uniaxial behaviour of suction caissons in · uniaxial behaviour of suction caissons in soft...
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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
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
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Centre for Offshore Foundation Systems The University of Western Australia
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
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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
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Centre for Offshore Foundation Systems The University of Western Australia
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
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Centre for Offshore Foundation Systems The University of Western Australia
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
7.3.2.2 Post-consolidation radial effective stress........................................ 7-30
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
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Centre for Offshore Foundation Systems The University of Western Australia
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
7.4.2.2 Post-consolidation radial effective stress........................................ 7-45
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
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9.2 FUTURE WORK.......................................................................................... 9-6
REFERENCES.........................................................................................................Ref-1
xi Notations
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 1 1-2 Introduction
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 1 1-3 Introduction
Centre for Offshore Foundation Systems The University of Western Australia
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.,
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Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 1 1-5 Introduction
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 1 1-6 Introduction
Centre for Offshore Foundation Systems The University of Western Australia
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,
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Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
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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:
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Centre for Offshore Foundation Systems The University of Western Australia
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
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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,
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Centre for Offshore Foundation Systems The University of Western Australia
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:
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Centre for Offshore Foundation Systems The University of Western Australia
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
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Centre for Offshore Foundation Systems The University of Western Australia
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
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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
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Centre for Offshore Foundation Systems The University of Western Australia
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.
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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
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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.
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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).
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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
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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
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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)
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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
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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
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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
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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.
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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.
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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
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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.
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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
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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
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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.
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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
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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.
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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.
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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.
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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
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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
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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
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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
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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
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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).
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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,
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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
Chapter 3 3-9 Experimental Apparatus and Soil Properties
Centre for Offshore Foundation Systems The University of Western Australia
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
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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
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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
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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
Centre for Offshore Foundation Systems The University of Western Australia
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
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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
Chapter 4 4-3 Performance of Total Pressure Transducers
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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
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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).
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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.
Chapter 4 4-6 Performance of Total Pressure Transducers
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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
Chapter 4 4-7 Performance of Total Pressure Transducers
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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
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Centre for Offshore Foundation Systems The University of Western Australia
(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
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Centre for Offshore Foundation Systems The University of Western Australia
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
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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
Chapter 4 4-11 Performance of Total Pressure Transducers
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 4 4-12 Performance of Total Pressure Transducers
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 4 4-13 Performance of Total Pressure Transducers
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 4 4-14 Performance of Total Pressure Transducers
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 4 4-15 Performance of Total Pressure Transducers
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 5 5-2 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 5 5-3 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 5 5-4 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 5 5-5 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 5 5-6 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 5 5-7 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 5 5-8 Interface Characteristics by Ring Shear Tests
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 6 6-2 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-3 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-4 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-5 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-6 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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.
Table 6.2 Variation of the resistance ratio with the number of cycles of T-bar test
in LOC kaolin clay
Number of cycle B13TB1
(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
Number of cycle B13TB2
(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.
Chapter 6 6-7 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
Subsequently a sensitivity of 4 - 5 was adopted for the sensitive clay used in this
research.
Table 6.3 Variation of the resistance ratio with the number of cycles of T-bar test
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
Chapter 6 6-8 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-9 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-10 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-11 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-12 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-13 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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 &
Chapter 6 6-14 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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)
Chapter 6 6-15 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-16 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-17 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-18 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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,
Chapter 6 6-19 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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,
Chapter 6 6-20 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-21 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-22 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-23 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-24 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-25 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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)
Chapter 6 6-26 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-27 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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).
Chapter 6 6-28 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-29 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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,
Chapter 6 6-30 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-31 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 6 6-32 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-33 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-34 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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,
Chapter 6 6-35 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-36 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 6 6-37 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 6 6-38 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 6 6-39 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 6 6-40 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 6 6-41 Axial Capacity of Caissons in Clay
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-2 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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)
Chapter 7 7-3 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-4 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-5 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-6 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-7 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-8 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-9 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-10 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-11 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-12 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-13 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-14 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-15 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-16 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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 Analysis of Radial Stresses during Installation
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
Chapter 7 7-17 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-18 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-19 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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%.
Chapter 7 7-20 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-21 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-22 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-23 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-24 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-25 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-26 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Measured or derived 7.3 1.8* 5.5* 1.08 0.25 0.83 5.4 1.2 4.2
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
Chapter 7 7-27 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-28 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-29 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-30 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
7.3.2.2 Post-consolidation radial effective stress
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
Chapter 7 7-31 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
7.3.3 Radial Stress Changes and Shaft Friction during Pullout
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
Chapter 7 7-32 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-33 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Measured 0.73 - 0.96
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 Analysis of Radial Stresses during Installation
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
Chapter 7 7-34 Radial Stress Changes around Caissons
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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.
Chapter 7 7-35 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-36 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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=
Chapter 7 7-37 Radial Stress Changes around Caissons
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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,
Chapter 7 7-38 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-39 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-40 Radial Stress Changes around Caissons
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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
Measured or derived 6.3 0.8* 5.5* 0.86 0.11 0.75 6.4 0.8 5.6
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.
Chapter 7 7-41 Radial Stress Changes around Caissons
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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.
Chapter 7 7-42 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Measured or derived 6.7 0.8* 5.9* 0.92 0.11 0.81 6.0 0.7 5.3
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
Chapter 7 7-43 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-44 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 7 7-45 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
7.4.2.2 Post-consolidation radial effective stress
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
Chapter 7 7-46 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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.
7.4.3 Radial Stress Changes and Shaft Friction during Pullout
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
Chapter 7 7-47 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 7 7-48 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Measured 0.65 - 0.75
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
Chapter 7 7-49 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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,
Chapter 7 7-50 Radial Stress Changes around Caissons
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-2 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-3 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-4 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-5 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-6 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-7 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-8 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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,
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Centre for Offshore Foundation Systems The University of Western Australia
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.
The variation of σr – u0 with embedment of the caisson during the early stage of
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.
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Centre for Offshore Foundation Systems The University of Western Australia
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
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Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-12 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-13 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
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Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-15 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-16 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
–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.
The variation of σr – u0 with embedment of the caisson during the early stage of
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
Chapter 8 8-17 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 8 8-18 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 8 8-19 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 8 8-20 Sustained Loading and Cyclic Loading
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 9 9-2 Conclusions and Future Work
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 9 9-3 Conclusions and Future Work
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 9 9-4 Conclusions and Future Work
Centre for Offshore Foundation Systems The University of Western Australia
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.
Chapter 9 9-5 Conclusions and Future Work
Centre for Offshore Foundation Systems The University of Western Australia
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
Chapter 9 9-6 Conclusions and Future Work
Centre for Offshore Foundation Systems The University of Western Australia
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:
Chapter 9 9-7 Conclusions and Future Work
Centre for Offshore Foundation Systems The University of Western Australia
• 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
Centre for Offshore Foundation Systems The University of Western Australia
REFERENCES
AG. (2002). Suction pile analysis code: AGSPANC, Version 4.1, Advanced
Geomechanics, Perth.
