Frontiers in Offshore Geotechnics: ISFOG 2005 – Gourvenec & Cassidy (eds)© 2005 Taylor & Francis Group, London, ISBN 0 415 39063 X

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1 INTRODUCTION

The purpose of this paper is to review recent researchwork on the design of suction caisson foundations foroffshore wind turbines. Most of the relevant work hasbeen conducted at, or in co-operation with, the univer-sities of Oxford and Aalborg, so we report heremainly the work of our own research groups.

Suction caissons have been extensively used asanchors, principally in clays, and have also been used asfoundations for a small number of offshore platformsin the North Sea. They are currently being consideredas possible foundations for offshore wind turbines. Asdiscussed by Houlsby and Byrne (2000) and by Byrneand Houlsby (2003), it is important to realise that theloading regimes on offshore turbines differ in severalrespects from those on structures usually encounteredin the offshore oil and gas industry. Firstly the structuresare likely to be founded in much shallower water: 10 mto 20 m is typical of the early developments, althoughdeeper water applications are already being planned.Typically the structures are relatively light, with amass of say 600t (vertical deadload 6 MN), but in proportion to the vertical load the horizontal loadsand overturning moments are large. For instance thehorizontal load under extreme conditions may beabout 60% of the vertical load.

An important consideration is that, unlike the oil andgas industry where large one-off structures dominate,many relatively small and inexpensive foundations arerequired for a wind farm development, which mightinvolve anything from 30 to 250 turbines.

The dominant device used for large scale windpower generation is a horizontal axis, 3-bladed tur-bine with the blades upwind of the tower, as shown inFigure 1. The details of the generator, rotational speedand blade pitch control vary between designs. Mostoffshore turbines installed to date generate 2 MW ratedpower, and typically have a rotor about 80 m in diam-eter with a hub about 80 m above mean sea level. Thesize of turbines available is increasing rapidly, andprototypes of 5 MW turbines already exist. Theseinvolve a rotor of about 128 m diameter at a hub heightof about 100 m. The loads on a “typical” 3.5 MW tur-bine are shown in Figure 2, which is intended to giveno more than a broad indication of the magnitude ofthe problem.

Note that in conditions as might be encountered inthe North Sea, the horizontal load from waves (say3 MN) is significantly larger than that from the wind(say 1 MN). However, because the latter acts at amuch higher point (say 90 m above the foundation) itprovides more of the overturning moment than thewave loading, which may only act at say 10 m abovethe foundation. Using these figures the overturningmoment of 120 MNm would divide as 90 MNm dueto wind and 30 MNm due to waves.

Realistic combinations of loads need to be con-sidered. For instance the maximum thrust on the turbineoccurs when it is generating at the maximum allow-able wind speed for generation (say 25 m/s). At higherwind speeds the blades will be feathered and providemuch less wind resistance. It is thus unlikely that themaximum storm wave loading would occur at the

Suction caissons for wind turbines

Guy T. Houlsby1, Lars Bo Ibsen2 & Byron W. Byrne1

1Department of Engineering Science, Oxford University, UK2Department of Civil Engineering, Aalborg University, Denmark

ABSTRACT: Suction caissons may be used in the future as the foundations for offshore wind turbines. Wereview recent research on the development of design methods for suction caissons for these applications. Wegive some attention to installation, but concentrate on design for in-service performance. Whilst much can belearned from previous offshore experience, the wind turbine problem poses a particularly challenging combin-ation of a relatively light structure, with large imposed horizontal forces and overturning moments. Monopodor tripod/tetrapod foundations result in very different loading regimes on the foundations, and we consider bothcases. The results of laboratory studies and field trials are reported. We also outline briefly relevant numericaland theoretical work. Extensive references are given to sources of further information.

same time as maximum thrust. Turbine designers mustalso consider important load cases such as emergencybraking. It is important to recognise that the design of aturbine foundation is not usually governed by consider-ations of ultimate capacity, but is typically dominatedby (a) considerations of stiffness of the foundationand (b) performance under fatigue loading.

An operational wind turbine is subjected to har-monic excitation from the rotor. The rotor’s rotationalfrequency is the first excitation frequency and is com-monly referred to as 1P. The second excitation fre-quency to consider is the blade passing frequency, oftencalled 3P (for a three-bladed wind turbine) at threetimes the 1P frequency.

Figure 3 shows a representative frequency plot of aselection of measured displacements for the VestasV90 3.0 MW wind turbine in operational mode. Thefoundation is a suction caisson. The measured data,monitoring system and “Output-Only Modal Analysis”used to establish the frequency plot are described inIbsen and Liingaard (2005). The first mode of the struc-ture is estimated, and corresponds to the frequencyobserved from idling conditions. The peak to the left ofthe first natural frequency is the forced vibration from

the rotor at 1P. To the right of the first natural frequencyis the 3P frequency. It should be noted that the 1P and3P frequencies in general cover frequency bands andnot just two particular values, because the Vestas windturbine is a variable speed device.

To avoid resonances in the structure at the keyexcitation frequencies (1P, 3P) the structural designerneeds to know the stiffness of the foundation with someconfidence, this means that problems of deformationand stiffness are as important as capacity. Furthermore,much of the structural design is dictated by consider-ations of high cycle fatigue (up to about 108 cycles),and the foundation too must be designed for theseconditions.

2 CASES FOR STUDY

The two main problems that need to be studied indesign of a suction caisson as a foundation are:

• installation;• in service performance.

In this review we shall discuss installation methodsbriefly, but shall concentrate mainly on design for inservice performance. The relevant studies involvetechniques as diverse as laboratory model testing,centrifuge model testing, field trials at reduced scale,

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Figure 1. Offshore tests in Frederikshavn, Denmark. Front:Vestas V90 3.0 MW turbine. Back: Nordex 2.3 MW turbine.

100 m

90 m6 MN

4 MN

h � 30 m

Figure 2. Typical loads on a 3.5 MW offshore wind turbine.

and a full-scale field installation. Complementing theseexperiments are numerical studies using finite elementtechniques, and the development of plasticity-basedmodels to represent the foundation behaviour.

Suction caissons may be installed in a variety ofsoils, but we shall consider here two somewhat ideal-ised cases: a caisson installed either in clay, whichmay be treated as undrained, or in sand. For typicalsands the combination of permeability value, size ofcaisson and loading rates leads to partially drainedconditions, although much of the testing we shall reportis under fully drained conditions. In this paper we reportmainly work on sands.

