INSTALLATION AND KEYING OF FOLLOWER EMBEDDED PLATE...
Transcript of INSTALLATION AND KEYING OF FOLLOWER EMBEDDED PLATE...
School of Civil and Resource Engineering
The University of Western Australia
INSTALLATION AND KEYING OF FOLLOWER EMBEDDED
PLATE ANCHORS
by
ADAM CHARLES LOWMASS
A thesis submitted for the degree of
MASTERS OF ENGINEERING SCIENCE
at
THE UNIVERSITY OF WESTERN AUSTRALIA
SCHOOL OF CIVIL AND RESOURCE ENGINEERING
2006
School of Civil and Resource Engineering
The University of Western Australia
ABSTRACT
The offshore oil and gas industry is moving into deeper water to meet the growing global
demand for hydrocarbons. Associated with this move to deep water is the need for more efficient
anchorage systems to moor floating facilities. Of the anchor concepts proposed in recent years,
the most promising utilise a follower to embed an initially vertical plate anchor, typically located
at the follower base. When the system has reached the design embedment depth, the plate anchor
mooring line is disengaged from the follower, leaving the follower free to be re-used for the next
installation. The mooring line attached to the vertically embedded plate anchor is tensioned
causing the plate anchor to rotate or ‘key’ to an orientation that is perpendicular to the direction
of loading. The offshore industry currently considers this keying process to be the main unknown in
relation to follower-embedded anchors.
This project contributes to the limited database of the behaviour of anchors during keying, in
particular quantifying the effects of eccentricity of loading from the plate on the vertical
displacement of the plate anchors during the keying process. Reduced scale model centrifuge
testing is used to facilitate the optical measurement of the rotation and displacement of the
various geometries of plate anchors through a soil/Perspex interface during keying. Additionally
the project has explored the effects of installation on the keying characteristics of Suction
Embedded Plate Anchors (SEPLA).
As well as the physical modelling aspect of the project, a numerical model is developed to
simulate the keying process, allowing accurate prediction of final embedment depth and anchor
orientation, and ultimately anchor load capacity.
This study has significantly enhanced the understanding of the keying process. In terms of the
practical application of embedded plates as anchors for floating offshore facilities, the influence
of padeye eccentricity ratio (e/B) on normalised embedment loss (∆ze/B) resulting from keying is
possibly the most important finding of the study. It indicates that current guidelines, stating
embedment loss during keying is twice the anchor height (B) in cohesive soils, are extremely
conservative given typically padeye eccentricities (e/B < 0.5). These results have indicated that
for typical embedded plate anchors the embedment loss is < 0.3B.
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ACKNOWLEDGMENTS
I would firstly like to thank Conleth O’Loughlin who has gone out of his way over the past two
years to ensure I had every possible advantage in completing this Masters paper. Even though his
schedule was always busy, he made himself available to answer any questions I had and to offer,
much appreciated, guidance. Mark Randolph and Christophe Gaudin also provided notable
guidance during the course of the last two years.
Throughout the project, Don Herley and Bart Thompson provided much assistance, not only
during testing but also in other areas of the project. They provided a relaxed, yet productive work
environment that made the long hours in the laboratory welcome. The workshop staff, namely
Gary Davies, Neil McIntosh, Alby Kalajzich, David Jones, Frank Tan, John Breen, Shane De
Catania and Wayne Galbraith all contributed considerably with last minute modifications to
testing apparatus and the manufacture of testing apparatus and models. Wenge Liu also provided
great IT support.
Thanks must also go to, PhD candidate and friend, Mark Richardson who assisted me during the
busy times. He helped in completing centrifuge tests, as well as providing advice throughout the
project. Mark always made himself available and was happy for me to approach him at all times.
Friends and family have also assisted me greatly, providing the much need emotional support.
Special thanks must go to my cousin Mark Norwell who continually motivated and encouraged
me to complete the project.
Last, and probably most importantly, I would like to thank my parents, Anna and Peter. Both
provided the much need support during the stressful times of the project. They always made
themselves available to help and offer advice whenever necessary.
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TABLE OF CONTENTS
Chapter 1 Introduction ......................................................................................................... 1-1
1.1 Outline...................................................................................................................... 1-1
1.2 Mooring Systems ...................................................................................................... 1-2
1.3 Anchoring Options.................................................................................................... 1-3
1.3.1 Gravity Anchors ....................................................................................... 1-3
1.3.2 Anchor Piles ............................................................................................. 1-3
1.3.3 Drag Embedment Anchors........................................................................ 1-4
1.3.4 Suction Caisson ........................................................................................ 1-4
1.3.5 Torpedo and Deep Penetrating Anchor (DPA) .......................................... 1-5
1.3.6 Follower Embedded Anchors.................................................................... 1-6
1.3.7 Anchor Option Comparison ...................................................................... 1-6
1.4 Research Objectives.................................................................................................. 1-8
1.5 Thesis Structure ........................................................................................................ 1-8
Chapter 2 Plate Anchor and Keying Background ............................................................... 2-1
2.1 Previous Studies........................................................................................................ 2-1
2.1.1 SEPLA ..................................................................................................... 2-1
2.1.2 Keying...................................................................................................... 2-1
2.1.3 Plate Capacity........................................................................................... 2-2
Chapter 3 Model Plate Anchor Testing ................................................................................ 3-1
3.1 Centrifuge Testing Apparatus.................................................................................... 3-1
3.1.1 Principles of Centrifuge Testing................................................................ 3-1
3.1.2 The Geotechnical Beam Centrifuge .......................................................... 3-2
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3.1.3 The Geotechnical Drum Centrifuge .......................................................... 3-4
3.2 Kaolin Clay............................................................................................................... 3-5
3.2.1 Soil Properties .......................................................................................... 3-5
3.2.2 Sample Preparation................................................................................... 3-6
3.3 Plate Anchor Keying Test Procedure......................................................................... 3-7
3.3.1 Soil Characterisation Tests........................................................................ 3-7
3.3.2 SEPLA Beam Centrifuge Tests................................................................. 3-8
3.3.3 Plate Keying Drum Centrifuge Tests........................................................3-11
3.3.4 Plate Keying Beam Centrifuge Tests........................................................3-14
Chapter 4 Experimental Results ........................................................................................... 4-1
4.1 SEPLA Tests ............................................................................................................ 4-1
4.1.1 Soil Characterisation Tests........................................................................ 4-1
4.1.2 SEPLA Capacities .................................................................................... 4-2
4.2 Keying Tests ............................................................................................................. 4-4
4.2.1 Soil Characterisation Tests........................................................................ 4-5
4.2.2 Drum Keying Tests................................................................................... 4-7
4.2.3 Beam Keying Tests................................................................................... 4-8
4.2.4 Keying Test Summary .............................................................................4-10
Chapter 5 Analytical Simulation .......................................................................................... 5-1
5.1 Background............................................................................................................... 5-1
5.2 Review of Numerical and Analytical Studies of Plate Anchors.................................. 5-1
5.3 Plasticity Concepts and the Yield Locus.................................................................... 5-2
5.4 Kinematic Anchor Analysis ...................................................................................... 5-3
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5.5 Results ...................................................................................................................... 5-4
Chapter 6 Discussion of Theoretical and Experimental Results.......................................... 6-1
6.1 SEPLA Tests ............................................................................................................ 6-1
6.2 Plate Keying ............................................................................................................. 6-2
6.2.1 Capacity ................................................................................................... 6-2
6.2.2 Keying...................................................................................................... 6-4
6.2.3 Comparison with Analytical Simulation.................................................... 6-7
Chapter 7 Conclusion and Further Research....................................................................... 7-1
7.1 Experimental Findings .............................................................................................. 7-1
7.1.1 SEPLA Testing......................................................................................... 7-1
7.1.2 Plate Keying Tests .................................................................................... 7-1
7.2 Recommendations for Future Development............................................................... 7-2
7.3 Concluding Statement ............................................................................................... 7-3
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NOMENCLATURE
A plate area
B plate height
bf depth
Cc compression index
Cs swelling index
cu local undrained shear
cv coefficient of consolidation
d diameter
dv padeye embedment loss
e loading eccentricity
Fmax peak load
g gravitational acceleration
Gs specific gravity
H initial plate embedment / horizontal loading
k shear strength gradient
L length
Le effective anchor length
Lf footing length
LL liquid limit
M mass / moment loading
N gravity scale factor
Nb non-dimensional T-bar factor
Nc, Ncy, Ncp non-dimensional breakout factor
P force per unit length
PL plastic limit
Qu ultimate uplift capacity
sc chain displacement
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su local undrained shear
T and t time
Ta padeye load
Tv dimensionless time factor
v pullout rate
V normalised velocity / vertical loading
Wa submerged anchor weight
z depth
∆ze anchor embedment loss
Greek
α soil adhesion factor;
β plate inclination to the vertical
γ weight of soil
θ plate inclination to the horizontal
θa inclination of the chain angle at anchor padeye
σv’ effective vertical stress
φ friction angle
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LIST OF FIGURES
Figure 1.1: Offshore production facilities (Courtesy: Minerals Management Services) ...........1-10
Figure 1.2: Offshore production facilities (Courtesy: Minerals Management Services) ...........1-10
Figure 1.3: Mooring types (Vryhof, 2000) ..............................................................................1-11
Figure 1.4: Drag anchors, sand and clay (Vryhof, 2000) .........................................................1-11
Figure 1.5: Suction caisson.....................................................................................................1-12
Figure 1.6: DPA and Torpedo anchor and installation schematic (Lieng et al. 1999 and
Medeiros, 2001) .................................................................................................1-12
Figure 1.7: SEPLA installation technique ...............................................................................1-13
Figure 1.8: SEA installation technique....................................................................................1-13
Figure 2.1: The SEPLA concept: � Suction installation, � Caisson retrieval, � Anchor
keying, � Mobilised anchor ................................................................................ 2-4
Figure 2.2: Failure mechanism of a plate (Merifield, 2002)...................................................... 2-4
Figure 2.3: Effect of overburden pressure on strip anchors (Merifield, 2002) ........................... 2-5
Figure 2.4: Nc comparison for Breakaway cases in weightless soil (Merifield, 2002) ............... 2-5
Figure 2.5: Upper Bound Nc values for horizontal anchors - inhomogeneous cohesive soil
(Merifield, 2002) ................................................................................................. 2-6
Figure 3.1: Geotechnical beam centrifuge...............................................................................3-17
Figure 3.2: Beam strongbox....................................................................................................3-17
Figure 3.3: Motor driven actuator ...........................................................................................3-18
Figure 3.4: T-bar penetrometer ...............................................................................................3-18
Figure 3.5: Drum centrifuge with clamshell removed..............................................................3-19
Figure 3.6: Keying test setup in sample box, @ 1 g ................................................................3-19
Figure 3.7: Digital camera cradle with trigger.........................................................................3-20
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Figure 3.8: Tool table actuator................................................................................................3-20
Figure 3.9: Sub-Terrain Oil impregnated Multiple Pressure Instrument (STOMPI) .................3-21
Figure 3.10: Mounted model SEPLA......................................................................................3-21
Figure 3.11: SEPLA test setup................................................................................................3-22
Figure 3.12: Experimental arrangement and test procedure (section view) ..............................3-23
Figure 3.13: Caisson with attachments (pneumatic valve hidden behind caisson guide) ..........3-24
Figure 3.14: SEPLA test, actuator setup..................................................................................3-24
Figure 3.15: Model plate anchors with various loading shafts attached (including two not
used in this study)...............................................................................................3-25
Figure 3.16: LCD attached to sample box ...............................................................................3-25
Figure 3.17: Drum keying test, loading arm............................................................................3-26
Figure 3.18: Drum keying test layout (White, 2003) ...............................................................3-26
Figure 3.19: Sample box held in place with brackets...............................................................3-27
Figure 3.20: Beam keying test configuration...........................................................................3-27
Figure 4.1: Clay, shear strength profile, SEPLA tests..............................................................4-12
Figure 4.2: Assumed load response during anchor keying and pullout for Test VE-ST1..........4-12
Figure 4.3: Dimensionless load displacement response for jacked SEPLA..............................4-13
Figure 4.4: Dimensionless load displacement response for suction embedded SEPLA ............4-13
Figure 4.5: Loss of embedment as a function of padeye load inclination), e/B = 0.66..............4-14
Figure 4.6: Keying test load orientations.................................................................................4-14
Figure 4.7: Clay, shear strength profile, drum tests box 2........................................................4-15
Figure 4.8: Clay, shear strength profile, drum tests box 3........................................................4-15
Figure 4.9: Clay, shear strength profile, drum tests box 4........................................................4-16
Figure 4.10: Clay, shear strength profile, drum tests box 5......................................................4-16
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Figure 4.11: Clay, shear strength profile, drum tests box 6......................................................4-17
Figure 4.12: Drum tests, average shear strength profiles .........................................................4-17
Figure 4.13: Clay, shear strength profile, beam tests box 1 .....................................................4-18
Figure 4.14: Clay, shear strength profile, beam tests box 2 .....................................................4-18
Figure 4.15: Clay, shear strength profile, beam tests box 3 .....................................................4-19
Figure 4.16: Clay, shear strength profile, beam tests box 4 .....................................................4-19
Figure 4.17: Clay, shear strength profile, beam tests box 5 .....................................................4-20
Figure 4.18: Clay, shear strength profile, beam tests box 6 .....................................................4-20
Figure 4.19: Beam tests, average shear strength profiles .........................................................4-21
Figure 4.20: Stages of keying, drum test e/B = 0.17 ................................................................4-21
Figure 4.21: Stages of keying, drum test e/B = 0.5 ..................................................................4-22
Figure 4.22: Stages of keying, drum test e/B = 1.0 ..................................................................4-22
Figure 4.23: Plate anchor rotation for drum tests, L = 80mm anchors & vertically loaded .......4-23
Figure 4.24: Plate anchor rotation for drum tests, L = 30mm anchors & vertically loaded .......4-23
Figure 4.25: Stages of keying, drum test b3a30e15 .................................................................4-24
Figure 4.26: Plate anchor rotation for beam tests, e/B = 0.25...................................................4-24
Figure 4.27: Plate anchor rotation for beam tests, e/B = 0.5.....................................................4-25
Figure 4.28: Plate anchor rotation for beam tests, e/B = 0.75...................................................4-25
Figure 4.29: Plate anchor rotation for beam tests, e/B = 1 .......................................................4-26
Figure 4.30: Plate anchor rotation for beam tests, e/B = 1.5.....................................................4-26
Figure 4.31: Nc vs. loss of embedment, e/B = 0.25 ..................................................................4-27
Figure 4.32: Nc vs. loss of embedment, e/B = 0.5 ....................................................................4-27
Figure 4.33: Nc vs. loss of embedment, e/B = 0.75 .................................................................4-28
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Figure 4.34: Nc vs. loss of embedment, e/B = 1.......................................................................4-28
Figure 4.35: Nc vs. loss of embedment, e/B = 1.5 ....................................................................4-29
Figure 4.36: Plate inclination vs. Nc, e/B = 0.25 ......................................................................4-29
Figure 4.37: Plate inclination vs. Nc, e/B = 0.5 ........................................................................4-30
Figure 4.38: Plate inclination vs. Nc, e/B = 0.75 ......................................................................4-30
Figure 4.39: Plate inclination vs. Nc, e/B = 1...........................................................................4-31
Figure 4.40: Plate inclination vs. Nc, e/B = 1.5 ........................................................................4-31
Figure 4.41: Loss in plate anchor embedment during keying...................................................4-32
Figure 5.1: The yield locus and plasticity potential function (Bransby and O'Neill, 1999)........ 5-6
Figure 5.2: V-H-M yield locus for rectangular fluke (Bransby and O'Neill, 1999) ................... 5-6
Figure 5.3: Kinematic analysis sign convention ....................................................................... 5-7
Figure 5.4: Analysis flowchart for kinematic anchor simulation using yield locus.................... 5-7
Figure 5.5: Loss in plate anchor embedment during keying – analytical simulation.................. 5-8
Figure 5.6: Normalised embedment loss vs. normalised load -analytical simulation................. 5-8
Figure 5.7: Angle of inclination vs. normalised embedment loss - analytical simulation .......... 5-9
Figure 5.8: Plate inclination vs. normalised load – analytical simulation.................................. 5-9
Figure 6.1: Test Nc comparison with Merifield et al. (2003)..................................................... 6-9
Figure 6.2: Keying analysis, e/B = 0.25 ................................................................................... 6-9
Figure 6.3: Keying analysis, e/B = 0.5 ....................................................................................6-10
Figure 6.4: Keying analysis e/B = 0.75 ...................................................................................6-11
Figure 6.5: Keying analysis, e/B = 1.0 ....................................................................................6-12
Figure 6.6: Keying analysis, e/B = 1.5 ....................................................................................6-12
Figure 6.7: Keying mechanisms..............................................................................................6-14
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Figure 6.8: Combined loading paths for high and low eccentricity plate anchors.....................6-14
Figure 6.9: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.25 ............6-15
Figure 6.10: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.5............6-15
Figure 6.11: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.75 ..........6-16
Figure 6.12: Plate inclination vs. Nc comparison with analytical simulation, e/B = 1.0............6-16
Figure 6.13: Plate inclination vs. Nc comparison with analytical simulation, e/B = 1.5............6-17
Figure 6.14: Plate anchor rotation comparison with analytical simulation, e/B = 0.25 .............6-17
Figure 6.15: Plate anchor rotation comparison with analytical simulation, e/B = 0.5 ...............6-18
Figure 6.16: Plate anchor rotation comparison with analytical simulation, e/B = 0.75 .............6-18
Figure 6.17: Plate anchor rotation comparison with analytical simulation, e/B = 1.0 ...............6-19
Figure 6.18: Plate anchor rotation comparison with analytical simulation, e/B = 1.5 ...............6-19
Figure 6.19: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.25 ......6-20
Figure 6.20: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.5 ........6-20
Figure 6.21: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.75 ......6-21
Figure 6.22: Nc vs. loss of embedment comparison with analytical simulation, e/B = 1.0 ........6-21
Figure 6.23: Nc vs. loss of embedment comparison with analytical simulation, e/B = 1.5 ........6-22
Figure 6.24: e/B vs. ∆ze/B .......................................................................................................6-22
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CHAPTER 1 INTRODUCTION
1.1 OUTLINE
Since the mid-1980s there has been rapid development of oil and gas reserves located in deep
and ultra-deepwater (> 500 m and > 1500 m respectively, Colliat, 2002). This has precipitated a
shift to floating facilities as the preferred method of extracting hydrocarbons. Bottom supported
structures such as steel jacket and gravity base structures are generally only economically viable
in waters less than 400 m deep (Aubeny et al., 2001). Although compliant towers provide an
option for water depths up to 750 m, recent developments in floating facilities have seen their
popularity diminish. Currently the most viable options for production facilities in waters greater
than 400 m include: Tension Leg Platforms (TLPs); SPAR Platforms; and Floating Production,
Storage and Offloading (FPSO) facilities. Figure 1.1 and Figure 1.2 show examples of these
different facilities.
