consolidation (m)

Embedment at failure (m)

By uplift capacity By σr – u0

σ rf

(kPa)

su

(kPa)

α

B11SOC 14.32 14.25 14.27 20.72 7.92 0.83

B12SOC 14.50 14.44 14.45 23.77 8.41 0.90

B12SCC 14.40 13.51 14.37 22.05 8.10 0.86

The external shaft friction ratio (α) during pullout of the caisson after consolidation can

thus be estimated by Equation 2.27; the external α values obtained for the above three

tests are listed in Table 7.5, with an average value of 0.86. As stated before, such an α

value is considered to be an upper bound. Taking this value, together with the deduced

lower bound αext values given in Table 6.13, suggests a most likely range for the

external α from 0.77 (as estimated from the sealed pullout tests) to 0.86. Recalling that

an average α of 0.7 was obtained by assuming equal α values on both sides of the

caisson in Chapter 6, the internal α can thus be back-figured to be in the range of

0.54 - 0.63. This is around 70% of the external α, possibly due to the longer drainage

paths for the re-consolidation of the soil inside the caisson, as described previously.

There is a reduction in radial effective stress as the caisson is loaded to failure (see

Figure 7.29). Therefore, estimates of the radial effective stress at failure, σ′rf, can be

derived by assuming that there is a reduction relative to the radial effective stress after



consolidation, σ′rc. The shaft friction from the MTD and CEM methods discussed

above may be obtained by Equation 2.26, by using the reduction factor, K. In fact, the

reduction factor can be obtained from the analysis of the radial stress changes measured

by the TPT and the pullout capacity when the caisson is loaded to failure.

In test B11SOC (see Figure 7.29), after consolidation the embedment (z) of the caisson

was 14.32 m (zTPT = 7.12 m), and σ′rc was 28.02 kPa. When the caisson was loaded to

failure the corresponding σ′rf was 20.72 kPa and the embedment was 14.27 m

(zTPT = 7.07 m). This means that the σ′r decreased 7.30 kPa during such a short distance

of shearing. Applying a linear reduction in radial effective stress due to the upward

movement of 0.05 m in the caisson, the real decrease in radial effective stress, ∆σ′r, was

7.08 kPa. Therefore, the reduction factor should be (1 – ∆σ′r / σ′rc) = 1 – 7.08/28.02

= 0.75. Similar analysis was carried out on test B12SOC (see Figure 7.30), for which

after consolidation the zTPT was 7.30 m (z = 14.50 m), and the σ′rc was 29.72 kPa; when

loaded to failure, σ′rf was 23.77 kPa and zTPT was 7.25 m (z = 14.45 m), thus the

reduction factor can be obtained as 0.81. The average reduction factor from these tests

is 0.78, which is close to that of 0.80 for open-ended piles suggested by Jardine &

Chow (1996). Therefore, a reduction factor, K, of 0.80 was used in Equation 2.26.

According to Equation 2.26, the external α values predicted by MTD and CEM, were

then derived as 1.60 and 0.80 respectively, based on post-consolidation radial effective

stress ratios of 1.10 (MTD) and 0.53 (CEM).

API RP2A (1993) recommends using Equations 2.24 and 2.25 for predicting αext for

open-ended piles after consolidation, with the constraint that 1α ≤ . If such a routine is

adopted to predict the external α for suction caissons in this study, using su/σ v0 = 0.18

as stated in Table 3.3, the external α is obtained as unity.

The CEM prediction appears reasonable in light of the measured values, while the MTD

method gives too high a value - as previously noted, largely due to extrapolation well

outside the database used to calibrate the approach. The NGI approach leads to an α

value of 0.65 (Andersen & Jostad, 2002), which is somewhat lower than that measured,

although the NGI prediction is just a recommended value when specific soil information

is not available. A summary of the measured and predicted external α values during

pullout of caissons after consolidation in NC clay is given in Table 7.6. A comparison

of profiles of the external shaft friction during pullout, between measurements and

various theoretical predictions, is shown in Figure 7.32.



Table 7.6 Measured and predicted external shaft friction ratio α during pullout of

the caisson after consolidation in NC clay

Methods α

Measured 0.77 - 0.86

MTD approach 1.60

NGI method 0.65

CEM 0.80

API (RP2A) 1.0

7.2.3.2 Immediate pullout

Other vertical pullout tests, immediately after installation, were performed on the model

suction caissons. The radial stress changes immediately before and after full

installation for a typical test B2JOI are shown in Figure 7.33. During pullout the radial

total stress recorded by the TPTs decreased by 2.9 kPa when the caisson was pulled out

from its full embedment of 14.17 m (immediately after installation), to 14.13 m (when

the movement just commenced). Although the measured stress during the following

pullout was slightly larger than that during installation, the overall values were very

close. A comparison between the measured radial stress for the immediate pullout test

(test B2JOI) and the pullout test after consolidation (test B11SOC) is shown in Figure

7.34. A large difference can be seen between these two tests, indicating that the

reduction in radial stress after consolidation is mainly due to the relaxation of the soil

during consolidation, rather than the influence of cross-sensitivity. This suggests that

for pullout tests after consolidation, the external α derived from the radial stress at

failure within the early stage of pullout is reasonable. For immediate pullout, the excess

pore pressure has not dissipated yet, which makes it impossible to derive the external α

value from the measured radial stress.

7.3 EXPERIMENTS IN LOC CLAY

To investigate the radial stress changes around caissons in lightly overconsolidated

(LOC) clay, and to compare measurements and predictions of the behaviour of caissons

in a different environment, tests were undertaken in kaolin clay with an OCR of 1.5.



The sample was consolidated at 180 g and tested at 120 g with details described in

Chapter 3. Caissons were installed either by jacking or by suction, with the radial stress

changes recorded by two transducers on the external wall of the caisson. It should be

noted that all tests in LOC clay were carried out on model caisson 2 (see Figure 3.7), for

which the model distance between the TPTs and the caisson tip is 40 mm (representing

4.8 m prototype at 120 g). The elevation where the TPTs are located here is lower than

that of model caisson 1 tested in NC clay. The measured external radial stresses during

installation of the caisson are analysed below.


7.3.1.1 σri , σ ri and ∆ui during suction installation: test B13SCC

Before the installation started, the caisson tip was hung just above the mudline, the

distance from the TPTs and the mudline was thus 40 mm in model scale (or prototype

4.8 m at 120 g). During suction installation in LOC clay, the caisson was first installed

by self-weight penetration to around half its length (i.e. ~60 mm at model scale, or 7.2

m at prototype scale). The TPTs entered the soil before the syringe pump was initiated

to effect suction installation.

A plot of the recorded radial total stress (σri) versus the embedment of the caisson

(expressed as the depth of the tip) during suction installation of test B13SCC is shown

in Figure 7.35. During penetration up to 4.8 m, the TPTs remained in water and

recorded the hydrostatic pressure. Once the TPTs entered the clay at a depth of 4.8 m,

the gradient suddenly changed, but the readings continued to increase almost linearly

with penetration depth, except for some obvious decreases at 12.11 m and 13.72 m (see

Figure 7.35). Variations of the applied syringe pump pressure and the measured radial

total stress with respect to penetration depth are shown in Figure 7.36. Judging from

the applied syringe pump pressure, the self-weight penetration (jacked installation)

ended at a penetration depth of 6.97 m (although the caisson moved slightly further due

to inertia), and the suction installation started at 7.31 m. In this transition period, the

variation in radial total stresses was insignificant, which confirms that (a) the vertical

length that the suction can influence is less than 4.8 m, and (b) any cross-sensitivity due

to changing from jacked installation to suction installation is trivial.

The measured radial total stress decreased slightly at the depth of 12.11 m (see Figure

7.35), when the TPTs entered the suction-affected area (since it took 4.8 m further



movement for the TPTs to reach the elevation of the caisson tip, after the initiation of

suction installation at 7.31 m). This variation, again, resulted from a reduction in the

moving speed of the caisson (see Figure 7.37). It can be seen that between depth (z) of

6.44 and 7.31 m, the average penetration rate (v, in model scale) of the caisson

decreased to 0.53 mm/s, when the load cell reading was tracking the target stress-hold

before suction installation started. The corresponding normalised velocity, V (V = vt/cv,

v is the velocity, t is the wall thickness, 0.5 mm, cv is 0.076 mm2/s, or 2.4 m2/year), was

3.5, and the penetration is thus partly drained according to Randolph (2004). When the

preset stress-hold was reached, the system was switched from jacked installation to

suction installation between 6.97 - 7.31 m, which caused a slight time delay of 1.1 s (or

0.18 days prototype time) (see Figure 7.37). Consolidation due to the time delay and

reduced speed of penetration thus caused a reduction in the measured σri at the depth of

12.11 m (4.8 m deeper than 7.31 m where suction installation started).

Another reduction in the radial total stress occurred at the depth of 13.72 m; the

reduction is not very large, and it corresponded to a surge in the syringe pump pressure

(see Figures 7.36, 7.37). Before 13.72 m, the caisson was installed at a velocity (in

model scale) of 1.80 mm/s, which corresponds to a V of 11.8; the installation was thus

undrained. However, after 13.72 m the speed decreased to 0.18 mm/s (see Figure 7.37),

which corresponds to a V of 1.8, and thus was partly drained. As discussed in Chapter

6, the soil plug reached the caisson lid at 13.62 m (see Figure 6.42), which subsequently

blocked the connection to the syringe pump and caused a sharp increase in the applied

underpressure while the system was trying to achieve further penetration. This caused

the penetration rate to reduce at 13.72 m. Therefore, excess pore pressures generated

during installation started to dissipate in this period (after 13.72 m); this is considered to

be the major reason for the corresponding drop in the measured radial total stress.

More details of the stress changes during installation of test B13SCC are shown in

Figure 7.38, where the radial total stress relative to the hydrostatic pressure during

installation, σri – u0, is plotted against the depth of the TPT (zTPT) below the soil. Here

zTPT is calculated as z – 4.8 m, and z is the embedment of the caisson tip.

At zTPT = 2.17 - 2.51 m (corresponding to z = 6.97 - 7.31 m), σri – u0 experienced a

small decrease of ~0.5 kPa, which was caused by slight consolidation during the time

delay when the operating system changed from jacked installation (self-weight

penetration) to suction installation. At zTPT = 6.97 m, the TPTs left the jacking-affected



area in the soil, and the measured σri – u0 was 57.88 kPa. When the TPTs entered the

suction-affected area at zTPT = 7.31 m, σri – u0 dropped to 55.99 kPa. Considering an

increase in lateral earth pressure of around 2.7 kPa due to an increase in the depth, the

actual decrease in σri – u0 was 4.6 kPa (see Figure 7.38) in the transition area. This

decrease is less than 10% of the measured stress in that region, and is considered to be

the result of 1) slight consolidation during the time delay in starting suction installation,

and 2) consolidation when the penetration rate reduced to 0.53 mm/s (in model scale)

between 6.44 and 7.31 m (see Figure 7.37).

During the subsequent penetration in the suction-affected area (after zTPT of 7.31 m),

σri – u0 continued to increase. In this test, the vertical range of the suction-affected area

in the soil is larger than 1.51 m, which is long enough to display the trend of variation.

During suction installation, the gradient of the radial total stress acting on the external

wall of the caisson is slightly lower than that during self-weight penetration (see Figure

7.38), probably due to a smaller penetration rate during suction installation (v = 1.80

mm/s in model scale) versus that of 2.77 mm/s during self-weight penetration (see

Figure 6.39), although both were estimated as being undrained. The slight change

indicates that the difference between the patterns of soil flow beneath the caisson tip is

very small for the two types of installation. This agrees with the observation in Chapter

6, which shows that when suction installation started, the applied underpressure was

well below the allowable value (see Figure 6.42), and thus no significant inward flow of

soil was induced by suction installation.

After zTPT = 8.92 m, σri – u0 decreased 9.3 kPa within 0.18 m of penetration, where the

velocity of installation decreased to 0.18 mm/s (in model scale). In fact, the suction

installation had essentially ceased at zTPT = 8.92 m (or z = 13.72 m), due to the soil plug

contacting the caisson lid at 13.62 m (see Figure 6.42), after which consolidation took

place and caused the reduction of σri – u0.

In the suction-affected area, stresses at the turning point (zTPT = 8.92 m) before the

penetration rate decreased are chosen for analysis. At zTPT = 8.92 m, σri – u0 was

70.13 kPa, corresponding to a stress gradient of 7.9 kPa/m. The value of su was

13.47 kPa (see Figure 6.46), corresponding to an undrained strength gradient, dsu/dz, of

1.51 kPa/m, which will be used in the following analysis. It should be noted that this

strength gradient is slightly different from that for the whole length of the caisson.

Therefore, the measured gradient of σri – u0 was ~5.2su for test B13SCC.



The excess pore pressure generated on the external wall of the caisson during

installation, ∆ui, can be derived from the measured radial total stress, using Equation

7.3. The α value during installation of test B13SCC was 0.42 (see Table 6.7), using

dsu/dz of 1.51 kPa/m as mention before, the average shaft friction gradient, αsu, is thus

0.63 kPa/m. The residual interface friction angle between the LOC clay and the caisson

obtained from the ring shear tests shown in Chapter 5 is 18.1° (see Figure 5.1). The

radial effective stress can thus be derived from Equation 7.2 as 1.9 kPa/m, or 1.3su.

Substituting this result into Equation 7.3 allows the excess pore pressure gradient to be

calculated as 6.0 kPa/m, or ~4.0su.

Following the same routine as that used for analysing radial stress changes for caissons

in NC clay, theoretical predictions by SPM, CEM, NGI and MTD will be presented, and

compared with the measurements (or derivations), as discussed below.

The NGI prediction (Andersen & Jostad, 2002) for ∆ui can be obtained from Equation

2.10. The lateral earth pressure coefficient at rest, K0, is 0.70 according to Chapter 3.

Substituting St = 2 - 2.5 (see Table 3.3) into Equation 2.10, the NGI method gives a

gradient of excess pore pressure of 3.4 - 3.8 kPa/m, or (2.3 - 2.5)su, which is 56 - 64%

of the value derived from measurements. According to Equation 2.11, the NGI

prediction of the radial effective stress is 1.8 - 2.3 kPa/m, or (1.2 - 1.5)su. The predicted

σri – u0 by the NGI method is 5.7 kPa/m, or 2.6su, which is 72% of the measured values.

The significant difference between the NGI prediction and the derived excess pore

pressure acting on the external wall of the caisson during suction installation suggests

that the assumption of 100% inward soil flow at the caisson tip under suction is not

supported by the test results. This difference in LOC clay, however, seems to be

slightly smaller than for caisson tests in NC clay. It should be noted that the effect of

OCR was not considered for the NGI method.

The CEM prediction of ∆ui can be obtained by Equation 2.14, using the parameters

presented previously in Table 3.3, i.e. St = 2 - 2.5, K0 = 0.7, and ρ = 0.066,

G/su = 100 – 150, the gradient of ∆ui may be estimated as 5.7 - 6.9 kPa/m, or

(3.8 - 4.6)su. The CEM predictions of σ ri and σri – u0 can be derived from Equations

2.11 and 2.15, and the results are 1.8 - 2.3 kPa/m, or (1.2 - 1.5)su for σ ri, and 7.5 - 9.2

kPa/m, or (5.0 - 6.1)su for σri – u0. It can be seen that the CEM prediction of the excess

pore pressure is close to the gradient of 6.0 kPa/m derived from measurements, with a

difference of –5 - 15%.



A comparison between the measured σri – u0 and predictions by the NGI method and

CEM during installation of test B13SCC is shown in Figure 7.39. It should be noted

that the NGI prediction was shown for the suction-affected area. Also shown in the

graph is the predicted trend of measured σri – u0 in the suction-affected area, following

the assumption of Andersen & Josod (2002) that the effect of self-weight penetration

will decrease linearly to zero within a depth of one diameter (3.6 m here) below the

self-weight penetration depth. At zTPT = 7.31 m when the TPT entered the

suction-affected area, the measured σri – u0 was 55.99 kPa. It decreased slightly, but it

remained just above the lower bound CEM prediction during subsequent penetration in

the suction-affected area, After TPTs entered the suction-affected area at zTPT = 7.31m,

the measured σri – u0 should develop along the thick green arrow in Figure 7.39,

reaching 58.86 kPa at 8.92 m when suction installation essentially stopped, under the

assumption of one diameter transition zone proposed by Andersen & Jostad (2002).

The obvious difference between the gradients of the measured σri – u0 and the NGI

assumption, however, suggests that the assumed transition may not occur.

After zTPT = 8.92 m when the moving speed (in model scale) of the caisson changed

from 1.80 mm/s to 0.18 mm/s, the measured σri – u0 experienced another decrease,

amounting to 9.3 kPa within a slow movement of 0.18 m until the end of penetration.

For the LOC clay with an OCR of 1.5, the prediction by SPM (Whittle & Baligh, 1988)

gives ∆ui = 1.21σ'v0 or 8.7 kPa/m, and σ ri = 0.37σ'v0 or 2.7 kPa/m for an open-ended

pile with d/t = 40. There are no existing solutions in SPM for an open-ended caisson

with d/t = 60. According to Equation 2.14 in the CEM, the ∆ui generated for an area

ratio ρ = 0.066 (d/t = 60) is around 93% of the ∆ui for ρ = 0.1 (d/t = 40). By assuming

the same degree of reduction of ∆ui with ρ as that in the CEM, the SPM predictions can

be adjusted (approximately) for the actual caisson d/t of 60, resulting in ∆ui = 1.13σ′v0,

σ ri = 0.34 σ′v0, and σri – u0 = 1.47σ′v0. Therefore, the SPM predicts ∆ui as ~8.0 kPa/m

(or ~5.3su), σ ri ~ 2.5 kPa/m (or 1.7su) and σri – u0 ~ 10.5 kPa/m (or 7.0su), during the

caisson installation of test B13SCC in LOC clay. The SPM over-predicts by 33% both

the excess pore pressure and radial total stress relative to the hydrostatic pressure,

although the predicted radial effective stress matches well the derived value from

measurements.

According to Equations 2.16 - 2.18, and using YSR = 1.5, h = 4.8 m, deq = 0.92 m (thus

h/deq = 5.2), the MTD method predicts σri – u0, ∆ui and σ ri as 20.8 kPa/m (13.8su), 17.0



kPa/m (11.3su), 3.8 kPa/m (2.5su), respectively. The MTD method over-predicts the

measured σri – u0 and the derived ∆ui by 163% and 183% respectively. According to

the MTD prediction, (σri – u0)/σ′v0 is proportional to YSR0.41⋅(h/deq)-0.2, with a factor of

3.4 (see Equation 2.18). For the measurements in test B13SCC in the LOC clay, this

factor is 1.3, which is different from that of 1.5 measured in NC clay (see section

7.2.1.1). Both these values are obviously lower than the ratio of 3.4 given by Chow

(1997). Therefore, the relationship between σri – u0, OCR and h/deq obtained from

installation of the small closed-ended piles does not match that for thin-walled suction

caissons with the same equivalent diameter.

The measured σri – u0 during installation is compared with various theoretical

predictions discussed above for test B13SCC in LOC clay, as shown in Figure 7.40.

7.3.1.2 σri , σ ri and ∆ui during suction installation: test B13sus

Variations of the measured radial total stress during another suction installation test,

B13sus, are shown in Figure 7.41. The profile in Figure 7.41 is similar to the

observation in test B13SCC (see Figure 7.35). A quasi-linear relationship exists

between the measured radial total stress and the depth once the TPTs entered the soil at

4.8 m depth, except for the decreases at 12.77 m and 13.59 m.

Variations of the applied syringe pump pressure and the radial total stress versus depth

of installation are plotted in Figure 7.42. Jacked installation ended at 7.60 m, and

suction installation started at 7.97 m. The penetration rate (v, in model scale) of jacking

reduced to an average value of 0.55 mm/s, when the system was trying to reach the

planned stress-hold (see Figure 6.43). The corresponding normalised velocity, V, was

3.6, and thus penetration was partly drained in this transition zone. The time delay

when starting suction was 3.7 s (or 0.62 days prototype time) (see Figure 7.43).

The measured σri reduced slightly when the TPTs entered the suction-affected area as

the caisson penetrated 12.77 m (a further 4.8 m from 7.97 m where suction installation

started) (see Figure 7.42). This reduction was caused by consolidation due to both the

time delay and the reduced moving speed of the caisson.

It can be seen in Figures 6.43 and 7.43 that at z = 13.52 m the soil plug hit the top of the

caisson, and caused the syringe pump pressure to increase suddenly. Subsequently,

suction installation became very difficult after 13.59 m. Before 13.59 m, the caisson

was installed at a speed of 1.90 mm/s (in model scale), which corresponds to an



undrained penetration (V = 12.5) (see Figure 7.43). After 13.59 m, the installation

velocity (in model scale) decreased to 0.20 mm/s, which corresponds to a partly drained

penetration (V = 1.3). Accordingly, dissipation of excess pore pressures caused the

measured radial total stress to decrease (see Figures 7.42, 7.43).

The radial total stress relative to the hydrostatic pressure during installation, σri – u0, is

plotted against the depth of TPT (zTPT) within the soil in Figure 7.44. When jacking

(self-weight penetration) ended and suction started at zTPT = 2.80 - 3.17 m

(corresponding to z = 7.60 - 7.97 m), some trivial fluctuation occurred in the reading

due to consolidation during the time delay and slow penetration. At zTPT = 7.60 m when

the TPTs left the jacking-affected area, σri – u0 was 65.44 kPa. When the TPTs entered

the suction-affected area at zTPT = 7.97 m, σri – u0 decreased to 65.35 kPa, with an actual

reduction of 3 kPa, considering an increase in depth between these two points. Within a

discernible length of 0.82 m for the suction-affected area in this test, the measured

σri – u0 continued to increase at a similar gradient as that under jacked installation.

After the penetration rate (in model scale) reduced at zTPT = 8.79 m (corresponding to

z = 13.59 m), σri – u0 decreased 5.8 kPa within 0.13 m of slow movement, until the

installation finished. At the turning point zTPT = 8.79 m, the σri – u0 was 74.09 kPa, and

su was 15.03 kPa (see Figure 6.46), with a corresponding strength gradient, dsu/dz, of

1.71 kPa/m. Therefore, the gradient of the measured σri – u0 was 8.4 kPa/m, or ~4.7su.

The excess pore pressure during caisson installation, ∆ui, can be derived from the

measured σri by Equation 7.3. The α value is 0.38 (see Table 6.7) during installation of

test B13sus, with dsu/dz being 1.71 kPa/m, the average shaft friction gradient, αsu, is

obtained as 0.65 kPa/m. Using δr of 18.1°, the radial effective stress can thus be

derived from Equation 7.2 as 2.0 kPa/m, or 1.2su. Substituting this result into Equation

7.3 allows the gradient of excess pore pressure (∆ui) to be calculated as 6.4 kPa/m, or

~3.7su.

By using K0 = 0.70, St = 2 - 2.5, γ′ = 7.18 kN/m3 in Equation 2.10, the NGI method

predicts the gradient of excess pore pressure as 3.1 - 3.6 kPa/m, or (1.8 - 2.1)su, which is

48 - 56% of the measurement. According to Equation 2.11, the NGI prediction of the

radial effective stress is 2.1 - 2.6 kPa/m, or (1.2 - 1.5)su. The value of σri – u0 predicted

by the NGI method is 5.7 kPa/m or 3.3su, which is 68% of the measurement.

By adopting St = 2 - 2.5, K0 = 0.7, ρ = 0.066, and G/su = 100 - 150, the CEM prediction

of the gradient of ∆ui can be calculated by Equation 2.14, and the result is



6.1 - 7.4 kPa/m, or (3.6 - 4.3)su. According to Equation 2.11 and Equation 2.15, the

CEM prediction of the gradient of σ ri is 2.1 - 2.6 kPa/m, or (1.2 - 1.5)su, and that of

σri – u0 is 8.2 - 10.0 kPa/m or (4.8 - 5.8)su. It can be seen that the CEM prediction of the

excess pore pressure is close to 6.4 kPa/m derived from measurements, with a

difference of –3 - 16%.

Predictions of the NGI method and the CEM for σri – u0 are compared to the

measurements, as shown in Figure 7.45. When TPTs left the jacking-affected area at

zTPT = 7.60 m and entered the suction-affected area at zTPT = 7.97 m, the measured

σri – u0 decreased slightly, but still remains close to the lower bound CEM prediction,

and is obviously larger than the NGI prediction (with no transition zone assumed). The

trend of the measured σri – u0 was obviously different from that of the NGI prediction，

even allowing for a transition zone of one diameter in the soil below the point where

suction was initiated.

According to the analysis for test B13SCC, the SPM predicts ∆ui ~ 8.1 kPa/m (or 4.7su),

σ ri ~ 2.5 kPa/m (or 1.5su) and σri – u0 ~ 10.6 kPa/m (or 6.2su), during the installation

phase of test B13sus. The SPM over-predicts by ~27% on both ∆ui and σri – u0, while

the prediction of σ ri is close to the deduction from measurements.

The MTD predictions of σri – u0, ∆ui and σ ri can be calculated by Equations 2.16 - 2.18.

Using YSR = 1.5, h = 4.8 m and deq = 0.92 m (thus h/deq = 5.2), the MTD method

predicts σri – u0, ∆ui and σ ri as 20.7 kPa/m (12.3su), 17.1 kPa/m (10.0su) and 3.8 kPa/m

(2.3su), respectively. The MTD method over-predicts the measured σri – u0 and the

derived ∆ui by 146% and 167% respectively; this difference is consistent with that

obtained in test B13SCC.

Comparisons of the measured σri – u0 and theoretical predictions mentioned above for

test B13sus in LOC clay are shown in Figure 7.46.

7.3.1.3 σri , σ ri and ∆ui during suction installation: test B13cyc

Test B13cyc was installed by suction in the same box as tests B13SCC and B13sus, but

at a later time. The overall gradient of undrained soil strength, dsu/dz, in test B13cyc,

was 1.77 kPa/m, which is larger than the two former tests (see Figure 6.46). The depth

where the soil plug contacted the caisson lid (see Figure 6.44) is lower than those in the

other two tests, perhaps due to the higher strength ratio, su/σ′v0.



The measured radial total stress during installation (by suction) of test B13cyc is shown

in Figure 7.47. The profile is similar to those observed in tests B13SCC and B13sus

reported previously. Jacked installation ended at 7.69 m and suction installation started

at 7.85 m (see Figure 7.48). The penetration rate (in model scale) decreased to

0.65 mm/s (V = 4.3, partly drained) in this transition area; the time delay for starting

suction was 2.9 s (or 0.48 days) (see Figure 7.54). Consolidation due to the time delay

caused a reduction in the measured radial total stress, as the TPTs left the

jacking-affected area (z = 12.49 m) and entered the suction-affected area. At 12.76 m

another reduction occurred, this was due to consolidation when the penetration rate (in

model scale) decreased from 1.89 mm/s (V = 12.4, undrained) to 0.21 mm/s (V = 1.4,

partly drained) (see Figure 6.41), after the soil plug hit the caisson lid and stopped the

real suction installation (see Figure 6.44).

Variations of σri – u0 with the embedment of TPTs for test B13cyc are shown in Figure

7.49. When entering the suction-affected area from the jacking-affected area (between

zTPT = 7.69 m and 7.85 m), the actual decrease in the σri – u0 was 1.9 kPa, which is

rather small, and is caused by consolidation due to the time delay and reduced speed

when changing the installation system. After zTPT = 7.96 m, σri – u0 decreased again,

also due to consolidation when the penetration rate reduced to 0.21 mm/s (in model

scale, V = 1.4, partly drained).

It can be seen in Figure 7.49 that σri – u0 was 66.38 kPa when the TPTs entered the

suction-affected area (at zTPT = 7.85 m), with an overall stress gradient of 8.5 kPa/m.

According to Table 6.7, the α value was 0.40 during installation of test B13cyc, and γ′

was 7.21 kN/m3. The value of su was 11.36 kPa at 7.85 m, showing a dsu/dz of 1.45

kPa/m. Therefore, σ′ri can be calculated by Equation 7.2 as 1.8 kPa/m (or 1.2su), and

∆ui can be derived as 6.7 kPa/m (or 4.6su) by Equation 7.3. Following the same routine

of analysis as in test B13SCC, predictions of σri – u0, ∆ui and σ ri can be made from the

NGI method, CEM, SPM and MTD method, with values shown in Table 7.7.

Comparison between the measured σri – u0 and predictions from the NGI method and

the CEM are shown in Figure 7.50. It can be seen that measurements in the

suction-affected area were close to the lower bound CEM solution. After zTPT = 7.89 m,

the measurements deflected from the general trend by showing a decrease of 8.3 kPa

until the caisson stopped moving. However, the measured value needs 12.3 kPa of

further reduction to reach the average NGI prediction. The comparison shows that the



observed decrease in the measured radial total stress acting on the external wall at the

end of installation was the result of dissipation in ∆ui, due to the reduced speed at that

stage.

Comparison between the measured σri – u0 and various theoretical predictions is plotted

in Figures 7.51.

Table 7.7 Average measured (or derived*) and predicted stresses around the

caisson during installation in LOC clay (OCR = 1.5) (test B13cyc, γ =

7.21 kN/m3, dsu/dz = 1.45 kPa/m above the TPT)

0ri uσ −

(kPa/m)

riσ′

(kPa/m)

i∆u

(kPa/m) v0

0ri

σuσ

′−

v0

ri

σσ′′

v0

i

σ∆u

′

u

0ri

suσ −

u

ri

sσ′

u

i

s∆u

Measured or derived 8.5 1.8* 6.7* 1.2 0.2 0.9 5.9 1.2 4.6

SPM 10.6 2.5 8.1 1.5 0.3 1.1 7.3 1.7 5.6

MTD 21 3.9 17.1 2.9 0.5 2.4 14.5 2.7 11.8

NGI (LB) 5.8 1.8 3.6 0.80 0.25 0.50 4.0 1.2 2.5

NGI (UB) 5.8 2.2 4.0 0.80 0.31 0.55 4.0 1.5 2.8

CEM (LB) 7.4 1.8 5.6 1 0.2 0.8 5.1 1.2 3.9

CEM (UB) 9 2.2 6.8 1.2 0.3 0.9 6.2 1.5 4.7

Note: ‘LB’ for ‘Lower Bound’, ‘UB’ for ‘Upper Bound’, ‘*’ for derivation from measurements

7.3.1.4 σri , σ ri and ∆ui during jacked installation: test B13JCC

An individual jacked installation test B13JCC was carried out in the same LOC sample.

The caisson was installed by displacement control at a velocity of 2 mm/s, which is

close to the installation speed during suction installation. The normalised velocity was

thus 13.2, and therefore undrained penetration. The measured radial total stress σri

(averaged from the reading of the two TPTs) versus depth of the caisson is shown in

Figure 7.52; also plotted in the graph is σri measured during test B13SCC, of which the

soil strength gradient is close to that of test B13JCC. It can be seen in the graph that the

measured σri was close between jacked installation and suction installation. The

gradient of σri − u0 was 7.3 kPa/m from the jacked installation test B13JCC, and is close

to 7.9 kPa/m measured in suction installation test B13SCC. The close match (with a



difference of 4%) proved directly the similarity of the pattern of soil flow at the caisson

tip in LOC clay between the two types of installation.

Table 7.8 Average measured (or derived*) and predicted stresses around the

caisson during installation in LOC clay (OCR = 1.5) (test B13JCC, γ =

7.15 kN/m3, dsu/dz = 1.42 kPa/m above the TPT)

0ri uσ −

(kPa/m)

riσ′

(kPa/m)

i∆u

(kPa/m) v0

0ri

σuσ

′−

v0

ri

σσ′′

v0

i

σ∆u

′

u

0ri

suσ −

u

ri

sσ′

u

i

s∆u


SPM 10.5 2.4 7.9 1.47 0.34 1.1 7.4 1.7 5.6

MTD 20.8 3.8 16.8 2.91 0.53 2.35 14.6 2.7 11.8

NGI (LB)** 6.0 1.7 3.8 0.84 0.24 0.53 4.2 1.2 2.7

NGI (UB)** 6.0 2.2 4.3 0.84 0.31 0.60 4.2 1.5 3.0

CEM (LB) 7.3 1.7 5.6 1.02 0.24 0.78 5.1 1.2 3.9

CEM (UB) 8.9 2.2 6.7 1.24 0.31 0.94 6.3 1.5 4.7

Note: ‘LB’ for ‘Lower Bound’, ‘UB’ for ‘Upper Bound’, ‘*’ for derivation from measurements,

‘**’ assuming suction installation, not jacking.

The undrained strength gradient dsu/dz at the final embedment of TPTs (zTPT = 9.19 m)

in test B13JCC was 1.42 kPa/m (see Figure 6.46), while α was 0.42 and γ′ was 7.15

kN/m3 (see Table 6.7). Using similar steps as before, σ′ri can be obtained as 1.8 kPa/m

(or 1.3su); ∆ui is then 5.5 kPa/m, or 3.9su, which is very close to that (~4.0su) derived

from suction installation test B13SCC. The match of the excess pore pressure generated

on the external wall of the caisson during the two types of installation suggests that the

proportion of soil displaced outside at the caisson tip is similar for both.