We shall consider two significantly different loadingregimes, which depend on the nature of the structuresupporting the wind turbine. Most offshore wind tur-bines to date have been supported on a “monopile” – asingle large diameter pile, which in effect is a directextension of the tubular steel tower which supports theturbine. Some turbines have been supported on circulargravity bases. An obvious alternative is to use a singlesuction caisson to support the turbine, and we shall callthis a “monopod” foundation, Figure 4(a). The mono-pod resists the overturning moment (usually the mostimportant loading component) directly by its rotationalfixity in the seabed.

As turbines become larger, monopod designs maybecome sufficiently large to be uneconomic, and analternative is a structure founded on three or four

smaller foundations: a “tripod” or “tetrapod”, Figure4(b). In either of these configurations the overturningmoment on the structure is resisted principally by“push–pull” action of opposing vertical loads on the

77

1-100

-80

-60

-40

-20

0

20

Frequency Domain Decomposition - Peak PickingAverage of the Normalized Singular Values of

Spectral Density Matrices of all Data Sets.dB | 1.0 / Hz

Frequency

First mode

3P

1P

Figure 3. Frequency plot of measured displacements for a wind turbine in operational mode.

(a)

D

DLL

s

(b)

Figure 4. Caisson foundations for a wind turbine, (a)monopod, (b) tripod/tetrapod.

upwind and downwind foundations. Alternativesusing asymmetric designs of tripod, and thoseemploying “jacket” type substructures are also underconsideration.

3 NORMALISATION PROCEDURES

A number of studies have been conducted at differentscales and it is necessary to compare the results fromthese various studies. To do this it is appropriate tonormalise all the results so that they can be repre-sented in non-dimensional form. This procedure alsoallows more confident extrapolation to full scale.

The geometry of a caisson is shown in Figure 5.The outside radius is R (diameter Do), skirt length is Land wall thickness t. In practice caissons may alsoinvolve stiffeners on the inside of the caisson, thesebeing necessary to prevent buckling instability duringsuction installation, but we ignore these in a simplifiedanalysis. Geometric similarity is achieved by requiringsimilar values of L/2R and t/2R.

The sign convention for applied loads and dis-placements is shown in Figure 6.

The rotation of the caisson � is already dimension-less, and we normalise the displacements simply bydividing by the caisson diameter, to give w/2R and u/2R.

In sand it is straightforward to show that, for similarvalues of dimensionless bearing capacity factor, theloads at failure would be proportional to �� and to R3.We therefore normalise vertical and horizontal loads asV/2�R3�� and H/2�R3��, where we have included thefactor 2� to give the normalisation factor a simplephysical meaning: it is the effective weight of a cylinderof soil of the same diameter of the caisson, and depthequal to the diameter. In a similar way we normalisethe overturning moment as M/4�R4��.

Use of the above normalisation is appropriate forcomparing tests in sands with similar angles of fric-tion and dilation. We recognise that these angles bothdecrease slightly with pressure and increase rapidlywith Relative Density (Bolton, 1986). This means thatcomparable tests at smaller scales (and thereforelower stress levels) will need to be at lower RelativeDensities to be comparable with field tests.

In clay the vertical capacity is proportional to arepresentative undrained shear strength su and to R2,so we normalise loads as V/�R2su and H/�R2su, andthe moment as M/2�R3su.

In order to be comparable, tests at different scaleswill need the profile of undrained strength with depthto be similar. If the strength profile is fitted by a sim-ple straight-line fit su � suo � �z, then this requiressimilar values of the factor 2R�/suo.

Scaling of results using the above methods shouldgive satisfactory results in terms of capacity. For clays itshould also lead to satisfactory comparisons in terms

of stiffness, provided that the clays being comparedhave similar values of Ir � G/su. This condition isusually satisfied if the clays are of similar compos-ition and overconsolidation ratio. For sands, however,an extra consideration needs to be taken into account.The shear modulus of a sand does not increase in pro-portion to the stress level, but instead can reasonablybe expressed by:

(1)

where g and n are dimensionless constants, and pa isatmospheric pressure (used as a reference pressure).The value of n is typically about 0.5, so that the stiffnessis proportional roughly to the square root of pressure.

Comparing rotational stiffnesses on the basis of aplot of M/4�R4�� against � effectively makes theassumption that the shear stiffness is proportional to2R��, which may be regarded as a representativestress level. Since in fact the stiffness increases at alower rate with stress level, this comparison will result

78

V’

hL

t

Mudline

Di

Do

Figure 5. Geometry of a caisson foundation.

Figure 6. Loading and displacement conventions for acaisson foundation (displacements exaggerated).

in larger scale tests giving lower apparent normalisedstiffness. This effect can be reduced by multiplying the� scale by the dimensionless factor ( pa/2R��)1�n, whichcompensates for the stiffness variation with stresslevel.

Thus we recommend that to compare both stiffnessand capacity data for sands one should plot M/4�R4��against �(pa/2R��)0.5 (assuming n � 0.5) for momenttests, and V/2�R3�� against (w/2R)(pa/2R��)0.5 forvertical loading tests. A fuller description of these nor-malisation procedures is given by Kelly et al. (2005a).

4 INSTALLATION STUDIES

The principal difference between installation of a suc-tion caisson for an offshore wind turbine and for previ-ous applications is that the turbines are likely to beinstalled in much shallower water. There is a popularmisconception that suction caissons can only beinstalled in deep water, where a very substantial headdifference can be established across the lid of thecaisson. In shallow water the net suction that can beachieved is indeed much smaller (being limited by theefficiency of the pumps, as the absolute pressureapproaches zero), but the suctions that can be achievedare nevertheless sufficient for installation in most cir-cumstances. Only in stiff clays is it likely that somepossible caisson designs, which might otherwise besuitable as far as in-service conditions are concerned,could not be installed by suction in shallow water.

In Table 1 we list the main instances where cais-sons have been installed in shallow water, as appro-priate to wind turbine installations. The water depthshw are approximate only. In addition to the field testslisted, a large number of small scale model tests ofinstallation have been carried out at Oxford University(on caissons of 0.1 m to 0.4 m diameter), the Universityof Western Australia (UWA), Aalborg and elsewhere.