Tension Leg Platforms (TLPs) consist of a semi submerged hollow structure, moored to the
seabed by vertical tendons. The structure’s excess buoyancy keeps the tendons taut even under
extreme storm loading conditions. TLPs have been used in deepwater fields up to 1500 m, since
the mid-1980s (Aubeny et al., 2001), and are recognised as the most stable of all deepwater
floating systems. The ability to disconnect moorings, allowing use at alternate sites, is a major
advantage of a semi submersible platform.
The SPAR platform has recently emerged as a popular deepwater alternative within the oil and
gas industry. It comprises a truncated cylinder, with soft tanks in the bottom and hard tanks in
the top that supports a platform by means of excess buoyancy controlled by ballast within the
tanks. The tanks within the cylinder can have the level of ballast varied to maintain the required
draft with a change in top load. SPAR moorings are usually taut mooring lines set at an angle but
vertical lines, similar to TLPs, are sometimes utilised. The advantage of having a cylinder is that
it allows the riser (the tubular sections used for drilling and hydrocarbon collection) to run down
a centre well, partially shielding them from the wave and current loads. SPARs are relatively
insensitive to deck loads, easy to transport and once moored are very stable (Aubeny et al.,
2001). SPARs in water depths of up to 1700 m currently exist, although the technology can
extend their use to water depths as great as 3000 m.
A Floating Production, Storage and Offloading (FPSO) facility is a tanker-based system moored
to the seafloor. It can consist of either converted tankers or new, specially designed vessels for
the specific purpose of handling hydrocarbons collected from nearby sub-sea wells. As the name
suggests FPSOs have the ability not only to produce but also to store and offload hydrocarbons.
The offloading function allows the FPSO to continue production without having to move from
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its mooring at a given reserve. Smaller shuttle tankers connected to the offloading hose can
transport the hydrocarbons produced. The ability to store the produced hydrocarbon in the hull of
the vessel, like normal tankers, allows FPSOs to operate without costly pipeline networks. The
FPSO is also highly mobile, allowing the facility to move from reserve to reserve when required.
Mooring lines attached to a turret, built as part of the tanker, is the common method for
anchoring FPSOs. The FPSO is then free to rotate about this turret allowing the vessel to
orientate itself relative to the wind and current so that the total environmental load on the facility
can be minimised. Currently FPSOs are highly popular within the oil and gas industry due to
their shorter development and implementation time, resulting in reduced time between the
discovery and production of a field.
1.2 MOORING SYSTEMS
During the past decade especially, the move to deeper waters has seen mooring systems change
from catenary moorings to taut leg moorings (Figure 1.3) and the introduction of synthetic
mooring lines. The major difference between the two systems is the angle at which the mooring
line meets the seabed. Catenary moorings arrive horizontally at the seabed resulting in high
lateral loads imposed on the anchor governing their design. Taut-leg mooring designs on the
other hand lead to much higher angles of inclination and hence the vertical holding capacity or
uplift resistance of the anchor system generally controls the anchor design.
Catenary moorings rely on the weight of the mooring chain to provide the majority of the
restoring force, and in deepwater the weight and horizontal spread of the mooring system both
become excessive. Taut-leg moorings reduce the length of mooring line required and the
elasticity (and pre-tensioning) of the mooring line provides the restoring force. The major
advantage of taut-leg mooring systems over catenary moorings is they have a smaller footprint
i.e. the mooring radius of the taut-leg mooring will be smaller than a catenary for a similar
application (Vryhof, 2000). A further advantage of the taut leg moorings is that they are better
for load sharing between adjacent lines than catenary moorings, therefore providing a more
efficient system. Taut leg moorings also allow for better control of floating facilities motion as
the lines have sufficient elasticity to absorb the vessels wave motions without over loading. In
very deepwater, however, the weight of steel mooring lines may become too large for the
facility’s payload. Furthermore, efficiency of the lines reduces as more of the line tension
capacity will be used in keeping the wire taut.
Synthetic mooring lines provide a competitive alternative to steel moorings, as they are not only
much lighter but they also resist corrosion in the highly corrosive environment of the ocean.
Their real potential lies in taut moorings where mooring lines combining low weight and low
elastic modulus, with good durability characteristics, enable efficient mooring systems for a
whole range of water depth and environmental conditions to be developed (TTI Ltd., 2003).
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1.3 ANCHORING OPTIONS
As the operational water depth of offshore facilities increases, the cost of anchoring the facilities
used for the production, storage and offloading of hydrocarbons increases exponentially. The
move towards synthetic, taut-leg, mooring systems has necessitated development of reliable
anchors capable of withstanding the large vertical forces. In addition, with production reaching
record depths and 40 deepwater projects scheduled to come on-stream between now and 2008
(DeLuca 2005) the need for an economical anchoring method for these floating facilities has
become paramount.
1.3.1 Gravity Anchors
Gravity anchors are the simplest type of anchor and the easiest to design and construct. Their
holding capacity is a combination of self-weight and the friction between the base of the anchor
and the seabed. Also called deadweight anchors, they are generally a hollow box located on, or
in, the seabed filled with high-density material, such as rock or iron ore. This anchoring method,
however, has a limited practical size and therefore holding capacity, resulting in them being
limited to relatively shallow waters. They are also relatively inefficient for their size under
tension loads, compared to other anchoring methods.
1.3.2 Anchor Piles
Piles are the most common form of anchor or foundation system used in shallower waters. They
are hollow steel pipes embedded into the seabed by piling hammers or vibrators depending on
the nature of the seabed. The pile’s holding capacity is mainly generated through friction along
the pile/soil interface, thus to achieve sufficient holding capacity they must be driven deep into
the seabed. Drilled and grouted piles are another method of piling, involving drilling holes into
the seabed, inserting steel piles and forcing grout into the space in and around the pile. Grouted
driven piles are piles driven into the seabed by a hammer with grout forced out of holes in the
pile wall. Grouted piles increase the frictional area hence mobilising a higher capacity. However,
the degree of additional capacity is difficulty to quantify.
The ability to accurately position piles and the existence of well-established methods for
assessing their horizontal and vertical capacities are major advantages that this anchorage
method affords. Although piles can be driven in waters up to 2500 m deep
(http://www.menck.com/, 2006), they are a comparatively expensive method of anchorage at
these depths.
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1.3.3 Drag Embedment Anchors
Drag anchors used in the offshore industry consist of a fluke (providing the primary resistance)
rigidly attached to a shank, which connects the fluke to the mooring line. The fluke/shank angle
is set to optimise performance in the particular soil type (Figure 1.4), with larger angles used in
soft soils (clays) and smaller angles used in stronger, dilatant soils (sands). Irrespective of the
fluke/shank angle, installation of conventional drag anchors, as their name suggests, involves
dragging them through the soil to obtain a desired embedment; in soft clays, this may be several
times the length of the anchor fluke. The drag distance to obtain the desired level of embedment
depends on the soil conditions and can be a few hundred metres, which may create design
problems in terms of planning site investigation, or ensuring appropriate mooring chain lengths.
In addition, their trajectory during embedment is difficult to determine, resulting in their final
embedment, and hence final capacity, being difficult to assess.
Until recently, drag anchors have not been able to withstand the vertical loads encountered
during deepwater anchoring. This led to the development of specific designs of drag anchors to
withstand vertical loading, and a resulting class of Vertically Loaded Anchors (VLAs). Two
recent designs, the Stevmanta, developed in 1996 by Vryhof Anchors, and the Near Normal
Load Anchor (DENNLA), developed by Bruce Anchors can sustain significant vertical loading.
Both Bruce and Vryhof have developed tracking devices for their respective anchors in an
attempt to solve the problem associated with determining an anchor’s final embedment.
However, the performance of the tracking devices is by no means perfect (Ehlers et al., 2004)
and does not overcome the challenges with the drag method of installation. Bruce is currently
working on a real time tracking system that would be capable of displaying the anchor’s
trajectory during installation giving the final anchor embedment with more accuracy and
reliability.
The advantage of drag embedment anchors is the relative ease of recovery allowing reuse in
other moorings, making them highly suited to short to medium term projects. Additionally, they
are a very efficient anchoring method, being able to withstand high loads in comparison to their
weight.
1.3.4 Suction Caisson
Suction caissons, depicted in Figure 1.5 are the most widely used anchorage method in
deepwater. A suction caisson is a capped cylinder which is lowered to the seabed, partly
embedding under self-weight. Water is then pumped out from within the caisson creating a
pressure differential between the inside and the outside of the caisson, causing the caisson to
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embed further into the soil. A recent industry study has addressed the design and performance of
suction caissons (Andersen et al., 2005).
A major advantage of suction caissons is that they can be installed in a relatively short time,
without heavy installation equipment such as underwater hammers, making them an attractive
alternative to piles. They can be used as foundations for fixed structures or as anchors for
floating structures, individually or as multi-cell groups. Suction caissons, as with piles, resist
vertical loads partly by friction along the caisson/soil interface, but the smaller length to diameter
ratio compared with typical piles leads to an increased reliance on (reverse) end-bearing. The
offshore industry is in the process of refining the prediction models for caissons so that current,
conservative assumptions and large associated safety factors can be optimised, making them a
more efficient anchoring method. A disadvantage is that caissons cannot easily penetrate hard
layers within the seabed, which may necessitate site-specific tests to ensure their suitability for a
given site.
1.3.5 Torpedo and Deep Penetrating Anchor (DPA)
Both the Torpedo anchor (patented by Petrobras in 1996) and the Deep Penetrating Anchor
(DPA conceptualised by Lieng et al. in 1999) concepts are very similar. They consist of a large
dart or arrow shaped anchor installed by dropping from a pre-determined height above the
seabed, using the kinetic energy gained during the fall to achieve their embedment (schematic in
Figure 1.6). The embedded anchor then performs similarly to a pile. This installation method is
highly suitable for deepwater as installation costs are much less sensitive to water depth.
The potential advantages of dynamically embedded anchors include:
• they are of simple design, resulting in cheap and easy fabrication;
• they are simple to install accurately, quickly and independent of water depth, reducing
the number of hours a full spread of anchors takes to install.
However, prediction of embedment and hence the holding capacity of these anchors is still
unreliable although work is being done to develop more accurate prediction models.
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1.3.6 Follower Embedded Anchors
Suction Embedded Plate Anchor (SEPLA)
The SEPLA is a patented (Technip Offshore Moorings, Inc.) plate anchor that uses a modified
suction caisson, termed a follower, to install the plate to the desired depth. The SEPLA consists
of a rectangular, flat fluke that has a keying flap running the full length across the top edge of the
fluke. This flap is mounted by way of a hinge allowing it to rotate with respect to the fluke once
the mooring line is tensioned, increasing the bearing area to encourage keying of the anchor
plate, and avoiding extraction back up the installation path. Two steel plates, forming the anchor
shank, connect the mooring line to the anchor.
Installation of the caisson occurs in the same manner as for a suction caisson, embedding the
SEPLA to the design depth. Removal of the caisson, for reuse, then leaves the SEPLA
embedded. Tensioning the mooring line attached to the SEPLA then causes the plate anchor to
rotate or ‘key’ to its optimal load bearing orientation. A major issue with the SEPLA is
concerned with this keying process. During keying, the plate moves vertically and horizontally in
addition to rotating and thus leads to uncertainties associated with the final embedment depth
(and thus capacity) during operation. Figure 1.7 displays the installation process.
Suction Embedded Anchor (SEA)
The Suction Embedded Anchor (SEA) is a recent development in deepwater anchoring,
developed by Suction Pile Technology Offshore (SPT). The SEA is made up of two half shells
mounted at the base of a reusable suction caisson. Pictured in Figure 1.8, installation occurs by
embedding the caisson then removing the caisson follower, leaving the SEA embedded, similar
to the SEPLA. The two shells are then forced to open, or key, sideways by pulling on the ‘cheek-
plate’ that is initially located between the shells. Once the shells have rotated 90 degrees, they
form a horizontal plate of semi-circular cross-section, capable of resisting vertical and inclined
loads. Prototype testing has shown that this new concept has considerable merit, although also
suffers to some extent from uncertainty in the degree of embedment loss occurring during keying
of the anchor. Although not currently used, the SEA has strong potential in the field of
deepwater anchors.