Al-Khafaji, Z., Hossain, M.K., Audibert, J.M.E., Clukey, E.C., Templeton, J.S. & de
Jong, P.R. (2003). Suction caisson foundation design for vortex-induced vibration
loading. Proc. of the 35th Annual Offshore Tech. Conf., Houston, USA, Paper
OTC15239.
Allersma, H.G.B., Kirstein, A.A., Brinkgreve, R.B.J. & Simon, T. (1999). Centrifuge
and numerical modelling of horizontally loaded suction piles. Proc. of the 9th Inter.
Offshore and Polar Engrg. Conf., ISOPE '99, Brest, France, 1, 711-717.
Andersen, K.H., Dyvik, R., Schroeder, K., Hansteen, O.E. & Bysveen, S. (1993). Field
tests of anchors in clay II: Predictions and interpretation. J. of Geotech. Engrg,
ASCE, 119 (10), 1532-1549.
Andersen, K.H., Andersen, L., Jostad, H.P. & Clukey, E.C. (2004). Effect of skirt-tip
geometry outside suction anchors in soft clay. Proc. 24th Int. Conf. Offshore Mech.
and Arctic Engrg., OMAE’04, Paper OMAE2004-51564, Vancouver, Canada.
Andersen, K.H., Jeanjean, P., Luger, D. & Jostad, H.P. 2003. Centrifuge tests on
installation of suction anchors in soft clay. Int. Symp. on Deepwater Mooring
Systems, Houston, Texas, USA, 13-27.
Andersen, K.H. & Jostad, H.P. (1999). Foundation design of skirted foundations and
anchors in clay. Proc. of the 31th Annual Offshore Tech. Conf., Houston, USA.
OTC Paper No. 10824.
Andersen, K.H. & Jostad, H.P (2002). Shear strength along outside wall of suction
anchors in clay after installation. Proc. 12th Int. Offshore and Polar Engng. Conf.,
ISOPE’2002, 785-794.
Andersen, K.H. & Jostad, H.P. (2004). Shear strength along inside wall of suction
anchor skirt wall in clay. Proc. of the 36th Annual Offshore Tech. Conf., Houston,
USA, Paper OTC16844, 1-13.
Ref-2 References
Centre for Offshore Foundation Systems The University of Western Australia
Andersen, K.H., Murff, J.D., Randolph, M.F., Erbrich, C., Jostad, H.P., Hansen, B. C.
Aubeny. C. & Sharma, P. & Supachawarote, C. (2005). Suction anchors for
deepwater applications. Proc. of International Symposium on Frontiers in Offshore
Geotechnics, ISFOG’05, Perth, Australia, 3-30.
API (1993). RP2A: Recommended practice for planning, designing and constructing
fixed offshore platforms. API Rec. Practice 2A. American Petroleum Institute,
Dallas, Texas.
Aubeny, C.P., Murff, J.D. & Roesset, J.M. (2001). Geotechnical issues in deep and ultra
deep waters. The International Journal of Geomechnics. 1(2), 225-247.
Audibert, J.M.E., Clukey, E. & Huang, J. (2003). Suction caisson installation at horn
mountain - A Case History. Proc. of the 13th Int. Offshore and Polar Engng. Conf.,
ISOPE’2003, Honolulu, USA, 1309-1316.
Azzouz, A.S. & Morrison, M.J. (1988). Field measurements on model pile in two clay
deposits. J. of Geotech. Engrg, ASCE, 114(1), 104-121.
Budhu, M. (1983). Nonuniformities imposed by simple shear apparatus. Canadian
Geotechnical Journal, 20, 125-137.
Baligh, M. M. (1985). Strain Path Method. J. of Geotech. Engrg.. 111(9), 1108-1136.
Baligh, M.M. (1986). Undrained deep penetration. Géotechnique, 36(4), 471-485;
487-501.
Baligh, M.M., Azzouz, A.S., Chin, C.T. (1987). Disturbance due to ‘ideal’ tube
sampling. J. of Geotech. Engrg.. 113(7). 739-757.
Bea, R.G. & Audibert, J.M.E. (1979). Performance of dynamically loaded pile
foundations. Boss’79, Second International conference on Behavior of Offshore
Structures, Imperial College, London, England, August, 28-31.
Bishop, A.W., Green, G.E., Garga, V.K., Andresen, A. & Brown, J.D. (1971). A new
ring shear apparatus and its application to the measurement of residual strength.
Géotechnique, 21(4), 273-328.
Bond, A.J. (1989). Behaviour of Displacement Piles in Overconsolidated Clays. Ph.D.
thesis, University of London, Imperial College, London.
Ref-3 References
Centre for Offshore Foundation Systems The University of Western Australia
Bond, A.J. & Jardine, R.J. (1991). Effects of installing displacement piles in a high
OCR clay. Géotechnique, 41(3). 341-363.
Bromhead, E.N. (1979). A simple ring shear apparatus. Ground Engineering, 12(5),
40-44.
Brown, S.F. & Pell, P.S. (1967). Subgrade stress and deformation under dynamic load.
J. of Soil Mech. and Found. Div., ASCE., 93(1), 17-46.
Brummund, W.F. & Leonards, G.A. (1973). Experimental study of static and dynamic
friction between sand and typical construction materials. ASTM Journal of Testing
and Evaluation, 1, 162-165.
Cao, J., Phillips, R., Popescu, R., Al-Khafaji, Z. & Audibert, J.M.E. (2002a).
Penetration resistance of suction caissons in clay. Proc. of 12th Int. Offshore and
Polar Engng. Conf., ISOPE’2002, Kitakyushu, Japan, 800-806.
Cao, J., Phillips, R., Popescu, R., Audibert, J. & Al-Khafaji, Z. (2002b). Excess pore
pressures induced by installation of suction caissons in NC clays. Proc. Int. Conf.
on Offshore Site Investigation and Geotechnics, Society for Underwater
Technology, London, 405-412.
Cao, J., Phillips, R., Popescu, R., Audibert, J.M.E. & Al-Khafaji, Z (2002c). Numerical
analysis of the behavior of suction caissons in clay. Proc. of 12th Int. Offshore and
Polar Engng. Conf., ISOPE’2002, Kitakyushu, Japan, 795-799.
Chen, W. & Randolph, M.F. (2004a). Radial stress changes around caissons installed in
clay by jacking and by suction. Proc. of 14th Int. Offshore and Polar Engng. Conf.,
ISOPE’2004, Toulon, France, 493-499.
Chen, W. & Randolph, M.F. (2004b). Radial stress changes and axial capacity for
suction caissons in soft clay. (submitted to Géotechnique).
Chen, W. & Randolph, M.F. (2005). Centrifuge tests on axial capacity of suction
caissons in clay. Proc. Int. Symp. on Frontiers in Off-shore Geotechnics,
ISFOG’05, Perth, Australia, 243-249.
Chin, C.T. (1986). Open-ended pile penetration in saturated clays. Ph.D. Thesis,
Massachusetts Institute of Technology.