The largest completed installation in shallow water isthat of a prototype suction caisson, shown in Figure 7,installed in the offshore research test facility inFrederikshavn, Denmark. The prototype has a diameterof 12 m and a skirt length of 6 m. The operational waterdepth is 4 m, and as the site is in a basin, no wave or iceloads are applied. As seen in Figure 7 the suction cais-son was installed in only 1 m of water in the basin. Thesteel construction has a mass of approximately 140 t,and the caisson was placed in late October 2002. Theinstallation period was about 12 hours, with the soilpenetration time being 6 hours. A computer systemwas used to control the inclination, suction pressure andpenetration rate. Det Norske Veritas (DNV) has certi-fied the design of the prototype in Frederikshavn to Blevel. The Vestas V90 3.0 MW turbine was erected onthe foundation in December 2002. The developmentof the design procedure for the bucket foundation is

described in Ibsen and Brincker (2004). An even largerinstallation is currently in progress at Wilhelmshaven,Denmark.

There are two main ways of predicting firstly theself-weight penetration of the caisson and secondly

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Table 1. Installations in shallow water.

hw D LSite Soil (m) (m) (m) Ref.

Wilhelmshaven Sand 6.0 16.0 15.0 InstallationApril 2005

Frederikshavn Sand 1.0 12.0 6.0 30

Frederikshavn Sand0.2

2.0 2.0–4.0 4.0

Sandy Haven Sand 0.5 4.0 2.5 23

Tenby Sand 2.0 2.0 2.0 23

Burry Port Sand 0.5 2.0 2.0 –

Luce Bay Sand0.2

3.0 1.5271.5 1.0

Bothkennar Clay0.2

3.0 1.5261.5 1.0

Figure 7. Installation of the prototype foundation at thetest site in Frederikshavn: (a) during installation, (b) at theend of installation.

the suction required to achieve full installation. Thefirst method (Houlsby and Byrne, 2005a, b) involvesuse of adaptations of pile capacity analysis, in whichthe resistance to penetration is calculated as the sumof an end bearing term on the rim and friction on theinside and outside. In sands the seepage pattern set upby the suction processes alters the effective stressregime in a way that aids installation.

The calculation has been implemented in a spread-sheet program SCIP. Figure 8 shows for example acomparison between variation of measured suction ina model test installation with tip penetration of thecaisson (Sanham, 2003), and the SCIP calculation.

The other approach involves use of CPT data toinfer directly the resistance Rd to penetration of thecaisson. The required suction ureq to penetrate thecaisson to depth d is calculated as:

(2)

where G�(d) is the self-weight of the caisson at pene-tration depth d (reduced for buoyancy), and Asuc is thearea inside the caisson, where the suction is applied.

The penetration resistance is calculated from thefollowing expression, which is based on calibrationagainst measured data:

(3)

where qt is the corrected cone resistance and fs thesleeve friction at depth z. Kt is a coefficient relating qtto the unit tip resistance on the rim. This resistance isadjusted for the reduction due to the applied suctionby the expression:

(4)

where kt is an empirical coefficient relating qt to thetip resistance during static penetration of the caisson,rt is the maximum reduction in tip resistance. �ucrit isthe critical suction resulting in the critical hydraulicgradient icrit � 1 along the skirt. �t is an empiricalfactor.

Kout and Kin are coefficients relating fs to the unitskin friction on the outside and inside of the skirt. Thewater flow along the skirt changes the skin friction.For the inside skin friction the coefficient reduces theskin friction when suction is applied, whereas on theoutside the skin friction is increased. The coefficientsare established as:

(5a,b)

where �out and �in are empirical coefficients relatingfs to the unit skin friction during static penetration ofthe caisson. rout and rin are the maximum changes inskirt friction. �out and �in are empirical factors.

The required suction ureq to penetrate the prototypein Frederikshavn was predicted using equation (2). Theresult of the analysis is shown in Figure 9. The lowerline represents ureq calculated from the CPT tests. Thecurved line represents the limiting suction upip whichwould cause piping to occur. umax is the theoreticalmaximum net suction, limited by the possibility ofcavitation within the caisson, as the absolute pressureapproaches zero, so that umax � 100 kPa above waterlevel and increases linearly with the water depth, asshown by Figure 9. umax is used to calculate the access-ible net suction, which is limited by the efficiency ofthe pumps, upump. As is seen, the suction in shallow

80

0

50

100

150

200

250

0 500 1000 1500 2000 2500 3000

Suction, s (Pa)

Pen

etra

tion,

h (

mm

)

SCIP ResultsExperimental Result

Figure 8. Comparison of SCIP with model test.

Figure 9. Suction required for installation at Frederikshavn.

water can be limited either by the suction causing piping or by the accessible net suction available fromthe pumps.

The suction upip causing piping has been studied atthe test site in Frederikshavn by installation tests on2 � 2 m and 4 � 4 m caissons. Figure 10 shows a4 � 4 m caisson where the limiting suction upip hasbeen achieved, and soil failure by piping has occurred.The soil outside of the skirt is sucked into the caissonand the penetration of the caisson cannot proceed.

If a tripod or tetrapod structure is to be installed,then levelling of the structure can be achieved by sep-arately controlling the suction in each of the caissons.For a monopod structure, however, an alternative strat-egy has to be adopted. Experience suggests that forinstallation in either clay or sand, the level of the caissonis rather sensitive to the application of eccentric loads(moments), especially in the early stages of installation.This offers one possibility for controlling the level ofthe caisson: by use of an eccentric load that can beadjusted in position to keep the caisson level.

An alternative strategy, which has proven to behighly successful for installation in sand, is to divide therim into sections and to control the pressures at the skirttip in each section individually. By applying pressureover one segment of the caisson rim the upwardhydraulic gradient within the caisson can be enhancedlocally, thus encouraging additional downward move-ment for that sector. By controlling the pressures at anumber of points the caisson may be maintained level.

This method would not be applicable in clays. Onepossibility, as yet untried at large scale, for control-ling level in clays would be to use a segmented cais-son in which the suctions in the different segmentscould be controlled independently.

Some preliminary small scale tests suggest that thisapproach might be successful in sand too (Coldicott,2005). Figure 11 shows the volumes of water pumpedfrom the two halves of a 400 mm diameter caissonsplit by a diametral vertical wall. About 60% of the

water pumped represents the volume displaced by thedescending caisson, whilst about 40% representsseepage beneath the caisson rim. Figure 12 shows thatduring the installation the suctions developed in thetwo halves were (as would be expected in a uniformmaterial) almost equal.

5 CAISSON PERFORMANCE: MONOPOD

A large number of tests have been devoted to studyingthe performance of a caisson under moment loadingat relatively small vertical loads, as is relevant to thewind turbine design. Some details of the test pro-grammes are given in Table 2.