1.3.7 Anchor Option Comparison
Table 1.1 gives a comparison of the current deepwater anchorage solutions, modified from a
similar table presented in Ehlers et al. (2004).
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Table 1.1: Comparison of Anchorage Methods
Anchor Advantage Disadvantage
Suction
Caisson
� Simple to install accurately with
respect to location, orientation,
and penetration
� Leverages design experience with
driven piles.
� Well developed design and
installation procedures.
� Anchor with the most experience
in deepwater for mooring Mobile
Offshore Drilling Unit (MODUs)
and permanent floating facilities.
� Large & Heavy.
� Requires Remotely operated
vehicle (ROV) for installation.
� Design requires soil data from
advanced laboratory testing.
� Concerns with holding capacity
in layered soils
� Lack of formal design guidelines
� Limited data on setup time for
uplift.
VLA/NNLA
� Low weight.
� Small.
� Well developed design and
installation procedures.
� Requires drag installation,
keying, proof loading, 2 or 3
Anchor Handling Vessel (AHV)
and a ROV.
� No experience with anchoring
permanent floating facilities
outside Brazil.
� Difficult to ensure installation to
the required embedment and
orientation.
� Extensive soil survey required.
SEPLA
� Uses proven suction caisson
installation methods.
� Cost of anchor element is the
lowest of all the deepwater
anchors.
� Provides an accurate measure of
embedment and position of the
anchor.
� Design based on well developed
procedures for plate anchors.
� Patented installation method.
� Installation time greater than for
a caisson.
� Requires keying and proof
loading.
� Requires an ROV.
� Limited field load tests.
SEA
� Uses proven suction caisson
installation methods.
� Provides an accurate measure of
embedment and position of the
anchor.
� Design based on well developed
� Requires an ROV.
� Little to no experience in the
operation of the SEA.
� Requires keying and proof
loading.
� Requires an ROV.
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procedures for plate anchors. � Installation time greater than for
a caisson.
Torpedo
anchor/DPA
� Simple to design.
� Simple and economic to fabricate.
� Robust and compact design makes
handling and installation simple
and economic with only one AHV
required and no ROV.
� Accurate to position with no
requirements for proof loading.
� Patented installation method.
� No experience outside Brazil.
� Lack of documented installation
and design methods with
verification agencies.
� Unknown orientation once
embedded.
1.4 RESEARCH OBJECTIVES
This project contributes to the limited database of the behaviour of anchors during keying, in
particular quantifying the effects of eccentricity of loading from the plate on the vertical
displacement of the plate anchors during the keying process. Reduced scale model centrifuge
testing has been used to facilitate the optical measurement of the rotation and displacement of the
various geometries of plate anchors through a soil/Perspex interface during keying. Additionally
the project has explored the effects of installation on the keying characteristics of SEPLAs. As
well as the physical modelling aspect of the project, a numerical model has been developed that
simulates the keying process. This allows accurate prediction of final embedment depth and
anchor orientation, and ultimately anchor load capacity.
From this research two papers have been published O’Loughlin et al. (2006) and Gaudin et al.
(2006). Both take results and the initial analysis, presented later in this thesis later and presented
them as a collaborated work. Copies of both can be found in the appendices.
1.5 THESIS STRUCTURE
Chapter 2 gives an overview of research previously conducted pertaining to SEPLA installation
and plate anchor keying.
Chapter 3 provides a detailed description of the experimental testing program, including
apparatus used and testing procedures.
Chapter 4 presents a summary of the experimental results with preliminary analysis provided.
Chapter 5 presents an analytical simulation model developed to predict plate anchor keying
behaviour, including a brief description of similar models.
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Chapter 6 presents a detailed comparison between the experimental results (Chapter 4),
theoretical solutions (Chapter 2) and results obtained from the analytical simulation in Chapter 5.
Chapter 7 summaries major finding of the research and provides suggestions for further studies.
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Figure 1.1: Offshore production facilities (Courtesy: Minerals Management Services)
Figure 1.2: Offshore production facilities (Courtesy: Minerals Management Services)
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Figure 1.3: Mooring types (Vryhof, 2000)
Figure 1.4: Drag anchors, sand and clay (Vryhof, 2000)
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Figure 1.5: Suction caisson
Figure 1.6: DPA and Torpedo anchor and installation schematic (Lieng et al. 1999 and Medeiros, 2001)
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Figure 1.7: SEPLA installation technique
Figure 1.8: SEA installation technique
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CHAPTER 2 PLATE ANCHOR AND KEYING BACKGROUND
2.1 PREVIOUS STUDIES
2.1.1 SEPLA
Aside from a limited number of reduced scale laboratory and field tests reported by Wilde et al.
(2001), very little performance data exists on SEPLA behaviour (Aubeny et al. 2001). Wilde et
al. (2001) indicated a wide range of vertical displacements during anchor keying of 0.5 – 1.7
times the plate height, for an anchor with a normalised loading eccentricity, e/B = 0.5. These
results are from testing conducted to failure at embedment ratios of 4 – 10, considered deep
failure conditions (Rowe & Davis, 1982 and Song & Hu, 2005). The original concept is the
investigation of quasi-vertical loading during keying. In practice, however, keying in the field is
carried out at angles as low as 30 – 35o to the horizontal.
The paucity of current performance data and the potential economic benefits of SEPLAs
necessitate further investigation of the effect of the installation procedure and plate anchor
keying processes on short and long-term anchor capacity.
2.1.2 Keying
Of the proposed anchor concepts in recent years, the most promising utilise a follower to embed
an initially vertical plate anchor, typically located at the follower base. The Suction Embedded
Plate Anchor (SEPLA) is a developed example of these new anchorage methods and has been
utilised to moor offshore structures in the Gulf of Mexico and West Africa. Dove et al. (1998)
illustrates the installation and keying process of the SEPLA in Figure 2.1.
The offshore industry has raised two main concerns regarding SEPLAs and other follower
embedded plate anchor concepts: the first is predetermining the amount of proof load required to
complete the keying process and the second is the reduction in embedment depth (and hence
anchor capacity in normally consolidated clay) due to keying. Additionally the keying process
will disturb the soil in the immediate vicinity of the plate resulting in a loss of strength in soil
adjacent to the plate (Randolph et al. 2005). Recovery of this strength may occur in time due to
consolidation but the embedment loss during keying is unrecoverable. Combined with the
typically increasing shear strength profile with depth of offshore clays, this will translate to an
unrecoverable loss in potential anchor capacity and is thus a crucial element to quantify.
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Very little has be done to quantify the loss of embedment. However, current guidelines specified
by the US Naval Civil Engineering Laboratory guidelines (NCEL, 1985) state that embedment
loss is twice the anchor height in cohesive soils. Contrary to this, Foray et al. (2005) have
reported embedment losses up to 2.5 times the anchor height, and as mentioned Wilde et al.
(2001) present a range of embedment loss of 0.5 - 1.7 times the plate height.
A recent numerical analysis (Song et al. 2005), showed a loss of embedment, in uniform strength
clay, of 0.6 times the anchor height (B) for vertically loaded strip anchors with loading
eccentricity of 0.625B. Centrifuge testing conducted in transparent synthetic “clay” confirmed
this result. In addition, when the loading was inclined at 45o the computed embedment loss
reduced to 0.25B.
The very limited database on embedment loss during plate anchor keying is a concern. With the
disconcertingly large range of results, being 0.25 - 2.5 times the plate height, the degree of
uncertainty associated with the capacity of these concepts is troubling. This has resulted in a
current lack of confidence in a potentially, highly economical anchorage solution.
2.1.3 Plate Capacity
Over the past four decades considerable attention has been paid to the capacity of plate anchors
under monotonic vertical loading conditions, with notable contributions from Vesic (1971), Das
(1978, 1980), Rowe & Davis (1982) and Merifield et al. (2001, 2003). Song & Hu (2005)
summarise the various approaches and findings to date.
There are different approaches for determining plate capacity depending on soil properties,
loading conditions and assumed failure mechanism (shown in Figure 2.2), the majority of which
are detailed in Merifield (2002). Merifield (2002) shows in Figure 2.3, that the ultimate anchor
capacity increases linearly with overburden pressure (γH/su) before reaching a limiting value, at
which point the anchor failure is considered deep. To be considered deep at failure the plates
embedded at an H/B ratio greater than 4 must have an overburden pressure greater than 5.8. For
clays used in this study, with shear strength proportional to depth, the overburden pressure
becomes γ’/k (effective weight of soil/shear strength gradient), and is ~ 6.5. Thus, the plates used
during testing satisfy the deep failure criteria.
Given deep or localised failures, the ultimate uplift capacity (Qu) of a horizontal plate is:
acuu WANsQ += (2.1)
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where su is the undrained shear strength at mid-anchor embedment (also referred to in previous
work as cu); Nc the non-dimensional breakout factor; A is the plate area and Wa is the submerged
anchor weight (dry anchor weight minus weight of displaced soil).
To determine the Nc values of strip plates several studies have been conducted using finite
element (FE) analysis, limit equilibrium and plasticity theories. Das & Singh (1994) conducted
research in uniformly consolidated clays. His results suggest for square and circular anchors, Nc
increases with H/D (where D is the anchor diameter/breadth) up to about 9 and then remains
constant for deep anchor conditions thereafter (Das & Singh, 1994). Martin and Randolph (2001)
determined exact Nc solutions for deeply embedded smooth and rough circular plates of 12.42
and 13.11 respectively.
Merifield et al. (2003) conducted further studies into the Nc values for strips anchors and
determined upper bound values by means of a finite element analysis. They showed that Nc
values were dependant on whether the soil/anchor interface at the rear of the anchor could
sustain adequate tension to ensure contact between the two. Figure 2.4 shows the difference
between breakaway and no breakaway cases.
From these studies, Merifield’s upper bound solution for a no breakaway, horizontal anchor in
inhomogeneous purely cohesive weightless soil of Nc ~ 12 (shown in Figure 2.5) is the most
suitable for comparison to the results presented later.
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Ocean seabed
� � � �
Figure 2.1: The SEPLA concept: ���� Suction installation, ���� Caisson retrieval, ���� Anchor keying, ���� Mobilised
anchor
Figure 2.2: Failure mechanism of a plate (Merifield, 2002)
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Figure 2.3: Effect of overburden pressure on strip anchors (Merifield, 2002)
Figure 2.4: Nc comparison for Breakaway cases in weightless soil (Merifield, 2002)
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Figure 2.5: Upper Bound Nc values for horizontal anchors - inhomogeneous cohesive soil (Merifield, 2002)
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CHAPTER 3 MODEL PLATE ANCHOR TESTING
A series of model SEPLA tests and plate anchor keying tests adjacent to a Perspex window, were
conducted on the drum and beam centrifuges at the University of Western Australia (UWA), in
order to determine the effect of installation on the holding capacity of a SEPLA and the
orientation of a plate anchor during keying. All tests were performed in normally consolidated
kaolin clay in an attempt to replicate soil conditions offshore. Soil characterisation tests
performed in conjunction with the anchor tests assisted in the interpretation of data obtained
during anchor testing.
3.1 CENTRIFUGE TESTING APPARATUS
Both the fixed beam and drum centrifuges at UWA were utilised in gathering the data for this
study. Below are the principles of physical modelling in the centrifuge and the centrifuge test
apparatus.
3.1.1 Principles of Centrifuge Testing
Centrifuge testing is an extremely useful tool and now a common method for the study and
analysis of geotechnical materials and problems (Taylor, 1995). The primary aim of centrifuge
modelling is to achieve stresses and strains on a reduced scale model representative of those
experienced on an equivalent prototype. By spinning a model at high rpm the model is exposed
to an artificial gravitational field which has the effect of increasing the self-weight of the soil. By
this means the stresses in a model with linear dimensions scaled as 1:N, become identical to
those in the prototype, provided the centrifuge acceleration level is N times gravity.
For the present application the ability to produce a clay sample with a linearly increasing
strength profile, which is hard to do by any other means than self-weight consolidation, is
probably the most significant advantage centrifuge modelling has over testing at 1 g.
Schofield (1980) and Taylor (1995) provide a detailed discussion of geotechnical centrifuge
modelling.
Table 3.1 summaries the various centrifuge scaling factors.
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Table 3.1: Centrifuge Modelling Scaling Factors
Parameter Dimensions Scale Factor
(Model:Prototype)
Acceleration LT -2
1:N -1
Length L 1:N
Area L2 1:N
2
Stress ML-1
T -2
1:1
Strain - 1:1
Force MLT -2
1:N 2
Mass ML 2T
-2 1:N
3
Time (diffusion) T 1:N 2
Velocity LT -1
1:1
Time (dynamic events) T 1:N
One particular advantage afforded by model testing concerns tests associated with consolidation.
To achieve the same degree of consolidation in a model as in the equivalent prototype, the
scaling relationship for the time factor is 1: N 2
(model: prototype) (Taylor, 1995), for example
one year of prototype consolidation compares to a model test of approximately 52 minutes and
34 seconds at 100 g.
3.1.2 The Geotechnical Beam Centrifuge
The fixed beam geotechnical centrifuge (Figure 3.1), located in the civil engineering laboratory
at UWA, is an Acutronic Model 661 centrifuge. It is a 40g-tonne machine, meaning that at its
maximum acceleration level of 200 g (approximately 340 rpm with a platform velocity of 64
m/s, Randolph et al., 1991) it has a maximum payload of 200 kg, or at 100 g it has a payload of
400 kg. The Model 661 has a swinging platform at a radius of 1.8 m on which test packages are
mounted, with a nominal working radius of 1.55 m.
A large movable weight located opposite the testing platform counter balances the package,
minimising unbalanced loads acting on the centrifuge pedestal. The machine operates in an air-
conditioned room specially designed to allow constant temperature during long tests. Randolph
et al. (1991) provide a detailed description of the facility.
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Test Strongbox
All SEPLA tests where performed in normally consolidated kaolin clay contained within a
strongbox (Figure 3.2). The strongbox is a sealed, rectangular aluminium box with internal
dimensions, 650 mm by 395 mm and an internal depth of 325 mm. Drainage holes at various
locations around the box permitted saturation or drainage of the test sample ‘in flight’. An
external standpipe is connected to one of the drainage holes allowing the water level to be
maintained. The top flanges of the box have holes at 25 mm intervals to allow mounting of
testing apparatus such as actuators, cameras and lights during testing.
Motor Driven Actuator
An electronically driven actuator (Figure 3.3), developed at UWA allows a range of beam
centrifuge tests to be undertaken, including T-Bar penetrometer and anchor tests. The actuator
has two degrees of freedom, allowing combined vertical and horizontal movement to a
maximum of 250 mm and 180 mm respectively. Two 30-Volt DC variable speed motors, each
with a maximum velocity of 3 mm/sec, allow vertical and horizontal loads of up to 10 kN and 2
kN respectively to be applied.
Data Acquisition and Control Software
The package is monitored ‘in-flight’ by an on-board ‘flight computer’, which allows for high-
speed data acquisition, with the data returning to the control room via a high-speed wireless
network (O’Loughlin et al., 2004). This monitoring allows the user to view real time data,
ranging from pore pressures within a sample, to the temperature in the room during centrifuge
operation. The control computers allow the operation of various instruments mounted on the
centrifuge, while monitoring the data returned and the progress of the test.