Ref-4 References
Centre for Offshore Foundation Systems The University of Western Australia
Chow, F.C. (1997). Investigations into the behaviour of displacement piles for offshore
structures. Ph.D. thesis, Univ. of London (Imperial College).
Clayton, C.R.I., Siddique, A. & Hopper, R.J. (1998). Effects of sampler design on tube
sampling disturbance - numerical and analytical investigations. Géotechnique,
48(6), 847-867.
Clukey, E.C. & Morrison, M.J. (1993). A centrifuge and analytical study to evaluate
suction caissons for TLP applications in the Gulf of Mexico. ASCE Spec. Publ. in
Design & Perform. of Deep Found.: Piles & Piers in Soil & Soft Rock.
Clukey, E. C., Morrison, M. J., Gariner, J., & Corté, J. F., (1995). The response of
suction caissons in normally consolidated clays to cyclic TLP loading conditions.
Proc. of the 27th Annual Offshore Tech. Conf., Houston, USA, Paper OTC 7796,
909–918.
Clukey, E.C. & Phillips, R. (2002). Centrifuge model tests to verify suction caisson
capacities for taut and semi-taut legged mooring systems. Proc. Int. Conf. on Deep
Offshore Tech., DOT 2002, New Orleans, USA, Nov., 1-16.
Clukey, E. C., Templeton, J. S., Randolph, M. F. & Phillips, R. (2004). Suction caisson
response under sustained loop current loads. Proc. of the 36th Annual Offshore
Tech. Conf., Houston, USA, Paper OTC16843, 1-9.
Clukey, E. C. (2005). Suction caisson soil displacement during installation. Proc. Int.
Symp. on Frontiers in Off-shore Geotechnics, ISFOG’05, Perth, Australia, 229-234.
Colliat, J.-L. (2002). Anchors for deepwater to ultradeepwater moorings. Proc. of the
34th Annual Offshore Tech. Conf., Houston, USA, Paper OTC14306, 2695-2703.
Colliat, J.-L. & Dendani, H. (2002). Girassol: Geotechnical design analysis and
installation of the suction caisson. Proceedings, SUT International Conference on
Offshore Site Investigation and Geotechnics, London, November. 26-28.
Coop, M.R. & Wroth, C.P. (1989). Field studies of an instrumented model pile in clay.
Géotechnique, 39(4), 679-696.
Dendani, H. & Colliat, J.-L. (2002). Girassol: Design analysis and installation of the
suction anchors. Proc. of the 34th Annual Offshore Tech. Conf., Houston, USA,
Paper OTC14240, 1869-1875.
Ref-5 References
Centre for Offshore Foundation Systems The University of Western Australia
Deng, W. & Carter J. P (2000). A theoretical study of the vertical uplift capacity of
suction caissons. Proc. of 10th Int. Offshore and Polar Engng. Conf., ISOPE’2000,,
Seattle, USA, 342-349.
Dewoolkar, M.M., Ko., H.Y. & Pak, R.Y.S. (1998). Suitability of total stress gauges for
soil pressure measurements. Kimura, Kusakabe & Takemura (eds.), Proc. Inter.
Conf. Centrifuge 98. Rotterdam, Balkema. 129-134.
Desai, C.S., Drumm, E.C. & Zaman, M.M. (1985). Cyclic testing and modelling of
interfaces. J. of Geotech. Engrg., ASCE, 111(6), 793-815.
Dyvik, R., Andersen, K. H., Hansen, S. B., Christophersen, H. P. (1993). Field tests of
anchors in clay. I: Description. J. of Geotech. Engrg., ASCE, 119(10), 1515-1531.
Egan, D. & Merrifield, C.M. (1998). The use of miniature earth pressure cells in multi-
gravity environment. Kusakabe & Takemura (eds.), Proc. Inter. Conf. Centrifuge
98. Rotterdam, Balkema. 55-60.
Ehlers, C.J., Young, A.G. & Chen, J.W. (2004). Technology Assessment of Deepwater
Anchors. Proc. of the 36th Annual Offshore Technology Conference, Houston,
Paper OTC 16840, 1-17.
Einav, I. and Randolph, M.F. (2005). Combining upper bound and strain path methods
for evaluating penetration resistance. Int. J. Num. Methods in Eng., in press.
El-Gharbawy, S.L. (1998). The Pullout capacity of suction caisson foundations for
tension leg platforms. PhD Dissertation, The University of Texas at Austin.
El-Gharbawy, S.L. & Olson, R.E. (1999). The cyclic pullout capacity of suction caisson
foundations. Proc. of the 9th Inter. Offshore and Polar Engrg. Conf., ISOPE '99,
Brest, France, 2, 660-667.
El-Gharbawy, S.L., Olson, R.E. & Scott, S.A. (1999). Suction anchor installations for
deep Gulf of Mexico applications. Proc. of the 31st Annual Offshore Technology
Conference, Houston, Paper OTC 10992, 747-754.
Elkhatib, S. (1997). In situ assessment of the shear strength of soils. Final year
undergraduate Honours thesis, The University of Western Australia.
Ref-6 References
Centre for Offshore Foundation Systems The University of Western Australia
Erbrich, C. & Hefer, P.A (2002). Installation of Laminaria suction caissons - a case
history. Proc. of the 34th Annual Offshore Tech. Conf., Houston, Paper OTC 14240,
2157-2170.
Fahey, M. & Lee Goh, A. (1995). A comparison of pressuremeter and piezocone
methods of determining the coefficient of consolidation. Proc. 4th Int. Symp. on
Pressuremeters. Qubec, Balkema, Rotterdam, 153-160.
Fuglsang, L. D. & Steensen-Bach, J. O. (1991). Breakout resistance of suction piles in
clay. Int. Conf. Centrifuge 91, Boulder Colorado, Proceedings, 153-159.
Gibson, R.E. & Anderson, W.F. (1961). In situ measurements of soil properties with the
pressuremeter. Civ. Eng., Public Wks Rev., 56, 615-618.
Hogervorst, J.R., (1980). Field trials with large diameter suction piles. Proceedings of
the 12th Annual Offshore Technology Conference, Houston, USA, Paper OTC 3817,
217–224.
House, A. (2002). Suction caisson foundations for buoyant offshore facilities. PhD
Thesis, The University of Western Australia.
House, A., Randolph, M.F. & Borbas, M.E. (1999). Limiting aspect ratio for suction
caisson installation in clay. Proc. 9th Int. Offshore and Polar Engineering
Conference - ISOPE ’99, Brest, France, 676-683.
House, A. & Randolph, M.F. (2001). Installation and pull-out capacity of stiffened
suction caissons in cohesive sediments. 11th Int. Offshore and Polar Engng. Conf.,
ISOPE’2001, 574-580.
House, A.R., Oliverira, J.R.M.S. & Randolph, M.F. (2001). Evaluating the coefficient
of consolidation using penetration tests. Int. J. of Physical Modelling in
Geotechnics, 1(3). 17-25.
Hu, Y., Randolph, M.F. & Watson, P.G (1999). Bearing response of skirted
foundations. J. of Geotech. and Geoenviron. Engrg., ASCE, 125(11), 924-935.