Theoretical predictions by SPM, CEM, NGI and MTD for the external stress changes

around the caisson during installation of test B13JCC are also calculated using the

process stated previously, and are listed in Table 7.8. The measured σri – u0 and

theoretical predictions are compared in Figures 7.53. It can be seen in the graph that

below the embedment of 6 m, the measured value is close to the lower bound CEM

prediction, and is obviously higher than the NGI predictions. The derived ∆ui is close



to the lower bound CEM prediction, and is 36% larger than the upper bound NGI

prediction, as expected, given that the installation was entirely by jacking.

7.3.1.5 Summary

Based on the above analysis, it can be concluded that during the penetration of suction

caissons in LOC clay, the radial stress changes on the external wall of the caisson are

similar for both jacked installation and suction installation. Reduction of the radial total

stress observed during suction installation is caused by a decrease in the penetration rate

and thus partial consolidation. Comparison between the measured external radial total

stresses (or derived excess pore pressures) during installation and theoretical predictions

suggests that the assumption all the soil particles displaced by the caisson wall move

inside the caisson during suction installation is untrue. This agrees with the result

derived from the soil heave during suction installation (see Chapter 6). A simple form

of cavity expansion method (CEM) provides reasonable predictions of the external

radial stress changes during installation of suction caissons. The NGI method

under-predicts the results, the SPM over-predicts the measurements, while the MTD

method tends to introduce large over-predictions. These results are consistent with

those observed in the NC clay.

Analysis will be continued on the external radial stress changes during the subsequent

consolidation after installation.

7.3.2 Relaxation of Radial Stresses during Consolidation

After the caisson was installed to the target depth, for jacked installation the axial force

was reduced to the nominal self-weight of 30 N, while for suction installation the

syringe pump was stopped immediately, maintaining the self-weight load.

7.3.2.1 t50 and t90

Variations of the depth of the caisson tip and the axial force during consolidation in

LOC clay are shown in Figure 7.54 (with units in model scale) for suction test B13SCC.

After 1 hour (model time, representing 1.7 years at 120 g) of consolidation, the

observed settlement of the caisson was 0.20 mm (24 mm at prototype scale), which is

slightly larger than that (0.12 - 0.13 mm) in NC clay, possibly due to a larger stress-hold



value applied here. Judging from the shape of the embedment curve with time, the

settlement was not completed after one hour of consolidation (in model scale), since the

depth of the caisson tip was still increasing, probably due to ongoing secondary

consolidation in the soil. This indicates that more time should be allowed to achieve

full consolidation of the soil sample prior to installing the caisson.

Variations of depth of the caisson tip and the applied axial force during consolidation of

test B13JCC are depicted in Figure 7.55. Altogether it settled 0.18 mm (22 mm at

prototype scale). This value is very close to that observed in the suction test B13SCC,

after the same consolidation time. Again, the settlement of the caisson developed

almost linearly with time, without reaching a stable state at the end of consolidation

time, indicating that secondary consolidation was still ongoing.

Results for suction installation tests B13sus and B13cyc are plotted in Figures 7.56 and

7.57, respectively. In these two tests, the settlement reached a quasi-stable state after 1

hour of consolidation (in model scale). Therefore, times for 50% (t50) and 90% (t90) of

consolidation can be derived from the settlement, or at least they can be viewed as

lower bound values. For test B13sus, times of t50 and t90 were around 517 seconds and

2101 seconds, corresponding to prototype times of 2.9 months and 11.7 months. In test

B13cyc, times for t50 and t90 were 736 seconds and 2562 seconds, corresponding to 4.1

months and 14.2 months at prototype scale (see Table 7.9). The average t50 and t90

derived from the measured settlement of caissons in LOC clay were thus 3.5 and 13.0

months, respectively.

Table 7.9 Measured 50% and 90% consolidation time in LOC clay

By embedment z By average (σr – u0 )

t50 t90 t50 t90

Test

Model

(s)

Proto.

(month)

Model

(s)

Proto.

(month)

Model

(s)

Proto.

(month)

Model

(s)

Proto.

(month)

B13SCC - - - - 291 1.6 2755 15.3

B13JCC - - - - 986 5.5 2053 11.4

B13sus 517 2.9 2101 11.7 28 0.2 1980 11.0

B13cyc 736 4.1 2562 14.2 208 1.2 1891 10.5

Average 627 3.5 2332 13.0 378 2.1 2244 12.1



Variations of the external radial stress of the caisson during consolidation were

monitored by the two TPTs located at L/3 (40 mm at model scale, where L is the length

of the caisson) from the tip of the caisson. In tests B13SCC, B13JCC, B13sus and

B13cyc, the radial total stress relative to the hydrostatic pore pressure, σr – u0, decreased

gradually with the consolidation time in LOC clay (Figures 7.58 - 7.61), as observed in

NC clay. For test B13SCC, σr – u0 did not reach a stable state at the end of the

consolidation time; analysis of the plots indicates that a further relaxation of ~2 kPa

would have occurred, if more time were allowed for consolidation. For the other three

tests, the expected state was reached, since the trend was almost flat by the end of the

consolidation period. Reductions in σr – u0 during consolidation were 12.06, 13.24,

15.20, 9.53 kPa, respectively for the above four tests, with very close values for the

jacked caisson (13.24 kPa) and the suction caisson (averaged at 12.26 kPa). The

decrease in the σr – u0 amounts to around one fifth of the initial value immediately after

installation. The relaxation ratio is therefore smaller compared to the tests in NC clay.

Times corresponding to 50% and 90% consolidation in terms of σr – u0 are listed in

Table 7.9. According to previous analysis, these times could be viewed as lower bound

values. It took around 2 months and 12 months to complete 50% and 90% radial

consolidation in LOC clay respectively (prototype scale). These measured t50 and t90 are

very close to those (3 months and 11 months, respectively) obtained in NC clay (see

Table 7.3).

Taking cv as 2.4 m2/yr, and ch ~ 3cv (Fahey & Lee Goh, 1995), and T50 ~ 1 and T90 ~ 10

(Randolph, 2003), the theoretical t50 and t90 would be 1.4 months and 14 months,

respectively. The theoretical 90% consolidation time is in reasonable agreement with

the measured value shown in Table 7.9, for caissons installed either by jacking, or by

suction. Again these measured times are much larger than the corresponding times of

~1 day (50%) and 6 days (90%) for the NC kaolin clay, and ~2 days (50%) and 20 - 40

days (90%) for the LOC clay in the Gulf of Mexico suggested by the NGI method (see

Table 2.4) (Andersen & Jostad, 2002). The significant consolidation times outside the

caisson suggests some outward motion of the soil particles during suction installation in

LOC clay.




The final radial effective stress ratio, Kc, defined as σ'rc/σ'v0, and the coefficient Hi,

defined as (σri – u0)/σ'v0, are derived from the measurements, and are summarised in

Table 7.10 for tests B13SCC, B13JCC, B13sus and B13cyc (see Figures 7.58 - 7.61).

Table 7.10 Measured radial stress changes, Hi and Kc values during consolidat-

ion in LOC clay

Test i,TPTz

(m) 0ri uσ −

(kPa) v0

0ri

σuσ

′− c,TPTz

(m) 0rcrc uσσ −=′

(kPa) v0

rc

σσ

′′

0ri

rc

uσσ−′

B13SCC 9.10 61.57 0.95 9.12 49.51 0.76 0.80

B13JCC 9.19 61.26 0.93 9.20 48.02 0.73 0.78

B13sus 8.92 68.31 1.07 8.94 52.57 0.82 0.77

B13cyc 8.73 59.50 0.95 8.79 52.38 0.83 0.87

Average

(by suction) 8.92 63.13 0.99 8.95 51.49 0.80 0.81

Average

(by jacking) 9.19 61.26 0.93 9.20 48.02 0.73 0.78

Average

(all) 8.99 62.66 0.97 9.01 50.62 0.78 0.81

It can be seen in Table 7.10 that the average Hi is 0.97, while the average final radial

effective stress ratio, Kc, is 0.78. Comparison between caissons installed by jacking and

by suction shows a difference of less than 10% for both the measured Hi and Kc. The

derived Kc value is slightly higher (11%) than the estimated in situ earth pressure

coefficient, K0, of 0.70. If more consolidation time were allowed in these tests,

especially the earlier ones, the difference between the observed Kc and the theoretical

K0 value might be even smaller. Table 7.10 gives an average ( )0rirc uσσ −′ of 0.81 in

these four tests, the average stress relaxation for caisson tests in LOC clay was thus

19%, which was lower than for full-displacement piles suggested by Lehane & Jardine

(1994). The stress relaxation seems to decrease with the increase in OCR of the soil,

since a relaxation of 37% was observed for tests in NC clay. This is reasonable, since



clay with higher OCR tends to generate less excess pore pressure during penetration,

thus causing less relaxation during subsequent consolidation.

According to CEM (Randolph, 2003), Kc can be estimated by Equation 2.20, with

R = 1.5, λ = 1 and µ = 5. Using the CEM predictions shown previously for the four

tests in LOC clay, 0vri σσ ′′ and 0vi σ∆u ′ average at 0.28 - 0.35 and 0.83 - 1.00,

respectively. The corresponding lower and upper bound Kc values can be obtained as

0.68 and 0.79, with an average value of 0.74. The CEM predictions of Kc differ from

the measured value by –13 - 1% only. This agreement is consistent with that reported

previously in NC clay.

The MTD prediction of Kc can be obtained using Equation 2.19. Considering h/deq =

5.2 for caisson 2, YSR = 1.5 and St = 2 - 2.5 for the LOC clay tested here, the calculated

lower bound and upper bound Kc values are 1.39 and 1.46, respectively. The average

Kc predicted by MTD, however, over-predicts the measured value by more than 85%.

This is consistent with its obvious over-prediction of radial stresses during installation.

Such a large discrepancy suggests that MTD is inappropriate to analyse the stress

changes around thin-walled suction caissons in LOC clay.

The NGI method does not give expressions for estimating Kc, and thus will not be

discussed here.


During monotonic uplift after consolidation, the radial stress dropped 3 kPa within

0.03 m of extraction for the suction-installed test B13SCC (see Figures 7.62), and 4 kPa

within 0.01 m for the jacked installation test B13JCC (see Figure 7.63). Comparison of

the average measurements from the above two tests shows that the radial stress

decreased first, then increased slightly after reaching its lowest value, before decreasing

almost linearly with the movement of the caisson during most of the following

extraction from the clay (Figure 7.64). The profiles are very similar between the two

tests installed by different methods. During pullout, the gradient of radial total stress

versus depth is lower compared to that during installation. This indicates that extensive

relaxation has occurred in the surrounding clay after consolidation, and suggesting that

the excess pore pressure generated is rather limited during extraction. When the TPTs

re-entered the water from the soil, their readings coincided with the hydrostatic pressure



curve; this agreement again proved the reliability of the TPT readings throughout the

whole process.

Variations of the radial total stress relative to hydrostatic pressure, σr – u0, with

movement of the caisson during the early stage of the pullout are shown in Figures 7.65

and 7.66 for tests B13SCC and B13JCC, respectively. Taking the excess pore pressure

as zero when the caisson is loaded to failure, σ rf can be calculated as σrf – u0, and the

results are 43.68 kPa and 37.94 kPa for tests B13SCC and B13JCC. The corresponding

depths at failure for the above two tests are shown in Table 7.11. Also shown are the

derived external α values from the measured σ rf by Equation 2.27. The α values thus

derived are 1.05 and 0.86 for tests B13SCC and B13JCC, with an average value of 0.96.

It should be noted that this α value is an upper bound. Recalling the lower bound α of

0.73 derived from the uplift capacity in the last chapter, the external α can be estimated

as 0.73 - 0.96, with an average value of 0.85.

Table 7.11 Upper bound external shaft friction ratio α when caisson is loaded to

failure in LOC clay

Test Depth of tip Depth of TPT σ rf su α

(m) (m) (kPa) (kPa)

B13SCC 13.50 8.70 43.68 13.6 1.05

B13JCC 13.75 8.95 37.94 14.0 0.86

Average 13.82 9.02 42.40 14.1 0.96

According to the estimated Kc values, α values during pullout can be predicted by the

CEM and MTD approaches, using Equation 2.26 shown in the Chapter 2. The

reduction factor is taken as the same value as that in NC clay, 20%, since it is

considered to be independent of the oversonsolidation ratio of the soil (Jardine & Chow,

1996). The α value thus predicted by CEM is 0.74 - 0.86, with an average value of

0.80, and 1.53 - 1.61 by the MTD, with an average value of 1.57. The API prediction

can be obtained by using Equation 2.24, since su/σ v0 = 0.23 for the LOC kaolin clay,

giving external α = 1. The NGI method presents an α value of 0.65 for general cases

when specific soil conditions are unavailable (Andersen & Jostad, 2002), according to

Table 2.5. A summary of the measured and predicted αext values during caisson pullout



in LOC clay are shown in Table 7.12. Compared to the average measured α value of

0.85 in LOC clay as stated above, the average CEM prediction under-predicts by around

6%; the MTD approach over-predicts by 86%; the API prediction is some 20% higher,

while the NGI method under-predicts by 20 - 30%.

Table 7.12 Measured and predicted external shaft friction ratio α during pullout

of the caisson after consolidation in LOC clay

Method α


API (RP2A) 1.0

NGI method 0.65

CEM 0.74 - 0.86

MTD approach 1.57

7.4 EXPERIMENTS IN SENSITIVE CLAY

Radial stress changes around the caisson were measured during installation,

consolidation and pullout of caissons in sensitive clay. As described previously, the

sensitivity of these samples was 4 - 5, with an OCR of 1. All the tests in sensitive clay

were installed by suction at 120 g, using the model caisson 2; the soil properties have

been listed in Table 3.3. The distance between TPTs and the caisson tip is 40 mm at

model scale (or 4.8 m prototype length), as described previously. The test results are

discussed below.


7.4.1.1 σri , σ ri and ∆ui during installation: test B14cyc

Before the installation started, the caisson tip was just above the mudline, the distance

between the TPTs and the mudline was thus 40 mm in model scale (or prototype 4.8 m

at 120 g). As described in Chapter 6, the caisson was first installed by self-weight

penetration to more than half the length (60 - 70 mm model scale, or 7.2 - 8.4 m in

prototype scale) of the caisson. Located 40 mm (4.8 m at prototype scale) from the



caisson tip, the TPTs entered the soil before the suction installation commenced. The

recorded radial total stress σri versus penetration of the caisson for test B14cyc is shown

in Figure 7.67. Variations of the radial total stress during installation are similar in

shape to those observed in NC and LOC clays. In the first 4.8 m of penetration, the

TPTs recorded the hydrostatic pressure. The TPTs entered the clay at 4.8 m depth, and

the gradient of the TPT readings suddenly surged; subsequently the readings increased

almost linearly with penetration depth, except for some decrease at 12.25 m, 14.20 m

and 15.11 m. It has been pointed out in Chapter 6 that although the syringe pump was

stopped before 14.4 m (the full length of the caisson), obvious secondary settlement

occurred when the load cell was trying to track the target stress-hold value; the caisson

was thus overdriven to a depth of 15.24 m.

Variations of the applied syringe pump pressure and the measured radial total stress

with respect to penetration depth are shown in Figure 7.68. It can be seen that

self-weight penetration (jacked installation) ended at the penetration depth of 7.21 m,

and suction installation commenced at 7.45 m. The time delay was 1 s (or 0.2 days

prototype time), and the penetration rate reduced to 0.81 mm/s (in model scale), which

corresponds to a normalised velocity, V (V = vt/cv, cv is 0.063 mm2/s, or 2.0 m2/year),

of 6.5, thus partly drained (see Figure 7.69), according to the limits recommended by

Randolph (2004). When the TPTs entered the suction-affected area at 12.25 m depth

(12.25 m = 7.45 m + 4.8 m), the measured σri decreased slightly due to consolidation

during the time delay when starting suction installation, and reduced speed of

penetration.

At 14.20 m the radial total stress reduced slightly (see Figure 7.69) as the installation

speed reduced due to the soil plug blocked the pneumatic venting of the caisson cap,

and caused the installation essentially ceased. It can be seen that before 14.20 m, the

caisson was installed by suction at a velocity (in model scale) of 2.18 mm/s which

corresponds to a normalised velocity (V) of 17, thus undrained. While after 14.20 m the

velocity decreased to 0.09 mm/s, the corresponding V of 1.8 indicates a partly drained

penetration. Therefore, dissipation of the excess pore pressure caused the reduction in

σr. This is in agreement with observations for tests in NC and LOC clays. Another

slight decrease in σr occurred at 15.11 m, when the penetration velocity of the caisson

decreased further to 0.01 mm/s (V = 0.08, thus drained), and creep is considered to have

occurred in this stage.



Variations of σri – u0 versus embedment of TPTs (zTPT, 4.8 m less than embedment of

the caisson) in test B14cyc are shown in Figure 7.70. It can be seen in the graph that

σri – u0 decreased slightly between zTPT = 2.41 m and 2.65 m, caused by consolidation

due to the time delay when starting suction, and the reduced speed of penetration. At

the end of the jacking phase (zTPT = 7.21 m), the TPTs gave σri – u0 of 46.09 kPa. Once

suction had commenced at zTPT = 7.45 m, the measured σri – u0 rose to 46.68 kPa.

Considering an increase in lateral earth pressure of around 1.9 kPa due to the increase in

the embedment of TPTs, the actual σri – u0 decreased 1.3 kPa (see Figure 7.70) due to

consolidation in this transition area. In the subsequent suction-affected area, σri – u0

continued to increase at a similar gradient as for the jacking-affected area; this indicates

that the pattern of soil flow at the caisson tip is similar for these two types of installation.

The reduction of σri – u0 at zTPT = 9.40 m (or z = 14.20 m) is considered to be caused by

consolidation when penetration essentially stopped.

At the turning point (zTPT = 9.40 m) in the suction-affected area, the measured σri – u0

was 62.64 kPa, corresponding to a stress gradient of 6.7 kPa/m. While the

corresponding su of the soil was 11.63 kPa, with a dsu/dz of 1.24 kPa/m (see Figure

6.60). The measured gradient of σri – u0 was thus ~5.4su for test B14cyc.

The excess pore pressure generated outside the caisson wall during installation, ∆ui, can

be derived from σri from Equation 7.3. The α value during installation of test B14cyc is

0.15 (see Table 6.10); the average shaft friction gradient, αsu, is thus 0.19 kPa/m, while

the residual interface friction angle between the sensitive clay and the caisson obtained

from the ring shear tests is 11.7° (see Table 5.1). The radial effective stress can thus be

derived from Equation 7.2 as 0.92 kPa/m, or 0.7su. Substituting this into Equation 7.3

allows the excess pore pressure gradient to be calculated as 5.8 kPa/m, or ~4.7su, which

is the close to the average value of 4.6su obtained in NC clay, but is slightly larger than

the average value of 4.1su in LOC clay, indicating that the ∆ui generated during

penetration decreases mainly with the increase in OCR of the soil, but is less affected by

the soil sensitivity. Following the same routine as used for analysing tests in NC clay,

theoretical predictions by SPM, CEM, NGI and MTD will be presented, and compared

with those derived from measurements by the TPTs, as described below.

The value of ∆ui acting on the external wall of the caisson during suction installation

can be predicted by the NGI method (Andersen & Jostad, 2002) using Equation 2.10.

Substituting K0 = 0.55, St = 4 - 5 and γ′ = 7.30 kN/m3 into this formula, the NGI method



predicts a gradient of excess pore pressure of 3.4 - 3.7 kPa/m, or (2.7 - 3.0)su, which is

59 - 64% of the value derived from measurements. According to Equation 2.11, the

NGI prediction of σ ri is 1.2 - 1.5 kPa/m, or (1.0 - 1.2)su, while its prediction of σri – u0

is 4.9 kPa/m, or (4.0 - 4.5)su. This is 73% of the measurement, and the result is similar

to those in the NC and LOC clays.

The CEM prediction of ∆ui is obtained by Equation 2.14, using St = 4 - 5, K0 = 0.55,

ρ = 0.066, γ′ = 7.30 kN/m3 and G/su = 50 - 100 (a lower range than those for NC and

LOC clays, taking account of the soft nature of the sensitive clay). The gradient of ∆ui

may be estimated as 5.3 - 6.4 kPa/m, or (4.3 - 5.2)su. The CEM predictions of σ ri and

σri – u0 can be derived from Equation 2.11 and Equation 2.15, and the results are

1.2 - 1.5 kPa/m, or (1.0 - 1.2)su, for σ ri, and 6.5 - 7.9 kPa/m, or (5.2 - 6.4)su, for σri – u0.

It can be seen that the CEM prediction of the excess pore pressure is close to 5.6 kPa/m

derived from measurements, with a difference of –8 - 11%. This accuracy is consistent

with those in NC and LOC kaolin clays.

The measured σri – u0 is compared with the upper and lower predictions of the NGI

method and CEM during installation of test B14cyc in Figure 7.71. Also plotted in the

graph is the trend of the measured σri – u0 by following NGI’s assumption of a one

diameter transition zone (Andersen & Jostad, 2002). At zTPT = 7.45 m when the TPTs

entered the suction-affected area, the measured σri – u0 decreased slightly (~1.3 kPa)

due to consolidation, but it remained close to the lower bound CEM prediction. During

a further penetration of 1.95 m in the suction-affected area, the measured σri – u0 was

always close to the lower bound CEM prediction, and was well above the NGI

prediction. This comparison shows that the external radial total stress was essentially

unaffected by the applied suction. The measured σri – u0 and the NGI prediction, which

assumes a one diameter transition zone below the depth of starting suction installation,

was obviously different, suggesting that the approach may not be valid.

No direct solutions have been presented by the SPM on the stress changes for

open-ended piles in clay with a sensitivity of 4 - 5. However, expressions for Boston

Blue Clay (BBC), which has a sensitivity of 7 and OCR = 1 and thus are somewhat

similar to the sensitive clay used here, were put forward by Whittle & Baligh (1988).

Their SPM solution gives ∆ui = 1.20σ'v0 or 8.8 kPa/m, and σ′ri = 0.08σ′v0 or 0.58 kPa/m

for an open-ended pile with d/t = 40. According to CEM, ∆ui generated by an open-

ended pile with d/t = 60 (ρ= 0.066) is around 93% of that of the pile with d/t = 40 (ρ=



0.1). Adjusting this (approximately) for the actual caisson d/t of 60 would result in ∆ui

= 1.12σ'v0, σ′ri = 0.07σ'v0, and σri – u0 = 1.19σ'v0. Hence SPM predicts ∆ui as

~8.2 kPa/m (or ~6.7su), σ′ri ~ 0.5 kPa/m (or 0.4su) and σri – u0 ~ 8.7 kPa/m (or 7.1su),

during installation of test B14cyc in sensitive clay. The SPM over-predicts by 42% and

30% respectively the derived excess pore pressure and the measured radial total stress

relative to hydrostatic pressure, although its prediction of the radial effective stress

matches well with the derived value from measurements. It should be noted that the

SPM prediction in sensitive clay here is rough, due to the uncertainty of the difference

between the sensitive kaolin clay used in experiments and the Boston Blue Clay used

for analysis.

The MTD predictions of σri – u0, ∆ui and σ′ri can be obtained from Equations

2.16 - 2.18 (Lehane, 1992; Jardine & Chow, 1996). For model caisson 2 tested in

sensitive clay, the parameters needed for calculation are YSR = 1.0, h = 4.8 m,

deq = 0.92 m, with h/deq = 5.2. The MTD predictions of σri – u0, ∆ui and σ′ri are 17.9

kPa/m (14.4su), 14.2 kPa/m (11.5su) and 3.7 kPa/m (3.1su), respectively. The MTD

method over-predicts the measured σri – u0 and the derived ∆ui by 167% and 146%

respectively. The difference is similar to those in NC clay, showing the risk of applying

the MTD method to the analysis of suction caissons in sensitive clay.

For test B14cyc during caisson installation by suction in sensitive clay, comparisons of

the measured σri – u0 and predictions of the NGI method, CEM, SPM and the MTD

method are presented in Figure 7.72.

7.4.1.2 σri , σ ri and ∆ui during suction installation: test B14susa

Variations of the measured radial total stress during installation (by suction) of another

test, B14susa, are shown in Figure 7.73, which shows a similar profile as that of test

B14cyc (see Figure 7.67). The measured radial total stress increased quasi-linearly with

the depth (except for the decreases at 11.54 m and 14.20 m), once the TPTs entered the

soil after a caisson penetration of 4.8 m. Variations of the applied syringe pump

pressure and the radial total stress, versus depth of installation are plotted in Figure 7.74.

Self-weight penetration (jacked installation) ended at 6.34 m, while suction installation

started at 6.74 m, with a slight reduction in the radial total stress observed during

transition. The rate of jacked installation decreased towards the end of self-weight

penetration (see Figure 7.75) with a final rate of only 0.4 mm/s (in model scale). This,



together with the time delay of 5.1 s (or 0.85 days prototype time) for starting suction

installation, allowed some consolidation prior to commencing undrained installation by

suction. When the TPTs entered the suction-affected area at 11.54 m (equals 6.74 m +

4.8 m), a small reduction occurred in the measured σri. At z = 14.15 m, the measured

σri decreased again, corresponding to a surge in the applied syringe pump pressure (see

Figure 7.74). This reduction corresponds to a decrease in the installation speed, after

the soil plug contacted the caisson lid at 13.45 m (see Figure 6.58). Before 14.15 m, the

caisson was installed at a velocity (in model scale) of 1.34 mm/s, and V ~ 11, thus

undrained (see Figures 7.75 and 6.54). After 14.20 m the speed decreased to 0.02 mm/s,

which corresponds to a partly drained penetration (V = 0.2). Dissipation of the excess

pore pressures is considered to be the main reason for reduction of the measured σri.

The variation of σri−u0 is plotted against the embedment of TPT (zTPT) in soil in Figure

7.76. In the suction-affected area (zTPT > 6.74 m), the measured σri−u0 continued to

increase at a slightly smaller gradient than that in the previous jacked (self-weight)

installation. Again, this corresponds to a lower penetration rate in the suction-affected

area versus that in the jacked area (see Figure 7.76). Similar to tests in LOC clay,

consolidation is considered to partially account for the slight difference mentioned here.

At the turning point (zTPT = 9.35 m) below which the penetration rate reduces, σri−u0

was 60.44 kPa, with a stress gradient of 6.5 kPa/m, or 4.8su, since the corresponding su

was 12.62 kPa (see Figure 6.60) and thus dsu/dz was 1.35 kPa/m.

The excess pore pressure generated outside the caisson wall during installation, ∆ui, is

calculated by Equation 7.3. The α value was 0.18 (see Table 6.10) during installation of

test B14susa, the average shaft friction gradient, αsu, was thus 0.24 kPa/m (with dsu/dz

= 1.35 kPa/m). Using δr as 11.7° (see Table 5.1), the radial effective stress can thus be

derived from Equation 7.2 as 1.2 kPa/m, or 0.9su. Substituting this into Equation 7.3

allows the gradient of excess pore pressure, ∆ui, to be calculated as 5.3 kPa/m, or ~3.9su.

Using K0 = 0.55, St = 4 - 5, γ′ = 7.30 kN/m3 (see Table 3.3) in Equation 2.10, the NGI

method predicts the gradient of excess pore pressure as 3.2 - 3.6 kPa/m, or (2.4 - 2.7)su,

which is 60 - 68% of the measurement. According to Equation 2.11, the NGI prediction

of σ ri is 1.3 - 1.6 kPa/m, or (1.0 - 1.2)su. The predicted σri – u0 by the NGI method is

4.9 kPa/m or 3.6su, which is 75% of the measurement.

By adopting St = 4 - 5, K0 = 0.55, ρ = 0.066, γ′ = 7.3 kN/m3 and G/su = 50 - 100, the



CEM prediction of the gradient of ∆ui may be estimated by Equation 2.14, and the

result is 5.4 - 6.6 kPa/m, or (4.0 - 4.9)su. According to Equation 2.11 and Equation

2.15, the CEM predicts the gradient of σ ri as 1.3 - 1.6 kPa/m, or (1.0 - 1.2)su, and that

of σri – u0 as 6.7 - 8.2 kPa/m or (5.0 - 6.1)su. It can be seen that the CEM prediction of

the excess pore pressure is just above the measured value of 5.3 kPa/m.

Predictions of the NGI method and the CEM for σri – u0 are compared with the

measurements, as shown in Figure 7.77. The NGI prediction which assumes a

transition zone in the suction-affected area was also presented, as shown by the thick

arrow. When the TPTs entered the suction-affected area at zTPT = 6.74 m, the measured

σri – u0 decreased slightly due to partial consolidation, but is clearly larger than the NGI

prediction, showing that the influence of the applied suction during installation is very

small. The gradient of the measured σri – u0 is also very different from that of the NGI

prediction assuming a transition zone; this observation is consistent with the previous

test in sensitive clay and tests in LOC clay. Below zTPT = 9.35 m, the moving speed (in

model scale) of the caisson changed from 1.34 mm/s to 0.02 mm/s, accompanied by an

immediate decrease in the measured σri – u0. After decreasing 6.7 kPa within 0.57 m of

slow movement, the measured σri – u0 reached the average NGI assumption. However,

the reduction of σri – u0 at this moment is ascribed to the dissipation in ∆ui, rather than a

change to totally inward movement of the soil.

As discussed previously for caissons in sensitive clay, the SPM predicts ∆ui = 1.12σ'v0,

σ ri = 0.07σ'v0, and σri – u0 = 1.19σ'v0, which leads to ∆ui ~ 8.2 kPa/m (or 6.1su),

σ ri ~ 0.5 kPa/m (or 0.4su) and σri – u0 ~ 8.7 kPa/m (or 6.4su), during installation of test

B14susa. The SPM over-predicts ∆ui by 55% and σri – u0 by 34%, while the prediction

of σ ri is close to the derived value from measurements.

The MTD predictions of σri – u0, ∆ui and σ ri can be calculated by Equations 2.16 - 2.18.

Using YSR = 1.0, h = 4.8 m, deq = 0.92 m, giving h/deq = 5.2, the MTD method predicts

σri – u0, ∆ui and σ ri as 18.0 kPa/m (13.3su), 14.2 kPa/m (10.5su), 3.8 kPa/m (2.8su),

respectively. The MTD method over-predicts the measured σri – u0 and the derived ∆ui

by 168% and 177% respectively; the difference is consistent with that observed in test

B14cyc. Values of σri – u0 measured in test B14susa and various theoretical predictions

mentioned above are compared in Figure 7.78.



7.4.1.3 σri , σ ri and ∆ui during suction installation: test B14SCC

Test B14SCC was installed in the same box as tests B14cyc and B14susa, but at an

earlier time. The measured σri is averaged from the readings of two TPTs, as shown in

Figure 7.79. Due to lack of experience at the beginning of tests in the sensitive clay, the

self-weight penetration depth of 8.90 m (see Figure 7.80) was greater than expected

(7 - 8 m). Suction-installation started at 9.12 m, causing a reduced length of suction

installation. The measured σri experienced a slight reduction at 13.92 m when the TPTs

entered the suction-affected area, and this was caused by the partial consolidation

during the time delay (1.8 s model time, or 0.3 days prototype time) for initiating

suction installationm and the penetration with a reduced speed (see Figures 7.81 and

6.51). At the end of installation, upheaval of the soil plug led to contact with the

caisson lid and subsequently caused the penetration rate to decrease at 14.57 m, and

therefore the measured σri reduced due to consolidation (see Figure 7.81).

The measured σri – u0 is plotted against the embedment of TPTs in Figure 7.82 for test

B14SCC. At the point (zTPT = 9.77 m) where the penetration rate reduced afterwards,

σri – u0 was 61.44 kPa, showing that the gradient of σri – u0 in the suction-affected area

was 6.3 kPa/m, or 6.4su, since su was 9.67 kPa at 9.77 m depth (see Figure 6.60), and

thus dsu/dz was 0.99 kPa/m.

Table 7.13 Measured (or derived) and predicted stresses around the caisson

during installation in sensitive clay (St = 4 - 5) (test B14SCC, γ = 7.30

kN/m3, dsu/dz = 0.99 kPa/m above the TPT)

0ri uσ −

(kPa/m)

riσ′

(kPa/m)

i∆u

(kPa/m) v0

0ri

σuσ

′−

v0

ri

σσ′′

v0

i

σ∆u

′

u

0ri

suσ −

u

ri

sσ′

u

i

s∆u


SPM 8.7 0.5 8.2 1.19 0.07 1.12 8.8 0.5 8.3

MTD 17.9 3.8 14.1 2.45 0.52 1.93 18.1 3.8 14.2

NGI (LB) 4.9 0.9 3.7 0.67 0.12 0.51 4.9 0.9 3.7

NGI (UB) 4.9 1.2 3.9 0.67 0.16 0.54 4.9 1.2 3.9

CEM (LB) 5.9 0.9 5 0.81 0.12 0.68 6 0.9 5.1

CEM (UB) 7.1 1.2 5.9 0.97 0.16 0.81 7.2 1.2 6

Note: ‘LB’ for ‘Lower Bound’, ‘UB’ for ‘Upper Bound’, ‘*’ for derivations from measurements.



Using the α value of 0.16 and γ′ of 7.3 kN/m3 in Table 6.10, and dsu/dz of 0.99 kPa/m,

σ′ri was calculated as 0.8 kPa/m (or 0.8su), while ∆ui was derived as 4.5 kPa/m (or 5.6su),

following the calculation process stated in test B14cyc. Predictions of σri – u0, ∆ui and

σ ri from the NGI method, CEM, SPM and MTD method are shown in Table 7.13. The

measured σri – u0 and various theoretical predictions are compared in Figure 7.83. It

can be seen that the measurements were just above the lower bound CEM solution in

the suction-affected area, except at the end of installation where significant

consolidation occurred due to reduced penetration rate.