T-bar Penetrometer
The T-bar penetrometer (Figure 3.4) is a site investigation tool, developed at UWA by Stewart
and Randolph (1991) for the determination of the shear strength profile of soft clay samples in
the centrifuge. The T-bar comprises a 5 mm diameter cylinder, 20 mm in length connected at
right angles to the vertical drive shaft and a highly sensitive load cell located directly behind the
head. The load cell measures the resistance encountered when the T-bar moves through the clay
sample. The load cell measures the soil resistance during the T-bar’s penetration and extraction,
up to maximum undrained shear strengths of approximately 100 kPa. Stewart and Randolph
(1991, 1994) have described the development of the T-bar penetrometer tests in detail.
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3.1.3 The Geotechnical Drum Centrifuge
The drum centrifuge at UWA (Figure 3.5), capable of a maximum acceleration level of 485 g at
a radius of 0.6 m, has been in operation since 1999. Developed in collaboration with Professor
Andrew Schofield at Cambridge University, and the manufacturer Thomas Broadbent & Sons
(TBS), the centrifuge has two concentric shafts connected to a servomotor, allowing its central
tool table and outer channel to rotate independently. Decoupling and stopping the tool table
allows exchanging or modification of tools, on the table while allowing the sample to continue
spinning. For a full description of the drum centrifuge, refer to Stewart et al. (1998).
Sample Box and Drum Channel
For the purpose of the drum tests six small sample boxes, an example of which is shown in
Figure 3.6, were constructed to fit inside the drum channel (1.2 m diameter, 300 mm width
(vertically) and 200 mm depth (radially)). The sample boxes were made from aluminium plate
and have internal dimensions of 258 mm long, 160 mm deep and 80 mm wide. Seals between
adjoining aluminium plates ensure the boxes are water tight for the test while allowing an easy
interchange of the box sides. One side of the box is interchangeable with a Perspex window to
permit the viewing of tests via digital camera.
Digital Camera and Cradle
A Canon S50 digital camera, with a five mega-pixel resolution (2592 x 1944 pixels) and 1 Gb
memory card captured the plate’s orientation, adjacent to the Perspex window of the sample box.
A specially designed camera cradle, fitted with a trigger (seen in Figure 3.7) and operated from
the drum control room, allowed image capture to commence and conclude when desired with a
frequency of 0.5 Hz. The cradle has a slotted base allowing correct positioning of the camera for
focusing of the test site. Camera settings as detailed by White (2003) must be adopted to ensure
full-resolution pictures are taken.
Tool Table Actuator
The tool table actuator (depicted in Figure 3.8), fabricated in the civil engineering workshop at
UWA, has three axes of movement: vertical (across the width of the channel), radial (in and out
of the channel) and circumferential (rotational, around the channel) (Stewart et al., 1998). The
vertical and radial axes have a continuous load rating of 10 kN, while the circumferential axis
has a torque rating of 500 Nm (Stewart et al., 1998). Stewart et al. (1998) provide further details
on the tool table actuator.
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Data Acquisition and Control Software
The drum centrifuge data acquisition and control system comprises several interacting branches;
the channel drive control; tool table control; and data acquisition (Stewart et al., 1998). The
rotation of the channel is monitored and computer-controlled. The computer ensures that it is
safe before stopping or starting the machine. A second computer controls vertical and radial
movement of the tool table, with the rotational movement controlled by a third computer. The
drum is fitted with two on-board data acquisition systems, one each on the channel and the tool
table. The basic system can record 32 direct signals, half on each of the channel and the tool
table (Stewart et al., 1998). A fourth computer records while a fifth computer displays this data
in the drum, control room. For a more in-depth description of this setup, refer to Stewart et al.
(1998).
3.2 KAOLIN CLAY
The suite of tests conducted during the investigation used commercially available kaolin clay,
which was selected for the abundance of reliable data regarding its geotechnical properties, its
isotropic nature, and its relatively quick consolidation time.
3.2.1 Soil Properties
Due to the extensive studies previously conducted on kaolin clay, at UWA, there was no
requirement during this study to perform additional classification tests on kaolin. The properties
of the locally sourced Kaolin clay, reported by Stewart (1994), are in Table 4.3.
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Table 3.2: Properties of Normally Consolidated Kaolin Clay (Stewart, 1994)
Property
Specific Gravity, Gs 2.60
Liquid Limit, LL (%) 61
Plastic Limit, PL (%) 27
Compression Index, Cc 0.47
Swelling Index, Cs 0.1
Friction Angle, φ' (°) 23
Strength ratio from simple shear, (su/σ'v)NC 0.17
Coefficient of Consolidation, cv (m2/yr) 3.9
As the coefficient of consolidation value varies with effective stress (House et al. 2001), the
value of cv shown in Table 4.3 is an average calculated over a prototype depth of 35 m, assuming
an effective unit weight of 6 kN/m3.
3.2.2 Sample Preparation
Normally Consolidated Clay (Beam Strong Box – SEPLA Test)
Kaolin was prepared at 120 % water content by combining 50 kg of commercially available
kaolin powder with 60 kg of water in a drum mixer, and mixing for a minimum of 24 hrs.
Concurrently a strongbox was sealed and a layer of coarse sand, approximately 10 mm thick,
was place at the bottom of the box. Filter paper was placed on the sand, and the sand and filter
paper saturated with water.
On completion of the slurry and sand layer preparation, the strongbox was filled with kaolin
slurry and consolidated in the centrifuge by spinning to the desired acceleration level (145 g).
After approximately 12 hrs of spinning, the sample had consolidated considerably and required
more slurry to obtain the required sample height of 230 mm. Several such ‘top-ups’ were
required in order to obtain the correct sample height, the entire process taking several days to
complete. After the final top-up, the sample spun for 48 hrs to ensure consolidation of the sample
was complete and the excess pore pressures had dissipated.
A column of pore pressure transducers, mounted on a rigid bar (seen in Figure 3.9), allowed
monitoring of the pore pressures within the clay sample. The bar holds five pressure transducers
positioned exactly 50 mm apart along the length of its shaft, allowing for measurement of pore
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pressures at several depths within the sample. This provided knowledge of when the excess pore
pressures had dissipated sufficiently for tests to commence.
Normally Consolidated Clay (Sample Box in Drum – Keying Tests)
The Kaolin slurry (prepared in the same manner as previously described) was placed in boxes at
20 g. This was done by means of a nozzle attached to the tool tables positioned in each box in
turn. Once each box was full, the drum was ramped up to 100 g. Placement of water on the
samples at 100g ensured complete saturation during consolidation. Repeating this process
several times produced the desired sample height. Spinning for a minimum of 48 hrs, after the
final top, ensured full consolidation. Size restrictions prevented the use of pore pressure
transducers. A T-bar penetrometer test conducted prior to testing confirmed full consolidation.
Normally Consolidated Clay (Sample Box in Beam – Keying Tests)
Kaolin slurry was placed in each of the six sample boxes at 1 g, then the sample boxes where
placed in the bottom of a strongbox. Each sample box had a felt mat, covering the bottom and
running up two sides, fixed down with double sided tape and mat fasteners. This allowed two-
way drainage to occur. The full sample boxes, in the sealed strongbox, were then consolidated in
the beam centrifuge at 100 g. Similar to the preparation of the strongbox sample, each sample
box was ‘topped up’ several times to reach the desired sample height. The water height was
maintained for the duration of consolidation by continuously trickling water into the strongbox
and opening a hole in the side of the strongbox (above the top of the sample boxes).
3.3 PLATE ANCHOR KEYING TEST PROCEDURE
3.3.1 Soil Characterisation Tests
T-bar tests were conducted prior to and at the completion of testing in each sample. Their
purpose was to assess the initial shear strength profile of the soil and to determine whether the
strength profile changed during the course of testing. To achieve equivalent prototype undrained
conditions, the normalised velocity must be greater than 30 (House et al., 2001). These
normalised velocities were determined using Equation 4.9 below.
vc
vdV = (3.1)
where V is the normalised velocity; v the pullout rate; d the T-bar diameter; and cv the
coefficient of consolidation of 0.1 mm2/s.
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The rates for both the penetration and the extraction were 1 mm/s in order to obtain undrained
conditions, corresponding here to a V of 50.
3.3.2 SEPLA Beam Centrifuge Tests
Model Caisson and SEPLA
To investigate the installation effects on the holding capacity of SEPLAs, a 1:145 reduced scale
model caisson and SEPLA were fabricated (Gaudin et al. 2006). The SEPLA, made from 1 mm
thick stainless steel plate, was 35 mm square, roughly based on plate anchor dimensions
suggested by Wilde et al. (2001). Figure 3.10 shows the model SEPLA and caisson.
The aluminium model caisson had the following dimensions: total height 170 mm; external
diameter 30 mm; and wall thickness 0.5 mm, replicating a typical caisson of 24.65 m high and
4.35 m in diameter. In addition, there was a threaded collar placed on the top of the caisson, for a
guide rod to ensure vertical installation of the caisson. The top of the caisson had a nozzle to
attach a hose to pump water out then into the caisson for installation and extraction respectively.
Two threaded holes, one for a pneumatic valve and the other for a pore pressure transducer
(PPT), along with a mount for a second PPT located on the top of the caisson. The PPTs
measured the internal and external water pressure while the pneumatic valve allowed venting of
the caisson.
SEPLA Test Guide, Motor and Actuator Setup
A low profile motor (seen in Figure 3.11) was required to facilitate the extraction of the caisson
on completion of the SEPLA installation. An aluminium guide, mounted below the motor,
ensured vertical installation and extraction of the caisson. Using the actuator in conjunction with
the motor, both mounted on an extended strongbox, allowed caisson extraction and anchor
loading without having to ramp down the centrifuge.
Test Procedure
Each of the six caisson tests followed the installation procedure (Figure 3.12), as detailed below:
1. Attach the PPTs, pneumatic valve (attached to a compressed air line) and syringe pump
to the top of the caisson as seen in Figure 3.13. Set the SEPLA with anchor chain
attached in the slot at the base of the caisson.
2. Ensure that the SEPLA remains in the slot and embed the caisson ~40 mm by pushing it
in by hand. Attach the vertical guide to the caisson. Attach the anchor chain, through a
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pulley mounted at the base of the actuator, and a pulley mounted in the actuator to a load
cell, and the guide string to the motor, as shown in Figure 3.14.
3. Ensure there is enough slack in the anchor chain so the caisson can fully embed. Take all
the slack out of the guide string so that the caisson will not embed under self-weight
during ramping up of the centrifuge.
4. Ensure that the pneumatic valve is open before ramping up.
5. Once the centrifuge has reached 145 g, drive the motor at 5 mm/s so the guide string is
unwound, allowing the caisson to penetrate under its self-weight. Close the pneumatic
valve.
6. Drive the syringe pump at 1 mm/s, equating to a caisson installation speed of ~ 2 mm/s to
remove water from within the caisson until maximum embedment of the caisson.
Of the six caisson tests, two involved caisson extraction by reverse pumping according to the
following procedure:
1. Keep the pneumatic valve closed.
2. Drive the syringe in at 0.6 mm/s to force water back inside the caisson to remove the
caisson from the clay at a rate of ~ 1.2 mm/s, leaving the plate installed.
For the remaining four tests, vented extraction of the caisson was employed as described below:
1. Open the pneumatic valve.
2. Drive the motor to extract the caisson at 1 mm/s with the guide string. Ensure full
extraction so that it does not interfere with the anchor pull.
At completion of the caisson extraction and SEPLA installation, the anchor line was tensioned at
a rate of 0.1 mm/s, ensuring undrained conditions as described previously.
For the jacked tests, the procedure was much simpler:
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1. At 1 g, use a guide to push the SEPLA, with anchor chain attached, to the desired
embedment into the clay. Fix the chain to the actuator through the pulleys.
2. Ramp up and pull the anchor chain at a rate of 0.1 mm/s to ensure undrained behaviour.
Test Program
The loading angle for each test shown was varied between 45o and 55
o, to the horizontal,
depending on the test location and the position available for the actuator as detailed in Table 3.3.
Table 3.3: SEPLA Test Summary
Test Number Test Name Installation
Method
Caisson
Extraction
Method
Loading
Angle (o)
Test 1 VE-ST1 Suction Vented 50
Test 2 PE-ST2 Suction Reverse Pumping 50
Test 3 VE-ST2 Suction Vented 50
Test 4 VE-LT1 Suction Vented 45
Test 5 PE-LT1 Suction Reverse Pumping 45
Test 6 VE-LT2 Suction Vented 45
Test 7 JI-ST1 Jacked - 50
Test 8 JI-ST2 Jacked - 55
Test 9 JI-LT1 Jacked - 50
Test 10 JI-LT2 Jacked - 55
VE/PE – Vented/Pumped Caisson extraction
JI – Jacked Installation
ST – Short Term, plate loaded immediately after caisson extraction.
LT – Long Term, plate loaded after an extended anchor soak time.
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3.3.3 Plate Keying Drum Centrifuge Tests
Model Anchors
Investigation of anchor orientation during keying utilised two model plate anchors at a 1:100
reduced scale. The plates, machined from 2 mm thick stainless steel, were all 30 mm wide and
had lengths of 30 and 80 mm. The 30 mm plate modelled the characteristics of an anchor twice
its length while the 80 mm plate attempted to model an infinitely long anchor. As the anchors are
tested adjacent to the constrained Perspex, the soil displacements at the soil-Perspex interface
allows these models can be considered representative of a plate (of length 2L), giving full
prototype dimensions of B = 3 m and Le = 3 and ‘∞’ m respectively (Figure 3.15). They
accommodated an ‘O’ ring at the plate-Perspex interface, as the anchors are tested adjacent to the
constrained Perspex. The effective length (in terms of soil behaviour) of the first three plates
gave a model aspect ratio, Le/B = 2 and ‘∞’ as the loading shaft was at the centre and the plate
spaned the width of the box, to model an infinitely long strip anchor. The eccentricity (e) of the
anchor padeye or load attachment point was varied for each plate using four interchangeable
‘anchor shafts’ oriented perpendicularly to the plate. The lengths of these anchor shafts were 5,
15, 30 and 45 mm, corresponding to eccentricity ratios (e/B) of 0.17, 0.5, 1.0 and 1.5.
Testing Arrangement
To facilitate optical measurement of the plate keying process, the Perspex face was marked with
a grid enabling conversion from picture scale to model scale. The installation of the
camera/guide arrangement, in the channel, with the camera lens perpendicular to the Perspex,
ensured the best picture quality and minimised optical distortion. A Liquid Crystal Display
(LCD, Figure 3.16) was attached with double-sided tape to the Perspex within the picture frame.
The LCD displays the line of code copied to the data file, allowing accurate referencing of
pictures to the test data.
A loading arm, as seen in Figure 3.17, connected the anchor shaft to the tool table actuator. A
guide for the loading arm, designed to reduce the unsupported length of the loading arm from
550 mm to 200 mm minimised the loading arm’s movement away from the Perspex. The
placement of a ‘knee-joint’ at the anchor shaft/loading arm connection permitted in plane
rotation whilst restricting out of plane rotation, thereby reducing the chance of the plate coming
away from the Perspex (and hence being lost from sight) while allowing it to key. Two strain
gauges located just above the knee joint on the loading arm enabled axial load and bending
moment data to be collected. Figure 3.6 shows this set up.
To ensure that the keying load was applied in the same ‘vertical’ plane for the entirety of the test,
the sample box was placed in the channel with the inner face of the Perspex aligned with the
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direction of the gravitational field created by the centrifuge. Aluminium plates held the boxes in
position.