Huang, J., Cao, J.C., Jean, M.E. & Audibert, M. (2003). Geotechnical design of suction
caissons in clay. Proc. 13th Int. Offshore and Polar Engng. Conf., ISOPE’2003,
Honolulu, USA, 770-779.
Ref-7 References
Centre for Offshore Foundation Systems The University of Western Australia
Hvorslev, M. J. (1976). The changeable interaction between soils and pressure cells:
tests and reviews at the Waterways Experiment Station, Technical Report No.
S-76-7. Vicksburg, MO: US Army Engineer Waterways Experiment Station.
Infield Worldwide Offshore Energy Database. www.infield.com
Jardine, R.J. & Chow, F.C. (1996). New design methods for offshore piles, MTD
Publication, 96/103.
Jewell R.J. & Randolph M.F. (1988). Cyclic rod shear tests in calcareous sediment. Vol.
1, Engineering for Calcareous Sediments, Proc. Inter. Conf. on Calcareous
Sediments, Perth, 215–222.
Keaveny, J.M., Hansen, S.B., Madshus, C. & Dyvik, R. (1994). Horizontal capacity of
large scale model anchors. Proc. XIII ICSMFE. New Delhi, 2, 677-680.
Kelly, R. (2001). Development of a large diameter ring shear apparatus and its use for
interface testing. Ph.D thesis, University of Sydney.
Kjellman, W. (1951). Testing the shear strength of clay in Sweden. Géotechnique, 1,
225-232.
Kolk, H.J. & van der Velde, E. (1996). A reliable method to determine friction capacity
of piles driven into clay. Proc. of the 28th Annual Offshore Technology Con.,
Houston, Paper OTC 7993, 337-346.
Kulhawy, F.H. & Peterson, M.S. (1979). Behavior of sand-concrete interfaces. 6th Pan-
American conference on soil mechanics, 225-236.
Labuz, J.F. & Theroux, B. (2005). Laboratory calibration of earth pressure cells.
Geotechnical Testing Journal, 28(2), 1-9.
Lee. F.H., Juneja, A. & Tan, T.S. (2004). Stress and pore pressure changes due to sand
compaction installation in soft clay. Géotechnique, 54(1), 1-16.
Lehane, B.M. (1992). Experimental investigations of pile behaviour using instrumented
field piles. Ph.D. thesis, University of London, Imperial College, London.
Lehane, B.M. & Jardine, R.J. (1992). Residual strength characteristics of Bothkennar
clay. Géotechnique, 42(2), 363-367.
Ref-8 References
Centre for Offshore Foundation Systems The University of Western Australia
Lehane, B.M. & Jardine, R.J. (1994). Displacement pile behavior in a soft marine clay.
Canadian Geotechnical Journal, 31(2), 181-191.
Lehane, B.M., Jardine, R.J., Bond, A.J. & Chow, F.C. (1994). The development of shaft
friction on displacement piles in clay, Proc. 13th Int. Conf. on Soil Mech. and
Found. Eng., New Dehli, 2, 473-476.
Lemos, L.J.L. (1986). The effect of rate of shear on residual strength of soil. Ph.D
thesis, University of London (Imperial College).
Loez, B. (2002). Girassol: The Biggest FPSO in the World: as Seen by Its Contractor.
Proc. of the 34th Annual Offshore Technology Conf., Houston, 1899-1918.
Labuz, J.F. & Theroux. B. (2005). Laboratory calibration of earth pressure cells.
Geotechnical Testing Journal, 28(2), 188-196.
Luke, A.M., (2002). Axial capacity of suction caisson in normally consolidated
kaolinite. M.S.Thesis, The University of Texas at Austin.
Lunne, T. (2001). In situ testing in offshore geotechnical investigations. Proc. Int. Conf.
on In Situ Measurement of Soil Properties and Case Histories, Bali, 61-81.
Lunne, T., Randolph, M.F., Sjursen, M.A. and Chung, S.F. (2005). Comparison of cone
and T-bar resistance factors in a range of offshore and onshore soft sediments.
Proc. Int. Symp. on Frontiers in Off-shore Geotechnics, ISFOG’05, Perth, Australia,
981-989
Mayne, P.W. & Kulhawy, F.H. (1982). K0-OCR relationships in soils. J. Geotech. Eng.
Div.,ASCE, 108(GT6), 851-872.
McCarron, W.O. & Sukumaran, B. (2000). Ultimate capacities of suction caissons and
pile elements for deepwater applications. Proc. 10th Int. Conf. On Offshore and
Polar Engng, Seattle, 2, 466-469.
Mello, J.R.C., Moretti, M.J., Sparrevik, P., Schrøder, K. & Hansen, S.B. (1998). P19
and P26 moorings at the Marlim Field. The first permanent taut leg mooring with
fibre rope and suction anchors. Proc. Int. Conf. on Floating Production Systems,
FPS’98.
Ref-9 References
Centre for Offshore Foundation Systems The University of Western Australia
Morrison, M.J., Clukey, E.C. & Garnier, J. (1994). Behaviour of suction caissons under
static uplift loading. Proc. of the Inter. Conf. on Centrifuge Modelling - Centrifuge
‘94, Leung, Lee & Tan (eds.), Singapore, Balkema, Rotterdam, 823-828.
Murff, J.D. (1996). Geotechnical centrifuge in offshore engineering. Proceedings of the
28th Annual Offshore Technology Conference, Houston, 675-689.
Mustang Engineering (2004). www.mustangeng.com.
Newlin, J.A. (2003a). Suction anchor piles for the Na Kika FDS mooring system, part
1: site characterization and design. Deepwater Mooring Systems: Concepts, Design,
Analysis, and Materials, ASCE. Houston, USA, Oct., 2003, 28-54.
Newlin, J.A. (2003b). Suction anchor piles for the Na Kika FDS mooring system, part
2: installation performance. Deepwater Mooring Systems: Concepts, Design,
Analysis, and Materials, ASCE. Houston, USA, Oct., 2003, 55-57.
Offshore Technology Research Center (2001). Workshop report: design methodologies
and criteria for suction caissons for deepwater mooring applications. OTRC, USA.
Offshore Engineer (1996a). Taut leg tested in rig role. November 1996, 15-17.
Offshore Engineer. (1996b). Suction success for Shell. November 1996, 26.
O’Loughlin, C.D., Randolph, M.F. & Richardson, M. (2004). Experimental and
theoretical studies of deep penetration anchors. Proc. of the 36th Annual Offshore
Technology Conference, Houston, Paper OTC 16841.
Olson, R.E., Rauch, A.F., Luke, A.M., Maniar, D.R., Tassoulas, J.L. & Mecham, E.C.
(2003). Soil reconsolidation following the installation of suction caissons. Proc., of
the 35th Offshore Technology Conf., Houston, Paper 15263.
O’Neill, D., Pezzetti, G. & Manes, V. (1991). In Situ Penetration of a Large Scale
Instrumented Model Skirt Pile. Field Measurements in Geotechnics, 1991
Balkema, Rotterdam. ISBN 90 5410 0257.