7.4.1.4 σri , σ ri and ∆ui during suction installation: test B14sus

Variations of the radial total stress measured during the installation of test B14sus are

shown in Figure 7.84, while the applied syringe pump pressure during installation is

shown in Figure 7.85. Over-driving during self-weight penetration (to 9.27 m here

versus 7 - 8 m normally) also occurred in this test. Suction installation started at 9.40 m,

the distinguishable suction-affected area is only 0.25 m, ranging between 14.20 m (9.40

+ 4.8 m) and 14.45 m when σri decreased. The reduction of σri at 14.45 m (see Figure

7.84) was caused by consolidation once the installation speed decreased (see Figure

7.86).

Variations of σri – u0 with embedment of TPTs in the suction-affected area in test

B14sus gave a gradient of 6.7 kPa/m, with a corresponding dsu/dz of 1.11 kPa/m (see

Figure 6.60) in the soil. According to Table 6.10, α was 0.15 and γ′ was 7.3 kN/m3.

Using the calculation process described in test B14cyc, σ′ri was derived as 0.8 kPa/m (or

0.7su), while the deduced ∆ui was 5.9 kPa/m (or 5.3su). Predictions of the NGI method,

CEM, SPM and MTD method of σri – u0, ∆ui and σ ri follow the process described in

the previous test, and the results are presented in Table 7.14. The measured σri – u0 and

predictions from various theoretical predictions are compared in Figure 7.87 for test

B14sus. It can be seen that in the suction-affected area the measurements stayed

between the lower bound and upper bound CEM predictions, except at the end of

installation where they decreased significantly due to consolidation as the penetration

rate reduced.



Table 7.14 Measured (or derived*) and predicted stresses around the caisson

during installation in sensitive clay (St = 4 - 5) (test B14sus, γ = 7.30

kN/m3, dsu/dz = 1.11 kPa/m above the TPT)

0ri uσ −

(kPa/m)

riσ′

(kPa/m)

i∆u

(kPa/m) v0

0ri

σuσ

′−

v0

ri

σσ′′

v0

i

σ∆u

′

u

0ri

suσ −

u

ri

sσ′

u

i

s∆u


SPM 8.7 0.5 8.2 1.19 0.07 1.12 7.8 0.5 7.4

MTD 17.9 3.8 14.1 2.45 0.52 1.93 16.1 3.4 12.7

NGI (LB) 4.9 1.1 3.5 0.67 0.15 0.48 4.4 1.0 3.2

NGI (UB) 4.9 1.3 3.8 0.67 0.18 0.52 4.4 1.2 3.4

CEM (LB) 6.2 1.1 5.1 0.85 0.15 0.7 5.6 1.0 4.6

CEM (UB) 7.5 1.3 6.2 1.03 0.18 0.85 6.8 1.2 5.6

Note: ‘LB’ for ‘Lower Bound’, ‘UB’ for ‘Upper Bound’, ‘*’ for derivations from measurements.

7.4.1.5 Summary

In summary, based on the analysis of four caisson tests in the sensitive clay mentioned

above, it can be seen that the gradients of radial stress changes around caissons are

similar in the suction-affected area and the jacking-affected (self-weight penetration)

area. The slight reduction in the radial total stress when the TPTs left the

jacking-affected area and entered the suction-affected area was caused by consolidation

during the short time delay and reduced speed of penetration, while the reduction

observed at the end of installation was caused by dissipation of excess pore pressures as

the penetration essentially stopped, due to the soil plug contacting the caisson lid.

Suction installation appears to have little influence on the radial total stresses and

excess pore pressures acting on the external wall, and thus the mode of soil flow at the

caisson tip. Even allowing for a one diameter transition zone between self-weight

penetration and suction installation, the approach suggested by Andersen & Jostad

(2002) is not supported by the measurements. The NGI method tends to under-predict

the excess pore pressure and radial total stress acting on the external shaft of the caisson

during suction installation; a simple form of CEM gives reasonable predictions; the

SPM obviously over-predicts the measurements, probably due to the difference between



the soil utilised in its analysis and the sensitive kaolin clay used here, particularly the

low strength ratio of the kaolin compared with Boston Blue Clay; the MTD method

significantly over-predicts the test results, indicating that extrapolation from the results

of closed-ended piles to open-ended piles is improper.

7.4.2 Relaxation of Radial Stresses during Consolidation

Once the caisson was installed by suction to the target depth (or the caisson stopped

moving), the syringe pump was stopped immediately, while the self-weight stress-hold

of 8 - 12 N was maintained on the caisson for 1 hour at 120 g, or 1.7 years prototype

time.

7.4.2.1 t50 and t90

Variations of depth of the caisson tip and the axial force during consolidation in

sensitive clay are shown in Figure 7.88 for test B14cyc, with units in model scale. After

1 hour of consolidation at 120 g, the settlement of the caisson was 0.16 mm in model

scale (representing 19 mm at prototype scale). This is slightly larger than that observed

in NC clay, reflecting the larger sensitivity of the sensitive clay. The settlement

continued after 1 hour (in model scale), indicating that either primary or (certainly)

secondary consolidation had not yet finished at that moment. Similar graphs are plotted

in Figures 7.89 to 7.91 for tests B14susa, B14SCC and B14sus (units shown in model

scale). The overall settlements (in model scale) of the caissons during consolidation of

these four tests were respectively 0.16, 0.16, 0.17 and 0.12 mm, with an average

settlement of 0.15 mm (or 19 mm at prototype scale).

Variations of the radial stress around the caisson during consolidation were monitored

by the two TPTs located at L/3 (40 mm at model scale) from the tip of the caisson. The

measured radial total stress relative to the hydrostatic pressure, σr – u0, for tests B14cyc,

B14susa, B14SCC and B14sus are plotted against consolidation time (in model scale) in

Figures 7.92 to 7.95. The measured σr – u0 decreased gradually within the one hour of

consolidation at 120 g (representing 1.7 years prototype time); reducing 4.5, 2.8, 6.6

and 9.5 kPa, respectively in the above four tests, with an average reduction of 5.9 kPa.

This degree of relaxation is much smaller than those measured in the NC and LOC clay.

Times for 50% and 90% of the consolidation in terms of σr – u0 are derived and are

presented in Table 7.15. It should be noted that these derived t50 and t90 are lower



bound values, as it appears that full consolidation was not achieved within the 1 hour

period.

Table 7.15 Measured 50% and 90% consolidation time in terms of average

σr – u0 in sensitive clay

t50 t90 Test

Model

(s)

Prototype

(month)

Model

(s)

Prototype

(month)

B14cyc 1242 6.9 2949 16.4

B14susa 1599 8.9 2964 16.5

B14SCC 669 3.7 2559 14.2

*B14sus 148 8.4 1917 16.3

Average 1170 6.5 2824 15.7

Note: * is abnormal, not considered in the average

It took more than 6.5 months and 15.7 months respectively to complete 50% and 90%

of the consolidation in sensitive clay. These are slightly longer than the corresponding

times measured in NC clay (3 months for t50 and 11 months for t90) and LOC clay (2.1

months and 12.1 months), and can be attributed to a lower consolidation coefficient for

the sensitive soil. As stated in the previous chapter, analysis of the settlement during

consolidation suggests a cv of 2 m2/year for the sensitive clay. Taking ch ~ 3cv and T50

and T90 of ~1 and ~10, where T is defined as cht/d2eq, the theoretical t50 and t90 would be

1.7 months and 17 months, respectively. The predicted 50% consolidation time is much

lower than the measured value, perhaps due to the influence of the unexpectedly large

vertical movement of the caisson before stable settlement. The theoretical 90%

consolidation time (17 months), however, is just above the measured value (15.7

months, which is a lower bound). In the NGI method, Andersen & Jostad (2002) gave

t50 and t90 as ~2 days and 55 days (~ 2 months) for the sensitive clay from Offshore

Africa. Again, their prediction is much lower than the lower bound values derived from

the measurements here.




Test results in sensitive clay, including the measured final radial effective stress ratio,

Kc, defined as σ'rc/σ'v0, and the coefficient Hi, defined as (σri – u0)/σ'v0, are summarised

in Table 7.16 for tests B14cyc, B14susa, B14SCC and B14sus, with the depth shown in

prototype scale.

Table 7.16 Measured radial stress changes during consolidation in sensitive clay

Test i,TPTz

(m)

0ri uσ −

(kPa) v0

0ri

σuσ

′− c,TPTz

(m)

0rcrc uσσ −=′

(kPa) v0

rc

σσ

′′

0ri

rc

uσσ−′

B14cyc 10.44 59.34 0.78 10.46 54.98 0.72 0.92

B14susa 9.97 53.50 0.74 9.99 50.53 0.69 0.93

B14SCC 10.50 56.75 0.74 10.51 51.08 0.67 0.91

B14sus 9.79 59.58 0.83 9.80 50.10 0.70 0.84

Average 10.18 57.29 0.77 10.19 51.67 0.70 0.90

It can be seen in this table that the average Hi is 0.77, and the average final radial

effective stress ratio, Kc, is 0.70. The average ( )0rirc uσσ −′ is 0.90 in those four tests.

The average stress relaxation for caisson tests in sensitive clay is thus ~10%, which is

less than that for full-displacement piles suggested by Lehane & Jardine (1994), and is

less than those measured in NC (relaxation ~ 40%) and LOC (relaxation ~ 20%) clays

with lower sensitivity. The difference between Kc and the estimated in situ earth

pressure coefficient, K0, is 27%. If more time were allowed for consolidation after

penetration, Kc should be closer to the K0 value.

According to the CEM (Randolph, 2003), Kc can be estimated by Equation 2.20, with

R = 1.0, λ = 1 and µ = 5. Recalling the CEM predictions for riσ′ and i∆u stated

previously, the average 0vri σσ ′′ and 0vi σ∆u ′ can be obtained as 0.15 - 0.19 and

0.71 - 0.86, respectively. This results in corresponding lower and upper bound Kc

values of 0.47 and 0.52, with an average value of 0.50. The CEM prediction is around

67 to 74% of the measured Kc of 0.70 shown in Table 7.16. The difference is larger



than in the case of NC clay and LOC clay with lower sensitivity. Considering the fact

that σr was still reducing at the ‘end’ of the consolidation period, the difference should

be smaller with further consolidation.

The MTD prediction of Kc can be estimated by Equation 2.19. Since h/deq = 5.2,

R = 1.0 and St = 4 - 5, the calculated lower bound and upper bound Kc values are 1.01

and 1.06, respectively. The average Kc predicted by MTD over-predicts the measured

value by around 50%, which is much smaller compared with those for the NC clay and

the LOC clay with lower sensitivities.

No Kc value is available from the NGI method.


The radial stress changes for two monotonic uplift tests, B14SCC and B14susa, are

shown in Figure 7.96 and Figure 7.97. In test B14susa, the caisson was loaded

monotonically to failure first, then consolidation was allowed again before the final

sustained loading; the analysis performed here is focused on the monotonic loading

period. The radial total stress dropped 6.92 kPa within 0.43 m (or 0.12d) of extraction

for test B14SCC (Figure 7.96), and 3.73 kPa within 0.44 m (or 0.12d) in the monotonic

loading stage of test B14susa (Figure 7.97). After the radial total stress reached its

lowest value in both tests, it increased slightly and then decreased almost linearly with

further movement. As in tests in clay with lower sensitivity, the gradient of the radial

total stress during pullout was lower than that during installation, indicating that

relaxation had occurred in the clay after consolidation.

Variations of the radial total stress relative to the hydrostatic pressure, σr – u0, with the

movement of the caisson during the early stage of pullout, are shown in Figures 7.98

and 7.99 for these two tests. During pullout of test B14SCC, σr – u0 decreased from

51.08 kPa at 15.31 m (after consolidation) to 44.16 kPa at 14.89 m (when loaded to

failure) (see Figure 7.98). Considering the reduction in lateral earth pressure during the

upward movement of 0.43 m, the corrected radial total stress after consolidation

declined by 5.24 kPa. By assuming the excess pore pressure to be zero when the

caisson is loaded to failure, σ rf can be calculated as the radial total stress relative to

hydrostatic pressure. The reduction factor in the radial effective stress during

monotonic pullout of the caisson, can thus be derived as (1 – 5.24/51.08) = 0.90.

Similar to test B14SCC, the radial total stress after consolidation in test B14susa (see



Figure 7.99) actually decreased 1.83 kPa when loaded from 50.53 kPa at 14.79 m to

failure (at 14.34 m), the reduction factor is thus (1 – 1.83/50.53) = 0.96. Therefore, the

reduction factor (i.e. K value in Equation 2.26) in the radial effective stress when

caissons are loaded to failure in sensitive clay is 0.90 - 0.96, with an average value of

0.93. This is larger than that of 0.8 derived in the NC and LOC clays with lower

sensitivities. As such, it is suggested that an increased stress reduction factor, K, of 0.9

should be adopted for deriving σ′rf from σ′rc for loading tests in sensitive clay.

The external α values during pullout of caissons can also be derived from the measured

σ rf by Equation 2.27. For test B14susa loaded to failure (see Figure 7.99), zTPT = 9.59

m (or z = 14.39 m); σ′rf = 46.90 kPa, with dsu/dz = 1.35 kPa/m above the TPT, and a

residual friction angle of 11.7° (see Table 5.1), the external α can thus be calculated as

0.75. Similar analysis of test B14SCC (see Figure 7.98) results in an external α value

of 0.92. However, the caisson was over-driven in test B14SCC so that the embedment

of the caisson when loaded to failure was 14.89 m, which is larger than the full length

of the caisson. The result of test B14SCC is therefore not taken into consideration.

Recalling a lower bound α of 0.65 derived from the measured uplift capacity in Chapter

6, the measured external α for caissons during uplift in sensitive clay can be taken as

0.65 - 0.75 (averaged at 0.70). This value is lower than that obtained in NC clay

(0.77 - 0.86), and in LOC clay (0.73 - 0.96), perhaps due to the difference in sensitivity

of the soil.

Predictions of the CEM and MTD method for the external α values during uplift can be

obtained from Equation 2.26, by adopting K as 0.9 stated previously. Using an average

γ′ of 7.3 kN/m3, a δr of 11.7°, and an average dsu/dz of 1.17 kPa above the final

embedment of the TPTs for the four tests in sensitive clay, the CEM predicts the

external α as 0.58, while the MTD method predicts ~1.21. The API prediction of

external α can be obtained by Equation 2.24, but since for the sensitive clay su/σ v0 is

0.19, α is again calculated as unity; the NGI method predicts an α value of 0.65

(Andersen & Jostad, 2002). Compared to the average measured α values of 0.65 - 0.75

stated above, the CEM under-predicts by 11 - 23%; the MTD approach over-predicts by

~70%; the API prediction is some 40% higher; the NGI method predicts exactly the

measured value, although this agreement is inconsistent with its under-prediction of

around 30% for the excess pore pressure generated during installation. A summary of



the measured and predicted external α values during vertical pullout of sealed caissons

after consolidation in sensitive clay are shown in Table 7.17.

Table 7.17 Measured and predicted external shaft friction ratio α during vertical

pullout of the caisson after consolidation in sensitive clay

Method α


MTD approach ~ 1.2

NGI method 0.65

CEM ~ 0.58

API RP2A 1.0

7.5 CONCLUSIONS

Total pressure transducers were placed at different elevations on model suction caissons

to monitor variations of radial stress changes during installation, consolidation and

vertical pullout in NC, LOC and sensitive clays. Comparisons were made between

various theoretical methods and measurements (or derivations), for the radial total stress

acting on the external wall of the caisson during installation. Theoretical predictions

and measurements were also compared for the radial effective stress around the caisson

after consolidation. Upper bound shaft friction ratios of the caisson during vertical

uplift after consolidation were derived from the radial stress measured when the caisson

was loaded to failure. The following conclusions can be drawn from the above

analysis:

1. The radial total stress acting on the external caisson wall of the caisson varied

almost linearly with penetration depth in clay, with insignificant difference

between jacked installation and self-weight penetration followed by suction

installation. The magnitude of the radial stress changes, and the time-scale for

consolidation following installation, both suggest that significant outward motion

of the clay at the caisson tip occurred under suction installation.

2. Gradients of the measured external radial total stress in the suction-affected area

are slightly smaller than that in the jacking-affected area. The slight reduction in



the external radial total stress when the TPTs left the jacking-affected area and

entered the suction-affected area was caused by consolidation, due to the time

delay when launching suction installation, and reduced speed during penetration.

3. Reduction in the measured radial total stress at the end of suction installation was

caused by consolidation, when the upheaval of the soil plug led to contact with

the caisson lid and caused the penetration rate to essentially stop.

4. The suggestion by Andersen & Jostad (2002) that the caisson wall is

accommodated entirely by inward motion of the clay during suction installation

gave rise to external stress changes, consolidation times and external shaft

friction ratios that were all significantly lower than those measured, except for

the α value during pullout in sensitive clay.

5. Little difference existed between the external radial total stresses measured

during pullout immediately after installation, and that during installation.

Comparison with measurements after consolidation suggests that extensive

relaxation occurred in the radial total stress during consolidation.

6. A simple cavity expansion approach (Randolph, 2003) gives reasonable

predictions of stress changes and post-consolidation external shaft friction.

Consolidation times were also reasonably consistent with those deduced from

displacements and radial stress relaxation measured in the model tests.

7. The MTD framework for displacement piles (Jardine & Chow, 1996)

significantly over-predicts the measured stress changes and shaft capacity of the

model caissons. The method was developed from, and calibrated against,

measurements for piles with much higher embedment ratios (L/d), and

displacement ratios (for example, closed-ended or with area ratios in excess of

10%) compared with the caissons considered here. The poor agreement

illustrates the danger in extrapolating empirically-based methods outside the

database from which they are derived.

8. Solutions from the strain path method (SPM) developed for open-ended piles

(Whittle & Baligh, 1988) over-predict both the radial total stress and the excess

pore pressure generated during caisson installation.

9. For caissons in NC clay, an upper bound external shaft friction ratio of 0.86 was

derived from the TPT measurements during pullout after consolidation.

Recalling the lower bound external α derived from the measured axial capacity,



α ranges 0.77 - 0.86 in NC clay, 0.73 - 0.96 for LOC clay, and 0.65 - 0.75 in

sensitive clay, as summarised in Table 7.18.

Table 7.18 External shaft friction ratio α for suction caissons during installation

and vertical pullout in NC, LOC and sensitive clays

Soil Installation Uplift after consolidation

NC clay 0.35 - 0.45 0.77 - 0.86

LOC clay ~ 0.42 0.73 - 0.96

Sensitive clay ~ 0.16 0.65 - 0.75

It should be noted that during pullout of the caisson the adopted rate of 0.3 mm/s during

sealed pullout is different to that used for caissons in the field, and this may cause some

difference for the obtained external α values during pullout. Such a rate effect needs to

be taken into consideration. The derived external α values during pullout were based

on centrifuge tests on model caissons, and uncertainties exist when applying these

results directly to caisson design for soils with different properties, without

confirmation from large scale field tests.

Chapter 8 8-1 Sustained Loading and Cyclic Loading


8 SUCTION CAISSONS UNDER SUSTAINED LOADING AND

CYCLIC LOADING IN CLAY

Deep to ultradeep water facilities such as tension leg platforms (TLP) and

semi-submersible structures are subjected to both static loading, and cyclic or sustained

loadings from wind, waves, tides, storms, hurricanes, circulations and so on

(Clukey et al., 1995; Al-Khafaji et al., 2003). Among these loads, sustained loading

exerted by loop currents and vortex-induced vibrations (VIV) is an important

consideration (Clukey et al., 2004). Sustained loading often controls suction caisson

design, since it acts over a longer period, where ‘passive suction’ may not be relied

upon fully, resulting in some reduction in the uplift capacity (Huang et al., 2003;

Clukey et al., 2004). Another major concern is the cyclic loading originated from

waves, storms (Clukey et al., 1995), and tsunami. In extreme situations, storms or

hurricanes could exert high frequency cyclic loads on offshore platforms. These high

frequency cyclic loads are oscillatory loads that are applied so fast (seconds to minutes)

that the soil response is undrained (Clukey et al., 1995).

In this chapter, results are reported from centrifuge tests, that were carried out to

investigate the uplift capacity and radial stress changes around caissons, under either

sustained loading or cyclic loading in NC, LOC and sensitive clays.

8.1 SUSTAINED LOADING

Sustained loading tests, labelled ‘sus’ in the test names, were performed on the caissons

after consolidation for 1 hour at 120 g following suction installation. Model caisson 1

was used in test B12sus, while model caisson 2 was used in tests B13sus, B14sus and

B14susa.

8.1.1 Sustained Loading in NC clay

Test B12sus was performed on caisson 1 in NC clay, with sustained loading following

consolidation. After consolidation for 1 hour model time at 120 g (1.7 years prototype

time), increasing uplift force was exerted stage by stage on the caisson, whilst the

vertical displacement and radial stress changes were monitored by the displacement

transducer and TPTs, respectively. The variation of uplift pressure during sustained

loading with prototype time is shown in Figure 8.1, together with the corresponding



displacement of the caisson. The tensile sustained pressure, ∆pmin, which is calculated

as baseAP , decreased (the absolute value increased) from –150 kPa in increments of

–50 kPa, with each stage lasting for 5 minutes model time (representing 7.1 weeks

prototype time), as shown in detail in Table 8.1.

Table 8.1 Details of sustained loading packets in NC clay (test B12sus)

Packet No. ∆p

(kPa)

tmodel

(min)

tprototype

(day) ∆z

(m) (σr − u0)i

(kPa)

(σr − u0)end

(kPa)

∆σr,cross

(kPa)

∆σr,stage

(kPa)

∆σr

(kPa)

1 –150 5 50 0.02 24.57 21.25 –2.70

(–11%)

–1.52

(–6%) −4.22

(–17%)

2 –200 5 50 0.02 19.14 17.49 –2.11

(–11%)

–1.65

(–9%) −3.76

(–18%)

3 –250 0.25 2.5 0.05 16.22 17.42 0 +1.20

(+7%)

+1.20

(+7%)

Note: numbers in brackets are ratios with respect to σr – u0 at the beginning of that loading stage.

When the tension load was raised to –250 kPa, the caisson started to move so quickly

that such an increment step in tension (from –200 kPa to –250 kPa) was probably too

large (see Figure 8.1). In fact, the caisson could not resist such a large tension, so the

load reduced to a relatively stable pressure of −241 kPa, under which obvious upward

movement of the caisson occurred; this pressure is thus considered to be the real

capacity. When the tendency for the caisson to be pulled out became clear under the

sustained loading, the control system was switched from ‘load control’ to ‘displacement

control’, and the caisson was pulled out monotonically at a speed (in model scale) of

0.3 mm/s.

The net resistance measured during installation, sustained loading and uplift of the

caisson in test B12sus is presented in Figure 8.2. Also shown in Figure 8.2a is the axial

pressure measured in the monotonic loading test B12SCC, which was performed in the

same box. The axial capacity of the caisson under sustained loading is obviously lower

than that during monotonic loading. The corresponding T-bar test shows an average

undrained shear strength gradient, dsu/dz, of 1.23 kPa/m (see Figure 8.3). As shown in

Table 8.2, the normalised uplift capacity, defined as umin sp∆− (where us is the



average undrained shear strength at the maximum embedment, Lmax, of the caisson), is

27.2. Also shown in the table, for the purpose of comparison, is the monotonic sealed

pullout capacity of test B12SCC in the same box, for which the normalised uplift

capacity is 34.6. The normalised capacity under sustained loading is approximately

79% of that under monotonic loading; this capacity ratio is close to that of 75%

observed by House (2002). The difference is considered to be caused by some

dissipation of ‘passive’ suction and thus reduction in the reverse end-bearing capacity.

Table 8.2 Normalised uplift capacity, external α and Nc values of the caissons

under sustained loading and monotonic loading in clay

Test Clay dsu/dz Lmax su, tip ∆pmin α Nc

(kPa/m) (m) (kPa) (kPa) u

min

s∆p

−

B12sus NC 1.23 14.39 17.7 –241 27.2 0.68 9.0

B12SCC NC 1.17 14.41 16.9 –292 34.6 0.86 11.0

B13sus LOC 1.76 13.74 24.2 –350 29.0 0.70 9.3

B13SCC LOC 1.64 13.92 22.8 –389 34.1 1.05 10.1

B14sus Sensitive 1.33 14.60 19.4 –250 25.8 0.76 7.5

B14SCC Sensitive 1.16 15.32 17.8 –296 33.3 0.92 10.2

B14susa

(Sust. load.)

Sensitive 1.58 14.23 22.5 –208 18.5 0.54 5.5

B14susa

(Mono. Load.) Sensitive 1.58 14.79 23.4 –293 25.1 0.75 6.8

Note: α values are derived from the measured σr at failure.

Variations of the radial total stress versus caisson embedment for test B12sus during

installation, consolidation, sustained loading and uplift are shown in Figure 8.4, which

is somewhat similar to those measured during monotonic pullout. It should be noted

that model caisson 1 was used in test B12sus; the distance between the TPTs and the

caisson tip is 60 mm (model scale), or 7.2 m at prototype scale. Variations of σr – u0

(average from two TPTs) during consolidation for test B12sus are shown in Figure 8.5,

which show that σ′rc was 24.57 kPa at the end of consolidation.

Variations of σr – u0 with time during sustained loading are shown in detail in Figure



8.6, and are summarised in Table 8.1. (σr − u0)i and (σr − u0)end are respectively the

initial and final σr − u0 for each loading stage, while ∆σr, stage is the change in σr – u0

during each stage of applied sustained loading, and ∆σr, cross is the change in σr – u0 due

to cross-sensitivity of the pressure cells with the axial load. It can be seen that σr – u0

reduced 2.70 kPa immediately after the loading pressure of −150 kPa was applied, due

to cross-sensitivity. During the following 50 days prototype time of loading under -150

kPa, σr – u0 decreased gradually by around 1.52 kPa. Under ∆p = −200 kPa, σr – u0

declined abruptly by 2.11 kPa, then decreased gradually by 1.65 kPa during the

following 50 days prototype time. It is interesting to find that σr – u0 increased during

the last stage of loading (∆p = −250 kPa), probably due to adjustment in the soil at

failure. The overall reduction, ∆σr, was 7.15 kPa within these loading stages. The

change in σr – u0 due to creep during each loading stage is less than 9% of its initial

value at that stage, while the cross-sensitivity amounts to less than 11% of the initial

value. The excess pore pressure is thus considered to be small (less than 9% of the

initial σr – u0 during each loading stage) during sustained loading (see Table 8.1),

considering the corresponding velocity of movement of the caisson (~0.001 mm/s in

model scale; V = vd/cv = 0.4, thus drained).

Values of σr – u0 versus the caisson embedment are plotted in Figure 8.7. The caisson

was loaded from z = 14.39 m (after consolidation) to failure at 13.92 m (where the

residual state appeared). During the upward movement of 0.07 m, σr – u0 decreased

from 24.57 kPa to 15.54 kPa (see Figure 8.7). Assuming that the excess pore pressure

generated during the sustained loading is negligible, σ rf can be taken as the measured

radial total stress at failure relative to the hydrostatic pressure, σrf – u0, and the value

was 15.54 kPa. The α value during pullout can thus be calculated by

u

rrf

stanδσ

α⋅′

= (8.1)

As discussed previously, rδ was 17.6 (see Table 5.1), while Figure 8.3 shows that su

was 7.25 kPa at zTPT of 6.72 m (z = 13.92 m), α can thus be derived from Equation 8.1

as 0.68. The corresponding Nc value derived from the data shown in Table 8.2 is 9.0.

The α value is obviously lower than that of 0.86 for monotonic loading test B12SCC,

and the Nc value of 9.0 for sustained loading test B12sus is also clearly lower than 11.0

in monotonic loading test B12SCC. This shows that the pattern of sustained loading



applied not only reduced the shaft friction between the caisson wall and the NC clay

due to creep, but also reduced the end-bearing capacity significantly, by causing the

dissipation of ‘passive’ suction at the caisson tip. .

Unfortunately, at the end of the test, obvious drifting was been observed in the TPT

reading, which is considered to be the result of a slight leakage in the wires. Therefore,

two new TPTs were fabricated on a new model caisson (caisson 2) in the next box of

cyclic loading tests, to avoid leakage and cross-sensitivity.

8.1.2 Sustained Loading in LOC clay

Test B13sus was performed on model caisson 2 in the lightly overconsolidated (LOC)

clay, with an OCR of 1.5, as described previously. The variation of penetration

resistance with depth, and the T-bar results have been described in Chapter 6, with an

overall soil strength gradient dsu/dz of 1.76 kPa/m (see Figure 6.60) at full embedment

of the caisson. After a subsequent consolidation of 1 hour (model time) at 120 g, four

stages of vertical sustained loading were applied to the sealed caisson, with each stage

lasting for 5 minutes model time (or 50 days prototype time), until obvious

displacement was observed after which the system was switched to monotonic pullout.

Details of the loading stages (∆p) and the corresponding vertical displacement (∆z) of

the caisson are listed in Table 8.3.

Table 8.3 Details of sustained loading packets and vertical displacement in LOC

clay (test B13sus)

Packet No. ∆p

(kPa)

tmodel

(min)

tprototype

(day) ∆z

(m) (σr − u0)i

(kPa)

(σr − u0)end

(kPa)

∆σr,stage

(kPa)

1 –190 5 50 –0.01 54.40 52.67 –1.73 (–3%)

2 –275 5 50 –0.01 52.67 48.96 –3.71 (–7%)

3 –315 5 50 –0.01 48.96 47.76 –1.20 (–2%)

4 –350 5 50 –0.03 47.76 48.67 +0.91 (+2%)

Note: values in brackets are the ratio regarding the initial σr – u0 during that stage.

The overall resistance during installation, consolidation, sustained loading and pullout

of the caisson for test B13sus is shown in Figure 8.8a. The variation of the internal pore



pressure for test B13sus is shown in Figure 8.8b. Also shown in Figure 8.8a is the

pressure response of the monotonic loading test B13SCC in the same box. Despite a

higher soil strength gradient, the sustained capacity was obviously lower than the

monotonic capacity. The detailed time history of applied loading and resulting

displacement is presented in Figure 8.9. The caisson displaced rapidly under a

sustained loading stage of –350 kPa, which is considered to be the sustained capacity.

The normalised uplift capacity for test B13sus was 29.4, which is 86% of that in test

B13SCC ( us∆p− = 34.1) shown in Table 8.2. This difference again indicates that the

loss of negative pore pressures (or so called ‘passive’ suction) at the caisson base during

sustained loading led to a reduction in base resistance.

Radial total stress changes for test B13sus during installation, consolidation, sustained

loading and pullout of the caisson in LOC clay are plotted in Figure 8.10. It should be

noted that the two TPTs are located 40 mm (or 4.8 m prototype) from the tip of the

model caisson 2 used here. Comparison of σr between the sustained loading (test

B13sus) and monotonic loading (B13SCC) is shown in Figure 8.11. The profiles are

very similar for most part of the depth, except between 13 m and 14 m during loading,

where σr reduced significantly under sustained loading, compared to that during

monotonic loading.

Variations of σr – u0 with prototype time under sustained loading are shown in detail in

Figure 8.12 for test B13sus. During each stage of loading, σr – u0 decreased gradually

with loading time. It dropped 1.73 kPa during the loading stage of –190 kPa, 3.71 kPa

under the loading of –275 kPa, and 1.20 kPa under –315 kPa (see Table 8.3). During

the last stage (–350 kPa) when failure commenced, the reading decreased first, then

increased to a value which was 0.91 kPa higher than that at the beginning of this

loading stage, as shown in Figure 8.12. The increase in σr during the last loading stage

was considered to be caused by switching the system from ‘load control’ to

‘displacement control’. It can be seen in Table 8.3 that the reduction in radial stress was

generally less than 7% of the initial value of σr – u0 for each loading stage, and thus is

considered insignificant. .

The decreases in σr, however, were well in excess of that caused by the reduction of the

overburden pressure due to upward movement of the caisson. For example during the

first stage of loading, the caisson mobilised upwards a prototype distance of 0.012 m,

which could result in a radial stress decrease of only ~0.1 kPa. In addition to the



possible cross-sensitivity to axial loading, the major source for the decrease in radial

stresses during sustained loading may be attributed to the creep of the soil particles

surrounding the caisson.

The variation of σr – u0 with embedment of the caisson during the early stage of

monotonic pullout after sustained loading is shown in Figure 8.13. After consolidation,

σr – u0 was 54.40 kPa at an embedment of 8.94 m for the TPTs and 13.74 m for the

caisson. The minimum σr – u0 of 32.94 kPa occurred at z = 13.18 m (or zTPT = 8.38 m),

and this is considered to be the moment when failure occurred on the caisson shaft

(slightly later than the failure at the caisson tip). The change in σ′r (or σr – u0, since v ≤

0.0007 mm/s in model scale; V ≤ 0.3, thus drained) was 21.46 kPa during this 0.56 m

(or 0.018d) of extraction; by subtracting the reduction of 2.83 kPa due to the decrease in

embedment, the real reduction in σ′r was 18.63 kPa. This reduction is obviously larger

than that of 3 kPa during the monotonic pullout test B13SCC (see Figures 8.11), and the

shaft friction during pullout after sustained loading is therefore estimated to be lower

than that during monotonic pullout.