Test Procedure
Six sample boxes were prepared and the clay initially consolidated at 100 g. To allow optical
measurements of the tests only two boxes could remain in the channel (situated 180o apart). For
the remaining boxes, the surface water off was drained off and the clay covered with Glad Wrap
(thin clear plastic - to ensure the samples did not dry out). The samples were placed on the
laboratory bench. Once testing in the initial two test samples was complete, they were replaced
by two of the remaining boxes. Reconsolidation for between 24 and 48 hours then occurred,
prior to testing in the new samples.
Prior to any anchor tests, a T-bar penetrometer test was performed (rate 1 mm/s). The procedure
for a vertically loaded anchor keying test in the drum centrifuge was:
1. Remove a fully consolidated sample box from the channel.
2. Remove the thicker of the two aluminium sides from the box by carefully sliding it down
the face of the clay towards the bottom of the box. Using a sieve, coat the exposed clay
surface with modelling flock.
3. Attach the desired anchor shaft to the desired plate. Place the loading arm through the
guide and fix it to the anchor shaft. Ensure the loading arm is correctly oriented within
the guide.
4. Carefully push the plate into the clay from the side of the box in the desired location.
Ensure that the plate and loading arm rotate as little as possible during the installation and
the plate is as near vertical as possible. Leave the plate protruding from the clay by a
couple of mm.
5. Place the lightly greased Perspex side on the box, using it to push the plate the remainder
of the way into the clay (ensuring that the plate is visible for the start of the test).
6. Place the sample box with the installed anchor back into the channel, move the tool table
vertically, to ensure that the loading arm is perpendicular to the soil surface, and attach
the loading arm to the tool table actuator.
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7. Ensure the camera has the correct settings and is focused to capture the test site,
including the LCD display.
8. Reconsolidate the sample in accordance with the amount of time spent at 1 g detailed in
Table 3.4, remembering to fill both boxes in the channel with water on ramp up.
Table 3.4: Reconsolidation times
Time @ 1g Extra
Consolidation
< 1 hour 1 hour
1 – 2 hours 2 hour
2 – 3 hours 3 hour
3 – 4 hours 4 hour
< 4 hours 24 hours
9. Operate the tool table to pull the loading arm towards the centre of the drum at a rate of
0.1 mm/s (giving a dimensionless velocity of vB/cv ~ 30) ensuring undrained behaviour
(Finnie & Randolph, 1994). Stop the actuator before the plate meets the guide.
Figure 3.18 shows the final test layout.
Test Program
All anchor tests in the drum were vertically loaded at 0.1 mm/s and conducted at 100 g, in
normally consolidated kaolin clay. Load application was vertically over the padeye. Table 3.5
summaries the tests conducted.
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Table 3.5: Drum Keying Tests Summary
Test Number
Anchor
Dimensions
B x L (mm)
Eccentricity
e (mm) Le/B e/B
B2a80e30 30 x 80 30 ∞ 1
B3a30e5 30 x 30 5 2 0.17
B3a30e15 30 x 30 15 2 0.5
B4a30e30 30 x 30 30 2 1
B4a30e45 30 x 30 45 2 1.5
B5a80e5 30 x 80 5 ∞ 0.17
B5a80e45 30 x 80 45 ∞ 1.5
B6a80e15 30 x 80 15 ∞ 0.5
B6a80e30 30 x 80 30 ∞ 1
3.3.4 Plate Keying Beam Centrifuge Tests
Model Anchors
Keying tests in the beam involved the use of an 80 x 20 mm stainless steel plate anchor made to
span the sample boxes used in the drum tests. The 3 mm thick plate had rubber O-rings placed on
the edges that made contact with the sides of the sample box. A milled slot in the top of the plate
allowed attachment of the ‘anchor shafts’, similar to those used in the drum tests, giving
eccentricity ratios (e/B) of 0.25, 0.5, 1.0 and 1.5 (namely model dimensions of 5, 10, 20 and 30
mm respectively).
Test Procedure
Six normally consolidated kaolin clay samples were consolidated in the sample boxes at 100g.
Rearranging the sample boxes at the completion of consolidation allowed testing, while ensuring
the digital camera and cradle sat inside the strongbox opposite the Perspex window of the sample
box. The preparation of a test box and the test procedure was as follows:
1. Remove the thicker of the two aluminium sides from the box by carefully sliding it down
the face of the clay towards the bottom of the box. Coat the exposed clay surface with
modelling flock.
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2. Attach the desired anchor shaft to the plate. Attach the anchor chain and load cells (as for
the SEPLA tests) to the anchor shaft.
3. Carefully push the plate into the clay from the side of the box in the desired location.
Ensure that during the installation, the plate remains vertical and the chain taut in the clay
at the correct loading angle at the padeye. Leave the plate protruding from the clay by a
couple of mm.
4. Place the lightly greased Perspex side on the box using it to push the plate the remainder
of the way into the clay (ensuring that the plate is visible for the start of the test).
5. Place the sample box with the installed into the strongbox and secure in position with
specially made aluminium brackets (Figure 3.19).
6. Ensure the camera has the correct settings and the image is focused to capture the test
site, including the LCD, with the memory card empty to capture the whole test.
7. Reconsolidate the sample in accordance with the amount of time spent at 1 g (see Table
3.4). During reconsolidation feed water into the test box and ensure the overflow runs
into the strongbox to ensure the sample remains fully saturated.
8. Mount an actuator with the horizontal axes of the actuator running down the centre line
of the test box. Mount a load cell in the actuator and connect the anchor chain.
9. Before ramping up the centrifuge, drive the actuator to remove the majority of slack in
the anchor chain, ensuring that during this process the anchor is not disturbed.
10. Drive the actuator to pull the chain at a rate of 0.15 mm/s (giving a dimensionless
velocity of vB/cv ~ 30) ensuring undrained behaviour (Finnie & Randolph, 1994). Stop
the actuator before the plate hits the base of the actuator.
Figure 3.20 shows the final test setup.
Test Program
All anchor tests in the beam were performed at 100 g, at a rate of 0.15 mm/s normal to the plate
direction, in normally consolidated kaolin clay. Application of the load was vertical, over the
centre of the plate. Table 3.6 summaries the tests conducted.
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Table 3.6: Beam Keying Tests Summary
Test Number Eccentricity
e mm)
Loading Angle
(to the horizontal
- degrees)
e/B
Test 1* 30 90 1
Test 2 10 90 0.5
Test 3 10 90 0.5
Test 4 20 90 1
Test 5 20 90 1
Test 6 15 90 0.75
Test 7 10 70 0.5
Test 8 5 90 0.25
Test 9 10 60 0.5
Test 10 30 90 1.5
Test 11 15 90 0.75
* Note Test 1 used a 80 x 30 mm plate all other plates were 80 x 20 mm.
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Figure 3.1: Geotechnical beam centrifuge
Figure 3.2: Beam strongbox
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Figure 3.3: Motor driven actuator
Figure 3.4: T-bar penetrometer
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Figure 3.5: Drum centrifuge with clamshell removed
Figure 3.6: Keying test setup in sample box, @ 1 g
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Figure 3.7: Digital camera cradle with trigger
Figure 3.8: Tool table actuator
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Figure 3.9: Sub-Terrain Oil impregnated Multiple Pressure Instrument (STOMPI)
Figure 3.10: Mounted model SEPLA
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Figure 3.11: SEPLA test setup
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Actuator Load Cell
Retrieval of the caisson
Recovery of the chain slack
Self weight installation
Caisson
To the syringe
pump
Chain Load Cell
Anchor
Load Cell
Vertical Guide
Kaolin clay
Sand
Water
Winch
Pulley
system
Actuator
Pullout
Suction installation
100
65
Figure 3.12: Experimental arrangement and test procedure (section view)
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Figure 3.13: Caisson with attachments (pneumatic valve hidden behind caisson guide)
Figure 3.14: SEPLA test, actuator setup
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Figure 3.15: Model plate anchors with various loading shafts attached (including two not used in this study)
Figure 3.16: LCD attached to sample box
Not used.
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Figure 3.17: Drum keying test, loading arm
Figure 3.18: Drum keying test layout (White, 2003)
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Figure 3.19: Sample box held in place with brackets
Figure 3.20: Beam keying test configuration
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CHAPTER 4 EXPERIMENTAL RESULTS
4.1 SEPLA TESTS
The ten SEPLA tests, conducted in one strong box, focused on determining the influence of
installation on SEPLA ultimate holding capacity. Six of the tests involved suction installation
four having the caisson removed by the vented extraction method and two by the reverse
pumping method. In addition, there were four jacked installation tests.
4.1.1 Soil Characterisation Tests
The soil resistance acting on the T-bar during penetration, measured by a load cell situated
immediately behind the cylindrical head, allows the soil shear strength to be estimated as:
dN
Ps
b
u = (4.1)
where: su is the undrained shear strength; P the force per unit length acting on T-bar; Nb the bar
factor; and d the T-bar diameter.
Stewart and Randolph (1991) suggested using a value of 10.5 for Nb, intermediate between the
plasticity solution for fully smooth and fully rough cylinders. The shear strength gradient, k, for
normally consolidated clay typically increases linearly with depth, z, and may be approximated
as:
kzsu = (4.2)
Table 4.1 and Figure 4.1 show the shear strength profile for the clay used for the SEPLA tests.
Table 4.1: Summary of shear strength gradient for SEPLA tests
Test Number Shear Strength
gradient, k (kPa/m)
Tests 1 to 10 0.88
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4.1.2 SEPLA Capacities
As discussed previously, plate anchor capacity may be conveniently expressed in terms of the
dimensionless bearing capacity factor, N, which is a function of the area of the plate and the
undrained shear strength of the soil at the plate anchor embedment depth. In order to determine
N, it is necessary to account for and quantify the loss of embedment (and hence su) associated
with the plate anchor keying process. However, quantifying this loss in embedment is not
straightforward as the points at which the anchor keying starts and finishes are unknown.
Identification of these points requires consideration of the load build up at each part of the
anchor-chain system, by examination of the response of the three load cells, the setup for which
are shown in Figure 3.12.
Figure 4.2 shows the load-time response for test VE-ST1 together with three stages identified by
considering the offsets in the load-time response in addition to the shape of the load-time curves.
During Stage �, the chain slack is recovered and the only load is that measured by the actuator
load cell due to the chain weight and friction in the pulleys. During Stage �, the vertical chain
cuts through the clay and develops an inverse catenary shape. Load develops progressively,
firstly on the chain load cell and soon after on the anchor load cell. After overcoming the
frictional resistance along the soil-chain interface, the anchor capacity starts to mobilise and the
rate of load development increases on all load cells. During Stage �, the plate continues to rotate
until the projected area reaches a maximum value, at which point the load reaches a peak and
starts to drop off as the plate enters weaker soil.
The slight reduction in gradient partway through Stage � is characteristic of observations made
by Song et al. (2005) from numerical simulations of the anchor keying process. Both jacked
anchors and suction embedded anchors show similar load responses as that shown in Figure 4.2.
For all tests, the start of keying has been determined using the graphical construction method
shown in Figure 4.2, and has been taken as the origin of the chain displacement for estimating
subsequent loss of embedment of the anchor.
The inclination of the chain loading further complicates the quantification of loss of embedment
during keying. As the load inclination is not vertical, the measured chain displacement must be
resolved into vertical and horizontal components. To do this the inclination of the chain at the
anchor padeye must be determined. Although the experimental arrangement was such that the
load inclination at the mudline was approximately 45°, the inclination of the chain at the anchor
padeye, θa, is expected to be higher due to the assumed inverse catenary profile of the chain in
the clay. θa was determined using the catenary shape theory (Neubecker & Randolph, 1995) and
the load recorded at each extremity of the chain, assuming a 45° chain inclination at the mudline.
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Note that this calculation assumes that the chain reaches the inverse catenary profile before the
anchor starts to key and that the load inclination remains constant during the pullout. Any
horizontal displacement of the anchor during the pullout will invalidate this assumption; this
simplification has little consequence for determining the loss in anchor embedment.
The upward movement of the plate anchor’s padeye during keying, dv, was then calculated as sc
× sin θa, where sc is the chain displacement during Stage � (i.e. between the start of anchor
keying and the peak load). This enabled estimation of the anchor embedment at failure and the
corresponding undrained shear strength obtained from the most representative T-bar profile.
Experimental bearing capacity factors were then calculated using:
uult
As
FN max= (4.3)
where Fmax is the peak load recorded at the anchor load cell and su is the undrained shear strength
at the estimated anchor embedment depth at the peak load.
Table 4.2 summarises the experimentally determined Nult factors together with other relevant
measured quantities.
Figure 4.3 and Figure 4.4 show the dimensionless load-displacement responses for jacked and
suction embedded anchors respectively, with the vertical displacement of the anchor, dv,
normalised by the fluke breadth, B. Note that the loads are normalised using a single value of su,
corresponding to the value at the estimated depth for peak load. Figure 4.5 displays embedment
loss as a function of the load inclination at the padeye.
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Table 4.2: Summary of SEPLA measurements
Test
Name
Time*
(s)
Anchor
capacity,
Fmax, (N)
Chain
displacement
(mm)
Load
inclination**
(°)
Loss of
anchor
embedment,
dv, (mm)
Anchor
embedment
ratio+
su++
(kPa)
Bearing
capacity
factor,
Nult
VE-ST1 1100 176.8 54.5 57 45.69 2.99 12.3 11.8
PE-ST2 920 166.3 49.0 59 41.97 3.04 12.6 10.9
VE-ST2 1191 174.8 52.4 57 43.96 2.98 12.2 11.7
VE-LT1 2013 193.1 46.2 55 37.85 3.11 12.3 12.2
PE-LT2 2131 190.3 40.7 51 31.62 2.49 12.0 12.9
VE-LT2 1815 181.5 45.9 54 37.10 2.55 12.1 12.2
JI-ST1 150 169.4 54.7 60 47.36 2.79 11.2 12.3
JI-ST2 120 169.1 58.8 62 51.88 2.66 10.6 13.1
JI-LT1 2200 175.6 54.3 61 47.53 2.79 11.2 12.8
JI-LT2 2200 182.0 57.5 60 49.81 2.72 10.9 13.5
∗ Time between anchor installation and load applied to anchor
∗∗ Load inclination at padeye determined from chain relationships
+ Mid-depth of plate normalised by fluke breadth, at peak load
++ Undrained shear strength at peak load (taken at mid-depth of anchor)
4.2 KEYING TESTS
The focus of the nine drum keying tests; four on a 30 x 30 mm plate and five on a 30 x 80 mm
plate was to determine the orientation of the different plate anchors during keying. The two
plates were keyed adjacent to a Perspex window and loaded vertically (see Figure 4.6) with
different eccentricities, over five sample boxes. It was assumed that, for the drum tests, the plane
of the loading cable remained constant (and vertical), while for the beam tests again it was
assumed that the angle of the loading cable below the pulley remained constant.
The 11 beam keying tests, over six sample boxes, had the same objective as the drum tests. Test
1 employed the 30 x 80 mm plate used in the drum tests. However, during the test the plate bent
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and did not yield any useful results. The remaining ten tests used a 20 x 80 mm plate under a
combination of vertical (relative to the plate centreline, see Figure 4.6) and inclined loads.