Ooi, L. H. & Carter, J.P. (1987). Constant normal stiffness direct shear divice for static
and cyclic loading. Geotechnical Testing Journal, 10(1), 3-12.
Randolph, M.F., Carter, J. P. & Wroth, C.P. (1979). Driven piles in clay - the effects of
installation and subsequent consolidation. Géotechnique, 29(4), 361-393.
Ref-10 References
Centre for Offshore Foundation Systems The University of Western Australia
Randolph, M.F., Cassidy, M., Gourvenec, S. & Erbrich. C. (2005). Challenges of
offshore geotechnical engineering. The 16th International Soil Mechanics and
Geotechnical Engineering, ICSMGE., OSAKA, Japan. September, 123-176.
Randolph, M.F. & House, A.R. (2002). Analysis of suction caisson capacity in clay.
Proceedings of the 34th Annual Offshore Technology Conference, Houston,
2145-2155.
Randolph, M.F. & Houlsby, G.T. (1984). The limiting pressure on a circular pile loaded
laterally in cohesive soil. Géotechnique, 34(4), 613-623.
Randolph, M.F., Jewell, R.J., Stone, K.J.L. & Brown, T.A. (1991). Establishing a new
centrifuge facility. Proc. Int. Conf. on Centrifuge Modelling - Centrifuge 91,
Boulder, Colorado, 3-9.
Randolph, M. F. & Murphy, B. S. (1985). Shaft capacity of driven piles in clay, Proc.
17th Ann. Offshore Technol. Conf., Houston, 1, 371–378.
Randolph, M.F., O'Neill, M.P., Stewart, D.P. & Erbrich, C. (1998). Performance of
suction anchors in fine-grained calcareous soils. Proc. of 30th Annual Offshore
Technology Conference, Houston, Paper OTC 8831.
Randolph, M.F. & Wroth, C.P. (1979). An analytical solution for the consolidation
around a driven pile. Int. J. Num. and Anal. Methods in Geomechanics, 3(3),
217-229.
Randolph, M.F. (2003). Science and empiricism in pile foundation design. 43rd Rankine
Lecture, Géotechnique, 53(10), 847-875.
Randolph, M.F. (2004). Characterisation of soft sediments for offshore applications,
Keynote Lecture, Proc. 2nd Int. Conf. on Site Characterisation, Porto, 1, 209-231.
Randolph, M.F., Watson, P.G. & Fahey, M. (1999). Site characterization and
foundation design in soft sediments. Int. Conf. on Offshore and Nearshore Geotech.
Engrg. New Delhi., December, 35-45.
Ramsey, N., Jardine R., Lehane B. & Ridley, A. (1998). A review of soil-steel interface
testing with the ring shear apparatus. Offshore Site Investigation and Foundation
Behaviour’98, London, 237-257.
Ref-11 References
Centre for Offshore Foundation Systems The University of Western Australia
Rauch, A.F., Olson, R.E., Coffman, R.A. & El-Sherbiny, R.M. (2004). Measured
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
Centre for Offshore Foundation Systems The University of Western Australia
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
Centre for Offshore Foundation Systems The University of Western Australia
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
0 50 100 150 200 250Radial total stress (kPa)
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
Shearing distance (mm)
Forc
e (N
) Proving Ring A
Proving Ring B
0
2
4
6
8
10
12
14
16
18
0 20 40 60
Shearing distance (mm)
Inte
rfac
e fr
ictio
n an
gle
(o )
0
5
10
15
20
25
30
35
0 20 40 60
Shearing distance (mm)
Shea
r St
ress
, τ
(kPa
)
σ v = 100 kPa
δr = 15
(a) Force - shearing distance (b) Vertical settlement - shearing distance
(c) Shear stress - shearing distance (d) δ - shearing distance
Figure 5.2 Ring shear test S1-2 in NC clay (smooth platen)
-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
Figure 5.3 Ring shear test S1-3 in NC clay (smooth platen)
-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
0 1 2 3 4 5 6 7 8Shearing distance (mm)
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
Shearing distance (mm)
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
Shearing distance (mm)
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
(a) Force - shearing distance (b) Vertical settlement - shearing distance
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
Shearing distance (mm)
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
0 2 4 6 8 10 12 14 16Shearing distance (mm)
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
Shearing distance (mm)
Forc
e (N
)
Proving Ring A
Proving Ring B
σ v = 50 kPa
(a) Force - shearing distance (b) Vertical settlement - shearing distance
δr = 18.3o
δp = 19.4o
0
5
10
15
20
25
0 1 2 3 4 5 6Shearing distance (mm)
Inte
rfac
e fr
ictio
n an
gle
(o )
02468
101214161820
0 1 2 3 4 5 6Shearing distance (mm)
Shea
r St
ress
, τ (k
Pa)
σ v = 50 kPa
(c) Shear stress - shearing distance (d) δ - shearing distance
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
Shearing distance (mm)
Forc
e (N
)
Proving Ring A
Proving Ring B
σ v = 50 kPa
(a) Force - shearing distance (b) Vertical settlement - shearing distance
0
2
4
6
8
10
12
14
16
18
0 5 10 15
Shearing distance (mm)
Shea
r St
ress
, τ (k
Pa)
δp = 18.3o
δr = 17.9o
0
5
10
15
20
0 5 10 15
Shearing distance (mm)
Inte
rfac
e fr
ictio
n an
gle
(o )
σ v = 50 kPa
(c) Shear stress - shearing distance (d) δ - shearing distance
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
Shearing distance (mm)
Forc
e (N
)
Proving Ring A
Proving Ring B
σ v = 50 kPa
(a) Force - shearing distance (b) Vertical settlement - shearing distance
δr = 14.8o
δr = 11.3o
0
5
10
15
20
0 5 10 15
Shearing distance (mm)
inte
rfac
e fr
ictio
n an
gle
(o )
0
2
4
6
8
10
12
14
0 5 10 15
Shearing distance (mm)
Shea
r St
ress
, τ (k
Pa)
σ v = 50 kPa
(c) Shear stress - shearing distance (d) δ - shearing distance
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
Shearing distance (mm)
Ver
tical
Dis
plac
emen
t (m
m)
0
2
4
6
8
10
12
14
16
0 5 10 15
Shearing distance (mm)
Focr
e (N
)
Proving Ring A
Proving Ring B
σ v = 50 kPa
(a) Force - shearing distance (b) Vertical settlement - shearing distance
0
2
4
6
8
10
12
14
0 5 10 15
Shearing distance (mm)
Shea
r St
ress
, τ (k
Pa)
δp = 14.3o
δr = 12.1ο
0
5
10
15
0 5 10 15
Shearing distance (mm)
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
Undrained shear strength su measured by T-bar (kPa)
Dep
th o
f tip
(m)
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 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
Undrained shear strength su measured by T-bar (kPa)
Dep
th o
f tip
(m)
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 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)
(a) Strength profiles during installation and pullout
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
Undrained shear strength su measured by T-bar (kPa)
Dep
th o
f tip
(m)
su,ave=1.55 kPa/mNT-bar = 10.5
(a) Strength profiles during installation and pullout
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
Undrained shear strength su (kPa)
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
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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
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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
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0 50 100 150
Penetration resistance (kPa)
Dep
th o
f tip
(m)
Measured
Install:Nc=7.5,Alpha=0.38
Nominal installation: 14.05 m
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0 50 100 150 200
Pore pressure (kPa)
Dep
th o
f tip
(m)
PPTI
Hydrostatic
(a) Penetration resistance (b) Internal pore pressure
Figure 6.10 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B2JOC)
Nominal installation: 14.02 m
∆p = P/A
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0 50 100 150 200
Penetration resistance (kPa)
Dep
th o
f tip
(m)
Measured
Install:Nc=7.5,Alpha=0.40
Figure 6.11 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B4JCI)
∆p = P/A
Nominal installation: 13.87 m
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0 50 100 150
Penetration resistance (kPa)
Dep
th o
f tip
(m)
Measured
Install:Nc=7.5,Alpha=0.40
Nominal installation: 13.87 m
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0 50 100 150 200
Pore pressure (kPa)
Dep
th o
f tip
(m)
PPTI
Hydrostatic
(a) Penetration resistance (b) Internal pore pressure
Figure 6.12 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B6JOC)
∆p = P/A
Nominal installation: 14.38 m
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0 50 100 150
Penetration resistance (kPa)
Dep
th o
f tip
(m)
Measured
Install:Nc=7.5,Alpha=0.39
Figure 6.13 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B6JCC)
Nominal installation: 14.10 m
∆p = P/A
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0 50 100 150
Penetration resistance (kPa)
Dep
th o
f tip
(m)
Measured
Install:Nc=7.5,Alpha=0.38
Nominal installation: 14.10 m