Since su was 15.4 kPa at zTPT = 8.38 m (for z = 13.18 m) (see Figure 6.46) and rδ was

18.1 (see Table 5.1), the α value during pullout can be obtained by Equation 8.1 as

0.70. This α value is clearly smaller than that of 1.05 derived during the monotonic

loading. This is reasonable, since damage to the soil surrounding the caisson occurred

during creep under sustained loading, the interface strength was thus reduced. Based on

the α value derived from the TPT measurements, the Nc value of test B13sus can be

derived as 9.3, which is smaller than 10.1 for monotonic loading test B13SCC. It can

be seen that for caissons under sustained loading in the LOC clay, both the shaft friction

capacity and the end-bearing capacity were reduced, with the latter resulting from the

dissipation of ‘negative’ pore water pressures during the long-term loading process.

It can be seen that the Nc and external α values derived from sustained loading test

B13sus in LOC clay are both close to those obtained in sustained loading test B12sus in

NC clay.

8.1.3 Sustained Loading in Sensitive Clay

8.1.3.1 Pure sustained loading

Sustained loading test B14sus was undertaken in the sensitive sample with a sensitivity



of 4 - 5. The overall soil strength gradient dsu/dz was 1.33 kPa/m, and the

corresponding α value during installation was 0.15, according to the analysis in Chapter

6. After consolidation for 1 hour (model time) at 120 g, vertical sustained loading was

applied at a starting value of –140 kPa which lasted for 50 days prototype time. The

tensile pressure was then raised to –165 kPa (50 days), –205 kPa (200 days), –234 kPa

(150 days), –250 kPa (150 days) and –270 kPa (170 days), as stated in Table 8.4.

Table 8.4 Details of sustained loading packets in sensitive clay (test B14sus)

Packet No. ∆p

(kPa)

tmodel

(min)

tPrototype

(day) ∆z

(m) (σr − u0)i

(kPa)

(σr − u0)end

(kPa)

∆σr,stage

(kPa)

1 –140 5 50 0.00 50.10 46.44 −3.66 (−7%)

2 –165 5 50 –0.03 46.44 45.74 −0.70 (−2%)

3 –205 20 200 –0.10 45.74 43.82 −1.92 (−4%)

4 –234 15 150 –0.07 43.82 43.34 −0.48 (−1%)

5 –250 15 150 –0.08 43.34 42.29 −1.05 (−2%)

6 –270 17 170 –0.75 42.29 44.20 +1.91 (+5%)

Note: values in brackets are the ratio regarding the initial σr – u0 during that stage.

The overall resistance during the loading for test B14sus is shown in Figure 8.14a; also

shown in the graph, for the purpose of comparison, is the pressure response of the

monotonic loading test B14SCC in the same box. The variation of internal pore

pressure for test B14sus is presented in Figure 8.14b. Details of the sustained loads

applied vertically are shown in Figure 8.15. When the sustained tension load was raised

to –270 kPa, accelerated upward displacement of the caisson occurred immediately,

indicating that the caisson essentially reached failure before this loading stage. In fact,

obvious displacement occurred during the sustained loading stage of –250 kPa, which

was subsequently taken as the failure load, since, if the load had been sustained for a

longer time at this stage, a larger displacement would have taken place. The maximum

embedment of the caisson was 14.60 m for test B14sus, the normalised uplift capacity

was thus 25.8. This value is 77% of the normalised uplift capacity in test B14SCC

during monotonic loading shown in Table 8.2.

Plotted in Figure 8.16 is the variation of radial total stresses during installation,



consolidation, sustained loading and pullout of the caisson in sensitive clay for test

B14sus. Attention should be paid to the fact that model caisson 2 was used here and the

two TPTs were located 40 mm (or 4.8 m prototype) above the caisson tip. Comparison

of the σr between sustained loading (test B14sus) and monotonic loading (B14SCC) is

presented in Figure 8.17. The measured σr during sustained loading was obviously

lower than that at the same depth during monotonic loading, revealing the significant

influence of creep on the radial stress during sustained loading. Values of the measured

σr seem to vary directly with depth, but at a larger gradient than for monotonic loading.

Variations of σr – u0 under sustained loading for test B14sus are shown in Figure 8.18,

with values shown in Table 8.4. By the end of the loading stage of –250 kPa, the

overall sustained loading time had amounted to 600 days. During this period the

embedment of the caisson changed from 14.59 m to 14.31 m, and σr – u0 dropped 7.81

kPa in total, for which only 1.2 kPa could be attributed to the upward displacement

(0.28 m in prototype scale) of the caisson, as shown in Table 8.4. Due to much lower

shaft friction, the drop in σr – u0 in sensitive clay was smaller than that measured in clay

with lower sensitivity.


monotonic pullout, after sustained loading in test B14sus, is presented in Figure 8.19.

When sustained loading started after consolidation, σr – u0 was 50.10 kPa (zTPT = 9.79

m, or z = 14.59 m). The minimum σr – u0 of 40.76 kPa appeared at 14.18 m (zTPT =

9.38 m) and this is considered to be the moment when interface failure occurred.

Again, σ rf can be taken as σrf – u0 due to the very low extraction rate (and thus zero

excess pore pressure) at this moment. Since su was 11.16 kPa at 9.38 m (see Figure

6.60), and δr was 11.7 , the external α during pullout can thus be obtained as 0.76 by

Equation 8.1. The α value after sustained loading is clearly smaller than that of 0.92 for

test B14SCC which was under monotonic loading, showing the significant reduction of

shaft friction for caissons in sensitive clay under long-term loading. A corresponding

Nc value of 7.5 is derived from the data of test B14sus, as shown in Table 8.2. This

value is also obviously lower than that of 10.2 during monotonic loading test B14SCC

in the same box. This difference shows that ‘passive’ suction developed at the caisson

tip significantly reduced under sustained loading in sensitive clay, and thus caused an

obviously lower end-bearing capacity than under monotonic loading.



8.1.3.2 Sustained loading after monotonic loading

In test B14susa carried out in sensitive clay, the caisson was first monotonically loaded

to failure, as described in chapter 6, and then 1 hour (model time) of re-consolidation

was allowed at 120 g. Sustained loading was then applied again until failure occurred,

to investigate the capacity ratio for sustained loading relative to that under previous

monotonic loading.

The sustained loading started from –208 kPa (~70% of the monotonic capacity). The

axial pressure during both monotonic and sustained loadings in the same site for test

B14susa is shown in Figure 8.20a; the variation of internal pore pressure versus

embedment of the caisson is shown in the left side of Figure 8.20b. Variations of uplift

pressure and embedment with time during sustained loading are shown in Figure 8.21.

Continuous displacement occurred under the sustained loading stage of –208 kPa,

which was considered to be the sustained capacity. The maximum embedment for the

monotonic loading was 14.79 m, with an average soil strength gradient dsu/dz of 1.58

kPa/m; the corresponding normalised axial capacity was 25.1. During the following

sustained loading after reconsolidation, the embedment of the caisson was 14.23 m and

the normalised axial capacity was 18.5, which is 74% of the monotonic loading, with

details shown in Table 8.2. Such a capacity ratio is very close to the capacity ratio

(78%) of test B14sus relative to that of test B14SCC, and average at 76%.

The average capacity ratio derived from the above four vertical sustained loading tests

is 79% (see Table 8.2). This ratio is very close to the 81% reported independently by

Randolph & House (2002), but is significantly lower than the value of 87 - 101%

reported by Clukey & Phillips (2002). It should be noted that the aspect ratio of the

caisson used by Clukey & Phillips (2002) is 4.5 - 5, which is different from the one

tested here.

Radial total stress changes are plotted against the caisson embedment during

installation, sustained loading and pullout in test B14susa in Figure 8.22. Details of the

variation of σr – u0 under sustained loading for test B14susa are shown in Figure 8.23.

During the sustained loading of –208 kPa for 196 days prototype time, σr – u0 decreased

first, and then increased, with an overall drop of 6.95 kPa, of which 5.4 kPa can be

attributed to the 1.35 m (in prototype scale) of upward displacement of the caisson. The

drop in σr – u0 for test B14susa is larger than that measured in test B14sus, probably

because the former was tested between two used sites, and disturbance in sensitive clay



is larger than that in clay with lower sensitivity.

Figure 8.24 shows the variation of σr – u0 with embedment of the caisson during the

early stage of monotonic pullout after sustained loading for test B14susa. At the

beginning of sustained loading (or the end of consolidation), z was 14.65 m

(or zTPT = 9.85 m), and σr – u0 was 45.58 kPa. When the ultimate shaft friction

developed at z was 14.08 m (or zTPT = 9.28 m), σr – u0 (or σ′rf) was 33.80 kPa. The

corresponding su was 13.04 kPa for the TPTs (see Figure 6.60), the α value during

pullout can thus be obtained as 0.54 by Equation 8.1, and subsequently the

corresponding Nc value is derived as 5.5.

The average external α value for the two sustained loading tests B14sus and B14susa

(sustained loading stage) in sensitive clay is 0.65, which is significantly lower than that

of 0.84 averaged from monotonic loading tests B14SCC and B14susa (monotonic

loading stage). In fact, sustained loading can simultaneously cause two opposite effects

to occur in the soil surrounding the caissons: one is the relaxation of the radial stress

due to creep; the other is the increase in the radial effective stress due to dissipation of

shear-induced excess pore pressures. It appears that the former prevails over the latter,

thus causing the shaft friction during sustained loading to be significantly smaller than

that under monotonic loading in sensitive clay. This observation is consistent with

those for tests in the NC and LOC clays with lower sensitivity. The Nc value of 5.5

after sustained loading of test B14susa is also clearly smaller than that of 6.8 during

monotonic loading in the same test (see Table 8.2), showing a lower end-bearing

capacity under sustained loading, perhaps due to the dissipation of ‘passive’ suction at

the tip of the caisson.

It can be seen in Table 8.2 that during sustained loading, the Nc value in sensitive clay is

much lower than that in the clay with lower sensitivity. In fact, the interface friction

between the soil plug and the internal wall of the caisson in sensitive clay was

extremely low. This was evident as it was impossible to cut soil samples using the

sampler after the completion of the experiments, as the soil plug fell out of the sampler.

During sustained loading, the possible falling out of the soil plug from the caisson may

have contributed to the rather low reverse end-bearing capacity.

8.1.4 Summary



Based on the above analysis, the sustained capacity ratio (normalised sustained

capacity/normalised monotonic capacity), α value and the corresponding Nc values

during sealed pullout of caissons after sustained loading in various clays are

summarised in Table 8.5. It can be seen that the uplift capacity of caissons under

sustained loading in clay was around 75 - 85% of the capacity under monotonic loading,

although this range of reduction will depend on the caisson geometry and duration of

sustained loads. The reduction results from two parts, one is from the slightly reduced

shaft capacity under creep, with α being 0.65 - 0.70, and the other is the reduction of

end-bearing capacity due to dissipation of ‘passive’ suction at the tip, with Nc reduced

to 6.8 - 9, compared to the Nc value of 10 - 12 under monotonic loading. It can be seen

that α for caissons subjected to sustained loading in sensitive clay is slightly lower than

clays with lower sensitivity, while the Nc value in the sensitive clay is clearly lower

than in clays with lower sensitivity.

Table 8.5 Capacity ratio, α and Nc values during sealed pullout of caissons after

sustained loading in clay

Clay type Capacity ratio α Nc

NC clay (OCR = 1, St = 2 - 2.8) 79% 0.68 9

LOC clay (OCR = 1.5, St = 2 - 2.5) 85% 0.70 9

Sensitive clay (OCR = 1, St = 4 - 5) 76% 0.65 6.8

8.2 CYCLIC LOADING

In the NC, LOC and sensitive clays, cyclic loading tests with a sealed lid were carried

out on the caissons after consolidation at 120 g for 1 hour (representing 1.7 years

prototype time), following suction installation. Such tests were designated with ‘cyc’ in

the name.

8.2.1 Cyclic Loading in NC clay

Test B12cyc was conducted on model caisson 1 in NC clay. After consolidation, cyclic

loading packets were applied stage by stage to the caisson, with the vertical

displacement and the radial stress changes monitored by the displacement transducer



and TPTs, respectively. The minimum (or maximum in absolute value) applied load in

each packet was adopted as a certain ratio of the uplift capacity during monotonic

loading in the same box. The cyclic loading packet was then set to vary between such

values and zero. For example, the first loading packet of test B12cyc was varied

between –130 kPa and 0 for 50 cycles, with the maximum tension of –130 kPa being

45% of the monotonic pullout capacity in test B12SCC. The maximum tension for each

loading packet increased to –170 kPa and then to –220 kPa, until obvious vertical

displacement was observed. Then the system was switched to ‘displacement control’

and the caisson was pulled out at a velocity of 0.3 mm/s (in model scale). Loading

time, cycles, frequency and upward displacement (∆z, in prototype scale) for each

loading packet in test B12cyc are shown in detail in Table 8.6.

Table 8.6 Details of cyclic loading in NC clay (test B12cyc)

Packet No. ∆pmin

(kPa)

∆pmax

(kPa)

Cycles Frequency

(Hz) ∆z

(m)

∆σr

(kPa)

1 –130 0 50 0.50 –0.01 –6.05

2 –170 0 50 0.40 –0.01 –7.25

3 –220 0 50 0.25 –0.03 –9.25

Variations of the axial pressure during installation, consolidation, cyclic loading and

uplift in NC clay are shown in Figure 8.25. Also plotted is the axial response of a

monotonic pullout test, B12SCC, undertaken in the same box, for the purpose of

comparison. The pullout capacity under cyclic loading is obviously smaller than that

under monotonic loading. The profiles are also very different between these two tests,

in that the peak pressure was reached shortly after cyclic loading. While in monotonic

loading, a certain displacement was needed to develop the peak resistance. The

corresponding overall soil strength gradient dsu/dz was 1.23 kPa/m for test B12cyc (see

Figure 8.26). It should be noted that the overall time taken to apply the cyclic loading

at model scale have allowed some consolidation to occur, unlike in a prototype storm

condition. As such, the results presented here should be used with caution in assessing

the cyclic performance of prototype caissons.

Variation of the uplift pressure with prototype time during cyclic loading for test



B12cyc is presented in Figure 8.27, together with the corresponding vertical embedment

of the caisson. The uplift pressure versus embedment of the caisson during cyclic

loading is plotted in detail in Figure 8.28. Obvious vertical movement of the caisson

occurred within the loading packet of –220 kPa - 0, and failure was considered to have

taken place here. There was a sudden jump in the uplift pressure when the monotonic

pullout commenced, reaching –250 kPa in a short time before soon decreasing with

further movement of the caisson. Such an increase was considered to be caused by the

increase in displacement rate of the caisson. The normalised uplift capacity, defined as

umin sp∆− , is 24.9, as shown in Table 8.7. This capacity is around 72% of the

monotonic pullout capacity for test B12SCC in the same box, and is slightly lower than

that during sustained loading (test B12sus).

Table 8.7 Normalised uplifted capacity, α and Nc values and capacity ratio of the

caisson under cyclic loading in clay

Test Clay dsu/dz

(kPa/m)

Lmax (m) su, tip

(kPa) ∆pmin

(kPa) u

min

s∆p

− α Nc Capacity ratio

B12cyc NC 1.23 14.39 17.7 –220 24.9 0.72 7.7 72%

B13cyc LOC 1.76 13.60 22.8 –389 29.2 0.80 9.0 85%

B14cyc Sensitive 1.36 15.26 20.8 –255 24.5 0.65 6.4 74%

Radial total stress changes for test B12cyc during installation, consolidation, sustained

loading and pullout of the caisson in NC clay are plotted in Figure 8.29. It should be

noted that the two TPTs are located 60 mm (or 7.2 m prototype) from the tip of the

model caisson 1 used here. The variation of σr – u0 during the three loading packets

(from ∆pmin to 0) were –6.05 kPa, –7.25 kPa and –9.25 kPa, as shown in Figure 8.30

and Table 8.6. Between adjacent loading packets, there was a reduction in the

measured σr – u0. These are considered to be caused by cross-sensitivity.

Variations of σr – u0 with embedment of the caisson during the early stage of monotonic

pullout after cyclic loading are shown in Figure 8.31. After consolidation, σ′rc was 25.5

kPa, with an embedment of 14.36 m for the caisson (and zTPT = 7.16 m). The measured

σr – u0 vibrates with the applied cyclic loading, and the average σr – u0 decreased

gradually between adjacent packets. It can be inferred that the excess pore pressure



generated during such cyclic loading packets was less than 0.7 kPa, since the average

radial total stress only changed ~1.4 kPa after 2 packets of cyclic loading (see Figure

8.31), for which contribution from the variation of the embedment of the TPTs was less

than 0.1 kPa. Failure at the caisson shaft was considered to had occurred when σr – u0

reached the peak point at 14.29 m (or zTPT = 7.09 m) during pullout after cyclic loading,

and σrf – u0 (or σ′rf) was 18.13 kPa. Since su was 7.95 kPa at zTPT = 7.09 m (see Figure

8.26), the α value during pullout can be calculated as 0.72 by Equation 8.1. This α

value is clearly lower than 0.86 obtained in monotonic loading test B12SCC (see Table

8.2). During cyclic loading, the strength of the soil surrounding the caisson was

degraded, due to the relative displacement of the caisson under the repeated loads; on

the other hand however, there is a tendency for the resistance to increase due to a

greater rate of loading (Huang et al., 2003). It seems the former prevailed over the

latter for caissons under cyclic loading tests in NC clay.

The Nc value back-figured from the data shown in Table 8.7 is 7.7, which is much lower

than 11 obtained in monotonic pullout test B12SCC. This means that the extended time

scale of the cyclic loading has allowed significant dissipation of the ‘passive’ suction at

the end of the caisson, which prevented full end-bearing capacity from developing at the

caisson tip, although this extended time scale may not occur for the field case..

8.2.2 Cyclic Loading in LOC clay

Test B13cyc was performed on model caisson 2 in the lightly overconsolidated (LOC)

clay, with an OCR of 1.5. As in the NC clay, cyclic loading packets were applied stage

by stage to the caisson. The loading time, cycles, frequency and upward displacement

(∆z) for each loading packet are shown in detail in Table 8.8.

Variations of the axial pressure during installation, cyclic loading and uplift in NC clay

are shown in Figure 8.32, also plotted in which is that of test B13SCC under monotonic

loading. The comparison shows that the cyclic loading capacity was clearly lower than

that during monotonic loading. The corresponding soil strength profile of test B13cyc

is shown in Figure 8.33, with an overall dsu/dz of 1.76 kPa/m. Variation of uplift

pressure with prototype time during cyclic loading for test B13cyc is shown in Figure

8.34, along with the variation of embedment of the caisson. The uplift pressure versus

embedment of the caisson during cyclic loading is plotted in Figure 8.35. The caisson

was observed to move upwards gradually under cyclic loading. A clear accumulation of

the vertical movements of the caisson was noticed within the loading packet of



–350 kPa - 0; if more cycles were applied at this stage, failure should have occurred

here. No jump occurred in the uplift pressure when the monotonic pullout commenced.

As listed in Table 8.7, the normalised uplift capacity was 29.2, which is around 85% of

the normalised monotonic pullout capacity.

Table 8.8 Details of cyclic loading in LOC clay (test B13cyc)

Packet No. ∆pmin

(kPa)

∆pmax

(kPa)

Cycles Frequency

(Hz) ∆z

(m)

∆σr

(kPa)

1 –195 0 50 0.35 0.01 –0.3

2 –270 0 50 0.25 0.01 –2.6

3 –315 0 50 0.18 0.01 –2.7

4 –350 0 50 0.15 0.02 –4.0

5 –370 0 50 0.12 0.03 –4.4

Radial total stress changes during installation, consolidation, cyclic loading and pullout

of the caisson for test B13cyc in LOC clay are plotted in Figure 8.36. It should be noted

that the two TPTs are located 40 mm (or 4.8 m prototype) from the tip of the model

caisson 2 used here. Variations of σr – u0 under cyclic loading are shown in detail in

Figure 8.37 for test B13cyc, in which the maximum ∆σr for each loading packet (from

∆pmin to 0) was less than 5 kPa, which is less than that observed in test B12cyc. This

lower variation was due to the fact that two new TPTs were used in this test, and error

due to cross-sensitivity was reduced to a minimum. Basically, the TPT response

decreased almost linearly with the increase in the absolute value of the cyclic loading

(Figure 8.37). There was no sudden jump in the measured σr between adjacent loading

packets.


monotonic pullout after cyclic loading is shown in Figure 8.38. After consolidation,

embedment of the caisson was 13.59 m (zTPT = 8.79 m), and σ′rc was 53.01 kPa. When

failure occurred on the caisson shaft (i.e. the lowest σr – u0 was developed), σrf – u0 was

31.98 kPa, with a corresponding embedment of the caisson of 13.21 m (zTPT = 8.41 m).

Taking this stress as the radial effective stress at failure, and considering a

corresponding su of 13.01 kPa (see Figure 8.33), whilst δr = 18.1° from the ring shear



tests, the α value during pullout can thus be obtained as 0.80 by Equation 8.1. This α

value was obviously lower than the upper bound value of 1.05 measured for the

monotonic loading test B13SCC (see Table 8.1) in the same box. The corresponding Nc

value back-figured from the data shown in Table 8.7 was 9.0. This is also clearly lower

than 10.1 obtained in monotonic loading test B13SCC. Cyclic loading applied on

caissons in LOC clay caused not only lower Nc values due to the reduction of ‘passive’

suctions developed at the caisson tip, but also due to some softening of the soil from the

repeated loading process.

8.2.3 Tests in sensitive clay

Test B14cyc was performed in sensitive clay, for which the sensitivity was 4 - 5, after

consolidation for 1 hour at 120 g following the suction installation. The loading time,

cycles, frequency and upward displacement (∆z) for each loading packet are shown in

detail in Table 8.9.

Table 8.9 Details of cyclic loading packets in sensitive clay (test B14cyc)

Packet No. ∆pmin

(kPa)

∆pmax

(kPa)

Cycles Frequency

(Hz) ∆z

(m)

∆σr

(kPa)

1 –166 0 50 0.35 0.01 –1.4

2 –207 0 50 0.29 0.01 –0.8

3 –235 0 50 0.28 0.01 –1.0

4 –255 0 50 0.23 0.02 –1.9

Variations of the axial pressure during installation, consolidation, cyclic loading and

uplift in sensitive clay are shown in Figure 8.39, along with that of test B14SCC during

monotonic loading in the same box. The corresponding average soil strength gradient,

dsu/dz, is 1.36 kPa/m for test B14cyc (Figure 8.40). Considering the larger soil strength

gradient measured in test B14cyc, the normalised axial capacity is different for those

two tests, even though the absolute values are close.

Variation of the uplift pressure with time during the cyclic loading for test B14cyc is

shown in Figure 8.41, together with the embedment of the caisson. The uplift pressure

versus the embedment of the caisson during cyclic loading is plotted in Figure 8.42.



The caisson moved upwards gradually under the cyclic loading, and significant vertical

movements in the caisson occurred within the loading packet of –255 kPa - 0, during

which failure is considered to have occurred. No jump was observed in the axial

pressure when the monotonic pullout commenced.

As listed in Table 8.7, the normalised uplift capacity, usp∆− , was 24.5, which is

around 74% of that in the monotonic pullout. This ratio is close to that measured in the

NC clay, but is lower than that in the LOC clay.

Radial total stress changes for test B14cyc during installation, cyclic loading and

pullout of the caisson in sensitive clay are plotted in Figure 8.43. For caisson 2 used

here, the two TPTs were located 40 mm (or 4.8 m prototype) from the tip of the caisson.

Variations of σr – u0 under the cyclic loading are shown in detail in Figure 8.44 for test

B14cyc, in which the maximum ∆σr for each loading packet (from ∆pmin to 0) was less

than 2 kPa (see Table 8.9). This change is lower than that observed in test B13cyc

performed in LOC clay, possibly due to the softer nature of the sensitive clay. A

relatively linear decrease in the radial total stress occurred throughout the cyclic

loading. No sudden jump in the measured σr between adjacent loading packets was

observed.

The variation of σr – u0 with the embedment of the caisson during the early stage of

monotonic pullout after the cyclic loading is depicted in Figure 8.45. After

consolidation, σ′rc was 55.01 kPa, and zTPT was 10.46 m (z = 15.26 m). When ultimate

shaft friction developed on the caisson, σ′rf (or σrf – u0) was 42.4 kPa and the

corresponding embedment of the caisson was 14.82 m (zTPT = 10.02 m). Considering a

soil strength gradient, su, of 13.45 kPa at 10.02 m embedment of the TPTs (see Figure

8.40), and a δr of 11.7 during the long displacement shearing, the upper bound α value

during pullout can be obtained as 0.65 by Equation 8.1. This α value is clearly lower

than 0.92 obtained during monotonic loading test B14SCC, due to the degradation of

the soil strength of sensitive clay under cyclic loading. The Nc value back-figured from

the data shown in Table 8.7 is 6.4, which is also clearly lower than that of 10.2 during

monotonic pullout test B14SCC, showing that the reverse end-bearing capacity under

cyclic loading in sensitive clay is greatly reduced by the dissipation of ‘passive’ suction

at the caisson tip.



8.2.4 Summary

The capacity ratio (normalised cyclic capacity/normalised monotonic capacity), external

α and Nc values during pullout after cyclic loading in NC, LOC and sensitive clays are

summarised in Table 8.10. The cyclic capacity ratio varies between 72% - 85%

(average at 77%), which is much larger than those of 20% - 50% reported by

Clukey et al. (1995). The lower values in their tests may have been due to the lateral

component of loading, since the angle of loading was varied between ±6 during cyclic

loading. The α and Nc values under cyclic loading are significantly lower than those

under monotonic loading, showing that the shaft friction of the caisson reduced due to

the repeated loading, while the end-bearing capacity also decreased due to loss of

‘passive’ suction under the cyclic loading. Values of α and Nc for caissons tested

under cyclic loading in sensitive clay are lower than those for clays with lower

sensitivity.

Table 8.10 Capacity ratio, α and Nc values during sealed pullout of caissons after

cyclic loading in clay

Clay type Capacity ratio α Nc

NC clay (OCR = 1, St = 2 - 2.8) 72% 0.72 7.7

LOC clay (OCR = 1.5, St = 2 - 2.5) 85% 0.80 9.0

Sensitive clay (OCR = 1, St = 4 - 5) 74% 0.65 6.4

8.3 CONCLUSIONS

The axial capacity and radial stress changes for suction caissons subjected to sustained

loading and cyclic loading were tested in NC, LOC and sensitive clays. The capacity

ratios under sustained or cyclic loading relative to the monotonic loading were

investigated. The external α values during vertical pullout for these two types of

loading were analysed from the radial stress measured at failure, the corresponding

reverse end-bearing capacity factors, Nc, were derived.

For sustained loading, the ratios of capacity relative to that of the monotonic loading are



1) 79% in NC clay; 2) 85% in LOC clay and 3) 76% in sensitive clay. During sealed

pullout after sustained loading, values of the external α and Nc values are respectively

(0.68, 9), (0.70, 9.0) and (0.65, 6.8) in NC, LOC and sensitive clays.

For cyclic loading, the capacity ratio relative to that during the monotonic loading is 1)

72% for the NC clay; 2) 85% for the LOC clay and 3) 74% for the sensitive clay.

During pullout following cyclic loading, values of the external α and Nc values are

respectively (0.72, 7.7), (0.80, 9.0) and (0.65, 6.4) in NC, LOC and sensitive clays.

The resistances of caissons subjected to sustained loading and cyclic loading are

significantly less than that developed under short-term monotonic loading. Both Nc and

α values for caissons subjected to sustained or cyclic loading are much lower than those

during monotonic loading, since dissipation of the ‘passive’ suction at the caisson tip

reduces the reverse end-bearing capacity, while creep and repeated loading caused the

shaft friction to reduce. It is interesting to find that there is a trend for the α and

(especially) Nc values to decrease with the increase of sensitivity of the surrounding

soil, for caissons subjected to sustained loading or cyclic loading.

It should be noted that the excess pore pressures generated during pullout were not

measured and have not been taken into consideration here, when deriving the external α

values in loading. In fact, the time scale of the sustained loading and cyclic loading is

such that full dissipation of excess pore pressures appears possible. The results

presented here thus provide a preliminary evaluation of the shaft friction and thus the

end-bearing capacity of caissons under either sustained loading or cyclic loading.

Certainly, future efforts towards accurate measurements of the excess pore pressures

surrounding the caisson during sustained loading, cyclic loading and pullout of the

caisson may be helpful in improving the accuracy of estimation.

Chapter 9 9-1 Conclusions and Future Work


9 CONCLUSIONS AND FUTURE WORK

A series of centrifuge tests on suction caissons were carried out in this research, to

investigate the axial behaviour of suction caissons in soft deposits in deep and ultradeep

waters. Radial stress changes and axial capacity of suction caissons in clay were

monitored and analysed. Comparison of resistance was made during both installation

and vertical pullout, between caissons installed by jacking and by self-weight

penetration followed by suction installation. Several analytical models for predicting

the penetration resistance of the caisson were evaluated. Radial stress changes around

the caisson were also compared between these two types of installation. Various

theoretical solutions were compared with the measured radial total stress acting on the

external shaft of the caisson, both during installation and after consolidation. The

relationship between radial stresses and shaft friction of the caisson was analysed.

Upper and lower bound shaft friction ratios on the external wall of the caisson during

vertical pullout after consolidation were derived. Behaviour of the caissons under

vertical sustained loading and cyclic loading was investigated. This chapter discusses

the major conclusions obtained from this research, and provides recommendations for

future research.

9.1 MAIN FINDINGS

9.1.1 Interface Normal Stress Measurements in Clay in the Centrifuge

Miniature total pressure transducers (TPTs) were embedded within the thin-walled

caissons with an aspect ratio (L/d) of 4, to monitor radial stress changes on the external

wall. The performance of the pressure cells in clay was investigated through a series of

calibration tests, which were performed both in the laboratory using a modified triaxial

apparatus, and in the centrifuge. In the triaxial apparatus, the TPTs were evaluated

under undrained loading, drained loading, cyclic loading and sustained loading. A

linear relationship between the applied pressure and the cell readings was obtained; the

accuracy of TPTs was above 92.5%. The change in initial values of the TPTs in

different media was very small, being less than 1 kPa when moving from water to clay.

The cross-sensitivity of the TPTs, under the maximum axial loading of the centrifuge

tests, was less than 1.5 kPa. In the centrifuge, TPTs were calibrated by moving the

caisson quasi-statically in water, and the accuracy was larger than 96%; readings were



also very stable under sustained loading. Penetration tests undertaken in clay in the

centrifuge also showed that the pressure cells react correctly to the applied pressure in

high g conditions. As a result, for the design adopted in this research, TPTs give

reliable measurements of the radial stress acting on the external wall of the caisson,

under both short-term and long-term loading conditions.

9.1.2 Interface Friction Angle between Caisson and Clay

Interface friction between caisson and clay can be determined from the interface friction

angle δ and the normal effective stress. Ring shear tests were carried out in a

Bromhead-type ring shear apparatus to investigate peak and residual interface friction

angles, δp and δr. For caisson tests in the NC kaolin clay, δr was obtained as 17.6

(δp = 19.4°); for the lightly overconsolidated (LOC) clay, δr was 18.1 (δp = 18.9°),

while for the sensitive clay, δr reduced to a value of 11.7 (δp = 14.6 ). It is necessary to

remove the soil squeezed out between the concentric ring and the top platen of the

Bromhead-type ring shear apparatus during fast pre-shearing, to obtain correct residual

friction angles. Such a step was not included in the operation manual of the ring shear

apparatus, and is thus recommended to be added after creating the shearing surface,

after which the sample should be reconsolidated under the target vertical pressure.

9.1.3 Installation and Axial Pullout of Caisson

The behaviour of suction caissons during both installation and uniaxial pullout in NC,

LOC and sensitive clays has been investigated by means of a series of centrifuge model

tests. Both the axial capacity and the radial stress changes around caissons during

installation, consolidation and vertical pullout of the caisson were measured and

analysed.

9.1.3.1 Caisson installation

Caissons were installed in an undrained mode either by jacking or by self-weight

penetration followed by suction installation. No soil plug failure occurred inside the

caisson either when suction installation started, or when the soil plug contacted the lid

of the caisson, since the lowest factor of safety was larger than 1.5 during this period.

At the end of suction installation in NC, LOC and sensitive clays, the soil heave derived

from measurements was 0.94 m, 0.83 m and 0.86 m, respectively; these values were

less than the predicted soil heave, based on the assumption of 100% inward motion of



the soil at the caisson tip during suction installation. The corresponding percentages of

inward soil flow at the caisson tip relative to the embedded caisson volume are 45%,

42% and 46% for these three clays, showing that the mode of soil flow at the caisson tip

during suction installation is about evenly divided between inward and outward flow.

Using data from 1) penetration resistance, and 2) deduced soil plug heave, there was

found to be no consistent difference between the behaviour of caissons installed by

either method. Simple in use, the α method was proven to be effective in predicting the

shaft friction of the caisson during installation. A model based on the α method was

shown to be adequate for analysing the penetration resistance of suction caissons:

• Nc = 7.5 for tip resistance;

• For the external wall and the internal wall below the first internal stiffener,

α = 1/St is found to be valid, with α = 0.30 - 0.45 (averaged at 0.38) in NC clay,

0.4 - 0.5 (average value of 0.41) in LOC clay, and ~0.16 in sensitive clay;

• Above the first internal stiffener, the average shaft friction τi-a is ~ 0.5 kPa in NC

clay, τi-a ~ 1.0 kPa in LOC clay, while α = 1/St in sensitive clay.