4.2.1 Soil Characterisation Tests
The shear strength profiles of the normally consolidated kaolin clay samples from each of the
five drum sample boxes and the six beam sample boxes are in Figure 4.7 to Figure 4.19. Similar
to the SEPLA tests the shear strength profile for normally consolidated soil increases linearly
with depth, z. Table 4.3 presents the profiles for the sample boxes from the drum testing, note the
distinct gradient change part way through the sample for four of the boxes. Table 4.4 shows the
linear shear strength profiles for the beam test samples.
Interestingly the shear strengths in the drum tests are significantly lower (~30%) than those in
the beam tests even though both had similar consolidation periods. This lower shear strength is
also approximately 30% lower than typical profiles of 1 to 1.3 kPa/m (prototype) seen in other
centrifuge tests conducted at UWA in normally consolidated kaolin clay. A contributing factor
may be the drainage paths for the two sets of tests. The drum tests had no felt mat placed in the
sample box, resulting in one-way drainage, whereas the beam tests did, resulting in two-way
drainage. This might also explain the layering observed in the drum samples and not the beam
samples. The drum keying tests were therefore all conducted in partially consolidated clay, with
an average, shear strength gradient of 0.7 kPa/m, over the set of tests. The beam tests used fully
consolidated clay (1 kPa/m), except for tests five and six that had a partially consolidated
sample, approximately 10% below typical strengths.
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Table 4.3: Summary of drum sample box shear strengths
Box Number Test Number Shear Strength Gradient
(prototype depth, z)
B2 (Figure 4.7) B2a80e30 0≤ z ≤8 su = 0.77z
8< z ≤11 su = 0.33z + 5.8
B3a30e5 B3 (Figure 4.8)
B3a30e15 su = 0.71z
B4a30e30 B4 (Figure 4.9)
B4a30e45
0≤ z ≤10 su = 0.66z
10< z ≤12 su = 0.4z + 7.45
B5a80e5 B5 (Figure 4.10)
B5a80e45
0≤ z ≤8.5 su = 0.73z
8.5< z ≤11.5 su = 0.54z + 5.1
B6a80e15 B6 (Figure 4.11)
B6a80e30
0≤ z ≤7.5 su = 0.66z
7.5< z ≤12 su = 0.61z + 4.5
Table 4.4: Summary of beam sample box shear strengths
Box Number Test Number Shear Strength Gradient
(prototype depth, z)
Test 1* B1 (Figure 4.13)
Test 2 su = 1 z
Test 3 B2 (Figure 4.14)
Test 4 su = 1 z
Test 5 B3 (Figure 4.15)
Test 6 su = 0.9 z
Test 7 B4 (Figure 4.16)
Test 8 su = 1 z
B5 (Figure 4.17) Test 9 su = 1.2 z
Test 10 B6 (Figure 4.18)
Test 11 su = 1.1 z
* Test Failed
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4.2.2 Drum Keying Tests
The primary focus of the keying tests was to obtain images and data allowing the orientation of
the plate anchor during keying to be analysed. The photographic images worked very well
allowing accurate plotting of the plate’s orientation for all stages of keying. Unfortunately, the
strain gauges within the loading arm malfunctioned and did not yield any usable load data. The
results from the drum tests were presented at the Sixth International Conference on Physical
Modelling (O’Loughlin et al. 2006).
Stepping through the digital camera images for each test generated the relationship between plate
orientation (inclination to the horizontal (θ)) and embedment loss (∆ze) (normalised by anchor
breadth (B)). Figure 4.20 to Figure 4.22 show images at different stages in the keying process for
an 80 mm plate anchor with e/B = 0.17, 0.5 and 1.0 (Figure 4.20, Figure 4.21 and Figure 4.22
respectively). Evidently the loss in anchor embedment for e/B = 0.17 (Figure 4.20) is much
higher than either e/B = 0.5 or e/B = 1.0. This is made clearer by Figure 4.23, which plots the
orientation of the plate anchor (in degrees from the horizontal) against the loss in anchor
embedment (∆ze) normalised by the anchor breadth, B.
The most striking observation on Figure 4.23 is the keying response for e/B = 0.17. In this test,
the plate anchor initially undergoes large vertical displacement with minimal plate rotation. At
∆ze/B ~1.65 the plate is inclined at 61° to the horizontal (rotation = 26°). At this point, the rate of
vertical displacement reduces suddenly, whilst the plate anchor continues to rotate. The plate
reaches a final orientation of 18° at ∆ze/B ~2.08. In contrast, tests with e/B = 0.5, 1.0 and 1.5 are
typified by almost complete plate rotation with minimal loss in embedment. The final plate
inclinations of 24°, 28° (average of two tests) and 24o
respectively, correspond to a normalised
embedment loss, ∆ze/B, of 0.28, 0.151 (average of two tests) and 0.115 respectively. The original
hypothesis that the plate would rotate through 90°, ending up perpendicular to the direction of
loading, proved incorrect as the final inclination of the plate was typically between 20° and 30°.
Figure 4.24 presents the same data for the 30 mm plate, interestingly these plates key at very
similar ∆ze/B as the 80 mm plates (for e/B = 1.0 and 1.5; ∆ze/B ~ 0.214 vs. 0.151 and 0.14 vs.
0.115 respectively). The angle at which this occurs is slightly higher (for e/B = 1.0 and 1.5; 34o
vs. 28° and 38o vs. 24
o). For e/B = 0.17 it appears that the smaller plate keys with a smaller ∆ze/B
(1.66 vs. 2.08) but the angle it reaches is significantly higher (48o vs. 18
o) meaning that it does
not rotate greatly. The major discrepancy on Figure 4.24 is test b3a30e15 (e/B = 0.5) where the
plate keys to an angle of 2o and has a significantly large ∆ze/B ~ 0.83. This is considered to be
because the underside of the L = 30 mm plate did not sustain tension at the clay-anchor interface
during the keying process and a cavity formed in the wake of the displaced anchor (see Figure
4.25). This is in contrast to the other tests where the clay-anchor interface sustained tension as
shown in Figure 4.20 to Figure 4.22.
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Table 4.5 summarises the different ∆ze/B with the e/B ratio for each test as well as presenting
values of dv/B, where dv is the displacement measured by the actuator during the test
corresponding to the padeye embedment loss (discussed further in Chapter 6). The anchor
embedment ratio, also presented, shows that the majority of the tests occur in the range of 2.75 –
3, which is sufficient to consider the anchor as being deep at failure (Rowe & Davis, 1982 and
Song & Hu, 2005). Noticeably the tests with e/B = 0.17 had significantly lower embedment
ratios (1.24 and 1.12) and hence can be considered shallow failures.
Table 4.5: Summary of drum keying tests
Test
Name e/B dv/B ∆ze/B
Anchor
embedment
ratio+
B2a80e30 1 1.13 0.195 2.74
B3a30e5 0.17 1.72 1.66 1.24
B3a30e15 0.5 1.35 0.83 2.34
B4a30e30 1 1.06 0.214 2.95
B4a30e45 1.5 1.65 0.14 2.99
B5a80e5 0.17 2.33 2.08 1.12
B5a80e45 1.5 - 0.115 2.75
B6a80e15 0.5 0.79 0.28 2.75
B6a80e30 1 0.98 0.107 2.89
+ Anchor embedment ratio at depth at which rotation ceases
4.2.3 Beam Keying Tests
Displacement
Continuing from the tests conducted in the drum, these tests focused on the keying of the 80 mm
plate. The primary focus of these tests was to obtain, not only the digital pictures, but also load
data for direct comparison with these images. From this set of tests the load data obtained was
very good and easily comparable with the imagery obtained. In the images it was not always
possible to see the plate against the Perspex however, careful examination of the movement of
soil around the area of the plate made it relatively simple to assess the plate’s position.
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Once again, stepping through the images generated relationships between the plate orientation
(θ) and the embedment loss (∆ze). Figure 4.26 to Figure 4.30 show these relationships for the
different e/B ratios of 0.25, 0.5, 0.75, 1.0 and 1.5 respectively. From these it is clear that the
plates, regardless of e/B, all key before a half plate width loss of embedment. Noticeably for e/B
= 0.25 (Figure 4.26) a small amount of embedment is lost (∆ze/B = 0.2) prior to any rotation of
the plate, with a total embedment loss of 0.42B. This is more than twice the embedment loss than
for e/B = 0.5 (Figure 4.27) with an average loss of 0.17B. Test 6 (in Figure 4.33) and test 10 (in
Figure 4.35) are of particular interest. In both of these tests, the plate appears to increase in
embedment during keying, prior to reaching maximum capacity and is not observed in any of the
other tests. This is not yet understood and requires further investigation.
It is clear from Figure 4.26 to Figure 4.30, that the plates during these tests rotate further than
those in the drum tests reaching final inclinations of almost 0o (rotation = 90
o) stopping at
approximately 2o for all the vertically loaded tests. The inclined tests also rotate to an angle very
close to that of the load orientation (Figure 4.27). Test 7 (70o load application to the horizontal)
rotates to 72o to the horizontal while test 9 (60
o load application to the horizontal) rotates to 62
o
to the horizontal, the same 2o short of a perpendicular orientation to the load observed in the
vertically loaded tests.
Along with ∆ze/B, Table 4.6 presents dv/B, which is the normalised displacement data from the
actuator. The actuator measures the displacement of the padeye rather than the centre of the
plate, which is the displacement referred to by all previous studies on plate anchor keying
(discussed further in Chapter 6). Table 4.6 also shows the anchor embedment ratio, with tests
occurring in the range of 4.4 – 5.6, which is again sufficient to consider the anchor as being deep
at failure (Rowe & Davis, 1982, Song & Hu, 2005)
Load
Unlike the drum tests, the beam tests yielded load data enabling comparison with the digital
images of the different tests. Table 4.6 presents the maximum load for each test while Figure
4.31 through Figure 4.35 and Figure 4.36 through Figure 4.40 show the normalised load (load /
area of anchor times shear strength at initial embedment depth) versus the loss of embedment,
∆ze/B, and plate inclination versus normalised load respectively. The Nc ranges from 9.02 to
11.51 with an average of approximately 10.5, consistent with theory as the actual embedment is
less than the initial embedment. Taking the embedment loss due to keying into account the Nc
range changes to 10.75 – 14.57, at an average of 13.14. This is similar to the theorised Nc range
of 12 – 13.
The plots of normalised load vs. plate inclination show that, other than the initial load increase,
the load does not increase until a certain degree of rotation has occurred. For e/B = 0.25 at 80o
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the load increases again and for e/B = 0.5, 0.75, 1.0 and 1.5 the plate reaches approximately 55o,
53o, 50
o and 38
o respectively prior to the load increasing again. This shows a trend of increasing
load eccentricities increasing the amount of rotation prior to an increase in load capacity.
Table 4.6: Summary of beam keying tests
Test
Name e/B
Load
inclination
(°)
dv/B ∆ze/B
Anchor
embedme
nt ratio+
Anchor
capacity
, Fmax,
(N)
su++
(kPa)
Nc1,
Initial
z
Nc2,
Actual
z
Test 2 0.5 90 0.99 0.19 5.56 189.04 11.5 10.27 12.7
Test 3 0.5 90 1.24 0.15 4.85 156.15 10 9.76 12.5
Test 4 1 90 1.42 0.1 5.11 184.32 10.4 11.08 13.9
Test 5 1 90 1.3 0.14 4.96 165.57 9.18 11.27 14.4
Test 6 0.75 90 1.28 -0.04 4.97 160.17 11.5 11.12 13.9
Test 7 0.5 70 0.57 0.08 4.42 135.64 9 9.02 11.0
Test 8 0.25 90 1.07 0.52 4.83 164.91 10.5 9.82 13.6
Test 9 0. 5 60 0.58 0.18 4.73 179.49 12 9.35 10.7
Test 10 1.5 90 1.9 -0.09 5.58 225.12 12.32 11.42 13.7
Test 11 0.75 90 1.04 0.13 5.02 208.69 11.33 11.51 14.5
+ Anchor embedment ratio at final keyed depth
++ Undrained shear strength at max embedment
1 With shear strength at initial embedment
2 With shear strength at mid-point of actual plate @ max load
4.2.4 Keying Test Summary
The loss in anchor embedment has been determined for each of the tests considered here and is
plotted on Figure 4.41 against the eccentricity ratio, e/B. With the exception of the breakaway
test, Figure 4.41 indicates no discernible difference (in terms of embedment loss) between
anchors with L = 80 mm and L = 30 mm, nor any difference between anchors with B = 20 mm
and B = 30 mm. In addition Figure 4.41 shows that for tests with an e/B ratio greater than 0.5
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there is very little embedment loss during keying. Nc ranges also appear to closely correlate to
theory.
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Figure 4.1: Clay, shear strength profile, SEPLA tests
Figure 4.2: Assumed load response during anchor keying and pullout for Test VE-ST1
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Figure 4.3: Dimensionless load displacement response for jacked SEPLA
Figure 4.4: Dimensionless load displacement response for suction embedded SEPLA
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Figure 4.5: Loss of embedment as a function of padeye load inclination), e/B = 0.66
Figure 4.6: Keying test load orientations
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Figure 4.7: Clay, shear strength profile, drum tests box 2
Figure 4.8: Clay, shear strength profile, drum tests box 3
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Figure 4.9: Clay, shear strength profile, drum tests box 4
Figure 4.10: Clay, shear strength profile, drum tests box 5
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Figure 4.11: Clay, shear strength profile, drum tests box 6
Figure 4.12: Drum tests, average shear strength profiles
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Figure 4.13: Clay, shear strength profile, beam tests box 1
Figure 4.14: Clay, shear strength profile, beam tests box 2
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Figure 4.15: Clay, shear strength profile, beam tests box 3
Figure 4.16: Clay, shear strength profile, beam tests box 4
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Figure 4.17: Clay, shear strength profile, beam tests box 5
Figure 4.18: Clay, shear strength profile, beam tests box 6
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Figure 4.19: Beam tests, average shear strength profiles
Figure 4.20: Stages of keying, drum test e/B = 0.17
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Figure 4.21: Stages of keying, drum test e/B = 0.5
Figure 4.22: Stages of keying, drum test e/B = 1.0
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Figure 4.23: Plate anchor rotation for drum tests, L = 80mm anchors & vertically loaded
Figure 4.24: Plate anchor rotation for drum tests, L = 30mm anchors & vertically loaded
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Figure 4.25: Stages of keying, drum test b3a30e15
Figure 4.26: Plate anchor rotation for beam tests, e/B = 0.25
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Figure 4.27: Plate anchor rotation for beam tests, e/B = 0.5
Figure 4.28: Plate anchor rotation for beam tests, e/B = 0.75
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Figure 4.29: Plate anchor rotation for beam tests, e/B = 1
Figure 4.30: Plate anchor rotation for beam tests, e/B = 1.5
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Figure 4.31: Nc vs. loss of embedment, e/B = 0.25
Figure 4.32: Nc vs. loss of embedment, e/B = 0.5
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Figure 4.33: Nc vs. loss of embedment, e/B = 0.75
Figure 4.34: Nc vs. loss of embedment, e/B = 1
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Figure 4.35: Nc vs. loss of embedment, e/B = 1.5
Figure 4.36: Plate inclination vs. Nc, e/B = 0.25
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Figure 4.37: Plate inclination vs. Nc, e/B = 0.5
Figure 4.38: Plate inclination vs. Nc, e/B = 0.75
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Figure 4.39: Plate inclination vs. Nc, e/B = 1
Figure 4.40: Plate inclination vs. Nc, e/B = 1.5
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Figure 4.41: Loss in plate anchor embedment during keying
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CHAPTER 5 ANALYTICAL SIMULATION
5.1 BACKGROUND
As mentioned in Chapter 1 a major issue concerning follower embedded plate anchors involves
their keying process. The inability to quantify the keying displacement and ultimately the final
capacity of the plate is one of the limitations of this method of anchorage. This study included
preliminary development of an analytical simulation that could predict the anchor keying process
under monotonic load conditions. The particular aim was to develop a method for accurately
determining the embedment loss and rotation of a plate in homogeneous, cohesive soil with
different shear strength gradients.