5.1 Sand: field tests

The largest test involves the instrumented Vestas V903.0 MW prototype turbine at Frederikshavn, Denmark.The caisson is installed in a shallow 4 m depth lagoonnext to the sea, and the turbine is fully operational.The only significant difference between this installa-tion and an offshore one is that the structure is notsubjected to wave loading.

The test program involving the prototype (turbineand caisson) is focusing on long-term deformations,

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Figure 10. The limiting suction upip has been achieved andsoil failure by piping has occurred.

0

50

100

150

200

250

300

350

400

0 5 10 15 20 25 30 35

Volume, (10-3 m3)

Pen

etra

tion,

h (

mm

)

Cell 1Cell 2

TotalVolume

VolumeDisplaced

SeepageVolume

Figure 11. Volumes pumped from 2-cell caisson in sand.

0

50

100

150

200

250

300

350

400

0 1000 2000 3000 4000 5000 6000

Suction, s (Pa)

Pen

etra

tion,

h (

mm

)

Cell 1Cell 2

Figure 12. Suctions required for installation of 2-cell cais-son in sand.

soil structure interaction, stiffness and fatigue. Theprototype has been equipped with:

• an online monitoring system that measures thedynamic deformation modes of the foundation andthe wind turbine,

• a monitoring system that measures the long-timedeflection and rotation of the caisson,

• a monitoring system that measures the pore pres-sure along the inside of the skirt.

The online monitoring system that measures themodes of deformation of the foundation and wind tur-bine involves 15 accelerometers and a real-time data-acquisition system. The accelerometers are placed atthree different levels in the turbine tower and at onelevel in compartments inside the caisson foundation.The positions are shown in Figure 13, and the locationsand measuring directions are defined in Figure14.

Output-only Modal Analysis has been used to ana-lyze the structural behaviour of the wind turbine duringvarious operational conditions. The modal analysishas shown highly damped mode shapes of the foun-dation/wind turbine system, which the present aero-elastic codes for wind turbine design cannot model.Further studies are to be carried out with respect tosoil-structure interaction. A detailed description ofthe measuring system and the Output-Only ModalAnalysis is given by Ibsen and Liingaard (2005).

The static moment tests referred to in Table 2 at Sandy Haven and at Burry Port were relatively

straightforward, with very simple instrumentation, butthose at Frederikshavn test site and at Luce Bay weredetailed investigations.

The large scale tests at Frederikshavn is part of aresearch and development program concerning caissonfoundation for offshore wind turbines. The researchprogram is a co-operation between Aalborg Universityand MBD offshore power (Ibsen et al., 2003). The largescale tests are complemented by laboratory studies. Thelaboratory and large scale tests are intended to modelthe prototype in Frederikshavn directly. In order todesign a caisson foundation for offshore wind turbinesseveral load combinations have to be investigated. Eachload combination is represented by a height of load habove the foundation and a horizontal force H. Themoment at the seabed is calculated as M � hH. Table 3shows that the resulting loading height varies from10 m (for a wave force in shallow water) to 104.4 m(force from normal production of a 3 MW turbine in20 m of water). Scaling of the tests is achieved by:

(6)

where D is the diameter of the caisson and index mand p are for model and prototype. The values of theloading height in the test program are shown in Table 3.

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Table 2. Moment loading tests.

Site Soil D (m) L (m) Ref.

Frederikshavn Sand 12.0 6.0 –

Frederikshavn Sand 2.0 2.0 –

Sandy Haven Sand 4.0 2.5 –

Burry Port Sand 2.0 2.0 21

Luce Bay Sand 3.0 1.5 27

0.1 0.0–0.066 2, 40.15 0.05 2, 7

Oxford laboratory Sand 0.15 0.1 42, 430.2 0.1 34, 42, 430.2 0.2 11, 430.3 0.15 11, 42, 43

0.2 0.0–0.2 –

Aalborg laboratory Sand 0.3 0.0–0.30.4 0.0–0.4

Bothkennar Clay 3.0 1.5 26

Oxford laboratory Clay 0.2 0.1 340.3 0.15

UWA centrifuge 0.02 12(100g) Clay 0.06 0.03

0.06

Level IV: 89 m

Level III: 46 m

Level II: 13 m

Level I: 6 m

Figure 13. Sensor positions in tower and foundation.

The large scale tests at Frederikshavn employ load-ing by applying a horizontal load at a fixed height,under constant vertical load. A steel caisson with anouter diameter of 2 m and a skirt length of 2 m hasbeen used. The skirt is made of 12 mm thick steelplate. Figure 15 shows the caisson prior to installa-tion, and Figure 16 the overall test setup. Currently 10experiments have been conducted, but the testing pro-gram is ongoing. Each test has three phases:

1. Installation phase: The caisson is installed bymeans of suction. CPT tests are performed beforeand after installation of the caisson.

2. Loading phase: An old tower from a wind turbineis mounted on top of the caisson. The caisson isloaded by pulling the tower horizontally with awire. The combined loading (H, M) is controlled bychanging the height of loading.

3. Dismantling phase: The caisson is removed byapplying overpressure inside the bucket.

Figure 17 shows the moment rotation curve for atest on the 2 � 2 m caisson at Frederikshavn. The testis performed with hm � 17.4 m and a vertical load onthe caisson of 37.3 kN. The fluctuations in the curveare caused by wind on the tower.

Tests at Luce Bay were designed by OxfordUniversity and conducted by Fugro Ltd.. The momentloading tests were of two types. Firstly small ampli-tude (but relatively high frequency) loading wasapplied by a “Structural Eccentric Mass Vibrator”(SEMV) in which rotating masses are used to apply

83

Figure 14. Sensor mountings in the tower and foundationat Frederikshavn.

Table 3. Loading heights in the Aalborg test program.

FieldPrototype LaboratoryModel Model

Dp � 12 m Dm � 0.2 m 0.3 m 0.4 m 2.0 m

hp [m] hm [m] hm [m]

104.4 1.74 2.61 3.48 17.4069.6 1.16 1.74 2.32 11.6038.0 0.63 0.95 1.27 6.3320.0 0.33 0.50 0.67 3.3310.0 0.17 0.25 0.33 1.67

Figure 15. Caisson for large scale test at Frederikshavn.