• Soil sensitivity St can be determined from the in situ cyclic T-bar test.

It should be noted that varying the shaft friction above the first stiffener from zero to

full strength leads to a difference of only 10% for tests in the NC clay, 12% in the LOC

clay and nearly no change in the sensitive clay

Radial stress changes acting on the outside wall of the caisson during installation were

measured by TPTs instrumented at different distances (one at 7.2 m and another at 4.8

m in prototype scale) from the tip of the caisson, in NC, LOC and sensitive clays. The

measured radial total stress varied almost linearly with penetration depth in all three

clays, with insignificant differences found between jacked installation and self-weight

penetration followed by suction installation.

When the TPTs left the jacking-affected area and entered the suction-affected area, the

change in the measured radial total stress was very small, and was caused by the time

delay for initiating suction installation and the reduced speed of penetration. The

gradient of the measured radial total stress in the suction-affected area is slightly

smaller than that during previous jacked installation, suggesting that there is only minor

difference between the patterns of soil flow at the caisson tip for these two types of

installation (at leaset initially). Even allowing for a one diameter transition zone



between self-weight penetration and suction installation, the approach suggested by

Andersen & Jostad (2002) is not supported by the measurements. Reduction in the

measured radial total stress at the end of installation was caused by dissipation of excess

pore pressures when the penetration speed reduced as the soil plug contacted the caisson

lid.

The NGI method (Andersen & Jostad, 2002) under-predicts the radial total stress and

excess pore pressure around the caisson during installation. The MTD approach

(Jardine & Chow, 1996) significantly over-predicts these values. The solution from the

strain path method (SPM) developed for open-ended piles (Whittle & Baligh, 1988)

slightly over-predicts both the radial total stress and the excess pore pressure generated

during caisson installation. A simple form of cavity expansion method (CEM)

(Randolph, 2003) gives reasonable predictions of the radial total stress and excess pore

pressure during the penetration of thin-walled caissons. In contrast to the NGI method,

the CEM is based on the assumption that all the clay is pushed outside during caisson

installation. The obvious under-prediction of the NGI method and the slight over-

prediction of the CEM estimations on the measurements again suggest that a significant

part of soil displaced by the caisson tip during suction installation moves outside the

caisson.

9.1.3.2 Relaxation during consolidation

During consolidation after caisson penetration, the radial total stress around caissons

decreased (relaxed) gradually with the consolidation time, while the caisson settled

simultaneously in the soil. The measured stress relaxation during consolidation was

around 40%, 20% and 10% respectively in NC, LOC and sensitive clays; these values

are less than that reported for full-displacement piles. The times corresponding to 90%

consolidation, t90, were all found to be longer than 1 year in NC, LOC and sensitive

clays. These times were much longer than those (6 days for NC kaolin clay; ~40 days

for LOC clay, and ~60 days for sensitive clay) suggested by the NGI method (Andersen

& Jostad, 2002). The obvious consolidation observed here again suggests significant

outward movement of the soil particles during caisson penetration. The measured t90

matches well with theoretical predictions using the guidelines propose by Randolph

(2003). This has repercussions for the design of suction caissons, since in most

developments there are only short delays between installation and attachment of

mooring lines.



Tests in NC and LOC clays show that a close match exists in the radial effective

stresses after consolidation, σ rc, between caissons installed by jacking and by suction.

The MTD method significantly over-predicts the post-consolidation radial effective

stress σ rc. The CEM prediction of σ rc is close to the measurement.

9.1.3.3 Caisson pullout

Gradients of the measured external σr during vertical pullout after consolidation were

smaller than those during installation, due mainly to relaxation in the radial total stress

during consolidation. The vertical pullout capacity after consolidation was very similar

for caissons installed by jacking and by suction. This proved directly that there is no

discernible difference in the behaviour of caissons installed by either method. The

normalised uplift capacity, umin s∆p− , was around 34 for the sealed pullout tests after

consolidation in NC, LOC and sensitive clays. By adopting an upper bound value of 12

for the reverse end-bearing capacity factor Nc during sealed pullout, the lower bound

external α values derived from the axial capacity of the caisson were 0.77, 0.73 and

0.65 in NC, LOC and sensitive clays, respectively. The α values during pullout after

consolidation increased significantly (almost 100%) over those for pullout immediately

after installation. The high values of Nc may have been affected by the limited depth

(~ 1d) of clay below the caisson in the test.

Considering the upper bound α valued derived from the measured radial stress at failure,

the α values during pullout after consolidation were 0.77 - 0.86, 0.73 - 0.96 and

0.65 - 0.75, respectively for NC, LOC and sensitive clays. The NGI method

under-predicts the shaft friction during pullout of the caisson after consolidation, except

in sensitive clay; the MTD approach over-predicts the measurements, while the CEM

gives reasonable predictions in all three clays.

For a solid pile with the same equivalent diameter and surface roughness as the model

caissons, the α value was 0.48 during installation and 0.93 during vertical pullout after

consolidation in NC clay. These α values are both significantly higher than those

derived from caisson tests. This shows that the simple extrapolation of the shaft friction

ratios for solid piles to thin-walled suction caissons (or open-ended piles) with an

equivalent diameter will result in obvious over-prediction.

In summary, the magnitude of the radial stress changes, the internal soil heave during

installation, the time-scale for consolidation following installation and pullout capacity



after consolidation all suggest that significant outward soil flow at the caisson tip

occurred during suction installation. The suggestion by Andersen & Jostad (2002) that

the caisson wall is accommodated entirely by inward motion of the clay during suction

installation gives rise to stress changes, consolidation times and external shaft friction

ratios that were all significantly lower than those measured. The MTD framework for

displacement piles (Jardine & Chow, 1996) cannot be extrapolated to thin-walled, low

aspect ratio suction caissons, since it leads to over-prediction of stress changes and shaft

capacity. A simple cavity expansion approach (Randolph, 2003) gave reasonable

predictions of stress changes and post-consolidation external shaft friction.

9.1.4 Behaviour under Sustained Loading and Cyclic Loading

Capacities of caissons under either axial sustained loading or cyclic loading were

significantly lower than those under undrained monotonic axial loading. Under axial

sustained loading, the average pullout capacity ratios with respect to the monotonic

loading from tests in NC, LOC and sensitive clays were respectively 79%, 85% and

76%. The α values for sustained loading were derived from the σr measured at failure,

with a result of 0.65 - 0.70. The corresponding Nc values were 9 in NC and LOC clays,

and as low as 6.8 in sensitive clay. Both Nc and α values reduced significantly under

sustained loading, compared to those under monotonic loading. This trend shows that

creep under sustained loading reduced the shaft capacity, and dissipation of ‘passive’

suction beneath the caisson clearly reduced the reverse end-bearing capacity.

Under cyclic loading, the average capacity ratio was 77%. The α value was 0.65 - 0.80,

while the Nc values were 7.7, 9 and 6.4 in NC, LOC and sensitive clays, respectively.

The α values here were lower than those measured under monotonic loading, due to soil

damage around the caissons under repeated loading. The Nc values under cyclic

loading were clearly lower than those during monotonic loading, since dissipation of the

‘passive’ suction at the caisson tip resulted in reduced reverse end-bearing capacity and

thus pullout capacity.

9.2 FUTURE WORK

A number of suggestions are given below for future research:



• It is recommended to measure the radial stress changes on the internal wall of

the caisson, during installation, consolidation and pullout, so as to obtain a

thorough understanding of the stress changes on both sides of the caisson.

• Direct measurements of the excess pore pressure generated both outside and

inside the caisson would enhance the understanding of the stress distribution. In

such measurements, some means of minimising soil disturbance during

installation of the instrument itself seems essential in achieving the target. Pore

pressure sensors made of tiny fibre optics are a possible choice.

• During sustained loading, longer times for loading before the final pullout are

suggested, in order to obtain a better view of the axial capacity and radial stress

changes during creep of the soil.

• Clay samples with different values of over-consolidation ratio (OCR) and

various sensitivities and plasticity indices should be tested, in order to

investigate the influence of these factors on the behaviour of caissons, and thus

provide a database for design purposes.

• Inclined loading tests (under monotonic loading, sustained loading and cyclic

loading) at different loading angles on the caisson are recommended to be

pursued, so as to investigate the relationship between radial stresses acting on

shaft of the caisson and the inclined pullout capacity, and the effect of inclined

loading on the reverse end-bearing capacity

• Field tests to investigate both the axial capacity and radial stress changes around

suction caissons would be beneficial.

Ref-1 References


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horizontal capacity of suction caissons. Proc., of 36th Offshore Technology Conf.,

Houston, Paper 16161.

Renzi, R., Maggioni, W., Smits, F. & Manes, V. (1991). A centrifugal study on the

behaviour of suction piles. Proceedings of the International Conference on

Centrifuge Modelling - Centrifuge ‘91, Colorado, USA, H.-Y. Ko (ed.), Balkema,

Rotterdam, 169-176.

Roscoe, K.H. (1970). The influence of strains in soil mechanics. Géotechnique, 20(2),

129-170.

Schofield, A.N. (1980). Cambridge geotechnical centrifuge operations. Géotechnique,

30(3), 227-268.

Skempton, A.W. (1951). The bearing capacity of clays. Building Research Congress,

London, 1, 180–189.

Solhjell, E., Sparrevik, P., Haldorsen, K. & Karlsen, V. (1998). Comparison and back-

calculation of penetration resistance from suction anchor installation in soft to stiff

clay at the Njord and Visund fields in the North Sea. Proceedings of the Society for

Underwater Technology Conference, London, 325-349.

Sparrevik, P. (2002). Suction pile technology and installation in deep waters. Proc. of

34th Annual Offshore Technology Conference, Houston, USA, 2171-2179.

Steenfelt, J. S., Randolph, M. F. & Wroth, C. P. (1981). Instrumented model piles

jacked into clay. Proc. of the Int. Conf. on Soil Mech. and Found. Engng., 2, 857-

864.

Steensen-Bach, J. O. (1992). Recent model tests with suction piles in clay and sand.

Proc., 24th Offshore Technology Conference. Houston, 323-330.

Stewart, D.P. (1992). Lateral loading of piled bridge abutments due to embankment

construction. PhD Thesis, Department of Civil Engineering, The University of

Western Australia.

Stewart, D.P. & Randolph, M.F (1991). A new site investigation tool for the centrifuge.

Centrifuge’91, Balkema, 531-538.

Ref-12 References


Stewart, D.P. & Randolph, M.F. (1994). T-bar penetration testing in soft clay. J.

Geotech. Eng. Div., ASCE, 120(12), 2230-2235.

Støve, O.J., Bysveen, S. & Christophersen, H.P. (1992). New foundation systems for

the Snorre development. Proceedings of the 24th Annual Offshore Tech. Conf.,

Houston, Paper OTC 6882, 75-83.

Supachawarote, C., Randolph, M.F. & Gourvenec, S. (2004). Inclined Pull-out Capacity

of Suction Caissons. Proc. of 14th Int. Offshore and Polar Engng. Conf.,

ISOPE’2004, Toulon, France. 500-506.

Taylor, R.N. (1995). Centrifuges in modelling: Principles and scale effects.

Geotechnical Centrifuge Technology, Blackie Academic & Professional, 19-33.

Templeton III, J.S. (2002). The role of finite elements in suction foundation design

Analysis. Proceedings of the 34th Annual Offshore Tech. Conf., Houston, Paper

OTC 14235, 2135-2143.

Tika, T. (1989). The effect of fast shearing on the residual strength of soils. Ph.D.

thesis, University of London (Imperial College).

Tjelta, T.I., Guttormsen, T.R., & Hermstad, J. (1986). Large scale penetration test at a

deepwater site. Proc. 18th Offshore Technology Conference. Houston, Texas.

Tjelta, T.I. (2001). Suction Piles: their position and application today. Proc. of the 11th

Inter. Offshore and Polar Engrg. Conf., ISOPE’ 2001, Stavanger, Norway, 2, 1-6.

Vesic, A.S. (1972). Expansion of cavities in infinite soil mass. J. Soil Mech. & Found.

Div., ASCE, 96(2), 561-584.

Watson, P.G. (1999). Performance of skirted foundations for offshore structures, Ph.D.

thesis, The University of Western Australia, Australia.

Watson, P.G., Suemasa, N. & Randolph, M.F. (2000). Evaluating undrained shear

strength using the vane shear apparatus. Proc. 10th Int. Conf. on Offshore and Polar

Engng, ISOPE’2000, Seattle, USA, 2, 485-493.

Weiler, W.A. & Kulhawy, F.H. (1982). Factors affecting stress cell measurements in

soil. J. Geotech. Engng. Div., ASCE, 108(12), 1529-1548.

Ref-13 References


Whittle, A.J. (1992). Assessment of an effective stress analysis for predicting the

performance of driven piles in clays. Proc. Conf. on Offshore Site Investigation and

Foundation Behaviour, Society for Underwater Technology, Kluwer, 28, 607-643.

Whittle, A.J. & Baligh, M.M. (1988). The Behavior of Piles Supporting Tension Leg

Platforms, Final Report Phase III. Research report, Dept. of Civil Engineering,

MIT, USA.

Whittle, A.J., Germaine, J.T. & Cauble, D.F. (1998). Behavior of miniature suction

caissons in clay. Offshore Site Inv. Found. Behavior '98, SUT 1998, 279-300.

Wong, H.Y. (1974). Some design and performance considerations of diaphgram type

pressure cells using strain gauges. Géotechnique, 24(1), 93-99.

Yoshimi, Y. & Kishida, T. (1981). A ring shear apparatus for evaluating friction

between soil and metal surfaces. ASTM Geotechnical Testing Journal, 4(4),

145-152.

Zdravkovic.L. Potts, D.M. & Jardine, R.J (2001). A parametric study of the pullout

capacity of bucket foundations in soft clay. Géotechnique, 51(1), 55-67.

Figure 1.1 Recent deepwater activities around the world (After ExxonMobile’s website: http://www2.exxonmobil.com)

Figure 1.2 Gas fields in deepwater of Exmouth Plateau, Australia (after Australian Government’s website,

http://www1.industry.gov.au/acreagereleases/Data/north_ex_plat/images/figure6.pdf)

(a) TLP (type1) (b) TLP (type 2) (c) SPAR (d) FPSO

Figure 1.3 Various anchoring systems for deepwater platforms (after Offshore-technology’s website: http://www.offshore-technology.com)

Figure 1.4 Distributions of FPSOs around the world in 2004

(After Mustang Engineering’s website: http://www.mustangeng.com)

(a) Suction caissons for Laminaria FPSOs, Australia (after Offshore-technology’s website: http://www.offshore-technology.com)

(b) Suction caissons used by Delmar Systems Inc.

(after Delmar Systems Inc.’s website: http://www.delmarus.com/)

Figure 1.5 Suction caissons used in the offshore industry

Figure 1.6 Large loading angles for the deepwater anchoring system of FPSO

(after Offshore-technology’s website: http://www.offshore-technology.com)

θ > 40°

Figure 2.1 Load transfer during penetration of caissons

(a) Full attachment (b) No attachment (c) Partial attachment (Very soft) (Very stiff) (Soft)

Figure 2.2 Flow mechanism of the soil inside the caisson

Water

Qtot

Qtip

Qside

Qextras

(a) Necessary underpressure ∆un (b) Allowable underpressure ∆ua

Figure 2.3 Necessary and allowable underpressure during suction installation

Figure 2.4 Suction caissons as mooring anchors for Na Kika FDS: elevation (after Newlin 2003a)

∆un

Suction

Base failure

∆ua

Suction

Figure 2.5 Typical soil parameters (NE anchor group) for Na Kika FDS (after Newlin 2003b)

Figure 2.6 Suction anchor pile sketch for Na Kika FDS: elevation (after Newlin 2003b)

(a) ∆un ~ depth (b) Range of ∆ua and ∆un~ depth

Figure 2.7 Necessary and allowable underpressures for NE anchor group

(after Newlin 2003b)

Figure 2.8 Actual applied underpressure and flow rate for NE group installations

(after Newlin 2003b)

Figure 2.9 Time from start of suction operation to final penetration for each pile,

including mechanical and other downtime (major instances noted) (after Newlin 2003b)

(a) installed by jacking (b) installed by suction

Figure 2.10 Displacement vectors during installation by jacking and by suction

around the caisson tip (after Andersen & Jostad, 2002)

Figure 2.11 Measured excess pore pressure at 9.5 m depth and 0.71 m away from the outside wall of the caisson as function of penetration depth

(after Andersen et al., 2003)

Figure 2.12 Measured, required and allowable underpressures versus penetration

depth of the caisson (after Andersen et al., 2003)

(a) NGI method (b) CEM, MTD

Figure 2.13 Soil flow below the caisson tip during suction installation:

assumptions by NGI method and CEM, MTD

Figure 2.14 Expansion of the cylindrical cavity in clay (after Vesic, 1972)

∆u

Suction

σri

∆u

σri

All soil move inside

All soil move outside

∆ushear induced ∆ui

∆ushear

+

=

∆uexpansion

Suction

Figure 2.15 Soil movements due to pile installation (after Randolph & Wroth, 1979)

Figure 2.16 Equivalent diameter for suction caissons or open-ended piles

d

t

dt2eq

d =

Figure 2.17 Deviatoric strain paths during simple pile penetration (after Baligh, 1985)

(a) Deformation pattern (after Baligh et al., 1987)

(b) Deviatoric strain contours (c) Octahedral shear strain contours

Figure 2.18 Deformations, strain and octahedral shear strain contours during

undrained simple sampler (d/t = 40) penetration in saturated clays (after Baligh et al., 1987; Chin, 1986)

Figure 2.19 Pile configuration of MTD (after Lehane & Jardine, 1994)

Figure 2.20 ‘h/d’ effect on normalised installation radial total stress

(after Lehane & Jardine, 1994) (Note: R = d/2)

Figure 2.21 Variation of the normalised radial total stress Hi with OCR (YSR) (after Lehane, 1992)

(a) Installation radial total stresses (b) Equalised radial effective stress profiles

Figure 2.22 Variation of radial stress around piles during consolidation in clay (after Lehane & Jardine, 1994)

(a) Unsealed (b) Sealed (c) Sealed (base-vented)

Figure 2.23 Different failure modes under vertical loading of caissons

(after Randolph & House, 2002)

W

Externalshaft

frictionαesu

Internal shaft

frictionαisu

qu(annulus)

V

W+Wplug

V

qutension(full base)

V

qubase(full base)

W

(a) Plan view: d = 6 mm (b) Side view: t = 0.6 mm

Figure 3.1 Geometry of total pressure transducer (Kyowa PS-10KA)

60.0

441

5.52.5

830

0.50.5

60.0

41

56.5

15.5 Transducer

44

48

4

Drainage

56.5

7

Valve

7.0

Ø2.0

2.5

6.0

56.0

15.0

8.5

120

45.0

4 8.5

41

Ø3

7.0

(t=0.5)

Stiffener 56.5

15.5

47

Pad-eye

Stiffener (t=1.5)

PPT

Transducer I

Front, D=6.5L=10(arc)

Hole I

56.5

7

Front,D=6.0

ValveDrainage

6.0

48. 5

0.50.5

12.06.5

2.5

(t=0.5)

73 15

Stiffener

41 45.0 52

.0

L=10(arc)

4 44

88

30

Transducer II

2.96.0

6.0

2.9

Ø2.0

Rear , D=6.5mm Hole II

Stiffener (t=1.5)

Pad-eye

Rear, D=6.0

PPT

(a) Side (b) Front

Figure 3.2 Elevation view of the designed model caisson 1 (units in mm)

6 mm

0.6 mm

30

2Ø4

3.51220

Ø8

Ø8

3mm OD(capped)For suction tubeattachment

0.5

Figure 3.3 Plan view of the designed model caisson 1 (units in mm)

(a) side (b) Front

Figure 3.4 Elevation view of the instrumented model caisson 1 after fabrication

L=120mm

d=30mm

TPT

Pad-eye

(a) Inner side (b) TPT inside the stiffener

Figure 3.5 Details of inner side and TPTs in stiffener of model caisson 1

(e) Details at the TPT area: TPT2 (f) Details at the TPT area: TPT1

Figure 3.6 Details of two TPTs instrumented on model caisson 1

0.5mm

0.6mm

1.5mm

1.0mm

Adhesive

TPT flush to wall

h 2=1

5.5

mm

Stiff

ener

-ste

p 2

Caisson wall

h 1=7

m

m

Stiff

ener

-ste

p 1

0.75mm

Leads covered by epoxy,

connected to DAQ system

Pad -eye

Caisson wall

TPT Electrical connection

t = 0.5 mm

56. 5

120

Pad-eyeTransducer

(t=0.5)

Stiffener (t=1)

4

30

4

Drainage

Valve

PPT

48.5

54.5

42.5

23.5

5.5

Ø2

126

7

8 8

2.5

Ø2.5

444

10

40.0

16.5 Ø7.0

10.0

1.5

0.5

1.0

35.0

10.0

11.5

45.0

63.5

0.5

40

(a) Elevation view in design

(b) Fabricated caisson

Figure 3.7 Details of the model caisson 2

30 mm

120 mm TPT

40 mm

Pad-eye

Figure 3.8 Details of connection between model caisson 1 and the 2 kN load cell

Figure 3.9 Arrangement of model caisson 1 in centrifuge

2 kN load cell connected to actuator

Jacking leg Caisson 1

Electrical leads from TPTs

PPT TPT

To syringe pump

Drainage valve

Actuator

PPT

Caisson

TPT1

Drainage tube

Water & kaolin clay in strong-box

Figure 3.10 Triaxial apparatus modified for calibrating TPTs

Figure 3.11 Configuration of Geotechnical Digital Systems (GDS) controller

Wire connecting TPTs and DAQ system

Clay sample: with caisson and TPTs inside

Stainless cylinder

Bottom plate

Top-cap

Figure 3.12 Pressure generation and data acquisition system for calibration of TPTs in water

Figure 3.13 WF25850 Bromhead type ring shear apparatus in UWA

p

Computer

Water

Triaxial apparatus

GDS

Input pressure

Output signal

Caisson

TPT

Sleeve

Loading yoke Weight

Level

Concentric ring

Proving ring Dial gauge

Figure 3.14 Sand-blasted top platen (left) and concentric ring (right)

Figure 3.15 Details of vertical load applied and shear transferred to the top platen in ring shear test

Torque arm

Concentric ring

Bearing rod

Top platen

100 mm

70 mm Filling soil inside: 5 mm thick

Figure 3.16 Stress under centrifugal acceleration

Figure 3.17 UWA fixed beam geotechnical centrifuge

Sample ‘strong-box’ on swinging platform

Loading actuator

Direction of rotation

rmin+z

40 mm

Nominal radiussnom h

32401800r −−=

Normal accelerationNgrωa 2

n =⋅=

Axis of rotation (0, 0)

na

Angular velocity, ω

1800 mm

r0

rmin

hs

z

390 mm Stress similitudevpσvmσ =

Strong box

Figure 3.18 Strong-box with actuator (with servo motor)

Figure 3.19 HEDS-5640 Optical Encoders on the servo motor

Actuator

Displacement transducer

Servo-motor

Strong-box

Power supply

Load cell

Motor

Encoder

(a) Elevation view (b) Internal

Figure 3.20 Dual slip ring

Figure 3.21 Instrumentation on the centrifuge arm

Pneumatic outlet

Hydraulic outlet

Electrical circulate

Conjunction box

Outlets for instruments

Syringe pump

Bearing arm

Strong-box

(a) Position on centrifuge

(b) Section view (after House, 2002)

Figure 3.22 Syringe pump

(a) Plan view (b) Position on centrifuge

Figure 3.23 Load cell (±2 kN capacity)

Syringe pump

Platform

Load cell

Actuator

(a) Plan view (b) Details at tip

Figure 3.24 T-bar penetrometer

Figure 3.25 Druck pore pressure transducer (PPT) inside a external fitting

Load cell

5 mm

20 mm

Figure 4.1 Wheatstone bridge for the diaphragm type total pressure transducer

Figure 4.2 ‘Arching effect’ for pressure cells in clay

D

p (not to scale)

(c) Normal stress Distributions over TPT in clay

Transducer

t

'Arch'

Diaphragm

Transducer

(a) Before loading

Soil mass

(b) After loading

Soil mass

p diaphragmDeformed

Figure 4.3 Hysteresis loop for TPTs (after Weiler & Kulhawy, 1982)

TPT

Top Cap

Lead

Figure 4.4 Adaptation on lead connection for triaxial apparatus

(a) Caisson on bottom plate (b) Caisson inside plastic tube

Figure 4.5 Design for triaxial calibration in water

(a) loading (b) unloading

Figure 4.6 Calibration of TPTs in water

Top-cap

Plastic sleeve

TPT

Bottom plate

Stainless cylinder

k2 = 0.524 kPa/bit

k1 = 0.560 kPa/bit

0

50

100

150

200

250

300

350

0 200 400 600

Output (bit)

p app

(kPa

)

TPT1TPT2

k2 = 0.530 kPa/bit

k1 = 0.563 kPa/bit

0

50

100

150

200

250

300

350

0 200 400 600

Output (bit)

p app

(kPa

)

TPT1TPT2

(a) Caisson in the kaolin clay sample (dissected after test)

(b) Kaolin clay sample in membrane

Figure 4.7 TPTs calibrated in kaolin clay in undrained condition

TPT

(a) load (b) unload

(c) cyclic loading (d) sustained loading

Figure 4.8 Calibration of TPTs on caisson in kaolin clay (undrained condition)

0

50

100

150

200

250

300

350

0 100 200 300

p app (kPa)

p mea

(kPa

)

TPT1TPT2Theoretical

0

50

100

150

200

250

300

350

0 100 200 300

p app (kPa)

p mea

(kPa

)

TPT1TPT2Theoretical

0

50

100

150

200

250

300

350

0 100 200 300

p app (kPa)

p mea

(kPa

)

TPT1TPT2Theoretical

0

50

100

150

200

0 50 100 150 200

p app (kPa)

p mea

(kPa

)

TPT1TPT2Theoretical

(a) Variation of pressure during consolidation in clay (first sample)

(b) Volume change of clay during consolidation (first sample)

(c) Change of CAF during consolidation (first sample)

153 kPa

-20

0

20

40

60

80

100

120

140

160

180

1 10 100 1000 10000 100000 1000000

Time (s)

Pres

sure

(kPa

)

TPT1

applied pressure

before installation

caisson installation

fillwater ap

ply

wat

er p

ress

ure

undrained drained

Starting draining here

-1.E+04

0.E+00

1.E+04

2.E+04

3.E+04

4.E+04

5.E+04

6.E+04

7.E+04

8.E+04

1 10 100 1000 10000 100000 1000000

Time (s)

Vol

ume

(mm

3 )

Volume change of sample

CAF=1.032

Starting draining here

-1.00

-0.50

0.00

0.50

1.00

1.50

1 10 100 1000 10000 100000 1000000

Time (s)

CA

F

Cell Action Factor, CAF

Average CAF=0.992

Figure 4.9 Calibration of TPTs in the first clay sample in consolidation

(a) Variation of pressure during consolidation in clay (second sample)

(b) Volume change of clay during consolidation (second sample)

(c) Change of CAF during consolidation (second sample)

0.92

0.94

0.96

0.98

1.00

1.02

1.04

0 100000 200000 300000 400000 500000 600000

Time (s)

CA

F

CAF

AVERAGE CAF = 0.969

undrained

appl

y pr

essu

re

papp = 154 kPa

drained

0

20

40

60

80

100

120

140

160

180

0 100000 200000 300000 400000 500000 600000

Time (s)

TPT

out

put (

kPa)

Start draining here

0

10000

20000

30000

40000

50000

60000

70000

0 100000 200000 300000 400000 500000 600000

Time (s)

Vol

ume

chan

ge (m

m3 )

Volumechange

Lowest CAF = 0.925 at t = 125267 s

Lowest pmea = 142.45 kPa at t = 125267 s

Figure 4.10 Calibration of TPTs in the second clay sample in consolidation

in water

in air

0

5

10

15

20

1 10 100 1000 10000Time (s)

TPT

out

put (

kPa) TPT1

TPT2

in water

in air

0

5

10

15

20

25

1 10 100 1000 10000 100000Time (s)

TPT

out

put (

kPa) TPT1

TPT2

(a) From air to water (test A1) (b)From air to water (test A2)

in kaolin

in air

0

5

10

15

1 10 100 1000 10000Time (s)

TPT

out

put (

kPa) TPT1

TPT2

in kaolin

in air

0

5

10

15

20

25

1 10 100 1000 10000 100000Time (s)

TPT

out

put (

kPa) TPT1

TPT2

(c) From air to kaolin slurry (test K1) (d) From air to kaolin slurry (test K2)

Figure 4.11 Change of initial value of TPTs in different media

Figure 4.12 Cross-sensitivity of TPTs to axial loading on the caisson

P

Radial total stress σr

TPT

0

20

40

60

80

100

120

140

0 200 400 600 800 1000 1200 1400

Axial load (N)

TPT

resp

onse

(kPa

)

TPT1

TPT2

Figure 4.13 TPTs responses to axial loading on the caisson

Figure 4.14 Caisson bearing TPTs for calibration in water in centrifuge

Actuator

Caisson

TPT1

Water

0

1

2

3

4

5

6

7

8

9

0 20 40 60 80 100

Pressure (kPa)D

epth

of T

PT (m

)

TPT1

TPT2

hydrostatic

(a) TPTs’ response along depth

Average CAF1=0.992

Average CAF2=0.989

0123456

789

10

0.95 0.96 0.97 0.98 0.99 1.00 1.01 1.02

CAF

Dep

th o

f TPT

(m)

TPT1

TPT2

Average

(b) Cell Action Factor (CAF) along depth

Figure 4.15 Calibration of TPTs in water in centrifuge

0

10

20

30

40

50

60

70

80

90

100

0 20000 40000 60000 80000 100000

Time (s)

Mea

sure

d pr

essu

re (k

Pa)

TPT1, average=87.40 kPa

TPT2, average=88.29 kPa

Figure 4.16 TPT response under sustained loading in water at 120 g

z = 7.2 m

In soil

In water

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250Radial total stress (kPa)

Dep

th o

f tip

(m)

NC clay, caisson 1

Hydrostatic

Figure 4.17 TPT responses during caisson installation in NC clay in centrifuge

z = 4.8 m

In water

In soil

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

LOC clay, caisson 2

Hydrostatic

Figure 4.18 TPT responses during caisson installation in LOC clay in centrifuge

(a) Force - shearing distance (b) Vertical settlement - shearing distance

(c) Shear stress - shearing distance (d) δ - shearing distance

Figure 5.1 Ring shear test S1-1 in NC clay (smooth platen)

-0.6

-0.5

-0.4

-0.3

-0.2

-0.1

0

0.1

0.2

0 20 40 60

Shearing distance (mm)

Ver

tical

Dis

plac

emen

t (m

m)

0

5

10

15

20

25

30

35

40

0 20 40 60


Forc

e (N

) Proving Ring A

Proving Ring B

0

2

4

6

8

10

12

14

16

18

0 20 40 60


Inte

rfac

e fr

ictio

n an

gle

(o )

0

5

10

15

20

25

30

35

0 20 40 60


Shea

r St

ress

, τ

(kPa

)

σ v = 100 kPa

δr = 15




-0.01

0

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0 10 20 30

Distance (mm)V

ertic

al D

ispl

acem

ent (

mm

)

0

10

20

30

40

50

60

70

0 10 20 30

Distance (mm)

Forc

e (N

) Proving Ring A

Proving Ring B

0

5

10

15

20

25

30

0 10 20 30

Distance (mm)

Inte

rfac

e fr

ictio

n an

gle

(o )

0

10

20

30

40

50

60

0 10 20 30

Distance (mm)

Shea

r St

ress

, τ

(kPa

)

σ v = 100 kPa

δr = 26

(a) Force ~ shearing distance (b) Vertical settlement ~ shearing distance

(c) Shear stress ~ shearing distance (d) δ ~ shearing distance


-0.002

0

0.002

0.004

0.006

0.008

0.01

0.012

0 10 20 30

Distance (mm)

Ver

tical

Dis

plac

emen

t (m

m)

0

10

20

30

40

50

60

70

80

90

0 10 20 30Distance (mm)

Forc

e (N

)

Proving Ring A

Proving Ring B

0

5

10

15

20

25

30

0 10 20 30

Distance (mm)

Inte

rfac

e fr

ictio

n an

gle

(o )

0

10

20

30

40

50

60

70

80

0 10 20 30

Distance (mm)

Shea

r St

ress

, τ

(kPa

)

σ v = 125 kPa

δr = 24

(a) Peak friction angle δp ~ PI

(b) Residual friction angle δult (or δr) ~ PI

Figure 5.4 Correlation between δp , δr and PI for displacement piles (after Lemos, 1986, Tika, 1989 and Lehane, 1992)

Figure 5.5 Interface after shearing for sand-blasted platen

-0.0005

0

0.0005

0.001

0.0015

0.002

0.0025

0.003

0.0035

0.004

0.0045

0 1 2 3 4 5 6 7 8Shearing distance (mm)

Ver

tical

dis

plac

emen

t (m

m)

0

5

10

15

20

25


Forc

e (N

) Proving Ring A

Proving Ring B

(a) Force ~ shearing distance (b) Vertical settlement ~ shearing distance

δ p = 19.1o

δr = 17.5o

0

5

10

15

20

25

0 1 2 3 4 5 6 7 8


Inte

rfac

e fr

ictio

n an

gle

(o)

