installation and keying of follower embedded plate anchors · school of civil and resource...

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



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.



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.



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



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


4.1.2 SEPLA Capacities .................................................................................... 4-2

4.2 Keying Tests ............................................................................................................. 4-4


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



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



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



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



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



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 4.10: Clay, shear strength profile, drum tests box 5......................................................4-16



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 4.19: 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



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 6.7: Keying mechanisms..............................................................................................6-14



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 6.14: 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 6.19: 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 6.24: e/B vs. ∆ze/B .......................................................................................................6-22

Installation and Keying of Follower School of Civil and Resource Engineering

Embedded Plate Anchors The University of Western Australia

1-1

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



1-2

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).



1-3

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.



1-4

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



1-5

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.



1-6

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).



1-7

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


� Little to no experience in the

operation of the SEA.

� Requires keying and proof

loading.




1-8

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.



1-9

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.



1-10

Figure 1.1: Offshore production facilities (Courtesy: Minerals Management Services)

Figure 1.2: Offshore production facilities (Courtesy: Minerals Management Services)



1-11

Figure 1.3: Mooring types (Vryhof, 2000)

Figure 1.4: Drag anchors, sand and clay (Vryhof, 2000)



1-12

Figure 1.5: Suction caisson

Figure 1.6: DPA and Torpedo anchor and installation schematic (Lieng et al. 1999 and Medeiros, 2001)



1-13

Figure 1.7: SEPLA installation technique

Figure 1.8: SEA installation technique



2-1

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.



2-2

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)



2-3

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.



2-4

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)



2-5

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)



2-6

Figure 2.5: Upper Bound Nc values for horizontal anchors - inhomogeneous cohesive soil (Merifield, 2002)



3-1

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.



3-2

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.



3-3

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.



3-4

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.



3-5

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.



3-6

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



3-7

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.



3-8

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



3-9

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:



3-10

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.



3-11

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



3-12

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.



3-13

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



< 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.



3-14

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.



3-15

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.



3-16

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.



3-17

Figure 3.1: Geotechnical beam centrifuge

Figure 3.2: Beam strongbox



3-18

Figure 3.3: Motor driven actuator

Figure 3.4: T-bar penetrometer



3-19

Figure 3.5: Drum centrifuge with clamshell removed

Figure 3.6: Keying test setup in sample box, @ 1 g



3-20

Figure 3.7: Digital camera cradle with trigger

Figure 3.8: Tool table actuator



3-21

Figure 3.9: Sub-Terrain Oil impregnated Multiple Pressure Instrument (STOMPI)

Figure 3.10: Mounted model SEPLA



3-22

Figure 3.11: SEPLA test setup



3-23

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)



3-24

Figure 3.13: Caisson with attachments (pneumatic valve hidden behind caisson guide)

Figure 3.14: SEPLA test, actuator setup



3-25

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.



3-26

Figure 3.17: Drum keying test, loading arm

Figure 3.18: Drum keying test layout (White, 2003)



3-27

Figure 3.19: Sample box held in place with brackets

Figure 3.20: Beam keying test configuration



4-1

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.


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



4-2

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.



4-3

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.



4-4

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



4-5

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.


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.



4-6

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


B4a30e45

0≤ z ≤10 su = 0.66z

10< z ≤12 su = 0.4z + 7.45


B5a80e45

0≤ z ≤8.5 su = 0.73z

8.5< z ≤11.5 su = 0.54z + 5.1


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 6 su = 0.9 z


Test 8 su = 1 z

B5 (Figure 4.17) Test 9 su = 1.2 z


Test 11 su = 1.1 z

* Test Failed



4-7

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.



4-8

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.



4-9

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



4-10

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



4-11

there is very little embedment loss during keying. Nc ranges also appear to closely correlate to

theory.



4-12

Figure 4.1: Clay, shear strength profile, SEPLA tests

Figure 4.2: Assumed load response during anchor keying and pullout for Test VE-ST1



4-13

Figure 4.3: Dimensionless load displacement response for jacked SEPLA

Figure 4.4: Dimensionless load displacement response for suction embedded SEPLA



4-14

Figure 4.5: Loss of embedment as a function of padeye load inclination), e/B = 0.66

Figure 4.6: Keying test load orientations



4-15

Figure 4.7: Clay, shear strength profile, drum tests box 2




4-16





4-17


Figure 4.12: Drum tests, average shear strength profiles



4-18

Figure 4.13: Clay, shear strength profile, beam tests box 1




4-19





4-20





4-21

Figure 4.19: Beam tests, average shear strength profiles

Figure 4.20: Stages of keying, drum test e/B = 0.17



4-22





4-23

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



4-24

Figure 4.25: Stages of keying, drum test b3a30e15

Figure 4.26: Plate anchor rotation for beam tests, e/B = 0.25



4-25





4-26

Figure 4.29: Plate anchor rotation for beam tests, e/B = 1




4-27

Figure 4.31: Nc vs. loss of embedment, e/B = 0.25




4-28


Figure 4.34: Nc vs. loss of embedment, e/B = 1



4-29


Figure 4.36: Plate inclination vs. Nc, e/B = 0.25



4-30





4-31

Figure 4.39: Plate inclination vs. Nc, e/B = 1




4-32

Figure 4.41: Loss in plate anchor embedment during keying



5-1

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.



5-2

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.



5-3

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.



5-4

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,



5-5

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.



5-6

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)



5-7

Figure 5.3: Kinematic analysis sign convention

Figure 5.4: Analysis flowchart for kinematic anchor simulation using yield locus



5-8

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



5-9

Figure 5.7: Angle of inclination vs. normalised embedment loss - analytical simulation

Figure 5.8: Plate inclination vs. normalised load – analytical simulation



6-1

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



6-2

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,



6-3

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.



6-4

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.



6-5

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).



6-6

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.



6-7

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



6-8

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.



6-9

Figure 6.1: Test Nc comparison with Merifield et al. (2003)

Figure 6.2: Keying analysis, e/B = 0.25



6-10




6-11

Figure 6.4: Keying analysis e/B = 0.75



6-12





6-13

4

3

1

2

Transition

2

4

1

3

1

2

3

4

No Mechanism



6-14

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



6-15

Figure 6.9: Plate inclination vs. Nc comparison with analytical simulation, e/B = 0.25




6-16





6-17


Figure 6.14: Plate anchor rotation comparison with analytical simulation, e/B = 0.25



6-18





6-19





6-20

Figure 6.19: Nc vs. loss of embedment comparison with analytical simulation, e/B = 0.25




6-21





6-22


Figure 6.24: e/B vs. ∆ze/B



7-1

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:



7-2

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.



7-3

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.



7-1

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7-5

APPENDICES

installation and keying of follower embedded plate anchors · school of civil and resource...

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