Figure 16. Setup for combined loading of 2 � 2 m caissonat Frederikshavn (Back: prototype 3 MW Vestas wind tur-bine on the 12 � 6 m caisson).

inertial loads at frequencies up to about 12 Hz.Secondly larger amplitude, but lower frequency, cycleswere applied using a hydraulic jack. A diagram of theloading rig, which allowed both moment and verticalloading tests, is shown in Figure 18.

The SEMV test involve cycles of moment loadingat increasing amplitude as the frequency increases.Figure 19 shows the hysteresis loops obtained from aseries of these cycles at different amplitudes. As thecycles become larger the stiffness reduces but hys-teresis increases. The tests were interpreted (Houlsbyet al., 2005b) using the theory of Wolf (1994), whichtakes account of the dynamic effects in the soil, and theequivalent secant shear modulus for each amplitudeof cycling determined.

Figure 20 shows the moment rotation curves formuch larger amplitude cycling applied by the hydraulicjack. Again hysteresis increases and secant stiffnessdecreases as the amplitude increases. The unusual“waisted” shape of the hysteresis loops at very large

amplitude is due to gapping occurring at the sides ofthe caisson.

The secant stiffnesses deduced from both theSEMV tests and the hydraulic jacking tests are com-bined in Figure 21, where they are plotted against theamplitude of cyclic rotation. It is clear that the two

84

Figure 17. Moment-rotation test on 2 � 2 m caisson.

Figure 18. Field testing equipment, dimensions in mm. Water level and displacement reference frames not shown. (a)arrangement for jacking tests on 1.5 m and 3.0 m caissons, (b) alternative arrangement during SEMV tests. Labels indicate(A) A-frame, (B) concrete block, (C) caissons, (H ) hydraulic jacks, (L) load cells, (R) foundations of reaction frame, (V)SEMV, (W) weight providing offset load for SEMV tests.

-30

-20

-10

0

10

20

30

-0.00005 -0.000025 0 0.000025 0.00005

Rotation (radians)

Mom

ent (

kNm

)

6Hz

7Hz

8Hz

9Hz10H

Figure 19. Hysteresis loops from SEMV tests on 3 m caisson.

-600-500-400-300-200-100

0100200300400500

-0.08 -0.06 -0.04 -0.02 0 0.02 0.04 0.06 0.08

Rotation of caisson centre (2Rθ ) (m)

Mom

ent (

kNm

)

Figure 20. Hysteresis loops from hydraulic jacking testson 3 m caisson.

groups of tests give a consistent pattern of reductionof shear modulus with strain amplitude, similar tothat obtained for instance from laboratory tests.

5.2 Sand: laboratory tests

Turning now to model testing, a large number of testshave been carried out both at Aalborg and at Oxford.Almost all the model tests have involved “in plane”loading (in which the moment is about an axis perpen-dicular to the horizontal load). However, a test rig cap-able of applying full 6 degree-of-freedom loading hasrecently been developed by Byrne and Houlsby (2005).

The model tests at Aalborg are performed by thetest rig shown in Figure 22. The rig consists of a testbox and loading frame. The test box consists of a steelframe with an inner width of 1.6 m �1.6 m and aninner total depth of 0.65 m. The test box is filled withAalborg University Sand No 0. After each experimentthe sand in the box is prepared in a systematic way toensure homogeneity within the box, and between thedifferent test boxes. The sand is saturated by the waterreservoir shown in Figure 22. Before each experimentCPT-tests are performed to verify the density andstrength of the sand. The caisson is then installed andloaded with a constant vertical load. The vertical loadis kept constant through the experiment, while thehorizontal force is applied to the tower by the loadingdevice mounted on the loading frame, see Figure 21.The tower and the loading device are connected by awire. The combined loading (H, M) is controlled bythe height of loading h. The loading frame allows thepossibility of changing h from 0.1 m to 4.0 m abovethe sand surface (Table 3). The horizontal force H ismeasured by a transducer connected to the wire. Thedeformation of the foundation and the moment aremeasured with the measuring cell mounted on the topof the caisson, as shown by Figure 23.

Laboratory tests at Oxford University have used aversatile 3 degree-of-freedom loading rig designed byMartin (1994) and adapted by Byrne (2000) (see alsoMartin and Houlsby (2000) and Gottardi et al. (1999)).

The rig is shown in Figure 24, and is capable of apply-ing a wide range of combinations of vertical, horizontaland moment loading under either displacement orload control.

Typical moment loading tests involve applying afixed vertical load, and then cycling the rotation atincreasing amplitude. An example is given in Figure 25.

The first interpretation of such tests is to determinethe yield surface for a single surface plasticity model(see section 7.2 below, and also Martin and Houlsby

85

0102030405060708090

100

0.000001 0.00001 0.0001 0.001 0.01 0.1

∆θ (radians)

G (

MP

a)

JackingSEMVHyperbolic curve fit

Figure 21. Shear modulus against rotation amplitude.

Figure 22. The caisson test rig at Aalborg University.

Figure 23. The measuring cell connecting the caisson andthe tower.

(2001), Houlsby and Cassidy (2002), Houlsby (2003),Cassidy et al. (2004)). An example of the yield pointsobtained, plotted in the vertical load-moment plane, isgiven in Figure 26. Of particular importance is the factthat at very low vertical loads there is a significantmoment capacity, and that this extends even into thetensile load range. In these drained tests the ultimateload in tension is a significant fraction of the weight ofthe soil plug inside the caisson.

Sections of the yield surface can also be plotted inH-M space as shown in Figure 27, where the data herehave been assembled from many tests at different

stress levels. The flow vectors are also plotted in thisfigure, and show that in this plane (unlike the V-Mplane) associated flow is a reasonable approximationto the behaviour. Feld (2001) has observed similarshapes of a yield surface for a caisson in sand.

We now consider the possibility of scaling the resultsof laboratory tests to the field. The test at Frederikshavnshown in Figure 17 was on a caisson with a ratioL/2R � 1, at an M/2RH value of approximately 8.7,and with a value of V/2�R3�� of about 0.62. Using thedata from the Oxford laboratory on 0.2 � 0.2 m cais-sons this requires a vertical load of about 60N. In facta test had been carried out with L/2R � 1 and V �50 N. According to the scaling relationships discussedin section 3, the moment should be scaled accordingto R4�� (a factor of 6250) and the rotational displace-ment 2R� according to √R3����(a factor of 25). Figures26 and 27 suggest that for a vertical load of 60 Nrather than 50 N a moment capacity say 5% highermight be expected, and that for the higher value ofM/2RH a further increase of say 15% is appropriate.We therefore apply a factor of 7500 to the moments and25 to the rotational displacements. The result is shown

86

Figure 24. Three degree-of-freedom testing rig at OxfordUniversity.