0

2

4

6

8

10

12

14

16

18

20

0 1 2 3 4 5 6 7 8


Shea

r St

ress

, τ (k

Pa)

σ v = 50 kPaσ v = 50 kPa

(c) Shear stress ~ shearing distance (d) δ ~ shearing distance

Figure 5.6 Ring shear test S2-1 in NC clay (sand-blasted platen)

00.010.020.03

0.040.050.060.07

0.080.09

0.1

0 5 10 15 20Shearing distance (mm)

Ver

tical

dis

plac

emen

t (m

m)

0

5

10

15

20

25

0 5 10 15Shearing distance (mm)

Forc

e (N

)

Proving Ring A

Proving Ring B

σ v = 50 kPa


δ p = 19.4o

δ r = 17.7o

0

5

10

15

20

25

0 5 10 15Shearing distance (mm)

Inte

rfac

e fr

ictio

n an

gle

(o)

0

2

4

6

8

10

12

14

16

18

20

0 5 10 15


Shea

r St

ress

, τ (k

Pa)

σ v = 50 kPa

(c) Shear stress- shearing distance (d) δ - shearing distance

Figure 5.7 Ring shear test S2-2 in NC clay (sand-blasted platen)

Average δp=19.3

Average δr = 17.6

0

5

10

15

20

25


Inte

rfac

e fr

ictio

n an

gle

( o )

S2-1

S2-2

Figure 5.8 Comparison between ring shear tests S2-1 and S2-2 in NC clay

-0.001

-0.0005

0

0.0005

0.001

0.0015

0.002

0.0025

0.003

0.0035

0 1 2 3 4 5 6Shearing distance (mm)

Ver

tical

Dis

plac

emen

t (m

m)

0

5

10

15

20

25

0 1 2 3 4 5 6


Forc

e (N

)

Proving Ring A

Proving Ring B

σ v = 50 kPa


δr = 18.3o

δp = 19.4o

0

5

10

15

20

25


Inte

rfac

e fr

ictio

n an

gle

(o )

02468

101214161820


Shea

r St

ress

, τ (k

Pa)

σ v = 50 kPa


Figure 5.9 Ring shear test S3-1 (LOC clay)

-0.001

0

0.001

0.002

0.003

0.004

0.005

0.006

0.007

0 5 10 15

Shearing distance (mm)V

ertic

al D

ispl

acem

ent (

mm

)

0

5

10

15

20

25

0 5 10 15


Forc

e (N

)

Proving Ring A

Proving Ring B

σ v = 50 kPa


0

2

4

6

8

10

12

14

16

18

0 5 10 15


Shea

r St

ress

, τ (k

Pa)

δp = 18.3o

δr = 17.9o

0

5

10

15

20

0 5 10 15


Inte

rfac

e fr

ictio

n an

gle

(o )

σ v = 50 kPa


Figure 5.10 Ring shear test S3-2 (LOC clay)

-0.001

0

0.001

0.002

0.003

0.004

0.005

0.006

0 5 10 15

Shearing distance (mm)V

ertic

al D

ispl

acem

ent (

mm

)

0

2

4

6

8

10

12

14

16

18

0 5 10 15


Forc

e (N

)

Proving Ring A

Proving Ring B

σ v = 50 kPa


δr = 14.8o

δr = 11.3o

0

5

10

15

20

0 5 10 15


inte

rfac

e fr

ictio

n an

gle

(o )

0

2

4

6

8

10

12

14

0 5 10 15


Shea

r St

ress

, τ (k

Pa)

σ v = 50 kPa


Figure 5.11 Ring shear test S3-1 (sensitive clay)

-0.01

0

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0 5 10 15


Ver

tical

Dis

plac

emen

t (m

m)

0

2

4

6

8

10

12

14

16

0 5 10 15


Focr

e (N

)

Proving Ring A

Proving Ring B

σ v = 50 kPa


0

2

4

6

8

10

12

14

0 5 10 15


Shea

r St

ress

, τ (k

Pa)

δp = 14.3o

δr = 12.1ο

0

5

10

15

0 5 10 15


inte

rfac

e fr

ictio

n an

gle

(o )

σ v = 50 kPa

(c) Vertical settlement - shearing distance (d) δ - shearing distance

Figure 5.12 Ring shear test S3-2 (sensitive clay)

Figure 6.1 Plan view of test locations

0.94 kPa/m

0.74 kPa/mk=1.18 kPa/m

k

Install

k=1.26 kPa/m

1Pullout

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20 25

Undrained shear strength su (kPa)

Dep

th (m

)

FirstMiddleEnd

NT-bar = 10.5

Figure 6.2 T-bar tests in NC kaolin clay

100

90 140 140 140

185 115 140 210

650

9595

100

390

Tbar-A1

Tbar-A2

Tbar-C1

Tbar-C2

Test2 Test1 Test 3

Strong-box

140

Test 4

Tbar-B1

Tbar-B2

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20

Undrained shear strength su measured by T-bar (kPa)

Dep

th o

f tip

(m) 1st in1st out

2nd

NT-bar = 10.5

(a) Strength profiles during installation and pullout

0.000.100.200.300.400.500.600.700.800.901.001.101.201.301.401.50

1 2 3 4 5 6 7 8 9 10

No. of cycles

Rem

ould

ed r

atio

depth =10 m

depth =11 m

depth =12 mReduced to0.36 - 0.44 of intact strength

(b) Resistance ratios versus cycles of T-bar penetration Figure 6.3 Strength profiles and remoulded ratios of cyclic T-bar tests in NC

kaolin clay (B12TB1, OCR=1, 9 cycles)

1st out

2nd

1st in

0

2

4

6

8

10

12

14

16

-30 -20 -10 0 10 20 30


Dep

th o

f tip

(m)

NT-bar = 10.5


0.000.100.200.300.400.500.600.700.800.901.001.101.201.301.401.50

1 2 3 4 5 6 7 8 9 10 11 12

No. of cycles

Rat

io

depth = 10 m

depth = 11 m

depth = 12 m

Reduced to 0.42-0.45 of intact strength

(b) Remoulded ratios versus cycles of T-bar penetration

Figure 6.4 Strength profiles and remoulded ratios of cyclic T-bar tests in LOC clay (B13TB1, OCR = 1.5, 11 cycles)

2nd, in

1st, out 1st, in

0

2

4

6

8

10

12

14

16

-30 -20 -10 0 10 20 30 40


Dep

th o

f tip

(m)

NT-bar = 10.5


0.000.100.200.300.400.500.600.700.800.901.001.101.201.301.401.50

1 2 3 4 5 6 7 8 9 10 11 12

No. of cycles

Rat

io

depth =10 m

depth = 11 m

depth = 12 m Reduced to 0.40 -0.45 intact strength

(b) Remoulded ratios versus cycles of T-bar penetration

Figure 6.5 Strength profiles and remoulded ratios of cyclic T-bar tests in LOC

clay (B13TB2, OCR=1.5, 11 cycles)


0.00

0.10

0.20

0.30

0.40

0.50

0.60

0.70

0.80

0.90

1.00

1 2 3 4 5 6 7 8 9 10 11 12 13

No. of cycles

Rat

io Remoulded ~ 22 % of the undisturbed strength

(b) Remoulded ratios versus cycles of T-bar penetration (depth of 11 m)

Figure 6.6 Strength profiles and remoulded ratios of cyclic T-bar tests in sensitive

clay (B14TC1, 12 cycles)

0

2

4

6

8

10

12

14

16

-20 -10 0 10 20 30


Dep

th o

f tip

(m)

su,ave=1.55 kPa/mNT-bar = 10.5


0.00

0.10

0.20

0.30

0.40

0.50

0.60

0.70

0.80

0.90

1.00

1 2 3 4 5 6 7 8 9 10 11 12 13

No. of cycles

Rat

io Remoulded ~ 22 % of

the undisturbed strength

(b) Remoulded ratios versus cycles of T-bar penetration (depth of 11 m)

Figure 6.7 Strength profiles and remoulded ratios of cyclic T-bar tests in sensitive

clay (B14TC2, 12 cycles)

Tbar test in NC Kaolin clay (B14TC2)

0

2

4

6

8

10

12

14

16

-20 -10 0 10 20 30


Dep

th o

f tip

(m)

su,ave=1.61 kPa/mNT-bar=10.5 kPa/m

(a) Full attachment (b) No attachment (c) Partial attachment (Very soft) (Very stiff) (Medium)

Figure 6.8 Flow mechanism of the soil inside the caisson and shaft friction above the first stiffener

∆p = P/A

Nominal installation: 13.92m

02468

10121416

0 50 100 150

Penetration resistance (kPa)

Dep

th o

f tip

(m)

Measured

Install:Nc=7.5,Alpha=0.41

Nominal installation: 13.92m

02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic

(a) Penetration resistance (b) Internal pore pressure

Figure 6.9 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B2JOI)

τi-a=0 kPaτi-a=αi-a⋅su

αi-a=αi-b

τi-a=0.5 kPa for NC clay, and ~1 kPa for LOC clay

Water Water

Nominal installation: 14.05 m

∆p = P/A

02468

10121416

0 50 100 150


Dep

th o

f tip

(m)

Measured



02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic


Figure 6.10 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B2JOC)


∆p = P/A

02468

10121416

0 50 100 150 200


Dep

th o

f tip

(m)

Measured


Figure 6.11 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B4JCI)

∆p = P/A


02468

10121416

0 50 100 150


Dep

th o

f tip

(m)

Measured



02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



∆p = P/A


02468

10121416

0 50 100 150


Dep

th o

f tip

(m)

Measured


Figure 6.13 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B6JCC)


∆p = P/A

02468

10121416

0 50 100 150


Dep

th o

f tip

(m)

Measured



02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



0

2

4

6

8

10

12

14

16

0 20 40 60 80 100 120 140


Dep

th o

f tip

(m)

B2JOI

B2JOC

B4JCI

B6JOC

B6JCC

B8JOC

Nc=7.5, alpha=0.39

Figure 6.15 Penetration resistance during jacked installation in NC clay


0

2

4

6

8

10

12

14

16

-150 -100 -50 0 50 100 150Pressure (kPa)

Dep

th o

f tip

(m)

PPT-internalSyringe-pumpStress-holdNet pressureHydrostatic


Suction started at 6.87 m

0

2

4

6

8

10

12

14

16

-150 -100 -50 0 50 100 150

Pressure (kPa)

Dep

th o

f tip

(m)


Figure 6.16 Variation of pressures during suction installation in the old system (NC clay, test B3SCI)


0

2

4

6

8

10

12

14

16

-200 -150 -100 -50 0 50 100 150

Pressure (kPa)

Dep

th o

f tip

(m)


Figure 6.17 Variation of pressures during suction installation in the old system (NC clay, test B3SOI)


0

2

4

6

8

10

12

14

16

0 20 40 60 80 100 120 140 160

Pressure (kPa)

Dep

th o

f tip

(m)

PPT-internalStress-holdNet pressureHydrostatic

Figure 6.18 Variation of pressures during suction installation in the old system (NC clay, test B3SCC, DAQ data from syringe pump not received)

Nominalinstallation: 14.06 m

0

2

4

6

8

10

12

14

16

-100 -50 0 50 100 150 200

Pressure (kPa)

Dep

th o

f tip

(m)


Figure 6.19 Variation of pressures during suction installation in the new system (NC clay, test B9SOI, DAQ data from syringe pump not received)


0

2

4

6

8

10

12

14

16

-50 0 50 100 150

Pressure (kPa)

Dep

th o

f tip

(m)


Figure 6.20 Variation of pressures during suction installation in the new system (NC clay, test B10SOC, DAQ data for syringe pump not received)


zs=6.96 m

zplug=13.53 m

0

2

4

6

8

10

12

14

16

-250 -200 -150 -100 -50 0 50 100 150

Pressure (kPa)

Dep

th o

f tip

(m)

Load cellPPT-internalSyringe-pumpHydrostaticNet pressure

zfinal=13.90 m

Figure 6.21 Variation of pressures during suction installation in the new system (NC clay, test B10SCI)


zs=6.19 m

zplug=13.59 m

0

2

4

6

8

10

12

14

16

-250 -200 -150 -100 -50 0 50 100 150

Pressure (kPa)

Dep

th o

f tip

(m)


zfinal=13.93 m

Figure 6.22 Variation of pressures during suction installation in the old system (NC clay, test B10SCC)

zplug=13.36 m

zs = 4.95 m

0

2

4

6

8

10

12

14

16

-500 -450 -400 -350 -300 -250 -200 -150 -100 -50 0 50 100 150Pressure (kPa)

Dep

th o

f tip

(m)


zfinal=14.30 m


Figure 6.23 Variation of pressures during suction installation in the new system (NC clay, test B11SOC)


zfinal=14.39 m

zs = 5.62 m

zplug=13.36 m

0

2

4

6

8

10

12

14

16

-600 -500 -400 -300 -200 -100 0 100 200

Pressure (kPa)

Dep

th o

f tip

(m)


Figure 6.24 Variation of pressures during suction installation in the new system (NC clay, test B12SCC)

0

2

4

6

8

10

12

14

16

0 50 100 150


Dep

th o

f tip

(m)

B3SOI

B3SCI

B3SCC

B9SOI

B10SCI

B10SOC

B10SCC

B11SOC

B12SCC

Nc=7.5, alpha=0.38

Self-weight penetration

Suction installation

Figure 6.25 Penetration resistance for caissons installed by suction in NC clay

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250 300Time (s)

Dep

th o

f tip

(m) z=6.63 m, end of self-weight

penetration (jacking)

start of self-weight penetration (jacking)

v=1.50 mm/s z=6.87 m, start of suction

z=14.13 m, end of suction

v=1.23 mm/s

Time delay =113.9 s, or19 days in prototype

Figure 6.26 Depth of tip versus time during suction installation in the old system (test B3SCI, NC clay) (note: time and velocity are shown in model scale)

0123456789

101112131415

0 50 100 150 200Time (s)

Dep

th o

f tip

(m)

z=4.54 m, end of self-weight penetration (jacking)

start of jacking

v=1.50 mm/s z=4.95 m, start of suction installation

z=13.90 m, speed decreasesto 0.36 mm/s

v=1.77 mm/s during suction installation, V=10.8, undrained

v=0.12 mm/s during suction installation, V=0.73, partly drained

Figure 6.27 Depth of tip versus time during suction installation in the new system (test B11SOC, NC clay) (note: time and velocity are in model scale)

0

2

4

6

8

10

12

14

16

0 50 100 150


Dep

th o

f tip

(m)

B3SOIB3SCIB3SCCB9SOIB10SCIB10SOCB10SCCB11SOCB12SCCNc=7.5, alpha=0.39B2JOIB2JOCB4JCIB6JOCB6JCCB8JOC



Figure 6.28 Penetration resistance for caissons installed by jacking and by suction in NC clay

0

2

4

6

8

10

12

14

16

0 50 100 150


Dep

th o

f tip

(m)

B3SOI

B3SCI

B3SCC

B10SCI

B11SOC



Figure 6.29 Penetration resistance for caissons installed by suction with and

without a time delay in NC clay

(a) ∆un - depth (b) ∆ua , ∆uapp - depth

(c) Fs - depth (d) hs,pre - depth

02468

10121416

0 50 100 150

Necessary underpressure (kPa)

Dep

th o

f tip

(m)

02468

10121416

0.0 0.5 1.0 1.5Predicted soil heave (m)

Dep

th o

f tip

(m)

02468

10121416

0 100 200 300 400

Allowable underpressure (kPa)

Dep

th o

f tip

(m)

AllowableNecessaryApplied

02468

10121416

0 1 2 3 4 5

Factor of safety

Dep

th o

f tip

(m)

13.36 m, plug reached lid

1.43 m

13.36 m, plug reached lid, Fs=2.06

4.95 m, suction starts

Figure 6.30 Predicted ∆un, ∆ua, hs,pre, actual applied underpressure and actual Fs versus depth of suction installation (test B11SOC, NC clay)

0

2

4

6

8

10

12

14

16

0 20 40 60 80 100 120 140


Dep

th o

f tip

(m)

Nc=6.0Nc=6.5Nc=7.0Nc=7.5Nc=8.0Nc=8.5Nc=9.0Nc=9.5Nc=10.0Nc=11.0Nc=12.0Measured



α = 0.38

Figure 6.31 Back-figured penetration resistance of caissons with different Nc values in NC clay (α = 0.38)

0

2

4

6

8

10

12

14

16

0 20 40 60 80 100 120 140 160


Dep

th o

f tip

(m)

B10SOC*

B3SOC*

B6JOC*

B5JOC*

B2JOC*

Figure 6.32 Penetration resistance of re-installation in disturbed sites in NC clay

0

2

4

6

8

10

12

14

16

0 20 40 60 80 100 120 140

Penetration resistance (kPa)D

epth

of t

ip (m

)

B2JOI

B2JOC*

B10SOI

B10SOC*



B10SOC*

B2JOI

B10SOI

B2JOC*

Figure 6.33 Penetration resistance during original installation and re-installation in NC clay

(a) ∆un - depth (b) ∆ua, ∆uapp - depth


02468

10121416

0 1 2 3 4 5Factor of safety

Dep

th o

f tip

(m)

02468

10121416

0 50 100Necessary underpressure (kPa)

Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

02468

10121416

0 200 400 600Allowable underpressure (kPa)

Dep

th o

f tip

(m)


11.73 m, plug reaches lid

0.98 m


Figure 6.34 Predicted ∆un, ∆ua, hs,pre, actual applied underpressure and actual Fs versus depth of suction installation (test B10SOC*, NC clay)

∆p = P/A


02468

10121416

0 50 100 150 200


Dep

th o

f tip

(m)

Measured


Nominal installation:

13.89 m

02468

10121416

0 50 100 150 200

Pressure (kPa)

Dep

th o

f tip

(m)

Internal PPT

Hydrostatic

zfinal=13.99 m zfinal=13.99 m

Figure 6.35 Penetration resistance and the internal pore pressure during jacked installation in LOC clay (test B13JCC)

0

2

4

6

8

10

12

14

16

-500 -400 -300 -200 -100 0 100 200Pressure (kPa)

Dep

th o

f tip

(m)

Internal PPTLoad cellSyringe-pumpHydrostaticNet pressure


zs=7.31 m

zplug=13.62 mzfinal=13.90 m

Figure 6.36 Variation of pressures during suction installation in LOC clay (test B13SCC)

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100 200

Pressure (kPa)D

epth

of t

ip (m

)

Internal PPTLoad cellSyringe pumpNet pressureHydrostatic


zs=7.77 m

zplug=13.52 mzplug=13.72 m

Figure 6.37 Variation of pressures during suction installation in LOC clay (test B13sus)

0

2

4

6

8

10

12

14

16

-600 -500 -400 -300 -200 -100 0 100 200Pressure (kPa)

Dep

th o

f tip

(m)

Internal PPT

Syringe pump

Load cell

Hydrostatic

Net pressure


zs=7.85 m

zplug=12.12 m

zfinal=13.53 m

Figure 6.38 Variation of pressures during suction installation in LOC clay (test B13cyc)

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250 300Time (s)

Dep

th o

f tip

(mm

)

Start of self-weight penetration

v=2.77 mm/s

z = 7.31 m, start of suction installation

v=1.80 mm/s during suction installation, V=11.8: undrained

z = 13.72 m, v reduces to 0.18 mm/s, V=1.2: partly drained

z=6.44-7.31 m, v=0.53 mm/s, V=3.5: partly drained

z = 6.97 m, end of jacking, followed by 1.1 s of time delay before suction

Figure 6.39 Variation of the depth of the caisson tip with time during suction installation in LOC clay (test B13SCC) (Note: time and v are in model scale)

0

2

4

6

8

10

12

14

16

0 20 40 60 80 100 120 140Time (s)

Dep

th o

f tip

(m)


v=2.83 mm/sz=7.97 m, start of suction installation

Before z=7.52 m, by jacking, v=2.83 mm/s, V=19, undrained

z = 7.60 - 7.97 m, jacking ended and suction started, with 3.7 s of time delay

z=7.52 - 7.97 m, v=0.55 mm/s, V=3.6: partly drained

z=7.97 - 13.59 m, v=1.90 mm/s, model scale, V=12.5: undrained z=13.59 m, v reduces to 0.20 mm/s,

V=1.3: partly drained

Figure 6.40 Variation of the depth of the caisson tip with time during suction

installation in LOC clay (test B13sus) (Note: time and v are in model scale)

0123456789

101112131415

0 20 40 60 80 100 120 140 160 180 200

Time (s)D

epth

of t

ip (m

)


z=7.69 - 7.85 m, start of suction installation

z=7.69 - 7.85 m, change from jacking to suction, v=0.65 mm/s, V=4.3, partly drained. Time delay is 2.9 s

Before z=7.69 m, by jacking, v=2.74 mm/s, V=18, undrained

z=7.85 - 12.76 m, by suction, v=1.89 mm/s, V=12.4, undrained

z=12.76 m: speed reduces to 0.21 mm/s , V=1.4, partly drained

Figure 6.41 Variation of the depth of the caisson tip with time during suction installation in LOC clay (test B13cyc) (Note: time and v are in model scale)

(a) ∆un - depth (b) ∆ua and ∆uapp- depth


02468

10121416


Dep

th o

f tip

(m)

02468

10121416

0.0 0.5 1.0 1.5

Predicted soil heave (m)

Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

Allowable

Necessary

Applied

02468

10121416

0 50 100 150

Necessary underpressure (kPa)

Dep

th o

f tip

(m)


13.62 m, plug reaches lid, Fs,plug = 2.26

1.15 m

7.31 m, suction starts

.

Figure 6.42 Predicted ∆un, ∆ua, hs,pre and actual ∆uapp, actual Fs versus depth

during suction installation (test B13SCC, LOC clay)



02468

10121416


Dep

th o

f tip

(m)

02468

10121416

0 50 100 150Necessary underpressure (kPa)

Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

02468

10121416

0 100 200 300 400Allowable underpressure (kPa)

Dep

th o

f tip

(m)

Allowable

Necessary

Applied



1.12 m

Figure 6.43 Predicted ∆un, ∆ua, hs,pre and actual ∆uapp, actual Fs versus depth

during suction installation (test B13sus, LOC clay)

(a) ∆un - depth (b) ∆ua - depth


02468

10121416


Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

Allowable

Necessary

Applied



1.10 m

Figure 6.44 Predicted ∆un, ∆ua, hs,pre and actual ∆uapp, actual Fs versus depth during suction installation (test B13cyc, LOC clay)

∆p = P/A

0

2

4

6

8

10

12

14

16

0 50 100 150 200

Penetration resistance (kPa)D

epth

of t

ip (m

)

B13JCC

B13SCC

B13cyc

B13sus

Figure 6.45 Variation of penetration resistance with depth of caissons installed by jacking and by suction in LOC clay (Box13, OCR=1.5)

NT-bar = 10.5

0

2

4

6

8

10

12

14

16

-30 -20 -10 0 10 20 30 40


Dep

th o

f tip

(m)

B13SCC, k =1.64 kPa/m

B13JCC, k =1.64 kPa/m

B13sus, k =1.76 kPa/m

B13cyc, k =1.77 kPa/m

Figure 6.46 T-bar strength profiles of caisson tests installed by jacking and by suction in LOC clay (Box13, OCR=1.5)

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100 200Pressure (kPa)

Dep

th o

f tip

(m)

Internal PPTLoad cellSyringe-pumpHydrostaticNet resistance


zs=9.12 m

zplug=13.62 m

zfinal=15.30 m

Figure 6.47 Variation of pressures during suction installation in sensitive clay (test B14SCC)

0

2

4

6

8

10

12

14

16

-200 -100 0 100 200 300Pressure (kPa)

Dep

th o

f tip

(m)



zs=9.40 m

zplug=13.63 m

zfinal=14.59 m

Figure 6.48 Variation of pressures during suction installation in sensitive clay (test B14sus)

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100 200Pressure (kPa)

Dep

th o

f tip

(m)

Syringe pump

Load cell

Hydrostatic

Net resistance

Internal pore pressure

by self-weight

by suction

creep


zs=7.45 m

zplug=13.44 m

zfinal=15.24 m

Figure 6.49 Variation of pressures during suction installation in sensitive clay (test B14cyc)

0

2

4

6

8

10

12

14

16

-650 -550 -450 -350 -250 -150 -50 50 150

Pressure (kPa)

Dep

th o

f tip

(m)



zs=6.74 m

zplug=13.45 m

zfinal=14.77 m

Figure 6.50 Variation of pressures during suction installation in sensitive clay (test B14susa)

0123456789

10111213141516

0 50 100 150 200 250 300Time (s)

Dep

th o

f tip

(m)


v=2.75 mm/s, in jacking

z = 9.12 m, start of suction, time delay = 1.8 s

after z = 9.12 m, v=1.83 mm/s, V=14.5: undrained

z=14.57 m, radial stress decreases and moving speed reduces

v = 0.96 mm/s, V = 7.6, partly drained

Figure 6.51 Variation of the depth of the caisson tip with time during suction installation in sensitive clay (test B14SCC) (Note: time and v are in model scale)

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10111213141516

0 20 40 60 80 100 120 140 160

Time (s)

Dep

th o

f tip

(mm

) z = 9.40 - 14.45 m, v = 1.79 mm/s, V=14: undrained

v=2.94 mm/s


z = 9.40 m, suction started, time delay = 3 s

v=0.92 mm/s, V=7.3, partly drained

z = 14.45 m, v decreases to 0.04 mm/s, V=0.3: partly drained

Figure 6.52 Variation of the depth of the caisson tip with time during suction installation in sensitive clay (test B14sus) (Note: time and v are in model scale)

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10111213141516

0 100 200 300 400 500 600Time (s)

Dep

th o

f tip

(m)


v=2.68 mm/s

z=7.45 m, start of suction, time delay = 1 s

v=1.89 mm/s during suction installation, v = 2.18 mm/s, V = 17: undrained,

v= 0.81 mm/s, V=6.5, partly drained

z = 14.20 m, v reduces to 0.03 mm/s, V = 0.24: partly drained (creep)

Figure 6.53 Variation of the depth of the caisson tip with time during suction installation in sensitive clay (test B14cyc) (Note: time and v are in model scale)

0123456789

10111213141516

0 100 200 300 400 500Time (s)

Dep

th o

f tip

(m)


v=1.46 mm/s

z=6.74-14.20 m, start of suction, time delay = 5.1 s,

z=6.74-14.20 m, v = 1.34 mm/s, V=11: undrained, by suction

v = 0.4 mm/s, V=3.2: partly drained

z = 14.20 m, v reduces to 0.02 mm/s, V=0.2: undrained

Figure 6.54 Variation of the depth of the caisson tip with time during suction installation in sensitive clay (test B14susa) (Note: time and v are in model scale)



02468

10121416


Dep

th o

f tip

(m)

02468

10121416

0 20 40 60 80Necessary underpressure (kPa)

Dep

th o

f tip

(m)

02468

10121416

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6Predicted soil heave (m)

Dep

th o

f tip

(m)

02468

10121416

0 100 200 300 400Allowable underpressure (kPa)

Dep

th o

f tip

(m)




1.31 m

Figure 6.55 Predicted ∆un, ∆ua, hs,pre, actual applied ∆uapp and actual Fs versus depth during suction installation (test B14SCC, sensitive clay)



02468

10121416

0 1 2 3 4 5 6 7 8Factor of safety

Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)

02468

10121416


Dep

th o

f tip

(m)




1.14 m

Figure 6.56 Predicted ∆un, ∆ua, hs,pre, actual applied ∆uapp and actual Fs versus

depth during suction installation (test B14sus, sensitive clay)



02468

1012141618

0 1 2 3 4 5 6 7Factor of safety

Dep

th o

f tip

(m)

02468

1012141618

0 50 100Necessary underpressure (kPa)

Dep

th o

f tip

(m)

02468

1012141618

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4Predicted soil heave (m)

Dep

th o

f tip

(m)

02468

1012141618


Dep

th o

f tip

(m)



13.44 m, plug reaches lid, Fs,plug

= 4.611.20 m

Figure 6.57 Predicted ∆un, ∆ua, hs,pre, actual applied ∆uapp and actual Fs versus

depth during suction installation (test B14cyc, sensitive clay)



02468

10121416

0 1 2 3 4 5

Factor of safety

Dep

th o

f tip

(m)

02468

10121416

0 20 40 60 80Necessary underpressure (kPa)

Dep

th o

f tip

(m)

02468

10121416

0.0 0.5 1.0 1.5

Predicted soil heave (m)

Dep

th o

f tip

(m)

02468

10121416

0 200 400 600

Allowable underpressure (kPa)

Dep

th o

f tip

(m)




1.25 m

Figure 6.58 Predicted ∆un, ∆ua, hs,pre, actual applied ∆uapp and actual Fs versus depth during suction installation (test B14susa, sensitive clay)



0

2

4

6

8

10

12

14

16

0 50 100Penetration resistance (kPa)

Dep

th o

f tip

(m)

B14SCCB14susB14cycB14susa

Figure 6.59 Variation of penetration resistance with depth for caissons installed by suction in sensitive clay (Box 14, St = 4 - 5)

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25

Undrained shear strength su from T-bar tests (kPa)

Dep

th o

f tip

(m)

B14SCC, k = 1.16 kPa/m

B14sus, k = 1.33 kPa/m

B14cyc, k =1.36 kPa/m

B14susa, k = 1.58 kPa/m

NT-bar = 10.5

Figure 6.60 T-bar strength profiles for caissons installed by suction in sensitive clay (Box 14, St = 4 - 5)

Pullout

∆p = P/A

02468

10121416

-200 -100 0 100 200

Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic

(a) Axial pressure (b) Internal pore pressure

Figure 6.61 Installation and pullout pressure and internal pore pressure changes

during unsealed pullout (NC clay, test B2JOI)

∆p= P/A

02468

10121416

-200 -100 0 100 200


Dep

th o

f tip

(m)

02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic

(a) Axial pressure (b) Internal pore pressure Figure 6.62 Installation and pullout pressure and internal pore pressure changes

during unsealed pullout (NC clay, test B8JOI)

0

2

4

6

8

10

12

14

16

-150 -100 -50 0 50 100 150Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

0

2

4

6

8

10

12

14

16

0 50 100 150 200Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



during unsealed pullout (NC clay, test B9SOI)

0

2

4

6

8

10

12

14

16

-150 -100 -50 0 50 100 150

Axial capacity, ∆p (kPa)

Dep

th o

f tip

(m)

B2JOI

B8JOI

B9SOIInstall

Pullout

Figure 6.64 Installation and pullout pressure of unsealed pullout immediately after installation in NC clay

∆p=P/A

02468

10121416

-300 -200 -100 0 100 200

Axial pressure, ∆p (kPa)D

epth

of t

ip (m

)

02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



during unsealed pullout (NC clay, test B2JOC)

02468

10121416

-300 -200 -100 0 100 200


Dep

th o

f tip

(m)

02468

10121416

0 50 100 150 200

Pore pressure (kPa)D

epth

of t

ip (m

)

PPTI

Hydrostatic




02468

10121416

-300 -200 -100 0 100 200Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

02468

10121416

0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic





Figure 6.68 Installation and pullout pressure and internal pore pressure changes during unsealed pullout (NC clay, test B11SOC)

02468

10121416


Dep

th o

f tip

(m)

0

2

46

8

1012

14

16

0 50 100 150

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



during unsealed pullout (NC clay, test B12SOC)

∆p=P/A

02468

10121416

-300 -200 -100 0 100 200


epth

of t

ip (m

)

0

2

4

6

8

10

12

14

16

0 50 100 150

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic

∆p = P/A

Pullout Installation

0

2

4

6

8

10

12

14

16

-300 -200 -100 0 100 200


Dep

th o

f tip

(m)

B12SOCB11SOCB2JOCB8JOCB6JOC

Figure 6.70 Installation and pullout pressure of unsealed pullout tests after consolidation in NC clay

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25


Dep

th (m

)

B12SOCB11SOCB2JOCB6JOCB8JOCB8JOIB9SOIB2JOI

NT-bar = 10.5

su (LB) = 1.02 kPa/m

su (UB) = 1.30 kPa/m

Figure 6.71 Undrained strength profiles for various T-bar tests in NC clay

∆p = P/AInstallation

Real capacity

Pullout

'Pseudo' capacity Consolidation

0

2

4

6

8

10

12

14

16

-300 -250 -200 -150 -100 -50 0 50 100 150Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

B11SOC

B2JOC

B9SOI

B2JOI

Figure 6.72 Installation and unsealed pullout pressure with and without consolidation in NC clay

∆p = P/A

12.0

12.5

13.0

13.5

14.0

14.5

15.0

-250 -200 -150 -100 -50 0


Dep

th o

f tip

(m)

B11SOC

B2JOC

B9SOI

B2JOI

Figure 6.73 Axial pressure versus depth of tip during early stage of unsealed pullout in NC clay

0

2

4

6

8

10

12

14

16

-250 -200 -150 -100 -50 0 50 100 150 200Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

B10SOC*

B3SOC*

B6JOC*

B2JOC*

Figure 6.74 Comparison of the axial pressure of the unsealed pullout after consolidation in the disturbed sites in NC clay

0

2

4

6

8

10

12

14

16

-300 -200 -100 0 100 200

Axial pressure (kPa)

Dep

th o

f tip

(m)

B2JOC

B2JOC*

B11SOC

B10SOC*InstallPullout

Figure 6.75 Comparison of the unsealed pullout after consolidation in original and disturbed sites in NC clay