This chapter presents the method developed and the results of the preliminary simulation. With
further development, the simulation will assist in devising design specifications for follower
embedded plate anchors.
5.2 REVIEW OF NUMERICAL AND ANALYTICAL STUDIES OF
PLATE ANCHORS
The majority of studies concerning the uplift behaviour of embedded plate anchors in clay have
been limited to physical modelling or simple analytical solutions. O’Neill (2000) summaries
several of the numerical analyses conducted by Rowe & Davis (1982), Merifield, Sloan & Yu
(2001) and Colwill (1996).
In addition to these studies, Bransby & O’Neill (1999), O’Neill et al. (2001) and Elkhatib &
Randolph (2005) all discus the application of numerical studies to drag anchors in clay. They
discuss the use of FE analysis to investigate the interaction between anchor flukes and undrained
soil at failure. By examination of fluke-soil interaction, using yield loci and plastic potentials,
they approximate the drag anchor kinematics, for both rectangular and wedge shaped flukes.
Although previous studies were concerned with the application to drag anchors, the method and
parameters (for rectangular drag flukes), are suitable for the preliminary numerical analysis of
the keying response of an embedded plate anchor.
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5.3 PLASTICITY CONCEPTS AND THE YIELD LOCUS
The Bransby & O’Neill (1999) study of drag anchor fluke-soil interaction in clays can be
simplified and applied to the case of an embedded plate anchor. The introduction of a three-
dimensional, plastic yield locus allows failure analysis of the foundation to be quantified in terms
of the combined vertical (V), horizontal (H) and moment (M) loading. The yield locus, illustrated
in Figure 5.1, and expressed by a mathematical function of V, H and M, describes the combined
loading that will result in failure of an embedded footing, length, Lf, and depth, bf.
( ) 0,, =MHVf (5.1)
In addition to allowing the calculation of the capacity of the plate under these combined
conditions the locus also facilitates the calculation of plastic vertical, δv, horizontal, δh, and
rotational, δβ, displacements at failure, as seen in Figure 5.3 (O’Neill et al., 2001).
For follower embedded plate anchors the failure can be considered a ‘deep’ failure and as a
result the soil failure will be independent of the load direction and anchor orientation.
Additionally the deep condition ensures there will be no soil-anchor detachment, resulting in
plastic displacements governed by normality to the failure yield locus (O’Neill et al., 2001) and
allowing prediction of anchor displacement directions at failure.
The yield function proposed by Bransby & O’Neill (1999) was expressed as:
pnmq
HH
HH
MM
MM
VV
VVf
1
1max
1
1max
1
1max
1 1
−
−+
−
−+−
−
−= (5.2)
where the exponents q, m, n and p with offsets V1, H1 and M1 are quantified in Table 5.1. These
values have been derived from finite element analysis for an interface friction coefficient (α) of
0.4 between the plate and the soil. The yield locus for the rectangular fluke in V-H-M space is
shown in Figure 5.2.
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Table 5.1: Typical yield locus curve fitting parameters
Parameter
Rectangular Fluke
Lf/t = 20, α = 0.4
(Elkhatib & Randolph,
2005)
Rectangular Fluke
Lf/t = 7, α = 0.4
(Elkhatib & Randolph,
2005)
Rectangular Fluke
Lf/t = 7, α = 0.4
(Bransby & O’Neill,
1999)
Hmax/(Lfsu) 1.97 3.38 4.29
Vmax/(Lfsu) 11.58 11.78 11.87
Mmax/(Lf2su) 1.53 1.55 1.49
H1/(Lfsu) 0 0 0
V1/(Lfsu) 0 0 0
M1/(Lf2su) 0 0 0
m 1.52 2.58 1.26
n 5.31 3.74 3.72
p 1.01 1.09 1.09
q 2.75 1.74 3.16
5.4 KINEMATIC ANCHOR ANALYSIS
Application of the kinematic analysis for embedded plates is a simplified version of that applied
by O’Neill et al. (2001) on drag anchors. This preliminary analytical solution has ignored the
contribution of anchor self-weight, chain-soil interaction and shank-soil interaction but made
provision for their inclusions at a later stage. Figure 5.3 illustrates the geometry of the modelled
plate anchor. The model allows variation in soil strength profile although for the preliminary
simulation α has been assumed as 0.4, so yield locus parameters do not require variation.
The centre of the plate is assumed to be in undisturbed soil at an embedment depth da, below the
soil surface with the top face angled at β to the vertical. Following the flowchart in Figure 5.4,
Vmax is the resistance normal to the plate, Hmax is the resistance parallel to the plate and Mmax the
rotational resistance. V, H and M are calculated by multiplying the undrained shear strength, su,
and Lf with the appropriate yield locus parameter. Once V, H and M have been determined vary
Ta until f(V,H,M) = 0.
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The fluke movement is then determined using the equilibrium solution:
V
f
H
f
h
v
∂
∂
∂
∂=
δ
δ (5.3)
and
( ) H
f
LM
fLh
f
f
∂
∂
∂
∂=
δβ
δ (5.4)
Moving the plate an incremental distance, ∆v, in the direction parallel to the plate surface results
in an incremental displacement perpendicular to the plate surface, ∆h, determined by:
vH
f
V
fh ∆
∂
∂
∂
∂=∆ (5.5)
and rotational displacement ∆β determined by:
( ) ff L
v
H
f
LM
f ∆
∂
∂
∂
∂=∆β (5.6)
Choosing ∆v, then allows the simulation to update the anchors position and loop the analysis
procedure until the β reaches 90 degrees or the plate exits the soil. From the results obtained for
this simulation the plate’s trajectory and the loads acting on it can be determined.
5.5 RESULTS
Figure 5.5 shows the loss in plate anchor embedment during keying from the analytical
simulation (using the parameters from Table 5.1) compared with values measured during testing.
It is clear from this that the most appropriate set of parameters for application to embedded plate
anchors are those for Lf/t = 7 which closely match Lf/t = 10 for the rectangular plate used during
our physical modelling tests. It is also obvious that the results are sensitive to the Hmax/(Lfsu)
parameter, which is the only significant difference between the Elkhatib & Randolph and
Bransby & O'Neill input parameters for the case of Lf/t = 7.
Thus using Bransby and O’Neill (1999) parameters, a linearly increasing shear strength profile
equal to 1 kPa/m and α = 0.4 as inputs, the simulation produces graphs as shown in Figure 5.6,
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Figure 5.7 and Figure 5.8. Chapter 6 presents the comparison of these results with those obtained
during experimental modelling.
Assumptions and steps used to simplify this initial simulation included:
1. The anchor was assumed to be weightless;
2. The effect of the chain system was removed by applying Ta directly to the padeye; and
3. The effect of the soil-shank interaction was removed, as its effect on the overall problem
was considered negligible.
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Figure 5.1: The yield locus and plasticity potential function (Bransby and O'Neill, 1999)
Figure 5.2: V-H-M yield locus for rectangular fluke (Bransby and O'Neill, 1999)
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Figure 5.3: Kinematic analysis sign convention
Figure 5.4: Analysis flowchart for kinematic anchor simulation using yield locus
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Figure 5.5: Loss in plate anchor embedment during keying – analytical simulation
Figure 5.6: Normalised embedment loss vs. normalised load -analytical simulation
Symbols – experimental data
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Figure 5.7: Angle of inclination vs. normalised embedment loss - analytical simulation
Figure 5.8: Plate inclination vs. normalised load – analytical simulation
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CHAPTER 6 DISCUSSION OF THEORETICAL AND
EXPERIMENTAL RESULTS
6.1 SEPLA TESTS
With the sole exception of test PE-LT2, the shape of the load-displacement curves in Figure 4.3
and Figure 4.4 conform to that of Figure 4.2. Numerical analysis (Song et al. 2005) provides a
similar response, which is initially quite stiff as the anchor begins to rotate. This is followed by a
softer phase as the rotation angle increases and a final stiff response as the anchor capacity is
fully mobilised.
When considering the magnitude of the experimental Nult factors in Figure 4.3, Figure 4.4 and
Table 4.2, it should be noted that despite the significant loss of padeye embedment during keying
(0.9 - 1.5 times the plate height), the final anchor embedment ratio lies in the range 2.5 - 3.1,
sufficient to ensure deep failure (Rowe & Davis, 1982 and Song & Hu, 2005). Supporting this is
the observation that the clay surface appeared intact after the anchor reached the peak load,
indicating that the failure mechanism did not extend to the surface but was localised around the
anchor.
Figure 4.3 and Figure 4.4 allow for direct comparisons between jacked and suction embedded
anchors and between short and long-term capacity. This comparison leads to the following
comments:
• The load-displacement responses in Figure 4.3 and Figure 4.4 are typified by an inflection
point at a vertical translation of half the plate height (dv/H = 0.5). The corresponding
normalised load at inflection is approximately 5 for the jacked anchors compared with
approximately 3 for the suction embedded anchors. Although not completely understood, this
suggests that the suction installation process softens the soil near the anchor (i.e. at the
caisson tip), resulting in a lower proof load to initiate keying. This potential advantage
appears to be lost as the plate anchor embedment reduces during keying into undisturbed
clay.
• The loss of padeye embedment at the peak capacity is between 1.3B and 1.5B for jacked
anchors in comparison to 0.9B - 1.3B for suction embedded anchors, for loading angles
ranging between 51o and 62
o (to the horizontal) and an e/B ratio of 0.66. The loss of padeye
embedment deduced from these tests is within the wide range reported in the literature for
vertical load inclinations (0.65B – 2B). Although the embedment loss during keying appears
to be lower for suction embedded anchors than jacked anchors, plotting this data against the
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load inclination indicates a much stronger correlation, well described by a linear relationship
(see Figure 4.5). Extending the linear fit to these test data to limiting load inclinations of 40°
and 90° corresponds to an embedment loss of 0.4B - 2.8B. Whilst this range is in broad
agreement with that reported in the literature, the much lower loss in embedment for strip
anchors reported by Song et al. (2005), reproduced on Figure 4.5, suggests that the
embedment loss is not only a function of load inclination but may also be dependent on plate
geometry, or other facets not captured by the finite element analysis.
• Bearing capacity factors for jacked anchors are in the range 12.3 - 13.5. Although no
published numerical solutions are available for deeply embedded square plate anchors, values
of 13.11 (Martin & Randolph 2001) and 14.31 (Song & Hu 2005) have been obtained for
rough circular anchors. Applying the Nc,square/Nc,circle = 0.947 factor reported by Merifield et
al. (2003) gives Nult = 13.55 and Nult = 12.42 for rough square anchors, which is in excellent
agreement with the experimental range of Nult = 12.3 - 13.5. It is important to note that the
effect of load inclination on Nc is minimal for anchor embedment ratios greater than 3 (Song
et al. 2005).
• Bearing capacity factors for suction embedded anchors are in the range 10.9 - 12.9,
approximately 8 % lower than the corresponding range for jacked anchors. This range can be
further categorised into Nult = 10.9 - 11.8 for short-term capacity and Nult = 12.2 - 12.9 for
long-term capacity. This is in contrast to the jacked anchor test results for which there is no
discernable difference between short term and long term bearing capacity factors (although
the time taken to accelerate the centrifuge precludes short term measurement). The range of
Nult for long-term capacity of suction embedded anchors (12.2 - 12.9) agrees well with the
adjusted numerical range (12.42 - 13.55), indicating that the suction installation process
reduces the short-term capacity. It may be noted that non-dimensional time factors, T =
tch/D2, range from ~0.25 (prototype time of ~0.7 years) for the short-term tests to 0.5
(prototype time of 1.5 years) for the long-term tests.
6.2 PLATE KEYING
6.2.1 Capacity
Normalising the maximum load with respect to the undrained shear strength (at peak load) and
the anchor’s projected area allows comparison with theoretical breakout factors, Nc. Figure 6.1
and Table 6.1 show the normalised maximum loads (minus submerged anchor weight, Table 6.2)
and anchor embedment ratios (H/B) for each of the infinite strip (L/B = ∞) tests conducted on the
beam centrifuge. Test 11 (presented later in Figure 6.4) shows a typical load vs. keying (or plate
rotation) analysis. From this, it is clear that the maximum load coincides with the end of keying,
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or rotation of the plate. Figure 6.1 compares the test results with theoretical results of Merifield
et al. (2003). It shows that although the plate model in the test intended to model an infinitely
long plate the initial assessment indicates that it fits closely with an aspect ratio (L/B) of 2.
Table 6.1: Load analysis
Test
Name e/B (H/B)
+ Fmax, (N) Nc Nc
*
Corrected
Nc**
Test 2 0.5 5.56 189.04 10.62 9.32 8.76
Test 3 0.5 4.85 156.15 10.05 8.56 7.92
Test 4 1 5.11 184.32 11.28 9.79 9.18
Test 5 1 4.96 165.57 11.59 9.89 9.19
Test 6 0.75 4.97 160.17 11.05 9.41 8.72
Test 7 0.5 4.42 135.64 9.17 7.60 6.93
Test 8 0.25 4.72 164.91 10.90 9.39 8.73
Test 9 0. 5 4.82 179.49 9.70 8.44 7.90
Test 10 1.5 5.58 225.12 11.24 9.98 9.48
Test 11 0.75 5.02 208.69 11.81 10.46 9.90
+ Embedment ratio at peak load.
++ su at max load embedment.
∗ Nc with anchor weight (Table 6.2) subtracted only.
** Nc allowing for a frictional resistance of 10 N.
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Table 6.2: Model anchor submerged weights
e/B Weight @ 100g
(N)
0.25 22.72
0.5 23.21
0.75 23.76
1.0 24.31
1.5 25.31
An aspect to consider when comparing these results with previous studies is that the frictional
forces developed between the plate/Perspex interfaces are difficult to quantify. Tests conducted
at 1 g indicated that this frictional force is negligible, ranging between 0 and 10 N (prototype)
however, the influence of the increase g level and clay between the interfaces is unknown. By
correcting for the friction effect, an average Nc value of 8.98 is obtained with all H/B values
greater than 4, thus allowing each to be considered deep. This results in a range of Nc values for
the test being 8.67 to 9.28, ~ 30% lower than the upper bound Nc of ~ 12 for deep, infinitely long
strips presented by Merifield (2002).
The large difference could be because the anchor chain was straight at commencement of
loading, possibly resulting in the ‘zero’ load reading taken from the load cell already
incorporating the weight of the anchor. If this were the case, the Nc value would increase to 10.8,
which is 10% lower than the Merifield (2002) upper bound Nc of ~ 12.
6.2.2 Keying
Keying results from tests 6 and 10 exhibit strange embedment loss profiles (shown in Figure
4.28 and Figure 4.30). All the other tests showed a decrease in embedment as the load develops
while these two showed an increase in embedment. This unexplained phenomenon requires
additional research in order to determine its significance but is likely to be due to setup/technical
problems, almost certainly friction related problems at the anchor/Perspex interface.