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0 50 100 150 200
Pore pressure (kPa)
Dep
th o
f tip
(m)
PPTI
Hydrostatic
(a) Penetration resistance (b) Internal pore pressure
Figure 6.14 Variation of penetration resistance and the internal pore pressure with penetration depth (NC clay, test B8JOC)
0
2
4
6
8
10
12
14
16
0 20 40 60 80 100 120 140
Penetration resistance (kPa)
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
Nominal installation: 14.02 m
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
Nominal installation: 14.02 m
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)
PPT-internalSyringe-pumpStress-holdNet pressureHydrostatic
Figure 6.16 Variation of pressures during suction installation in the old system (NC clay, test B3SCI)
Nominal installation: 13.86 m
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)
PPT-internalSyringe-pumpStress-holdNet pressureHydrostatic
Figure 6.17 Variation of pressures during suction installation in the old system (NC clay, test B3SOI)
Nominal installation: 14.02 m
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)
PPT-internalStress-holdNet pressureHydrostatic
Figure 6.19 Variation of pressures during suction installation in the new system (NC clay, test B9SOI, DAQ data from syringe pump not received)
Nominal installation: 14.18 m
0
2
4
6
8
10
12
14
16
-50 0 50 100 150
Pressure (kPa)
Dep
th o
f tip
(m)
PPT-internalStress-holdNet pressureHydrostatic
Figure 6.20 Variation of pressures during suction installation in the new system (NC clay, test B10SOC, DAQ data for syringe pump not received)
Nominal installation: 13.78 m
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)
Nominal installation: 13.83 m
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)
Load cellPPT-internalSyringe-pumpHydrostaticNet pressure
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)
Load cellPPT-internalSyringe-pumpHydrostaticNet pressure
zfinal=14.30 m
Nominal installation: 14.11 m
Figure 6.23 Variation of pressures during suction installation in the new system (NC clay, test B11SOC)
Nominal installation: 14.12 m
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)
Load cellPPT-internalSyringe-pumpHydrostaticNet pressure
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
Penetration resistance (kPa)
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
Penetration resistance (kPa)
Dep
th o
f tip
(m)
B3SOIB3SCIB3SCCB9SOIB10SCIB10SOCB10SCCB11SOCB12SCCNc=7.5, alpha=0.39B2JOIB2JOCB4JCIB6JOCB6JCCB8JOC
Self-weight penetration
Suction installation
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
Penetration resistance (kPa)
Dep
th o
f tip
(m)
B3SOI
B3SCI
B3SCC
B10SCI
B11SOC
Self-weight penetration
Suction installation
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
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0 50 100 150
Necessary underpressure (kPa)
Dep
th o
f tip
(m)
02468
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0.0 0.5 1.0 1.5Predicted soil heave (m)
Dep
th o
f tip
(m)
02468
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0 100 200 300 400
Allowable underpressure (kPa)
Dep
th o
f tip
(m)
AllowableNecessaryApplied
02468
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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
Penetration resistance (kPa)
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
Self-weight penetration
Suction installation
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
Penetration resistance (kPa)
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*
Self-weight penetration
Suction installation
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
(c) Fs - depth (d) hs,pre - depth
02468
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0 1 2 3 4 5Factor of safety
Dep
th o
f tip
(m)
02468
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0 50 100Necessary underpressure (kPa)
Dep
th o
f tip
(m)
02468
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0.0 0.5 1.0 1.5Predicted soil heave (m)
Dep
th o
f tip
(m)
02468
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0 200 400 600Allowable underpressure (kPa)
Dep
th o
f tip
(m)
AllowableNecessaryApplied
11.73 m, plug reaches lid
0.98 m
11.73 m, plug reaches lid
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
Nominal installation: 13.89 m
02468
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0 50 100 150 200
Penetration resistance (kPa)
Dep
th o
f tip
(m)
Measured
Install:Nc=7.5,Alpha=0.42
Nominal installation:
13.89 m
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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
Nominal installation: 13.87 m
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
Nominal installation: 13.64 m
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
Nominal installation: 13.50 m
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)
Start of self-weight penetration
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
)
Start of self-weight penetration
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
(c) Fs - depth (d) hs,pre - depth
02468
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0 1 2 3 4 5Factor of safety
Dep
th o
f tip
(m)
02468
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0.0 0.5 1.0 1.5
Predicted soil heave (m)
Dep
th o
f tip
(m)
02468
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0 200 400 600Allowable underpressure (kPa)
Dep
th o
f tip
(m)
Allowable
Necessary
Applied
02468
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0 50 100 150
Necessary underpressure (kPa)
Dep
th o
f tip
(m)
13.62 m, plug reaches lid
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)
(a) ∆un - depth (b) ∆ua and ∆uapp- depth
(c) Fs - depth (d) hs,pre - depth
02468
10121416
0 1 2 3 4 5Factor of safety
Dep
th o
f tip
(m)
02468
10121416
0 50 100 150Necessary 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
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0 100 200 300 400Allowable underpressure (kPa)
Dep
th o
f tip
(m)
Allowable
Necessary
Applied
13.52 m, plug reaches lid
13.52 m, plug reaches lid, Fs,plug = 3.33
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
(c) Fs - depth (d) hs,pre - depth
02468
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0 1 2 3 4 5Factor of safety
Dep
th o
f tip
(m)
02468
10121416
0 50 100 150Necessary underpressure (kPa)
Dep
th o
f tip
(m)
02468
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0.0 0.5 1.0 1.5Predicted soil heave (m)
Dep
th o
f tip
(m)
02468
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0 200 400 600Allowable underpressure (kPa)
Dep
th o
f tip
(m)
Allowable
Necessary
Applied
12.12 m, plug reaches lid
12.12 m, plug reaches lid, Fs,plug = 2.15
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
Undrained shear strength su (kPa)
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
Nominal installation: 14.40 m
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)
Internal PPTLoad cellSyringe-pumpHydrostaticNet resistance
Nominal installation: 14.33 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
Nominal installation: 14.21 m
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)
Internal PPTLoad cellSyringe-pumpHydrostaticNet resistance
Nominal installation: 14.28 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)
Start of self-weight penetration
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)
0123456789
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
Start of self-weight penetration
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)
0123456789
10111213141516
0 100 200 300 400 500 600Time (s)
Dep
th o
f tip
(m)
Start of self-weight penetration
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)
Start of self-weight penetration
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)
(a) ∆un - depth (b) ∆ua and ∆uapp- depth
(c) Fs - depth (d) hs,pre - depth
02468
10121416
0 1 2 3 4 5Factor of safety
Dep
th o
f tip
(m)
02468
10121416
0 20 40 60 80Necessary underpressure (kPa)
Dep
th o
f tip
(m)
02468
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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
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0 100 200 300 400Allowable underpressure (kPa)
Dep
th o
f tip
(m)
AllowableNecessaryApplied
13.62 m, plug reaches lid
13.62 m, plug reaches lid, Fs,plug = 2.69
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)
(a) ∆un - depth (b) ∆ua and ∆uapp- depth
(c) Fs - depth (d) hs,pre - depth
02468
10121416
0 1 2 3 4 5 6 7 8Factor of safety
Dep
th o
f tip
(m)
02468
10121416
0 20 40 60Necessary 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 300Allowable underpressure (kPa)