∆p= P/A

02468

10121416

-400 -300 -200 -100 0 100 200


Dep

th o

f tip

(m)

B11SCCB12SOC

02468

10121416

-100 -50 0 50 100 150 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic


Figure 6.76 Installation and pullout pressure and internal pore pressure during sealed pullout after consolidation in NC clay (test B11SCC)

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20

Undrained shear strength su from T-bar tests(kPa)

Dep

th (m

)

su,ave =1.26 kPa/m

NT-bar = 10.5

Figure 6.77 Undrained shear strength profiles measured by T-bar (NC clay, test B11SCC)

∆p= P/A

02468

10121416

-400 -300 -200 -100 0 100 200


Dep

th o

f tip

(m)

B12SCC

02468

10121416

-150 -50 50 150

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20

Undrained shear strength su from T-bar tests (kPa)

Dep

th o

f tip

(m)

su, ave = 1.17 kPa/mNT-bar = 10.5 kPa/m

Figure 6.79 Undrained shear strength measured by T-bar (NC clay, B12SCC)

02468

10121416


Dep

th o

f tip

(m)

B3SCC

02468

10121416

-100 -50 0 50 100 150

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic



0

2

4

6

8

10

12

14

16

-15 -10 -5 0 5 10 15 20


Dep

th (m

)

su, ave =1.08 kPa/mNT-bar = 10.5

Figure 6.81 Undrained shear strength versus depth measured by T-bar (NC clay, test B3SCC)

∆p=P/A

0

2

4

6

8

10

12

14

16

-300 -200 -100 0 100 200


epth

of t

ip(m

)

B4JCC

Figure 6.82 Variation of the axial pressure of sealed pullout after consolidation (NC clay, test B4JCC)

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20


Dep

th (m

)

su, ave=1.13 kPa/mN T-bar = 10.5

Figure 6.83 Undrained shear strength versus depth measured by T-bar (NC clay, test B4JCC)

Figure 6.84 Small solid pile with equivalent diameter (deq = 7.68 mm) and roughness to that of the model caisson


0

2

4

6

8

10

12

14

16

-800 -600 -400 -200 0 200 400 600


Em

bedm

ent (

m)

Measured

Install:Nc=9,Alpha=0.45

Pullout:Nc=9,Alpha=0.90

Figure 6.85 Axial pressure for equivalent pile test in NC clay (Test B12pile1)

7.68 mm


0

2

4

6

8

10

12

14

16

-1000 -800 -600 -400 -200 0 200 400 600


Em

bedm

ent (

m)

Measured

Install:Nc=9,Alpha=0.52

Pullout:Nc=9,Alpha=0.96

Figure 6.86 Axial pressure for equivalent pile test in NC clay (Test B12pile2)

∆p = P/A

0

2

4

6

8

10

12

14

16

-300 -200 -100 0 100Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

B12SCI

B9SOI

B12SCC

0

2

4

6

8

10

12

14

16

-200 -100 0 100 200

Pressure (kPa)

Dep

th o

f tip

(m)

Hydrostatic

Internal PPT


Figure 6.87 Installation and pullout pressure and internal pore pressure of immediate sealed pullout test in NC clay (Suction installation, test B12SCI)

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20


Dep

th o

f tip

(m)

su, ave = 1.17 kPa/mNT-bar = 10.5 kPa/m

Figure 6.88 Undrained shear strength profiles measured by T-bar (NC clay, test B12SCI)

0

2

4

6

8

10

12

14

16

-300 -200 -100 0 100


Dep

th o

f tip

(m)

Figure 6.89 Installation and pullout pressure and of immediate sealed pullout test in NC clay (Suction installation, test B3SCI)

0

2

4

6

8

10

12

14

16

-15 -10 -5 0 5 10 15 20


Dep

th (m

)

su, ave =1.08 kPa/mNT-bar = 10.5

Figure 6.90 Undrained shear strength profiles measured by T-bar (NC clay, test B3SCI)

0

2

4

6

8

10

12

14

16

-300 -250 -200 -150 -100 -50 0 50 100 150


epth

of t

ip (m

)

∆p = P/A

Figure 6.91 Installation and pullout pressure and of immediate sealed pullout test in NC clay (Jacked installation, test B4JCI)

0

2

4

6

8

10

12

14

16

-15 -10 -5 0 5 10 15 20

Undrained shear strength su(kPa)

Dep

th (m

)

su, ave=1.13 kPa/mNT-bar = 10.5

Figure 6.92 Undrained shear strength profiles measured by T-bar (NC clay, test B4JCI)

∆p=P/A

02468

10121416

-500 -300 -100 100


Dep

th o

f tip

(m)

0

2

4

6

8

10

12

14

16

-100 0 100 200

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic


Figure 6.93 Installation and pullout pressure and internal pore pressure during sealed pullout after consolidation in LOC clay for jacked caisson (test B13JCC)

∆p=P/A

02468

10121416

-500 -300 -100 100


Dep

th o

f tip

(m)

02468

10121416

-100 -50 0 50 100 150

Pore pressure (kPa)

Dep

th o

f tip

(m)

PPTI

Hydrostatic


Figure 6.94 Installation and pullout pressure and internal pore pressure during sealed pullout after consolidation in LOC clay (test B13SCC)

∆p = P/A

0

2

4

6

8

10

12

14

16

-500 -400 -300 -200 -100 0 100 200


Dep

th o

f tip

(m)

B13SCC, OCR=1.5

B13JCC, OCR=1.5

B12SCC, OCR=1

Figure 6.95 Installation and pullout pressure during sealed pullout after consolidation for jacked caissons and suction caissons in LOC and NC clay

∆p=P/A

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100


Dep

th o

f tip

(m)

Penetration resistance

Install: Nc=7.5, k=1.16kPa/m, Alpha=0.16

Figure 6.96 Installation and pullout pressure during sealed pullout after consolidation for suction caisson in sensitive clay (test B14SCC)

∆p = P/A

0

2

4

6

8

10

12

14

16

-350 -250 -150 -50 50 150


epth

of t

ip (m

)

Figure 6.97 Installation and pullout pressure during sealed pullout after consolidation for suction caisson in sensitive clay

(Monotonic loading stage of test B14susa)

0

2

4

6

8

10

12

14

16

0 50 100 150 200

Radial total stress (kPa)D

epth

of t

ip (m

)

Radial total stress

HydrostaticIn water

In soil

u0σ ri+∆ui

σri

Figure 7.1 Variations of measured external radial total stress σri with depth

during jacked installation in NC clay (test B4JCC)

0

2

4

6

8

10

12

14

16

0 50 100 150 200Radial total stress (kPa)

Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui

σri

Figure 7.2 Variations of measured external radial total stress σri with depth during jacked installation in NC clay (test B4JCI )

0

2

4

6

8

10

12

14

16

0 50 100 150 200


epth

of t

ip (m

)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui

Figure 7.3 Variations of measured external radial total stress σri with depth during jacked installation in NC clay (test B5JOI)

0

2

4

6

8

10

12

14

16

0 50 100 150 200

Radial total stress (kPa)

Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui


during jacked installation in NC clay (test B6JOI)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


epth

of t

ip (m

)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui


during jacked installation in NC clay (test B8JOI)

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

B4JCCB4JCIB5JOIB6JOIB8JOIHydrostatic

In water

In soil

u0 σ ri+∆ui

σri

Figure 7.6 Comparison of measured external radial total stress σri during jacked installation in NC clay

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui

σri


during suction installation in NC clay (test B3SCC)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui

σri

Figure 7.8 Variations of measured external radial total stress σri with depth during suction installation in NC clay (test B3SCI)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


epth

of t

ip (m

)

Radial total stress

HydrostaticIn water

In soil

u0σ ri+∆ui

σri

TPTs entered suction-affected area at 12.15 m

Figure 7.9 Variations of measured external radial total stress σri with depth during suction installation in NC clay (test B11SOC)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui

Figure 7.10 Variations of measured external radial total stress σri with depth during suction installation in NC clay (test B12SOC)

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui

Figure 7.11 Variations of measured external radial total stress σri with depth during suction installation in NC clay (test B12SCI)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)

Radial total stress

Hydrostatic

In water

In soil

u0 σ ri+∆ui

Figure 7.12 Variations of measured external radial total stress σri with depth during suction installation in NC clay (test B12SCC)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


epth

of t

ip (m

)

Radial total stress

Hydrostatic

In water

In soil

u0 σ ri+∆ui

Figure 7.13 Variations of measured external radial total stress σri with depth during suction installation in NC clay (test B12cyc)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)

Radial total stress

HydrostaticIn water

In soil

u0 σ ri+∆ui


during suction installation in NC clay (test B12sus)

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

B3SCIB3SCCB11SOCB12SOCB12SCIB12SCCB12susB12cycHydrostatic

In water

In soil

u0 σ ri+∆ui

σri

Figure 7.15 Comparison of measured external radial total stress σri during suction installation in NC clay

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)


Jacked installation

Hydrostatic

In water

In soil

u0 σ ri+∆ui

σri

Figure 7.16 Comparison of measured external radial total stress σri during installation by jacking and by suction

0

1

2

3

4

5

6

7

8

0 10 20 30 40 50 60 70 80 90 100 110 120σri − u0 (kPa)

Dep

th o

f TPT

(m)

SuctionJackedCEM (Lower bound )CEM (Upper bound)SPMNGIMTD

St=2 - 2.8, G/su=100 - 150, K0=0.65, YSR=1.0, γ =6.85 kN/m3, k=1.18 kPa/m δr = 17.6o

Figure 7.17 External radial total stress relative to hydrostatic pressure σri – u0 during caisson installation in NC clay: predicted and measured (Note: upper and lower bound NGI predictions are the same)

0

1

2

3

4

5

6

7

8

0 10 20 30 40 50 60 70 80 90 100∆ui (kPa)

Dep

th o

f TPT

(m)

SuctionJackedCEM (Lower bound )CEM (Upper bound)SPMNGI (Lower bound)NGI (Upper bound)MTD

St=2 - 2.8, G/su=100 - 150, K0=0.65, YSR=1.0, γ =6.85 kN/m3, k=1.18 kPa/m δr = 17.6o

Figure 7.18 Derived external excess pore pressure ∆ui during caisson installation in NC clay: predicted and derived (from measurements)

Loa

d ce

ll (N

)

119.0

119.1

119.2

119.3

119.4

119.5

119.6

119.7

119.8

119.9

120.0

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)D

epth

of t

ip (m

m)

0

5

10

15

20

25

depth of tip

load cell

0.12 mmt50=1326 st90=3230 s

Figure 7.19 Caisson settlement and variation of axial force during consolidation in NC clay (test B11SOC) (units in model scale)

Loa

d ce

ll (N

)

120.0

120.1

120.2

120.3

120.4

120.5

120.6

120.7

120.8

120.9

121.0

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Dep

th o

f tip

(mm

)

0

10

20

30

40

50

60

depth of tip

load cell

t50=1079 st90=2818 s

0.13 mm

Figure 7.20 Caisson settlement and variation of axial force during consolidation in NC clay (test B12SOC) (units in model scale)

Loa

d ce

ll (N

)

119.80

119.85

119.90

119.95

120.00

120.05

120.10

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

Dep

th o

f tip

(mm

)

0

10

20

30

40

50

60

depth of tip

load cell

0.07 mm

t50=1163 s

t90=2596 s

Figure 7.21 Caisson settlement and variation of axial force during consolidation (NC clay, test B12SCC) (units in model scale)

0

10

20

30

40

50

60

70

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

σ r -

u 0 (k

Pa)

TPT1

TPT2

Average

t50=146 st90=1733 s

15.2 kPa

Figure 7.22 Variation of external σr – u0 during consolidation in NC clay (test B11SOC)

25

30

35

40

45

50

55

60

65

70

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

σ r −

u0 (

kPa)

TPT1

TPT2

Average

t50=521 s t90=2187 s

13.75 kPa

Figure 7.23 Variation of external σr – u0 during consolidation in NC clay

(test B12SOC)

25

30

35

40

45

50

55

60

65

70

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

σ r -

u 0 (k

Pa)

TPT1

TPT2

Average

t50 = 762 s t90 = 2068 s

21.78 kPa

Figure 7.24 Variation of external σr – u0 during consolidation in NC clay (test B12SCC)

σrc

In soil Installation

Consolidation

Pullout

In water

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

TPT1TPT2AverageHydrostatic

Figure 7.25 External radial total stress changes during installation, consolidation and pullout of caissons (NC clay, test B11SOC)

σrcConsolidation

In water

0

2

4

6

8

10

12

14

16


dept

h of

tip(

m)


Figure 7.26 External radial total stress changes during installation, consolidation and pullout of caissons (NC clay, test B12SOC)

Pullout

In soil

In water

σrc

Installation

Consolidation

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

TPT1TPT2Average (B12SCC)Hydrostatic

Figure 7.27 External radial total stress changes during installation, consolidation and pullout of caissons (NC clay, test B12SCC)

Pullout

Consolidation

InstallationIn soil

σrc

In water

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

B11SOCB12SOC

B12SCCHydrostatic

Figure 7.28 Average external radial total stress changes during installation, consolidation and pullout of caissons (Box 11 and 12, NC clay)

13.0

13.5

14.0

14.5

0 10 20 30σr − u0 (kPa)

Dep

th o

f tip

(m)

13.0

13.5

14.0

14.5

-300 -200 -100 0Uplift pressure, ∆p (kPa)

Dep

th o

f tip

(m)Fail here

z=14.25 m

Fail herez=14.27 mzTPT=7.07 mσ rf =20.72 kPaAfter consolidationz=14.32 mzTPT=7.07 mσ rc =28.02 kPa

Figure 7.29 External σr – u0 and uplift pressure at failure during unsealed pullout of caisson in NC clay (test B11SOC)

13.0

13.5

14.0

14.5

15.0

0 20 40σr − u0 (kPa)

Dep

th o

f tip

(m) 13.0

13.5

14.0

14.5

15.0

-300 -200 -100 0Uplift pressure, ∆p (kPa)

Dep

th o

f tip

(m)

Fail herez=14.45 mzTPT=7.25 mσ rf =23.77 kPa

After consolidationz=14.50 m, zTPT=7.30 mσ rc =29.72 kPa

Fail herez=14.44 m

Figure 7.30 External σr – u0 and uplift pressure at failure during unsealed pullout of caisson in NC clay (test B12SOC)

11.0

12.0

13.0

14.0

15.0

0 10 20 30σr − u0 (kPa)

Dep

th o

f tip

(m) 11.0

12.0

13.0

14.0

15.0

-400 -300 -200 -100 0Uplift pressure, ∆p (kPa)

Dep

th o

f tip

(m)Fail here

z=14.37 mzTPT=14.37 mσ rf =22.05 kPa

Fail herez=13.51 m

After consolidation,z=14.40 m, σ rc =25.80 kPa

Figure 7.31 External σr – u0 and uplift pressure at failure during sealed pullout of caisson in NC clay (test B12SCC)

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25 30

Shaft friction (kPa)

Dep

th (m

)

Shear strengthMeasured (lower bound)Measured (upper bound)MTDCEMAPINGI

Figure 7.32 Profiles for external shaft friction during pullout of the caisson in NC clay

13.2

13.3

13.4

13.5

13.6

13.7

13.8

13.9

14.0

14.1

14.2

14.3


Dep

th o

f tip

(m)

Installation

Pullout

Pullout

2.9 kPa

Figure 7.33 External radial total stress changes during immediate pullout of

caisson (NC clay, test B2JOI)

Pullout

Consolidation

InstallationIn soil

σrc

In water

0123456789

10111213141516


Dep

th o

f tip

(m)

B11SOC: after conso.

B2JOI: immediate

Hydrostatic

Figure 7.34 Comparison of external radial total stress changes during immediate

pullout and pullout after consolidation in NC clay

In water

In soil

0123456789

101112131415

0 50 100 150 200 250


epth

of t

ip (m

)

Average

TPT1

TPT2

Hydrostatic

z=12.11 m

z=13.72 m

Figure 7.35 Variation of measured external radial total stress σri during suction

installation in LOC clay (test B13SCC)

0123456789

101112131415

-500 -400 -300 -200 -100 0 100 200 300Pressure (kPa)

Dep

th o

f tip

(m)

Radial total stress

Syringe pump pressure

Insignificant change in gradient of radial stress when suction starts

z=12.11 m, TPTs enter the suction-affected area in soil, a small decrease exists

z=13.72 m, σri decreases and pump pressure surges due to moving speed reduces to 0.18 mm/s

z = 6.97-7.31 m, jacking ends and suction starts

z=4.8 m, TPTs leave water and enter soil

Figure 7.36 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in LOC clay (test B13SCC)

-500

-400

-300

-200

-100

0

100

200

300

150 160 170 180 190 200 210 220 230 240 250

Time, in model scale (s)

Oup

tput

in in

stru

men

tsEmbedment (z) of caisson in model scale (mm)

Embedment (z) of caisson in prototype scale (m)

Syringe pump pressure (kPa)


z=13.72 m, speed reduces, σri decreases and pump pressure surges

z=7.31-13.72 m, v=1.80 mm/s, V=11.8: undrained, by suction

z=13.72-13.92 m, v=0.18 mm/s, V=1.2: partly drained

z=6.44-7.31 m, v=0.53mm/s, V=3.5: partly drained

z = 6.97-7.31 m, jacking ends and suction starts, with 1.1 s of time delay

Note: velocity (v) is shown in model scale

Figure 7.37 Variations of syringe pump pressure, embedment of caisson in model and prototype scales, external radial total stress versus time (in model scale)

during suction installation in LOC clay (test B13SCC)

0

1

2

3

4

5

6

7

8

9

10

0 20 40 60 80 100σri − u0 (kPa)

Dep

th o

f TPT

(m)

In transition region, σri − u0 decreases 1.9 kPa. Considering an increase of 2.7 kPa due to depth increase, it reduces 4.6 kPa indeed, due to consolidation during time delay and slow movement

zTPT=8.92 m, 70.13 kPa, reduction due to speed drops from 1.8 to 0.18 mm/s (in model scale)

zTPT=6.97 m,TPTs leavejacking-affected area, σri − u0 = 57.88 kPa

zTPT=7.31 m, TPTs enters suction-affected area,σri − u0 = 55.99 kPa

suction-affected area: >1.51 m

zTPT=2.17-2.51 m, small change due to consolidation in time delay when jacking ends and suction starts

Figure 7.38 Variations of external σri – u0 versus depth of TPT during suction installation in LOC clay (test B13SCC)

0

1

2

3

4

5

6

7

8

9

10

0 20 40 60 80 100

σri − u0 (kPa)D

epth

of T

PT (

m)

Measured

CEM (lower bound)

CEM (upper bound)

NGI (if no transition zone)

zTPT = 6.97 m, TPTs leave jacking-affected area

9.3 kPa: due to reduced speed (partly drained)

zTPT = 7.31 m, σri−u0

decreases in the suction affected area, due to time delay when suction started, and reduced velocity of caisson.

zTPT=8.92 m, decrease due to consolidation

Linearly reduction in 1 D (3.6 m) of transition zone in suction area，σri−u0 = 58.86 kPa at 8.92 m, according to Andersen & Jostad (2002)

Figure 7.39 Measured external σri – u0 and predictions by NGI method and CEM versus depth of TPT during suction installation in LOC clay (test B13SCC)

0

1

2

3

4

5

6

7

8

9

10

0 50 100 150 200

σri − u0 (kPa)

Dep

th o

f TPT

(m

)

MeasuredMTD(Jardine & Chow 1996)SPM(Whittle & Baligh 1988)CEM (lower bound)CEM (upper bound)NGI (if no transition zone)

St=2 - 2.5, G/su=100 - 150, K0=0.70, YSR=1.5, γ =7.15 kN/m3, k=1.51 kPa/m, δr = 18.1o

Figure 7.40 Comparison of measured external σri – u0 and theoretical predictions during suction installation in LOC clay (test B13SCC)

0123456789

101112131415

0 50 100 150 200 250


epth

of t

ip (m

)

AverageTPT1TPT2Hydrostatic

z=12.77 m

z=13.59 m

Figure 7.41 Variation of the measured external radial total stress σri during suction installation in LOC clay (test B13sus)

0123456789

101112131415

-400 -300 -200 -100 0 100 200 300

Pressure (kPa)

Dep

th o

f tip

(m)

Syringe pump

Radial total stress

z=13.59 m, σri reduces due to reduced speed

z=12.77 m, TPTs enters the suction-affected area in soil, a small decrease exists

z = 7.60 - 7.97 m, slight change in σri when suction starts

z = 7.60 - 7.97 m, jacking ends and suction starts

Figure 7.42 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in LOC clay (test B13sus)

-400

-300

-200

-100

0

100

200

300

0 20 40 60 80 100 120 140


Out

put o

f ins

trum

ents

Embedment of caisson (z) in model scale (mm)Embedment of caisson (z) in prototype scale (m)Radial total stress (kPa)Syringe pump pressure (kPa)


z=7.97 - 13.59 m, v=1.90 mm/s, model scale, V=12.5: undrained, by suction

z=13.59 m, σri reduces and pump pressure surges, due to v slows to 0.20 mm/s at 13.52 m

z = 7.60 - 7.97 m, jacking ends and suction starts, with 3.7 s of time delay


Note: v in model scale


during suction installation in LOC clay (test B13sus)

0

1

2

3

4

5

6

7

8

9

10

0 20 40 60 80 100

σri − u0 (kPa)

Dep

th o

f TPT

(m)

In transition region, σri − u0 decreased 0.09 kPa, considering an increase of 2.9 kPa due to depth, it decreased 3 kPa indeed, due to consolidation


zTPT=8.79 m, 74.09 kPa, reduction caused by the change of speed from 1.9 mm/s to 0.20 mm/s (model scale)

zTPT=7.60 m, TPTs leave jacking-affected area,σri−u0 = 65.44 kPa

zTPT=7.97 m, TPTs enters suction-affected area,σri−u0 = 65.35 kPa

zTPT=2.80 - 3.17m, small change due to consolidation

Figure 7.44 Variations of external σri – u0 versus depth of TPT during suction installation in LOC clay (test B13sus)

0

1

2

3

4

5

6

7

8

9

10

0 20 40 60 80 100

σri−u0 (kPa)D

epth

of T

PT (m

)

Measured

CEM (lower bound)

CEM (upper bound)


z TPT = 7.97 m, σri−u0

decreases, due to consolidation in time delay and reduced rate of penetration.

z = 7.60 m, TPTs leave jacking-affected area

5.8 kPa: due to consolidation at lower speed

zTPT=8.79 m, decrease in radial stress

Linearly reduction in 1 D (3.6 m) of transition zone in suction area，σri-u0 = 65.7 kPa at 8.79 m, according to Andersen & Jostad (2002)

Figure 7.45 Comparison of measured external σri – u0 and predictions of NGI

method and CEM during suction installation in LOC clay (test B13sus)

0123456789

10

0 50 100 150 200

σri−u0 (kPa)

Dep

th o

f TPT

(m)


St = 2 - 2.5, G/su=100 - 150, K0 = 0.70, YSR=1.5, γ =7.21 kN/m3, k=1.71 kPa/m,δr = 18.1o

Figure 7.46 Comparison of measured external σri – u0 and theoretical predictions during suction installation in LOC clay (test B13sus)

In soil

In water

0123456789

101112131415

0 50 100 150 200 250


epth

of t

ip (m

)

AverageTPT1TPT2Hydrostatic z=12.76 m, σri

decrease, speed slows to 0.21 mm/s (model scale)

z=12.65 m, TPTs enters the suction-affected area in soil

z=13.47 m, further decrease, speed reduces to 0.21 mm/s

Figure 7.47 Variation of the measured external radial total stress σri during suction installation in LOC clay (test B13cyc)

0

2

4

6

8

10

12

14

16

-650 -550 -450 -350 -250 -150 -50 50 150 250

Pressure (kPa)

z (m

)

Radial total stress



Slight change when suction starts

z = 12.49 m, TPTs leave the jacking-affected area in soil

z = 13.47 m, decrease in radial stress due to moving speed slows to 0.02 mm/s at the end of installation

z = 12.76 m, decrease, speed slows to 0.21 mm/s

Figure 7.48 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in LOC clay (test B13cyc)

0

1

2

3

4

5

6

7

8

9

10

0 20 40 60 80 100

σri − u0 (kPa)D

epth

of T

PT (m

)

In transition region, σri - u0

decreased 1.1 kPa, considering an increase of 0.8 kPa due to depth, it decreased 1.9 kPa indeed.

zTPT = 7.96 m, reduction caused by reduction of speed from 1.89 mm/s to 0.21 mm/s (model scale)

zTPT=7.69 m, TPTs leave jacking-affected area, σri−u0 = 67.45 kPa, zTPT=7.80 m, speed slows to 0.21 mm/s (model scale)

zTPT = 7.85 m, enters suction-affected area, σri−u0 = 66.38 kPa

distinguishable suction-affected area: ~0.14 m

zTPT=2.89-3.05 m, change from jacking to suction installation

zTPT = 8.67 m, speed reduces to 0.02 mm/s

Figure 7.49 Variations of external σri – u0 versus depth of TPT during suction installation in LOC clay (test B13cyc)

0

1

2

3

4

5

6

7

8

9

10

0 20 40 60 80 100

σri−u0 (kPa)

Dep

th o

f TPT

(m)

Measured

CEM (lower bound)

CEM (upper bound)


zTPT = 7.69 m, TPTS leave jacking-affected area, v slows to 0.21 mm/s (model scale)


decreases, due to consolidation in time delay and penetration at reduced rate

8.3 kPa: due to changing to lower speed

zTPT = 7.96 m, v reduces to 0.21 mm/s (model scale)

Figure 7.50 Comparison of measured external σri – u0 and predictions from NGI method and CEM during suction installation in LOC clay (test B13cyc)

0123456789

10

0 50 100 150 200

σri−u0 (kPa)

Dep

th o

f TPT

(m)

Measured

MTD(Jardine & Chow 1996)

SPM(Whittle & Baligh 1988)

CEM (upper bound)


St=2 - 2.5, G/su=100 - 150, K0 = 0.70, YSR = 1.5, γ =7.21 kN/m3, k=1.45 kPa/m, δr = 18.1o

Figure 7.51 Comparison of measured external σri – u0 and theoretical predictions during suction installation in LOC clay (test B13cyc)

0

2

4

6

8

10

12

14

16


Dep

th o

f tip

(m)

B13JCC (by jacking)

B13SCC (by suction)

Hydrostatic

In water

In soil

u0 σ ri+∆ui

σri

Figure 7.52 Measured external radial total stress σri during jacked installation

(test B13JCC) and suction installation (test B13SCC) in LOC clay

0

1

2

3

4

5

6

7

8

9

10

0 50 100 150 200 250

σri – u0 (kPa)D

epth

of T

PT (m

)

Measured

CEM (Lower bound )

CEM (Upper bound)

SPM

NGI

MTD

St=2 - 2.5, G/su=100 - 150, K0 = 0.70, YSR = 1.5, γ =7.15 kN/m3, k = 1.42 kPa/m δr = 18.1o

Figure 7.53 Comparison of measured external σri – u0 and theoretical predictions for jacked installation test B13JCC in LOC clay

115.75

115.80

115.85

115.90

115.95

116.00

116.05

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

Dep

th o

f tip

(mm

)

0

50

100

150

Loa

d (N

)

Depth of tip

Load cell

0.20 mm(from 115.79 mm to 115.99 mm)

Figure 7.54 Depth of tip and variation of axial force during consolidation in LOC clay (test B13SCC) (units in model scale)

t50= 1934 sect90=3295 sec

116.40

116.45

116.50

116.55

116.60

116.65

116.70

0 1000 2000 3000 4000

Time (sec)D

epth

of t

ip (m

m)

0

50

100

150

Loa

d (N

)

Depth of tip

Load cell

t90=3295 sect50= 1934 sec

116.40

116.45

116.50

116.55

116.60

116.65

116.70

0 1000 2000 3000 4000

Time (sec)D

epth

of t

ip (m

m)

0

50

100

150

Loa

d (N

)

Depth of tip

Load cell

116.40

116.45

116.50

116.55

116.60

116.65

116.70

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)D

epth

of t

ip (m

m)

0

50

100

150

Loa

d (N

)

Depth of tip

Load cell

0.18 mm(from 116.46 mm to 116.64 mm)

Figure 7.55 Depth of tip and variation of axial force during consolidation in LOC clay (test B13JCC) (units in model scale)

t50' = 637.7 sect50 = 517.0 sec

(2.9 mths. Prot.) t90' = 2222 sect90 = 2101 sec

(11.7 mths. Prot.)