-100

-80

-60

-40

-20

0

20

40

60

80100

-2.0 -1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 2.0

Rotational Displacement, 2Rθ (mm)

Mom

ent L

oad,

M/2

R (

N)

Figure 25. Moment-rotation test on sand.

0

20

40

60

80

100

-160 -120 -80 -40 0 40 80 120Vertical Load, V (N)

Mom

ent

Load

, M

/2R

(N

)

Experiment, M/2RH = 1

Fitted Yield Surface

Soil Plug Weight

Figure 26. Experimentally determined yield surface in V-M plane.

-120

-80

-40

0

40

80

120

-180 -140 -100 -60 -20 20 60 100 140 180

Horizontal Load, H(N)Incremental Horizontal Displacement, du (mm)

Mom

ent L

oad,

M/2

R(N

)In

crem

enta

l Rot

atio

n, 2

Rdt

heta

(m

m)

V = -50 N

V = 0 N

V = 50 N

Figure 27. Yield surfaces and flow vectors in H-M space.

in Figure 28. It can be seen that after scaling themoment at a 2R� value of 0.04 m is about 120 kNm,compared to about 280 kNm measured in the field.Although there is a factor of about 2 between thesevalues, it must be borne in mind that there are a num-ber of possible causes of difference between the tests(e.g. the sand in the field test may be much denser),and also that a factor of 7500 has already been applied:a factor of 2 is relatively small by comparison.

5.3 Clay: field and laboratory tests

Less work has been carried out on clay than on sand.The large scale trials at Bothkennar (Houlsby et al.2005b) are complemented by laboratory studiesintended to model these trials directly, and thereforeadd confidence to the scaling of the results to proto-type size caissons (Kelly et al., 2005a).

At Bothkennar, moment loads were applied to a3 m �1.5 m caisson by two means. Small amplitude,but relatively high frequency (10 Hz) loading wasapplied by means of the SEMV device describedabove, and larger amplitude cycles, but at much lowerfrequency, were applied using a hydraulic jack. Inboth cases the loading was 4 m above the caisson, sothat hload/D � 1.33. The most important observationfrom these tests was the gradual reduction of secantstiffness (and increase in hysteresis) as the amplitudeof the load cycles increases.

The laboratory tests, specifically modelling thefield tests, involved just relatively low frequency load-ing. After the scaling relationships described in section 3were applied, there was a satisfactory agreementbetween laboratory and field data, especially at rela-tively small amplitudes of movement. As an example,Figure 29(a) shows the results (in dimensionless form)for rotation of the 3.0 m diameter caisson in the field,and Figure 29(b) the equivalent results, also in dimen-sionless form, from the small scale model test. The pat-tern of behaviour is remarkably similar in the two tests.

This sort of comparison is vital to establish confidencein the use of model testing to develop design guidelines.

6 CAISSON PERFORMANCE: TETRAPODOR TRIPOD

In the following, in which we consider multiple foot-ing designs to support the wind turbine, we shall referprincipally to a tetrapod (four footings) rather than atripod. As a tripod is perhaps the most obvious mul-tiple footing design to use, and has the obviousadvantage of simplicity, our preference for the tetra-pod deserves some explanation.

As is discussed below, prudent design of a multiplefooting structure will avoid tension being applied to any of the foundations (except under the most extremeof circumstances). This in effect dictates the separationof the foundations for a given overturning momentand weight of structure. Approximate calculationsindicate that the tetrapod structure is usually a morefavourable configuration to avoid tension, as it requiressomewhat less material. The differences are not large,and a tripod may be preferred in some circumstances,but we shall refer to a tetrapod, as this will probablybe more efficient. The important mechanism is thesame in both cases: the overturning moment is resistedby opposing “push–pull” action on the foundations.

87

-150

-100

-50

0

50

100

150

-0.06 -0.04 -0.02 0 0.02 0.04 0.06

Rotational Displacement, 2Rθ (m)

Mom

ent,

M (

kNm

)

Figure 28. Laboratory moment test scaled to field condi-tions for comparison with Figure 17.

-0.4

-0.3

-0.2

-0.1

0

0.1

0.2

0.3

-0.015 -0.01 -0.005 0

θ

θ

0.005 0.01

M/[s

u(2R

)3 ]M

/[su(

2R)3 ]

(a) field test

-0.4

-0.3

-0.2

-0.1

0

0.1

0.2

0.3

-0.015 -0.01 -0.005 0 0.005 0.01

(b) model test

Figure 29. Moment-rotation results presented in non-dimensional form for laboratory and field tests.

In Table 4 we list the tests that have been carried outon vertical loading of caissons relevant to the windturbine problem. In addition to these studies there area number of other relevant studies which have beendirected towards vertical loading of caissons for struc-tures in the oil and gas industry or for use as anchors.

6.1 Sand: field and laboratory tests

The simplest tests on vertical loading of caissons insand, which are relevant both to installation and to sub-sequent performance, simply involve pushing caissonsvertically into sand to determine the vertical load-dis-placement response. Figure 30 shows the results of a setof such tests on caissons of different L/D ratios, Byrneet al. (2003). It is clear from the figure that there is awell-established pattern. While the caisson skirt ispenetrating the sand there is relatively low verticalcapacity, but as soon as the top plate makes contact withthe sand there is a sudden increase in capacity. Theenvelope of the ultimate capacities of footings of differ-ent initial L/D ratios also forms a single consistent line.

Of most importance, however, is the performance ofthe caissons under cyclic vertical loading. Figure 31shows the results of tests on a 300 mm diameter cais-son subjected to rapid cyclic loading. Smallamplitudecycles show a stiff response, with larger cycles show-ing both more hysteresis and more accumulated dis-placement per cycle. The most important observationis that as soon as the cycles go into tension, a muchsofter response is observed, and the hysteresis loopsacquire a characteristic “banana” shape. Clearly thesoft response on achieving tension should be avoidedin design. Closer examination of the curves revealsthat the softening in fact occurs once the drained fric-tional capacity of the skirts has been exceeded, ratherthan simply the transition into tension.

Paradoxically, although additional accumulateddisplacement is observed once tension is reached, thisaccumulated displacement is downwards (not upwardsas one might expect because of the tensile loading).