Figure 6.2 through Figure 6.6, show variations of rotation, embedment and load development
obtained from the analyses. These show clearly that the maximum load coincides with
completion of plate rotation, but it is also important to analyse how the rotation changes as the
load builds up to the maximum value. Looking at extreme e/B ratios (0.25 and 1.5) differences in
the keying process for high and low eccentricities become apparent.
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For low eccentricities (e/B = 0.25), a load of 44 N is required to initiate plate rotation,
approximately 27% of the peak load. Prior to rotation commencing the plate embedment loss is
approximately 0.2B, approximately 40% of the total embedment loss for this case. For high
eccentricities (e/B = 1.5) a load of 11.4 N (6% of the peak load) is required which is significantly
lower than for low eccentricities. Additionally, there is no (observed) plate embedment loss prior
to rotation for high eccentricities, unlike the low eccentricity cases. Peak loads for the extreme
cases also differ significantly from 164.9 N to 225.1 N for low and high cases respectively, a
difference of ~ 27%. This is expected due to the large embedment loss difference between cases
(~ 0.5 and ~ 0 for low and high eccentricities respectively). Looking at the Nc values for the two
cases shows that this is in fact the case as they only differ by 6% (9.39 and 9.98 for low and high
eccentricities respectively).
Additionally, comparison of high and low eccentricities allows a prediction of failure
mechanisms, although verification of these by Practical Image Velocimetry (PIV, White et al.
2003) analysis is required. There are four stages suggested for the keying process from initial
load application to failure (or maximum load). The first stage is the application of load during
which there is no failure mechanism. The second stage is a purely rotational failure of the plate
around its centre, with no vertical or horizontal translation. There is then a transition stage where
the plate undergoes small amounts of vertical and horizontal movement while the dominant
displacement remains rotational, before the final stage where the plate is only displaced
vertically at which point a deep strip failure mechanism is mobilised.
Tests at low eccentricity did not show any purely rotational mechanisms. Embedment appeared
to reduce upon commencement of rotation, indicating the transition stage started immediately. At
high eccentricities, it is possible to identify each of the four stages (Figure 6.7).
Plate embedment loss during keying ranges from ∆ze/B = 0.0 – 2.1, with the upper and lower
limits corresponding to e/B = 0.17 and 1.5 respectively. The combined data from the two sets of
tests indicates that minimal embedment loss occurs during the keying process when e/B is
greater than 0.5. The practical advantages of having the padeye eccentricity greater than half the
plate width are two fold. Not only is there a significantly lower embedment loss during keying
but also the load required to initiate rotation is lower and although the load required to complete
plate rotation (~90o) increases slightly, the load required to reach 50% rotation (45
o) reduces
with increasing e/B ratios. Table 6.3 demonstrates the reducing load required for the initial 45o
rotation and compares it to the load required for ~90o of rotation.
The development of shear (H), normal (V) and moment (M) loads during keying must be
considered to explain keying behaviour. Once plastic deformation commences, with the
combination of V, H and M loads lying on the yield locus, the displacement of the plate will be
normal to the yield locus (Murff 1994; Bransby & Randolph 1998).
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As the plate is subjected to an eccentrically applied vertical load, the plate displaces and rotates.
For high eccentricity applied loads, the plate will be subjected to a high moment (M) and
commence rotation at a relatively low applied load. Additionally, with the plate initially vertical,
this low applied load will result in low H, hence the initial point on the yield locus for the high
eccentricities is at low H and high M (and zero Fn). During rotation, the effective eccentricity
will decrease causing M to reduce while V increases and H decreases.
Conversely, low eccentricities initially required a large applied load to generate sufficient
moment to initiate plate rotation. Subsequently the starting point on the yield surface is at higher
H and lower M (and zero V). This will result in slow plate rotation reducing M further while the
dominate force changes from H to V.
Normality requires the plastic displacements to be normal to the yield surface, thus the suggested
loading path reveals that displacements are principally normal to the plate. This shows that for
high eccentricities, very little embedment loss occurs during keying and the vertical
displacement of the plate will only occur once the plate is normal to the load or when keying
concludes, which is consistent with the test results. For low eccentricities, inspection of the yield
loci reveal that displacements are predominantly parallel to the plate, which corresponds with the
high embedment loss observed in the low e/B centrifuge tests. Figure 6.8 shows the low and high
eccentricity loading paths.
Table 6.3: Normalised load required for various stages of rotation
e/B Load to reach
45o (N)
% of final load
@ 45o
Load to reach
~ 90o (N)
0.25 57 35% 165
0.5 36 22% 165
0.75 32 17% 184
1.0 30 17% 175
1.5 22 10% 225
While comparing the two series of tests (drum and beam) it is important to remember the slight
load orientation differences. The drum tests applied loads vertically over the padeye while the
beam tests applied loads vertically over the plate’s centreline (shown in Figure 4.6). An
additional difference is that the drum tests used a rigid loading arm whilst the beam tests used a
flexible chain.
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In comparing the strip tests (Table 6.4) the combination of loading angle and rigid loading arm
used in the drum appears to limit the rotation of the plate, preventing it from reaching the near
normal load orientation to the load observed in the beam testing. However, it is interesting to
note that ∆ze/B for equivalent e/B ratios from the different loading methods are comparable.
Table 6.4: Drum vs. Beam tests
e/B
Final Keyed
Angle – Drum
(~ o)
Final Keyed
Angle – Beam
(~ o)
∆ze/B – Drum ∆ze/B – Beam
(Average)
0.17 18 - 2.08 -
0.25 - 0 - 0.42
0.5 24 0 0.28 0.17
0.75 - 0 - 0.05
1.0 25 0 0.107 0.12
1.5 28 0 0.115 -0.09
It is difficult to compare directly the results from this study with those of previous studies. Song
et al. (2005 & 2006) presented an embedment loss of 0.6B for square anchors (determined
experimentally) and 0.65B for strip anchors (determined numerically), while Wilde et al. (2001)
showed an embedment loss range of 0.5 to 1.7, for e/B ratios of 0.62 and 0.5 respectively when
vertically loaded. Problems arise when comparing the results with these studies because it is
uncertain whether the embedment loss reported is for the centre of the plate (∆ze/B) or the padeye
(dv/B). Assuming they are in terms dv/B correction for the padeye eccentricity is required to
obtain ∆ze/B results and thus allow comparison.
Assuming a correction is required and the plate’s final orientation is normal to the applied load,
then Song et al. (2005 & 2006) results would become 0B, for square anchors, and 0.05B for strip
anchors, for e/B = 0.62, while the results of Wilde et al. (2001) would become 0 – 1.2, for e/B =
0.5. These are now similar to the results from this study. Additionally, using the same
assumptions to correct the SEPLA test results (described in section 6.1) gives ∆ze/B = 0.24 – 0.84
for an e/B = 0.66. This indicates that the NCEL (1985) guidelines, which suggest embedment
loss is twice the anchor height in cohesive soils, are perhaps over cautious.
6.2.3 Comparison with Analytical Simulation
Initial comparison of physical modelling results with the analytical simulation (using Bransby &
O’Neill (1999) parameters described in Chapter 5) indicates a high correlation. Comparisons of
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four aspects are possible. Plate inclination vs. normalised load (Figure 6.9 to Figure 6.13) shows
the best agreement between the simulation and physical modelling results. The simulation
follows the path of the results from physical modelling closely until the final 10% of plate
rotation (~10o to 0
o) where the simulation over predicts the ultimate load. This could be due to
remoulding of clay adjacent to the anchor during keying in the physical tests weakening the clay.
This analytical simulation does not consider this phenomenon.
For ∆ze/B vs. plate inclination (Figure 6.14 to Figure 6.18) the simulation is closely matched by
the physical modelling plots with the exception of tests 6 and 10, which unexplainably show an
increase in embedment during keying. In addition, for e/B = 0.25 the simulation shows no
embedment loss prior to the start of rotation, which contrasts with what was observed in the
physical tests. In contrast, the simulated development of load with respect to embedment loss
(normalised load vs. ∆ze/B, Figure 6.19 to Figure 6.23) shows only moderate agreement with the
test results.
The most important aspect of comparison is e/B vs. ∆ze/B. Figure 6.24 shows very clearly that
the results from the analytical simulation agree with those obtained though physical modelling.
This allows the prediction of embedment loss for a given load eccentricity to be undertaken with
a greater degree of confidence, given that physical tests have been validated by theoretical or
analytical solutions.
One limitation of the simulation is that it does not yet predict the behaviour of plates once they
have achieved normality with the applied load, which can be seen by the simulation plots ending
abruptly and not showing the behaviour of the plate after the peak load, or 90o has been reached.
This aspect requires rectifying so that the full behaviour during keying and subsequent loading
can be simulated.
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Figure 6.1: Test Nc comparison with Merifield et al. (2003)
Figure 6.2: Keying analysis, e/B = 0.25
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Figure 6.3: Keying analysis, e/B = 0.5
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Figure 6.4: Keying analysis e/B = 0.75
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Figure 6.5: Keying analysis, e/B = 1.0
Figure 6.6: Keying analysis, e/B = 1.5
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4
3
1
2
Transition
2
4
1
3
1
2
3
4
No Mechanism
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Figure 6.7: Keying mechanisms
1.0
0.8
0.6
0.4
0.2
0.00.0 0.2 0.4 0.6 0.8 1.0
high eccentricity
loading path
low eccentricity
loading path
M/Mmax
= 0.99
M/Mmax
= 0
Fn/F
n,max
Fs/F
s,m
ax
Figure 6.8: Combined loading paths for high and low eccentricity plate anchors
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Figure 6.9: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.25
Figure 6.10: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.5
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Figure 6.11: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.75
Figure 6.12: Plate inclination vs. Nc comparison with analytical simulation, e/B = 1.0
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Figure 6.13: Plate inclination vs. Nc comparison with analytical simulation, e/B = 1.5
Figure 6.14: Plate anchor rotation comparison with analytical simulation, e/B = 0.25
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Figure 6.15: Plate anchor rotation comparison with analytical simulation, e/B = 0.5
Figure 6.16: Plate anchor rotation comparison with analytical simulation, e/B = 0.75
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Figure 6.17: Plate anchor rotation comparison with analytical simulation, e/B = 1.0
Figure 6.18: Plate anchor rotation comparison with analytical simulation, e/B = 1.5
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Figure 6.19: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.25
Figure 6.20: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.5
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Figure 6.21: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.75
Figure 6.22: Nc vs. loss of embedment comparison with analytical simulation, e/B = 1.0
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Figure 6.23: Nc vs. loss of embedment comparison with analytical simulation, e/B = 1.5
Figure 6.24: e/B vs. ∆ze/B
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CHAPTER 7 CONCLUSION AND FURTHER RESEARCH
7.1 EXPERIMENTAL FINDINGS
7.1.1 SEPLA Testing
A series of centrifuge tests conducted to investigate how suction installation, caisson retrieval
and plate anchor keying affect the behaviour of suction embedded plate anchors have been
completed. The main findings have been published in Gaudin et al. (2006). Those relating to the
keying of SEPLAs are:
1. Jacked plate anchor bearing capacity factors are in the range 12.3 - 13.5, which is in
excellent agreement with existing numerical solutions. Any soil strength reduction due to
driving the plate anchors was regained during accelerating the centrifuge and subsequent
short and long consolidation periods. By contrast, the range of bearing capacity factors
for suction embedded anchors is 10.9 - 12.9, with the lower values corresponding to
anchors tested after a short consolidation period. Evidently, the suction installation
process reduces short-term anchor capacity.
2. The tests reported here demonstrated a range of padeye embedment loss (dv/B) of 0.9B –
1.5B, corresponding to a plate centre embedment loss (∆ze/B) of 0.24B – 0.84B , which is
in agreement with values reported in the literature. However, it is demonstrated that the
loss of embedment is strongly correlated with the padeye chain inclination. The loss of
embedment and hence potential anchor capacity may be minimised by keying the plate
anchor at lower load inclination angles.
Considering the results obtained, it is evident that rectangular plate anchors (with the smaller
dimension in the vertical plane) will not lose as much embedment as an equivalent area square
anchor. However, rectangular plate anchors provide slightly less capacity per unit area than
square anchors (11.42 vs. 13 respectively Merifield et al. 2003), the difference being a function
of the plate aspect ratio (L/B). Evidently an optimal aspect ratio exists which maximises anchor
capacity whilst minimising the reduction in anchor embedment during keying.
7.1.2 Plate Keying Tests
Centrifuge model tests have successfully investigated the keying characteristics of embedded
plate anchors in normally consolidated clay. The following conclusions have been drawn:
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1. Plate embedment loss during keying ranges from ∆ze/B = 0.0 – 2.1, with upper and lower
limits corresponding to load eccentricities (e/B) of 0.17 to 1.5 respectively, with
embedment loss increasing sharply for e/B ratios < 0.25, for vertically applied loads
(shown in Figure 6.24).
2. During keying, the maximum load coincides with the completion of plate rotation. For
low load eccentricities (e/B = 0.25) peak loads are approximately 27% lower than for
high load eccentricities (e/B = 1.5), 164.9 N vs. 225.1 N respectively. The difference is
due to the difference in embedment loss during keying and thus the difference in shear
strength at failure.
3. The embedded strip plate anchors, for deep failure, have Nc values in the range 7.6 –
10.46, averaging 9.28. These results are ~ 30% lower than existing results for embedded
strip plates.
4. As expected, the load required to initiate plate rotation was significantly higher at low e/B
ratios than at high values. The load for low e/B is 44 N (~ 27% of peak load) compared
with 11.4 N (~ 6% of peak load) for high e/B. In addition, the load required to complete
50% of the rotation, i.e. 0 – 45o, is significantly higher for low e/B, being 35% of the
peak load for the low e/B case in comparison to 10% of the peak load for high e/B.
In addition, these physical modelling results were verified by results from a simple analytical
simulation that showed good agreement with physical test results. Most noticeably comparing
the analytical simulation results for load eccentricity (e/B) vs. embedment loss during keying
(∆ze/B) with the results from the physical modelling, as shown in Figure 6.24..
7.2 RECOMMENDATIONS FOR FUTURE DEVELOPMENT
Although this study has significantly enhanced the understanding of keying characteristics for
follower embedded plate anchors several aspects require further attention. The following areas
are required to complete understanding of the keying process and the factors contributing to
keying behaviour.
1. Incorporate different anchor geometries to find an optimal aspect ratio exists to maximise
anchor capacity.
2. Further investigation on the different installation methods and their effects on keying
behaviour and ultimate capacities to complement the SEPLA results.
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3. Quantifying the contribution of the post installation consolidation period to anchor
capacity, to assess the significance of the time between anchor keying and loading on the
ultimate anchor capacity.
4. Measurement of vertical anchor capacities under sustained monotonic and cyclic loading
conditions to develop understanding of embedded plate anchor behaviour under different
loading conditions.
5. More extensive investigation of the influence of load inclination on keying covering a
range of practical values, which may be as low as 30o to the horizontal.
6. Further development of the trajectory model (Chapter 5) to incorporate load angle,
varying soil strength profile, influence of anchor weight, influence of anchor/shank
interaction, influence of the chain, and load conditions.
7.3 CONCLUDING STATEMENT
In terms of the practical application of embedded plates as anchors for floating offshore
facilities, Figure 6.24 presents e/B vs. ∆ze/B, possibly the most important summary of results
from this study. It indicates that current guidelines, stating embedment loss during keying is
twice the anchor height (B) in cohesive soils, are extremely conservative given typically padeye
eccentricities (e/B < 0.5). These results have indicated that for typical embedded plate anchors
the embedment loss is < 0.3B.
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APPENDICES