114.34

114.36

114.38

114.40

114.42

114.44

114.46

114.48

114.50

114.52

0 500 1000 1500 2000 2500 3000 3500 4000Time (sec)

Dep

th o

f tip

(mm

)

0

50

100

150

Loa

d (N

)

Depth of tip

Load cellt50 = 517 s t90 = 2101 s

114.34

114.36

114.38

114.40

114.42

114.44

114.46

114.48

114.50

114.52

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

Dep

th o

f tip

(mm

)

0

50

100

150

Loa

d (N

)Depth of tip

Load cell0.12 mm(from 114.37 mm to 114.49 mm)

Figure 7.56 Depth of tip and variation of axial force during consolidation in LOC clay (test B13sus) (units in model scale)

0.20 mm

t50 = 736 s

t90 = 2562 s

113.00

113.05

113.10

113.15

113.20

113.25

113.30

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)D

epth

of t

ip (m

m)

0

50

100

150

Loa

d (N

)

Depth of tip

Load cell

Figure 7.57 Depth of tip and variation of axial force during consolidation in LOC clay (test B13cyc) (units in model scale)

t50 = 291 s

t90 = 2755 s 12.1 kPa

45

50

55

60

65

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

σr −

u0 (

kPa)

AverageTPT1TPT2

Figure 7.58 Variation of external σr – u0 during consolidation in LOC clay

(test B13SCC) (units in model scale)

13.2 kPa

t50 = 986 s

t90 = 2053 s

40

45

50

55

60

65

70

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

σ r −

u0 (

kPa)

Average

TPT1

TPT2

Figure 7.59 Variation of external σr – u0 during consolidation in LOC clay (test B13JCC) (units in model scale)

t50 = 28 s

t90 = 1980 s15.2 kPa

45

50

55

60

65

70

75

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

σr −

u0 (

kPa)

TPT1TPT2TPTave

Figure 7.60 Variation of external σr – u0 during consolidation in LOC clay (test B13sus) (units in model scale)

t90 = 1891 s 9.5 kPa

t50 = 208 s

45

50

55

60

65

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

σr −

u0 (

kPa)

TPT1TPT2Average

Figure 7.61 Variation of external σr – u0 during consolidation in LOC clay (test B13cyc) (units in model scale)

In water

In soil

PulloutInstallation

Consolidationσrc

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250

Radoal total stress (kPa)

Dep

th o

f tip

(m)

TPT1

TPT2

Average (B13SCC)

Hydrostatic

Figure 7.62 External radial stress changes during installation, consolidation and pullout in LOC clay (OCR = 1.5, test B13SCC)

In soil

σrc Consolidation

InstallationPullout

In water

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


epth

of t

ip (m

)

TPT1

TPT2

Average

Hydrostatic

Figure 7.63 External radial total stress changes during installation, consolidation and pullout in LOC clay (OCR = 1.5, test B13JCC)

PulloutInstallation

Consolidationσrc

In soil

In water

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


Dep

th o

f tip

(m)

B13JCC: by jacking

B13SCC: by suction

Hydrostatic

Figure 7.64 Comparison of external radial total stress changes during installation, consolidation and pullout between jacked caisson and suction caisson in LOC clay

(OCR=1.5)

13.0

13.1

13.2

13.3

13.4

13.5

13.6

13.7

13.8

13.9

14.0

0 10 20 30 40 50 60

σr−u0 (kPa)D

epth

of t

ip (m

)

Fail herez=13.50 mzTPT=8.70 mσ rf =43.68 kPa

After consolidationz=13.92 mzTPT=9.12 mσ rc = 49.51 kPa

Figure 7.65 External σr – u0 at failure during caisson pullout in LOC clay (test B13SCC)

12.6

12.8

13.0

13.2

13.4

13.6

13.8

14.0

14.2

0 10 20 30 40 50 60σr − u0 (kPa)

Dep

th o

f tip

(m) Fail here

z=13.75 mzTPT=8.95 mσ rf =37.94 kPa

After consolidationz=14.00 mzTPT=9.60 mσ rc = 48.02 kPa

Figure 7.66 External σr – u0 at failure during caisson pullout in LOC clay (test B13JCC)

0123456789

10111213141516


Dep

th o

f tip

(m)

Average

Hydrostatic

TPT1

TPT2

z = 4.8 m, TPT left water and entered soil

in soil

z = 7.21-7.45

z = 12.25 m

z = 14.20 m

z = 15.11 m, reduced

Figure 7.67 Variation of measured external radial total stress σri during suction installation in sensitive clay (test B14cyc)

0123456789

10111213141516

-400 -300 -200 -100 0 100 200 300

Pressure (kPa)

Dep

th o

f tip

(m)

Radial total stress


Slight change in gradient of radial stress when suction starts


z=14.20 m, σri decreases and pump pressure surges due to moving speed reduces


z=15.11 m, speed slows further

Figure 7.68 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in sensitive clay (test B14cyc)

-500

-400

-300

-200

-100

0

100

200

300

0 50 100 150 200 250 300 350 400 450


Oup

tput

in in

stru

men

tsEmbedment of caisson (z) in model scale (mm)

Embedment of caisson (z) in prototype scale (m)

Syringe-pump (kPa)


z=14.20 m, v slows and σri

reduces

z = 14.20 -15.24 m, v = 0.03 mm/s (model), V = 0.24: partly drained (creep)

z = 7.45 - 14.20 m, v = 2.18 mm/s, model scale, V = 17: undrained, by suction

Time delay = 1 s, speed reduces to 0.81 mm/s, V=6.5, partly drained


during suction installation in sensitive clay (test B14cyc)

0123456789

101112

0 20 40 60 80 100

σri−u0 (kPa)

Dep

th o

f TPT

(m)

zTPT = 9.40 m, σri−u0 = 62.64 kPa, then reduces due to speed decreases to 0.03 mm/s (model) scale)

zTPT = 7.21 m, TPTs leave jacking-affected area,σri−u0 = 46.09 kPa

zTPT = 7.45 m, TPTs enterssuction-affected area,σri−u0 = 46.68 kPa suction-affected

area: >1.96 mzTPT = 10.31 m, reduces due to speed reduces to 0.01mm/s (model scale)

zTPT = 2.41 - 2.65 m, small change when jacking ended and suction started

In transition region, σri − u0 increases 0.59 kPa. Considering an increase of 1.93 kPa due to depth, it decreases 1.34 kPa indeed, due to consolidation in time delay

Figure 7.70 Variations of external σri – u0 versus depth of TPT during suction installation in sensitive clay (test B14cyc)

0123456789

101112

0 20 40 60 80 100

σri−u0 (kPa)D

epth

of T

PT (m

)

Measured

CEM (lower bound)

CEM (upper bound)

NGI (lower bound)


3.2 kPa: due to reduced speed

zTPT = 7.45 m, TPTs enters suction-affected area, σri−u0 decreases, due to consolidation in time delay and penetration with reduced speed


zTPT = 9.40 m, speed decreases

z = 10.31 m,lower speed

Linearly reduction in 1 D (3.6 m) of transition zone in suction area，σri−u0 = 52.57 kPa at 9.4 m, according to Andersen & Jostad (2002)

Figure 7.71 Measured external σri – u0 and predictions by NGI method and CEM versus depth of TPT during suction installation in sensitive clay (test B14cyc)

0

2

4

6

8

10

12

0 50 100 150 200

σri−u0 (kPa)

Dep

th o

f TPT

(m)


St = 4 - 5, G/su = 50 - 100, K0 = 0.55, YSR = 1, γ = 7.3 kN/m3, k = 1.24kPa/m,δr = 11.7o

Figure 7.72 Comparison of measured external σri – u0 and theoretical predictions during suction installation in sensitive clay (test B14cyc)

In soil

In water

0123456789

10111213141516

0 50 100 150 200


epth

of t

ip (m

)

TPT1

TPT2

Hydrostatic

Average

z=11.54 m

z=14.15 m

Figure 7.73 Variation of measured external radial total stress σri during suction installation in sensitive clay (test B14susa)

0123456789

10111213141516

-600 -400 -200 0 200 400Pressure (kPa)

Dep

th o

f tip

(m)

Radial total stress


z=14.15 m, σri reduces due to moving speed reduces to 0.02 mm/s



z = 11.54 m, TPTs entering the suction-affected area in soil, a small decrease exists

Figure 7.74 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in sensitive clay (test B14susa)

-600

-500

-400

-300

-200

-100

0

100

200

300

0 100 200 300 400 500


Oup

tput

in in

stru

men

ts

Embedment (z) of caisson in model scale (mm)


Syringe pump pressure (kPa)


z=14.15 m, σri reduces when v reduces and pump surges

z=6.74-14.15 m, v = 1.34 mm/s, V=11: undrained, by suction

z = 14.15-14.77 m, v = 0.02 mm/s, V=0.2: undrained

Time delay = 5.1 s, v = 0.4 mm/s, V=3.2: partly drained


during suction installation in sensitive clay (test B14susa)

0

2

4

6

8

10

0 20 40 60 80 100σri−u0 (kPa)

Dep

th o

f TPT

(m)

In transition region, σri − u0 reduces 1.35 kPa. Considering an increase of 3.21 kPa due to depth, it reduces 4.56 kPa indeed, due to time delay and reduced speed

zTPT = 9.35 m, σri − u0

=60.44 kPa, then reduces due to speed change to 0.02 mm/s (model scale)

zTPT = 6.34 m, TPTs leaves jacking-affected area, σri−u0

= 44.66 kPa

zTPT = 6.74 m, TPTs enterssuction-affected area,σri−u0 = 43.31 kPa


zTPT = 1.54 - 1.94 m, small change when jacking ends and suction starts

Figure 7.76 Variations of external σri−u0 versus depth of TPT during suction installation in sensitive clay (test B14susa)

0123456789

1011

0 20 40 60 80 100

σri−u0 (kPa)D

epth

of T

PT (m

)

Measured

CEM (lower bound)

CEM (upper bound)



6.7 kPa: due to reduced speed


decreases due to consolidation in time delay and penetration at a reduced speed

zTPT=9.35 m, stress decreases due to speed reduces to 0.02 mm/s (model scale)

Linearly reduction in 1 D (3.6 m) of transition zone in suction area，σri−u0 = 50.2 kPa at 9.35 m, according to A & J (2002)

Figure 7.77 Measured external σri−u0 and predictions of NGI method and CEM versus depth of TPT during suction installation in sensitive clay (test B14susa)

0123456789

1011

0 50 100 150 200

σri −u0 (kPa)

Dep

th o

f TPT

(m)


St=4 - 5, G/su=50 - 100, K0=0.55, YSR=1, γ =7.3 kN/m3, k=1.35 kPa/m,δr = 11.7o

Penetration rate reduced

Figure 7.78 Comparison of measured external σri−u0 and theoretical predictions during suction installation in sensitive clay (test B14susa)

In water

In soil

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


epth

of t

ip (m

)


z=13.92 m, TPTs enter the suction-affected area in soil

z=14.57 m, σri decreases, speed reduces

Figure 7.79 Variation of measured external radial total stress σri during suction installation in sensitive clay (test B14SCC)

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100 200 300

Pressure (kPa)

Dep

th o

f tip

(m)

Radial total stress



z=13.92 m, TPTs enters the suction-affected area in soil, σri reduces slightly, due to time delay and reduced speed

Reduction in σri due to moving speed decreases after 14.57 m


Figure 7.80 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in sensitive clay (test B14SCC)

-400

-300

-200

-100

0

100

200

300

0 50 100 150 200 250 300


Oup

tput

in in

stru

men

tsEmbedment (z) of caisson in model scale (mm)


Syringe-pump (kPa)


z=14.57 m, radial stress decreases due to moving speed reduces

v=0.08 mm/s, V=0.7: partly drained (creep)

after z = 9.12 m, v=1.83 mm/s, model scale, V=14.5: undrained, by suction

Time delay = 1.8 s, v = 0.96 mm/s, V = 7.6, partly drained


during suction installation in sensitive clay (test B14SCC)

0

2

4

6

8

10

12

0 20 40 60 80 100σri−u0 (kPa)

Dep

th o

f TPT

(m) In transition region,

σri − u0 increases 1.4 kPa, considering an increase of 0.9 kPa due to depth, it increases 0.5 kPa indeed.

zTPT=9.77 m, σri − u0 = 61.44 kPa, then reduces since speed reduces to 0.08 mm/s (model scale)

zTPT=8.90 m, TPTs leave jacking-affected area,σri−u0 = 54.94 kPa

zTPT=9.12 m, TPTsenters suction-affected area,σri−u0 = 56.31 kPa


Figure 7.82 Variations of external σri−u0 versus depth of TPT during suction installation in sensitive clay (test B14SCC)

0

2

4

6

8

10

12

0 50 100 150 200

σri −u0 (kPa)D

epth

of T

PT (

m)


St=4 - 5, G/su=50 - 100, K0 = 0.55, YSR = 1, γ =7.30 kN/m3, k = 0.99 kPa/m, δr = 11.7o

Figure 7.83 Comparison of measured external σri−u0 and theoretical predictions during suction installation in sensitive clay (test B14SCC)

In water

In soil

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


Dep

th o

f tip

(m)

TPT1TPT2HydrostaticAverage


z=14.45 m, σri reduces due to moving speed reduces to 0.04 mm/s

z=14.20 m, TPTs enters the suction-affected area in soil

Figure 7.84 Variation of measured external radial total stress σri during suction installation in sensitive clay (test B14sus)

0

2

4

6

8

10

12

14

16

-200 -100 0 100 200 300Pressure (kPa)

Dep

th o

f tip

(m)

Radial total stress

Syringe pump pressureSlight change in gradient of radial stress when suction starts


z=14.45 m, σri reduces due to moving speed drops to 0.04 mm/s


Figure 7.85 Variation of syringe pump pressure and measured external radial total stress σri during suction installation in sensitive clay (test B14sus)

-200

-150

-100

-50

0

50

100

150

200

250

0 20 40 60 80 100 120 140 160


Oup

tput

in in

stru

men

ts

Embedment (z) of caisson in model scale (mm)


Syringe-pump (kPa)


z = 14.45 m, σri reduces due to installation speed decreases to 0.04 mm/s

z = 14.45-14.59 m, v=0.04 mm/s, V=0.3: partly drained

z = 9.40 - 14.45 m, v = 1.79 mm/s, model scale, V=14: undrained, by suction

Time delay = 3 s, speed reduced to 0.92 mm/s, V=7.3, partly drained


during suction installation in sensitive clay (test B14sus)

0123456789

10

0 50 100 150 200

σri−u0 (kPa)D

epth

of T

PT (m

)


St = 4 - 5, G/su = 50 - 100, K0 = 0.55, YSR = 1, γ = 7.3 kN/m3, k = 1.11 kPa/m, δr = 11.7o

Reduced speed

Figure 7.87 Comparison of measured σri−u0 and theoretical predictions during suction installation in sensitive clay (test B14sus)

0.16 mm

127.02

127.04

127.06

127.08

127.10

127.12

127.14

127.16

127.18

127.20

127.22

0 500 1000 1500 2000 2500 3000 3500

Time (s)

Dep

th o

f tip

(kPa

)

0

20

40

60

80

100

120

140

Loa

d (N

)

Depth of tip

Load

Figure 7.88 Caisson settlement and variation of axial force during consolidation in sensitive clay (test B14cyc) (units in model scale)

0.16 mm

123.08

123.10

123.12

123.14

123.16

123.18

123.20

123.22

123.24

123.26

123.28

0 500 1000 1500 2000 2500 3000 3500

Time (s)D

epth

of t

ip (m

m)

0

20

40

60

80

100

120

140

Loa

d (N

)Depth of tip

Load cell readings

Figure 7.89 Caisson settlement and variation of axial force during consolidation in

sensitive clay (test B14susa) (units in model scale)

0.17 mm

127.30

127.35

127.40

127.45

127.50

127.55

127.60

127.65

127.70

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

Dep

th o

f tip

(mm

)

0

20

40

60

80

100

120

140

Loa

d (N

)

Depth of tip

Load cell readings

Figure 7.90 Caisson settlement and variation of axial force during consolidation in sensitive clay (test B14SCC) (units in model scale)

0.12 mm

121.54

121.56

121.58

121.60

121.62

121.64

121.66

121.68

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

Dep

th o

f tip

(mm

)

0

20

40

60

80

100

120

140

Loa

d (N

)

Depth of tip

Load cell

Figure 7.91 Caisson settlement and variation of axial force during consolidation in sensitive clay (test B14sus) (units in model scale)

t50 = 1242 st90 = 2949 s

4.5 kPa

45

50

55

60

65

70

0 500 1000 1500 2000 2500 3000 3500

Time (s)

σ r −

u0 (

kPa)

TPT1TPT2Average

Figure 7.92 Variation of external σr – u0 during consolidation in sensitive clay (test B14cyc) (units in model scale)

t50 = 1599 s

2.8 kPa

t90 = 2964 s

45

47

49

51

53

55

57

59

0 500 1000 1500 2000 2500 3000 3500

Time (s)

σ r −

u0 (

kPa)

Average

TPT1

TPT2

Figure 7.93 Variation of external σr – u0 during consolidation in sensitive clay

(test B14susa) (units in model scale)

t50 = 669 s

t90 = 2559 s

6.6 kPa

40

45

50

55

60

65

70

0 500 1000 1500 2000 2500 3000 3500 4000Time (s)

σr −

u0 (

kPa)

AverageTPT1TPT2

Figure 7.94 Variation of external σr – u0 during consolidation in sensitive clay (test B14SCC) (units in model scale)

t50 = 148 s t90 = 1917 s

9.5 kPa

40

45

50

55

60

65

70

0 500 1000 1500 2000 2500 3000 3500 4000

Time (s)

σ r −

u0 (

kPa)

Average

TPT1

TPT2

Figure 7.95 Variation of external σr – u0 during consolidation in sensitive clay (test B14sus) (units in model scale)

Installation

Consolidation

In water

In soil

Pullout

σrc

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


Dep

th o

f tip

(m)

TPT1

TPT2

Hydrostatic

Average (B14SCC)

Figure 7.96 External radial total stress changes during installation, consolidation and pullout in sensitive clay (St = 4 - 5, test B14SCC)

In soil

σrc

Pullout

Installation

In water

0

2

4

6

8

10

12

14

16

0 50 100 150 200


epth

of t

ip (m

)

TPT1TPT2HydrostaticAverage (B14susa)

Figure 7.97 External radial total stress changes during installation, consolidation and pullout in sensitive clay (St = 4 - 5, test B14susa)

13.0

13.5

14.0

14.5

15.0

15.5

0 10 20 30 40 50 60

σr − u0 (kPa)

Dep

th o

f tip

(m) fail here

z =14.89 mzTPT = 10.09 mσ rf = 44.16 kPa

z =15.31 mσ rc = 51.08 kPa

Figure 7.98 External σr – u0 at failure during caisson pullout in sensitive clay (St = 4 - 5, test B14SCC)

13.0

13.2

13.4

13.6

13.8

14.0

14.2

14.4

14.6

14.8

15.0

0 10 20 30 40 50 60

σr − u0 (kPa)D

epth

of t

ip (m

)

Fail herez=14.39 mzTPT = 9.59 mσ rf = 46.90 kPa

Fail herez=14.79 mσ rc =50.53 kPa

Figure 7.99 External σr – u0 at failure during caisson pullout in sensitive clay (St = 4 - 5, Monotonic loading stage of test B14susa)

∆p = −250 kPa∆p = −241 kPa

∆p = −200 kPa

∆p = −150 kPa

Fail at this stage

∆p = P/A

-300

-250

-200

-150

-100

-50

00 50 100 150

Prototype time from start of loading (day)

Loa

ding

pre

ssur

e (k

Pa)

3.5

3.6

3.7

3.8

3.9

4.0

4.1

Em

bedm

ent (

Dia

met

ers)

Axial pressure

Displacement

Figure 8.1 Variations of axial pressure and embedment of the caisson during sustained loading in NC clay (test B12sus)


Figure 8.2 Axial pressure and internal pore pressure of the caisson during installation, consolidation, sustained loading and uplift in NC clay (test B12sus)

∆p = P/Abase

Fail here

02468

10121416

-400 -300 -200 -100 0 100 200


Dep

th o

f tip

(m)

B12sus, Sustainedloading,k = 1.23 kPa/m, -delt(p)/su,ave=27.2

B12SCC, Monotonicloading, k = 1.17 kPa/m,-delt(p)/su,ave=34.6

0

2

4

6

8

10

12

14

16

-100 -50 0 50 100 150

Internal pore pressure (kPa)

Dep

th o

f tip

(m)

Internal PPT

Hydrostatic

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20


Dep

th o

f tip

(m) su, ave=1.23 kPa/mNT-bar=10.5

Figure 8.3 Undrained shear strength versus the depth of T-bar in NC clay (B12sus)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)


Figure 8.4 External radial total stress changes during installation, sustained loading and uplift of the caisson in NC clay (test B12sus)

20

25

30

35

40

45

50

0 100 200 300 400 500 600 700Time (day)

σ r −

u0

(kPa

)

Average

σ rc = 24.57 kPa, after consolidation

Figure 8.5 External radial total stress changes versus prototype consolidation time in NC clay (test B12sus)

+1.20 kPa

−1.27 kPa

−2.70 kPa

−1.65 kPa−2.11 kPa

−1.52 kPa

0

5

10

15

20

25

30

0 50 100 150Prototype time from start of sustained loading (day)

Rad

ial t

otal

stre

ss (k

Pa)

3.5

3.6

3.7

3.8

3.9

4.0

4.1

Em

bedm

ent (

diam

eter

s)

TPT1TPT2Average TPTEmbedment

∆p = −150 kPa ∆p = −200 kPa ∆p = −250 kPa

Figure 8.6 Variations of external σr – u0 and embedment of caisson during sustained loading of the caisson in NC clay (test B12sus)

13.7

13.8

13.9

14.0

14.1

14.2

14.3

14.4

14.5

0 5 10 15 20 25 30σr−u0 (kPa)

Dep

th o

f tip

(m)

Fail herez =13.92 mzTPT = 6.72 mσ rf = 15.54 kPa After consolidation

z =14.39 mz TPT = 7.19 mσ rc = 24.57 kPa

Figure 8.7 Variation of external σr – u0 during uplift of the caisson subjected to sustained loading in NC clay (test B12sus)


Figure 8.8 Variations of axial pressure and internal pore pressure of the caisson

during installation, sustained loading and uplift in LOC clay (test B13sus)

∆p = P/Abase

0

2

4

6

8

10

12

14

16

-600 -400 -200 0 200


Dep

th o

f tip

(m)

B13sus, Sustained loading, k = 1.76 kPa/m, -delt(p)/su,ave=29.0

B13SCC, Monotonic loading,k = 1.64 kPa/m, -delt(p)/su,ave=34.1

0

2

4

6

8

10

12

14

16

-100 -50 0 50 100 150


Dep

th o

f tip

(m)

Internal pore pressure

Hydrostatic

∆p = P/A

Fail at this stage∆p = -190 kPa

∆p = -350 kPa

∆p = -275 kPa

∆p = -315 kPa

-400

-350

-300

-250

-200

-150

-100

-50

00 50 100 150 200 250

Prototype time from start of sustained loading (day)

Axi

al p

ress

ure,

∆p

(kPa

)13.30

13.35

13.40

13.45

13.50

13.55

13.60

13.65

13.70

13.75

Em

bedm

ent (

Dia

met

ers)

Axial pressure

Displacement

Figure 8.9 Variations of axial pressure and embedment of the caisson during sustained loading in LOC clay (test B13sus)

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


Dep

th o

f tip

(m)

TPT1

TPT2

Average

Hydrostatic

Figure 8.10 External radial total stress changes during installation, consolidation, sustained loading and uplift of the caisson in LOC clay

(test B13sus)

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


epth

of t

ip (m

)

Average (B13sus): Sustained loading

Average (B13SCC): Mono. loading

Hydrostatic

Figure 8.11 Comparison of variation of external σr with depth between sustained loading and monotonic loading in LOC clay (test B13sus and B13SCC)

+0.91 kPa(∆p=−350 kPa)

−1.5 kPa(∆p=−315 kPa)

−2.2 kPa(∆p=−275 kPa)

−3.1 kPa (∆p=−190 kPa)

20

25

30

35

40

45

50

55

60

65

70

0 50 100 150 200 250Prototype time from start of sustained loading (day)

σ r −

u0 (

kPa)

3.70

3.72

3.74

3.76

3.78

3.80

3.82

Em

bedm

ent (

Dia

met

ers)

TPT1TPT2Average TPTDisplacement

Figure 8.12 Variations of external σr – u0 and embedment of caisson during sustained loading in LOC clay (test B13sus)

10.0

10.5

11.0

11.5

12.0

12.5

13.0

13.5

14.0

0 10 20 30 40 50 60

σr−u0 (kPa)D

epth

of t

ip (m

) Fail herez =13.18 mzTPT = 8.38 m,σ rf = 32.94 kPa

After consolidationz =13.74 mzTPT = 8.94 m,σ rc = 54.40 kPa

Figure 8.13 Variation of external σr – u0 during uplift of the caisson subjected to sustained loading in LOC clay (test B13sus)


Figure 8.14 Variation of the axial pressure and internal pore pressure of the caisson during installation, sustained loading and uplift in sensitive clay

(test B14sus)

∆p = P/A

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100


Dep

th o

f tip

(m)

B14sus, Sustainedloading, k = 1.33 kPa/m,-delt(p)/su,ave=25.8B14SCC, Monotonicloading, k = 1.16 kPa/m,-delt(p)/su,ave=33.3

0

2

4

6

8

10

12

14

16

-150 -100 -50 0 50 100 150 200


Dep

th o

f tip

(m)

Internal porepressureHydrostatic

-140 kPa

-234 kPa

-165 kPa

-250 kPa

-205 kPa

Fail at this stage

∆p = P/A

-270kPa

-350

-300

-250

-200

-150

-100

-50

00 200 400 600 800


Axi

al p

ress

ure,

∆p

(kPa

)

3.0

3.2

3.4

3.6

3.8

4.0

4.2

Em

bedm

ent (

diam

eter

s)

Axial pressure

Displacement

Figure 8.15 Variation of the axial pressure and displacement of the caisson during sustained loading in sensitive clay (test B14sus)

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


Dep

th o

f tip

(m)

TPT1

TPT2

Average

Hydrostatic

Figure 8.16 External radial total stress changes during installation, sustained loading and uplift of the caisson in sensitive clay (test B14sus)

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


epth

of t

ip (m

)

Average of B14sus: Sustained loading

Average of B14SCC: Monotonic loading

Hydrostatic

Figure 8.17 Comparison of variation of external σr with depth between sustained loading and monotonic loading in sensitive clay

(tests B14sus and B14SCC)

∆p = −250 kPa, fail at this stage

−7.81 kPa

∆p = P/A

0

10

20

30

40

50

60

0 100 200 300 400 500 600 700 800Prototype time from start of sustained loading (day)

σ r −

u0

(kPa

)

3.0

3.2

3.4

3.6

3.8

4.0

4.2

Em

bedm

ent (

diam

eter

s)TPT1TPT2Average TPTDisplacement

Figure 8.18 External radial total stress changes and vertical displacement during sustained loading in sensitive clay (test B14sus)

13.7

13.8

13.9

14.0

14.1

14.2

14.3

14.4

14.5

14.6

14.7

0 10 20 30 40 50 60

σr−u0 (kPa)

Dep

th o

f tip

(m)

Fail herez =14.18 mzTPT = 9.38 mσ rf = 40.76 kPa After

consolidationz =14.59 mzTPT = 9.79 mσ rc = 50.10 kPa

Figure 8.19 Variation of external σr−u0 during monotonic uplift of the caisson after sustained loading in sensitive clay (test B14sus)


Figure 8.20 Axial pressure and internal pore pressure of the caisson during installation, monotonic and sustained loading in sensitive clay (test B14susa)

∆p = P/A

Monotonic loading

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100Axial pressure, ∆p (kPa)

Dep

th o

f tip

(m)

0

2

4

6

8

10

12

14

16

-50 0 50 100Internal pore pressure (kPa)

Dep

th o

f tip

(m)

Monotonic loading

Sustained loading after re-consolidation

Sustained loading after re-consolidation

∆p = −208 kPa

Fail at this stage

∆p = P/A

-300

-250

-200

-150

-100

-50

00 50 100 150 200 250


Axi

al p

ress

ure,

∆p

(kPa

)3.2

3.3

3.4

3.5

3.6

3.7

3.8

3.9

4.0

4.1

4.2

Em

bedm

ent (

diam

eter

s)

Axial pressure

Displacement

Figure 8.21 Variations of the axial pressure and displacement of the caisson during sustained loading in sensitive clay (test B14susa)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)


Figure 8.22 External radial total stress changes during installation, sustained loading and uplift of the caisson in sensitive clay (test B14susa)

∆p = −208 kPa

− 6.95 kPa

0

5

10

15

20

25

30

35

40

45

50

0 50 100 150 200 250


σ r −

u0 (

kPa)

3.2

3.3

3.4

3.5

3.6

3.7

3.8

3.9

4.0

4.1

4.2

Em

bedm

ent (

diam

eter

s)

TPT1TPT2Average TPTDisplacement

Figure 8.23 External radial total stress changes and vertical displacement during sustained loading in sensitive clay (test B14susa)

13.0

13.2

13.4

13.6

13.8

14.0

14.2

14.4

14.6

14.8

0 10 20 30 40 50

σr−u0 (kPa)

Dep

th o

f tip

(m)

Fail herez = 14.08 mzTPT = 9.28 mσ rf = 33.8 kPa

After consolidationz = 14.65 mzTPT = 9.85 mσ rc = 45.58 kPa

Figure 8.24 Variation of external σr – u0 during uplift of the caisson after sustained loading in sensitive clay (test B14susa)

∆p = P/A

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100 200


epth

of t

ip (m

)

B12cyc, Cyclic loading, k = 1.23 kPa/m, -delt(p)/su,ave=24.9

B12SCC, Monotonicloading, k = 1.17 kPa/m, -delt(p)/su,ave=34.6

Figure 8.25 Variation of the axial pressure of the caisson during installation, cyclic loading and uplift in NC clay (test B12cyc)

0

2

4

6

8

10

12

14

16

-20 -15 -10 -5 0 5 10 15 20


Dep

th o

f tip

(m) su, ave=1.23 kPa/mNT-bar = 10.5

Figure 8.26 Undrained shear strength during T-bar tests in NC clay (for test B12cyc)

−220 kPa

−170 kPa

−140 kPa

-300

-250

-200

-150

-100

-50

00 20 40 60 80 100

Prototype time from start of cyclic loading (day)

Upl

ift p

ress

ure,

∆p

(kPa

)3.92

3.93

3.94

3.95

3.96

3.97

3.98

3.99

4.00

4.01

4.02

Em

bedm

ent (

diam

eter

)

Uplift pressure

Embedment

Fail in this satge

Figure 8.27 Variation of the axial pressure and embedment of the caisson during cyclic loading in NC clay (test B12cyc)

-120 kPa-170 kPa

-220 kPa

14.24

14.26

14.28

14.30

14.32

14.34

14.36

14.38

-300 -250 -200 -150 -100 -50 0

Uplift pressure, ∆p (kPa)

Dep

th o

f tip

(m)

Figure 8.28 Uplift pressure versus embedment of the caisson during cyclic loading in NC clay (test B12cyc)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


epth

of t

ip (m

)

TPT1

TPT2

Hydrostatic

Average TPT

Figure 8.29 External radial total stress changes during installation, consolidation, cyclic loading and uplift of the caisson in NC clay (test B12cyc)

−7.25 kPa

−6.05 kPa

−9.25 kPa

15

17

19

21

23

25

27

29

31

33

35

0 20 40 60 80 100


σ r −

u0 (

kPa)

3.92

3.93

3.94

3.95

3.96

3.97

3.98

3.99

4.00

4.01

4.02

Em

bedm

ent (

diam

eter

)Average TPT

Embedment

Figure 8.30 Variations of external σr – u0 and embedment of the caisson during cyclic loading in NC clay (test B12cyc)

∆u ~1.4 kPa

−220 kPa

−170 kPa

−120 kPa

14.24

14.26

14.28

14.30

14.32

14.34

14.36

14.38

0 5 10 15 20 25 30σr − u0 (kPa)

Dep

th o

f tip

(m)

Fail herez =14.29 mzTPT = 7.09 mσ rf = 18.13 kPa

Average of 1st packet

Average of 2nd packet

Average of 3rd packet

After consolidationz =14.36 mzTPT=7.16 mσ rc = 25.5 kPa

Figure 8.31 Variation of external σr – u0 during uplift of the caisson after cyclic loading in NC clay (test B12cyc)

∆p = P/A

0

2

4

6

8

10

12

14

16

-500 -400 -300 -200 -100 0 100 200


Dep

th o

f tip

(m)

B13cyc, OCR=1.5, k=1.76kPa/m, -delt(pmin)/su,ave=29.2

B13SCC, OCR=1.5, k=1.64kPa/m, -delt(pmin)/su,ave=34.1

Figure 8.32 Variation of axial pressure of the caisson during installation, cyclic loading and uplift in LOC clay (test B13cyc)

su,ave=1.76 kPa/mNT-bar=10.5

0

2

4

6

8

10

12

14

16

-30 -20 -10 0 10 20 30 40

Undrained shear strength su (kPa)D

epth

of t

ip (m

)

Figure 8.33 Undrained shear strength versus the depth of T-bar in LOC clay (test B13cyc)

-370 kPa-350 kPa-315 kPa

-270 kPa

-195 kPa

-500

-450

-400

-350

-300

-250

-200

-150

-100

-50

00 50 100 150 200 250


Upl

ift p

ress

ure,

∆p

(kPa

)

3.66

3.68

3.70

3.72

3.74

3.76

3.78

Em

bedm

ent (

diam

eter

)

Uplift pressure

Embedment

Figure 8.34 Variation of the uplift pressure and embedment of the caisson during cyclic loading in LOC clay (test B13cyc)

13.44

13.46

13.48

13.50

13.52

13.54

13.56

13.58

13.60

13.62

-400 -300 -200 -100 0Uplift pressure, ∆p (kPa)

Dep

th o

f tip

(m)

Figure 8.35 Uplift pressure versus embedment of the caisson during cyclic loading in LOC clay (test B13cyc)

0

2

4

6

8

10

12

14

16

0 50 100 150 200


Dep

th o

f tip

(m)

TPT1

TPT2

HydrostaticAverage TPT

Figure 8.36 External radial total stress changes during installation, consolidation, cyclic loading and uplift of the caisson in LOC clay (test B13cyc)

20

25

30

35

40

45

50

55

0 50 100 150 200 250Prototype time from start of cyclic loading (day)

σr −

u0 (

kPa)

3.66

3.68

3.70

3.72

3.74

3.76

3.78

Em

bedm

ent (

diam

eter

)

Radial total stress

Embedment

Figure 8.37 Variations of external σr – u0 and embedment of the caisson under cyclic loading in LOC clay (test B13cyc)

12.8

12.9

13.0

13.1

13.2

13.3

13.4

13.5

13.6

13.7

0 10 20 30 40 50 60

σr−u0 (kPa)

Dep

th o

f tip

(m)

Fail herez = 13.21 mzTPT = 8.41 mσ rf = 31.98 kPa

Fail herez = 13.59 mzTPT = 8.79 mσ rc = 53.01 kPa

Figure 8.38 Variation of external σr – u0 during uplift of the caisson after cyclic loading in LOC clay (test B13cyc)

Fail here at a capacity ratio of 74 %

0

2

4

6

8

10

12

14

16

-400 -300 -200 -100 0 100


epth

of t

ip (m

)

B14cyc, cyclic loading, k= 1.36kPa/m, -det(pmin)/su, ave =24.6

B14SCC, monotonic loading, k=1.16kPa/m, -det(pmin)/su,ave =33.3

Figure 8.39 Variation of the axial pressure of the caisson during installation, cyclic loading and uplift in sensitive clay (test B14cyc)

0

2

4

6

8

10

12

14

16

-30 -20 -10 0 10 20 30


Dep

th o

f tip

(m)

su, ave=1.36kPa/mNT-bar = 10.5

Figure 8.40 Undrained shear strength versus the depth of T-bar in sensitive clay (test B14cyc)

-166 kPa

-207 kPa

-235 kPa-255 kPa

-300

-250

-200

-150

-100

-50

00 20 40 60 80 100 120 140


Upl

ift p

ress

ure,

∆p

(kPa

)

4.21

4.21

4.22

4.22

4.23

4.23

4.24

4.24

4.25

Em

bedm

ent (

diam

eter

)

Uplift pressure

Embedment

Figure 8.41 Variation of the axial pressure and displacement of the caisson during cyclic loading in sensitive clay (test B14cyc)

15.0

15.1

15.1

15.2

15.2

15.3

15.3

-300 -250 -200 -150 -100 -50 0Uplift pressure (kPa)

Dep

th o

f tip

(m)

Figure 8.42 Uplift pressure versus embedment of the caisson during cyclic loading in sensitive clay (test B14cyc)

0

2

4

6

8

10

12

14

16

0 50 100 150 200 250


epth

of t

ip (m

)

TPT1

TPT2

Average

Hydrostatic

Figure 8.43 External radial total stress changes during installation, consolidation, cyclic loading and uplift of the caisson in sensitive clay

(test B14cyc)

40

42

44

46

48

50

52

54

56

58

60

0 20 40 60 80 100 120 140


σr −

u0 (

kPa)

4.16

4.17

4.18

4.19

4.20

4.21

4.22

4.23

4.24

4.25

Em

bedm

ent (

diam

eter

)

Radial stress

Embedment

Figure 8.44 Variations of external σr – u0 (average) and embedment of the caisson during cyclic loading in sensitive clay (test B14cyc)

12.0

12.5

13.0

13.5

14.0

14.5

15.0

15.5

16.0

0 10 20 30 40 50 60

σr−u0 (kPa)D

epth

of t

ip (m

) Fail herez = 14.82 mzTPT = 10.02 mσ rf = 42.4 kPa

After consolidationz = 15.26 mzTPT = 10.46 mσ rc = 55.01 kPa

Figure 8.45 Variation of external σr – u0 with embedment of the caisson during

uplift of after cyclic loading in sensitive clay (test B14cyc)

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