The above observations mean that tension must beavoided in a prudent design of a tripod or tetrapod foun-dation for a wind turbine. However, in all but the shal-lowest of water, avoiding this tension means that eitherthe foundation must have a large spacing between thefootings, or that ballasting must be used. The lattermay in fact be a cost effective measure in deep water.

Some designers may wish to reduce conservatism byallowing for the possibility of tension under extremecircumstances. It is therefore useful to examine the ultimate tensile capacity under rapid loading. Figure 32shows the result of three such tests. The slowest test (at5 mm/s) is almost drained, and a very low capacity intension is indicated. The capacity in this case is simplythe friction on the skirts. The test at 100 mm/s (but zeroambient water pressure) shows a larger capacity, and itis straightforward to show that this is controlled bycavitation beneath the foundation. This means that atelevated water pressures (as in the third test) thecapacity rises approximately in step with to the ambientwater pressure, as correspondingly larger pressurechanges are required to cause cavitation. This problemis studied in more detail by Houlsby et al. (2005a).

It is important to note, however, that although ambi-ent water pressure increases the ultimate capacity, it

88

Table 4. Vertical loading tests.

Site Soil D (m) L (m) Ref.

Luce Bay Sand 1.5 1.0 27

0.05 0.0–0.1 110.1 0.0–0.066 2,50.15 0.05 2,5

Oxford Sand 0.15 0.1 34laboratory 0.2 0.133 34

0.28 0.18 25, 32, 33, 35

Bothkennar Clay 1.5 1.0 26

UWA 0.02centrifuge Clay 0.06 0.03 3(100g) 0.06

0

50

100

150

200

250

300

350

400

0 50 100 150Vertical Displacement, w (mm)

Ver

tical

Loa

d, V

(N

)

0 0.5 1 1.5 2 2.5 3 3.5

Normalised Vertical Displacement, w/D

Figure 30. Vertical load-penetration curves for caissons ofdifferent L/D ratios.

-400

-2000

200400

600

800100012001400

1600

200 210 220 230 240 250 260 270

Vertical Displacement (mm)

Ver

tical

Str

ess

(kP

a)

Figure 31. Cyclic vertical loading of model caisson.

has negligible influence on the tensile load at which aflexible response begins to occur.

Comparison of cyclic loading tests at different scalesand at different speeds shows that it is difficult to scalereliably the accumulated displacements, which reducewith larger tests and higher loading rates. However,when the scaling rules described earlier are applied,the shapes of individual hysteresis loops at differentscales and at different rates become remarkably simi-lar. Figure 33 shows a comparison, for instance, ofloops at three different load amplitudes from four dif-ferent tests. At each particular load amplitude the loopsfrom the different tests are very similar.

The accumulation of displacement after very largenumbers of cycles is difficult to predict, and so far fewdata are available. Rushton (2005) has carried out verti-cal loading tests to about 100 000 cycles on a modelcaisson in sand, using a simple loading rig whichemploys a rotating mass and a series of pulleys to applya cyclic load. A typical result is shown in Figure 34,on a caisson 200 mm diameter and 100 mm deep, withcycling between 210 � 260 N. The caisson is there-fore subjected (at the minimum vertical load) to a smalltension, but less than the frictional capacity of the skirts.The dimensionless accumulated vertical displacement

is seen in Figure 34 to increase approximately withthe logarithm of the number of cycles of loading (afterabout 1000 cycles). Note that even in this case wherethere is a tensile loading in part of the cycle, the netmovement is downwards. The displacement is of coursevery sensitive also to the amplitude of the cycling.

6.2 Clay: field and laboratory tests

Very few vertical loading tests relevant to the wind tur-bine problem have been completed on caissons in clay,although there have been a number of studies directedtowards suction caissons used as tension anchors, e.g.El-Gharbawy (1998), Watson (1999), House (2002).

At Bothkennar tests were carried out in whichinclined (but near vertical) loading was applied to a1.5 m diameter caisson (Houlsby et al., 2005b). Diffi-culties were encountered with the control of the loadsusing a hydraulic system, and the resulting load pathsare therefore rather complex, leading to difficulties ininterpretation. Further work on vertical loading in clayis required before definitive conclusions can be drawn,and in particular the issue of tensile loading in clayneeds attention. Some preliminary results (Byrne andCassidy, 2002), shown in Figure 35, show that the ten-sile response may be sensitive to prior compressiveloading. Footings loaded in tension immediately after

89

-450

-400

-350

-300

-250

-200

-150

-100

-50

0150 160 170 180 190 200 210

Vertical Displacement (mm)

Ver

tical

Str

ess

(kP

a)

5mm/s, 0kPa100mm/s, 0kPa

100mm/s, 200kPa

Direction ofmovement

Figure 32. Tensile capacity of model caisson pulled at dif-ferent rates and at different ambient pressures.

-1

0

1

2

3

4

5

0 0.01 0.02 0.03 0.04 0.05

[w/(2R)][pa/(2Rγ')]1/2

V/[γ

'(2R

)3 ]

1.5m Field

0.15m Suction

0.2m Pushed

0.15m Pushed

Figure 33. Hysteresis loops from tests at different scalesand rates.

-0.25

-0.02

-0.15

-0.10

-0.05

0.00

0.05

1 10 100 1000 10000 100000

Number of Cycles

[w/(

2R)]

[pa/

(2R

γ’)]1/

2

MinMax

Figure 34. Accumulated displacement during long termcyclic vertical loading on sand.

-80

-60

-40

-20

0

20

40

60

0 0.2 0.4 0.6 0.8 1

Normalised Displacement, (w + L)/D

Ver

tical

Str

ess,

V/A

(kP

a)

Test 1: Post Bearing Capacity

Test 2: Pre Bearing Capacity

Figure 35. Tension tests on caisson foundations in clay.

installation showed a stiff tensile response, whilstthose loaded after first applying a compressive load tofailure showed a more flexible tensile response.

7 NUMERICAL STUDIES

7.1 Finite element studies

A number of analyses of suction caissons for offshorewind farms have been carried out as part of commercialinvestigations for possible projects. A more detailedresearch project was carried out by Feld (2001).

Finite element analysis is particularly appropriatefor establishing the effects of design parameters on theelastic behaviour of caissons, and has been used byDoherty et al. (2004a, b) to determine elastic stiffnesscoefficients for caisson design which take into accountthe flexibility of the caisson wall as well as couplingeffects between horizontal and moment loading.