Dep
th o
f tip
(m)
AllowableNecessaryApplied
13.63 m, plug reaches lid
13.63 m, plug reaches lid, Fs,plug = 5.14
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)
(a) ∆un - depth (b) ∆ua and ∆uapp- depth
(c) Fs - depth (d) hs,pre - depth
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
0 200 400 600Allowable underpressure (kPa)
Dep
th o
f tip
(m)
AllowableNecessaryApplied
13.44 m, plug reaches lid
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)
(a) ∆un - depth (b) ∆ua and ∆uapp- depth
(c) Fs - depth (d) hs,pre - depth
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)
AllowableNecessaryApplied
13.45 m, plug reaches lid
13.45 m, plug reaches lid, Fs,plug = 3.17
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)
Suction installation
Self-weight penetration
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
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 6.62 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
(a) Axial pressure (b) Internal pore pressure
Figure 6.63 Installation and pullout pressure and internal pore pressure changes
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
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-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
(a) Axial pressure (b) Internal pore pressure
Figure 6.65 Installation and pullout pressure and internal pore pressure changes
during unsealed pullout (NC clay, test B2JOC)
02468
10121416
-300 -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)D
epth
of t
ip (m
)
PPTI
Hydrostatic
(a) Axial pressure (b) Internal pore pressure
Figure 6.66 Installation and pullout pressure and internal pore pressure changes
during unsealed pullout (NC clay, test B6JOC)
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
(a) Axial pressure (b) Internal pore pressure
Figure 6.67 Installation and pullout pressure and internal pore pressure changes
during unsealed pullout (NC clay, test B8JOC)
(a) Axial pressure (b) Internal pore pressure
Figure 6.68 Installation and pullout pressure and internal pore pressure changes during unsealed pullout (NC clay, test B11SOC)
02468
10121416
-300 -200 -100 0 100 200Axial pressure, ∆p (kPa)
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
(a) Axial pressure (b) Internal pore pressure
Figure 6.69 Installation and pullout pressure and internal pore pressure changes
during unsealed pullout (NC clay, test B12SOC)
∆p=P/A
02468
10121416
-300 -200 -100 0 100 200
Axial pressure, ∆p (kPa)D
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
Axial pressure, ∆p (kPa)
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
Undrained shear strength su (kPa)
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
Axial pressure, ∆p (kPa)
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
Axial pressure, ∆p (kPa)
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
(a) Axial pressure (b) Internal pore pressure
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
Axial pressure, ∆p (kPa)
Dep
th o
f tip
(m)
B12SCC
02468
10121416
-150 -50 50 150
Pore pressure (kPa)
Dep
th o
f tip
(m)
PPTI
Hydrostatic
(a) Axial pressure (b) Internal pore pressure
Figure 6.78 Installation and pullout pressure and internal pore pressure during sealed pullout after consolidation in NC clay (test B12SCC)
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
-300 -200 -100 0 100 200Axial pressure, ∆p (kPa)
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
(a) Axial pressure (b) Internal pore pressure
Figure 6.80 Installation and pullout pressure and internal pore pressure during sealed pullout after consolidation in NC clay (test B3SCC)
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 =1.08 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
Axial pressure, ∆p (kPa)D
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
Undrained shear strength su (kPa)
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
Pullout Installation
0
2
4
6
8
10
12
14
16
-800 -600 -400 -200 0 200 400 600
Axial pressure, ∆p (kPa)
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
Pullout Installation
0
2
4
6
8
10
12
14
16
-1000 -800 -600 -400 -200 0 200 400 600
Axial pressure, ∆p (kPa)
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
(a) Axial pressure (b) Internal pore pressure
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
Undrained shear strength su (kPa)
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
Axial pressure, ∆p (kPa)
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
Undrained shear strength su (kPa)
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
Axial pressure, ∆p (kPa)D
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
Axial pressure, ∆p (kPa)
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
(a) Axial pressure (b) Internal pore pressure
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
Axial pressure, ∆p (kPa)
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
(a) Axial pressure (b) Internal pore pressure
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
Axial pressure, ∆p (kPa)
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
Axial pressure, ∆p (kPa)
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
Axial pressure, ∆p (kPa)D
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
Radial total stress (kPa)D
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
Figure 7.4 Variations of measured external radial total stress σri with depth
during jacked installation in NC clay (test B6JOI)
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
u0 σ ri+∆ui
Figure 7.5 Variations of measured external radial total stress σri with depth
during jacked installation in NC clay (test B8JOI)
0
2
4
6
8
10
12
14
16
0 50 100 150 200Radial total stress (kPa)
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
Radial total stress (kPa)
Dep
th o
f tip
(m)
Radial total stress
HydrostaticIn water
In soil
u0 σ ri+∆ui
σri
Figure 7.7 Variations of measured external radial total stress σri with depth
during suction installation in NC clay (test B3SCC)
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
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
Radial total stress (kPa)D
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
Radial total stress (kPa)
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
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
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
Radial total stress (kPa)
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
Radial total stress (kPa)D
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
Radial total stress (kPa)
Dep
th o
f tip
(m)
Radial total stress
HydrostaticIn water
In soil
u0 σ ri+∆ui
Figure 7.14 Variations of measured external radial total stress σri with depth
during suction installation in NC clay (test B12sus)
0
2
4
6
8
10
12
14
16
0 50 100 150 200Radial total stress (kPa)
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
0 50 100 150 200Radial total stress (kPa)
Dep
th o
f tip
(m)
Suction installation
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
0 50 100 150 200Radial total stress (kPa)
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
0 50 100 150 200Radial total stress (kPa)
dept
h of
tip(
m)
TPT1TPT2AverageHydrostatic
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
0 50 100 150 200Radial total stress (kPa)
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
0 50 100 150 200Radial total stress (kPa)
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
165 170 175 180 185Radial total stress (kPa)
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
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0 50 100 150 200Radial total stress (kPa)
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
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0 50 100 150 200 250
Radial total stress (kPa)D
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)
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-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)
Radial total stress (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
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0 50 100 150 200 250
Radial total stress (kPa)D
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)
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-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
Time, in model scale (s)
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=13.59 - 13.72 m, v=0.20 mm/s, V=1.3: partly drained
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
z=7.52 - 7.97 m, v=0.55 mm/s, V=3.6: partly drained
Note: v in model scale
Figure 7.43 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 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
suction-affected area: >0.82 m
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)
NGI (if no transition zone)
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)
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 =7.21 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
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0 50 100 150 200 250
Radial total stress (kPa)D
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
Syringe pump pressure
z = 7.69 - 7.85 m, jacking ends and suction starts
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)
NGI (if no transition zone)
zTPT = 7.69 m, TPTS leave jacking-affected area, v slows to 0.21 mm/s (model scale)
zTPT = 7.85 m, σri−u0
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)
NGI (if no transition zone)
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
0 50 100 150 200Radial total stress (kPa)
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
Radial total stress (kPa)D
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
Radial total stress (kPa)
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)
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10111213141516
0 50 100 150 200 250Radial total stress (kPa)
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)
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10111213141516
-400 -300 -200 -100 0 100 200 300
Pressure (kPa)
Dep
th o
f tip
(m)
Radial total stress
Syringe pump pressure
Slight change in gradient of radial stress when suction starts
z=12.25 m, TPTs enters the suction-affected area in soil, a small decrease exists
z=14.20 m, σri decreases and pump pressure surges due to moving speed reduces
z = 7.21 - 7.45 m, jacking ends and suction starts
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
Time, in model scale (s)
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)
Radial total stress (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
Figure 7.69 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 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)
NGI (if no transition zone)
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 = 7.21 m, TPTs leave jacking-affected area
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)
MeasuredMTD(Jardine & Chow 1996)SPM(Whittle & Baligh 1988)CEM (lower bound)CEM (upper bound)NGI (if no transition zone)
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
Radial total stress (kPa)D
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
Syringe pump pressure
z=14.15 m, σri reduces due to moving speed reduces to 0.02 mm/s
z = 6.34 - 6.74 m, jacking ends and suction starts
Slight change in gradient of radial stress when suction starts
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
Time, in model scale (s)
Oup
tput
in in
stru
men
ts
Embedment (z) of caisson in model scale (mm)
Embedment (z) of caisson in prototype scale (m)
Syringe pump pressure (kPa)
Radial total stress (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
Figure 7.75 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 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
suction-affected area: >2.66 m
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)
NGI (if no transition zone)
zTPT = 6.34 m, TPTs leave jacking-affected area
6.7 kPa: due to reduced speed
zTPT = 6.74 m, σri−u0
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)
MeasuredMTD(Jardine & Chow 1996)SPM(Whittle & Baligh 1988)CEM (lower bound)CEM (upper bound)NGI (if no transition zone)
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
Radial total stress (kPa)D
epth
of t
ip (m
)
TPT1TPT2AverageHydrostatic
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
Syringe pump pressure
Slight change in gradient of radial stress when suction starts
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
z = 8.90-9.12 m, jacking ends and suction starts
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
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 (kPa)
Radial total stress (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
Figure 7.81 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 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
suction-affected area: >0.65 m
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)
MeasuredMTD(Jardine & Chow 1996)SPM(Whittle & Baligh 1988)CEM (lower bound)CEM (upper bound)NGI (if no transition zone)
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
Radial total stress (kPa)
Dep
th o
f tip
(m)
TPT1TPT2HydrostaticAverage
z = 9.27-9.40 m, jacking ends and suction starts
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=14.20 m, TPTs enters the suction-affected area in soil, a small decrease exists
z=14.45 m, σri reduces due to moving speed drops to 0.04 mm/s
z = 9.27 - 9.40 m, jacking ends and suction starts
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
Time, in model scale (s)
Oup
tput
in in
stru
men
ts
Embedment (z) of caisson in model scale (mm)
Embedment (z) of caisson in prototype scale (m)
Syringe-pump (kPa)
Radial total stress (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
Figure 7.86 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 sensitive clay (test B14sus)
0123456789
10
0 50 100 150 200
σri−u0 (kPa)D
epth
of T
PT (m
)
MeasuredMTD(Jardine & Chow 1996)SPM(Whittle & Baligh 1988)CEM (lower bound)CEM (upper bound)NGI (if no transition zone)
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
Radial total stress (kPa)
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
Radial total stress (kPa)D
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)
(a) Axial pressure (b) Internal pore pressure
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
Axial pressure, ∆p (kPa)
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
Undrained shear strength su (kPa)
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
Radial total stress (kPa)
Dep
th o
f tip
(m)
TPT1TPT2AverageHydrostatic
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)
(a) Axial pressure (b) Internal pore pressure
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
Axial pressure, ∆p (kPa)
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
Internal pore pressure (kPa)
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
Radial total stress (kPa)
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
Radial total stress (kPa)D
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)
(a) Axial pressure (b) Internal pore pressure
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
Axial pressure, ∆p (kPa)
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
Internal pore pressure (kPa)
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
Prototype time from start of sustained loading (day)
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
Radial total stress (kPa)
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
Radial total stress (kPa)D
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)
(a) Axial pressure (b) Internal pore pressure
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
Prototype time from start of sustained loading (day)
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
Radial total stress (kPa)
Dep
th o
f tip
(m)
TPT1TPT2AverageHydrostatic
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
Prototype time from start of sustained loading (day)
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
Axial pressure, ∆p (kPa)D
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
Undrained shear strength su (kPa)
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
Radial total stress (kPa)D
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
Prototype time from start of cyclic loading (day)
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
Axial pressure, ∆p (kPa)
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
Prototype time from start of cyclic loading (day)
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
Radial total stress (kPa)
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
Axial pressure, ∆p (kPa)D
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
Undrained shear strength su (kPa)
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
Prototype time from start of cyclic loading (day)
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
Radial total stress (kPa)D
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
Prototype time from start of cyclic loading (day)
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)