7.2 Plasticity models

An important tool for the analysis of soil-structureinteraction problems, particularly those involvingdynamically sensitive structures are “force resultant”models. In these the behaviour of the foundation isrepresented purely through the force resultants actingupon it, and the resulting displacements (see Figure 4).Details of stresses and deformations within the soilare ignored. The models are usually framed within thecontext of work-hardening plasticity theory. Examplesinclude models for foundations on clay (Martin andHoulsby, 2001) and on sand (Houlsby and Cassidy,2002). Overviews of the development of these modelsare given by Houlsby (2003) and Cassidy et al. (2004).

These models have been further developed specif-ically for the offshore wind turbine application. Thedevelopments include:

• Generalisation to full three-dimensional loadingconditions,

• Inclusion of special features to represent the cais-son geometry,

• Expression of the models within the “continuoushyperplasticity” framework to allow realistic descrip-tion of hysteretic response during cyclic loading.

A model with all these features is described byLam and Houlsby (2005). The fitting of cyclic data toa continuous hyperplastic model is discussed byByrne et al. (2002a).

8 OTHER CONSIDERATIONS

We have concentrated here on the design of caissonfoundations as far as capacity and stiffness are con-cerned for in-service conditions. However, there a

number of other issues which need to be addressed ina caisson design, and we mention them here briefly.

8.1 Scour

Scour is more important for caissons, since they arerelatively shallow, than for piles. The size of caissons,and the fact that part of the caisson inevitably pro-trudes above mudline level, creates rather aggressiveconditions for scour. The fact that the caissons may beinstalled in mobile shallowwater environments meansthat proper consideration of this problem is essential,especially in sands.

If the scour depth can be determined with suffi-cient confidence (e.g. from comprehensive model test-ing) then it may be possible to permit the scour tooccur, and simply allow for this in the design byensuring that the caisson is deep enough.

It is more likely, however, that scour protectionmeasures such as rock-dumping will need to beemployed. Practical experience suggests that such pro-tection must be placed very soon after caisson installa-tion, as scour can occur very rapidly. In highly mobileenvironments, significant scour can, for instance, occurdue to the currents in a single tide. Model testing indi-cates, however, that scour protection measures can beeffective in preventing further erosion (R. Whitehouse:private communication). For in-service conditionsregular monitoring for the possibility of scour wouldbe prudent.

8.2 Liquefaction

The transient pore pressures induced in the seabedcan induce liquefaction, especially if the seabed ispartially saturated due to the presence of gas (as canoccur in shallow seabeds, largely due to decay of organicmatter).

The problem is a complex one, but typically, at onestage in the wave cycle, the pore pressure in the seabedcan become equal to the overburden stress, and theeffective stress falls to zero. This problem is furthercomplicated by the presence of a structure, whichclearly modifies the pore pressure pattern that wouldoccur in the far field. Although some progress hasbeen made, the interactions are complex, and theoret-ical modelling of the problem is not straightforward.

8.3 Wave-induced forces

A quite different problem from liquefaction is alsorelated to the fact that the principal forces on thestructure are wave induced. As a wave passes the col-umn of the structure it exerts large horizontal forces(of the order of a few meganewtons for a large wave),which also cause overturning moments. However, at

90

the same time the wave causes a transient pressure onthe seabed, and on the lid of the caisson. Because thecaissons are in shallow water these pressures are quitelarge. The pore water pressure within the caisson isunlikely to change as rapidly as the pressure on thelid, so there will be pressure differentials across thelid of the caisson which result in net vertical forces,and overturning moments on the caisson.

The relative phase of the different sources of load-ing is important. As the crest of the wave just reachesthe structure, the wave kinematics are such that thehorizontal forces are likely to be largest. At this stagethe pressure on the upwave side of the caisson islikely to be larger than on the downwave side. The netresult is that the moment caused by the pressures onthe caisson lid opposes that caused by the horizontalloading, so this effect is likely to be beneficial to theperformance of the caisson. Little work has, however,yet been completed on the magnitudes of these effects.The problem is complicated by the fact that the kin-ematics of large (highly non-linear) shallow waterwaves is still a matter of research, as is their interactionwith structures.

9 CONCLUSIONS

In this paper we have provided an overview of theextensive amount of work that has been carried out onthe design of suction caisson foundations for offshorewind turbines. Further verification of the results pre-sented here is still required, and in due course it is hopedthat this will come from instrumented caisson foun-dations offshore. Our broad conclusions at present are:

• Suction caissons could be used as foundations foroffshore wind turbines, either in monopod or tri-pod/tetrapod layout.

• The combination of low vertical load and high horizontal load and moment is a particular featureof the wind turbine problem.

• Stiffness and fatigue are as important for turbinedesign as ultimate capacity.

• Monopod foundation design is dominated bymoment loading.

• Tripod/tetrapod foundation design is dominated byconsiderations of tensile loading.

• The moment-rotation response of caissons in sandhas been extensively investigated by model tests andfield trials, and modelled theoretically by finiteelement analyses and force resultant (yield surface)models.

• As amplitude of moment loading increases, stiff-ness reduces and hysteresis increases.

• Moment loading in clay has been less extensivelyinvestigated in the laboratory and field.

• Vertical loading in sand has been extensively inves-tigated in the laboratory and field.

• As the amplitude of vertical loading increases,stiffness reduces and hysteresis increases. Oncetension is reached there is a sudden reduction ofstiffness.

• Whilst high ultimate tensile capacities are possible(especially in deep water) this is at the expense oflarge movements.

• Application of scaling procedures for tests in bothsand and clay allows model and field tests to becompared successfully as far as stiffness and theshapes of hysteresis loops is concerned.

• Cumulative displacements after very many cyclesare harder to model.

• The design of caisson foundations also needs totake into consideration issues such as scour andliquefaction.

It is hoped that the conclusions above lead in duecourse to application of suction caissons as founda-tions for offshore wind turbines, thereby making animportant renewable energy source more econom-ically viable.

ACKNOWLEDGEMENTS

The work at Oxford University has been supported by the Department of Trade and Industry, the Engi-neering and Physical Sciences Research Council anda consortium of companies: SLP Engineering Ltd,Aerolaminates (now Vestas), Fugro Ltd, Garrad Hassan,GE Wind and Shell Renewables. An outline of theproject is given by Byrne et al. (2002b). The work ofRichard Kelly, Nguyen-Sy Lam and Felipe Villaloboson this project is gratefully acknowledged.

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