Field Testing of Suction Caissons at Bothkennar and Luce Bay

by

G.T. Houlsby, R.B. Kelly, J. Huxtable and B.W. Byrne

Report No. OUEL 2276/05

University of Oxford Department of Engineering Science Parks Road, Oxford, OX1 3PJ, U.K.

Tel. 01865 273162/283300

Fax. 01865 283301 Email Civil@eng.ox.ac.uk

http://www-civil.eng.ox.ac.uk/

Field Testing of Suction Caissons at Bothkennar and Luce Bay

G.T. Houlsby, R.B. Kelly, J. Huxtable and B.W. Byrne

This report consists of three papers that have resulted from a joint industry project investigating the application of suction caissons to offshore wind turbines. The first two papers report field scale testing at two different locations in Scotland. The final paper presents some scaling relationships for use in comparing laboratory testing to the field testing. a) “Field trials of suction caissons in clay for offshore wind turbine foundations.” Houlsby, G.T., Kelly, R.B., Huxtable, J. and Byrne, B.W. Abstract: A programme of testing of caisson foundations in clay at the Bothkennar test site is described. The tests are relevant to the design of foundations for offshore wind turbines, either in the form of monopod or tetrapod foundations. Records are presented for installation of the caissons, cyclic moment loading under both dynamic and quasi-static conditions, cyclic inclined vertical loading and for pullout of the caisson. Variation of stiffness of the foundation is observed, with high initial stiffness followed by hysteretic behaviour at moderate loads and degradation of response at high loads. Some implications for the design of wind turbine foundations are briefly discussed. b) “Field trials of suction caissons in sand for offshore wind turbine foundations.” Houlsby, G.T., Kelly, R.B., Huxtable, J. and Byrne, B.W. Abstract: A programme of testing on suction caisson foundations in an artificially prepared sand test bed near Luce Bay, in Scotland, is described. The tests are relevant to the design of either monopod or tetrapod foundations for offshore wind turbines. Records are presented for suction installation of the caissons, cyclic moment loading under both quasi-static and dynamic conditions to simulate the behaviour of a monopod foundation, and cyclic vertical loading and pullout of caissons to simulate one footing in a quadruped foundation. Variations of stiffness with loading level of the foundation are observed, with high initial stiffness followed by hysteretic behaviour at moderate loads and degradation of response at high loads. Some implications for the design of wind turbine foundations are briefly discussed. c) “A comparison of field and laboratory tests of caisson foundations in sand and clay.” Kelly, R.B., Houlsby, G.T. and Byrne, B.W. Abstract: Laboratory tests applying vertical and moment loads to suction caissons founded in sand and clay have been conducted to simulate an equivalent series of field tests. The caissons used in the laboratory were 0.15m, 0.2m and 0.3m in diameter, while those for the field tests were 1.5m and 3.0m diameter. The loads applied to the caissons in the laboratory tests were scaled from those in the field tests, and the models were loaded in a near identical manner to the field trials. The test results are presented in non-dimensional form for comparison. The non-dimensional laboratory moment test data were similar to the field data in most cases. The non-dimensional data from vertically loaded caisson tests in the laboratory and in the field show some differences, and possible reasons for these are discussed.

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Field trials of suction caissons in clay for offshore wind turbine foundations

G.T. Houlsby1, R.B. Kelly1, J. Huxtable2 and B.W. Byrne1 Keywords: bearing capacity, clay, dynamic test, foundations, stiffness

ABSTRACT

A programme of testing of caisson foundations in clay at the Bothkennar test site is described.

The tests are relevant to the design of foundations for offshore wind turbines, either in the form of

monopod or tetrapod foundations. Records are presented for installation of the caissons, cyclic

moment loading under both dynamic and quasi-static conditions, cyclic inclined vertical loading

and for pullout of the caisson. Variation of stiffness of the foundation is observed, with high initial

stiffness followed by hysteretic behaviour at moderate loads and degradation of response at high

loads. Some implications for the design of wind turbine foundations are briefly discussed.

INTRODUCTION

The offshore wind energy industry is a very rapidly expanding sector of vital economic

importance in the UK, and foundation costs are an important part of the costs of offshore wind

turbine installations (Byrne and Houlsby, 2003). Most current foundations for offshore wind

turbines are large “monopiles”, although some have been founded on gravity bases. However, with

the current expansion of the offshore wind energy industry, alternative foundation types are being

considered. One possibility is the use of “suction caisson” foundations (Houlsby and Byrne (2000),

Byrne and Houlsby (2003)). Suction caissons are now widely used as anchors for floating

structures, and have also been used offshore as foundations for a small number of fixed platforms

(Bye et al., 1995). They are large cylindrical structures, open at the base (see Figure 1). During

installation they cut a small distance into the seabed under their own weight, but are then installed

to their full depth (with the caisson lid flush with the seabed) by pumping out the water that is

trapped within the caisson. They can be installed in either clays or sands. The principal advantage

for the offshore wind application is that the caissons can be installed rapidly, using relatively

inexpensive equipment.

Methods for designing caisson foundations for offshore wind applications are in their infancy,

and in response to the need for design methods, a programme of research has been sponsored by the

DTI, EPSRC and a consortium of companies (see Acknowledgements) (Byrne et al. 2002). In this 1 Department of Engineering Science, Oxford University 2 Fugro Structural Monitoring Ltd.

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study, design methods are being developed principally on the basis of small-scale model testing

(Byrne and Houlsby (2002, 2004), Byrne et al. (2003), Kelly et al. (2003, 2004)), but an important

part of the research is a programme of intermediate scale field trials to check on the scalability of

the results. In this paper tests on 1.5m and 3.0m diameter caissons at the Bothkennar test site are

reported. Typical sizes of prototype caissons are discussed below. Other studies of design methods

for suction caissons for foundation (as opposed to anchor) applications have concentrated on

analytical procedures (e.g. Bransby and Randolph, 1998) or finite element analysis (e.g. Gourvenec

and Randolph, 2003).

A key feature of offshore wind turbine structures is that (for their size) they are relatively light

(with a mass of the order of 600t for a 3.5MW turbine structure), yet they are subjected to large

horizontal forces and overturning moments from wind and waves (Byrne and Houlsby, 2003). The

horizontal load may, for instance, be of the order of 65% of the vertical load. Thus the challenge to

the foundation engineer is to carry large (and repetitive) horizontal loads and overturning moments,

but relatively little vertical load. Two main structural configurations using caissons are being

considered: either a “monopod” consisting of a single large caisson (typically 20m to 25m in

diameter for a modern large turbine structure), or a “tetrapod” in which the load is transferred

through a truss structure to four smaller caissons, see Figure 2 (preliminary calculations indicate

that the obvious alternative of a tripod is a less favourable configuration). Each of the smaller

caissons might be say 6m to 8m in diameter. For the monopod the most important load on the

foundation is the overturning moment. In the case of the tetrapod the moment loading is principally

carried by “push-pull” action by opposing footings, and it is the variation of vertical load (and in

particular the possibility of tension on a footing) that is most important. In both cases the design

objective is to select an appropriate diameter D and depth h of the caisson, and in the tetrapod case

the spacing s must also be determined (see Figure 2).

The testing programme described below includes tests directed towards the design of both the

monopod and tetrapod. Data were obtained from the installation phases for each caisson. Loading of

the caissons was by means of a combination of dead weights, hydraulic jacks and inertial loading

from a “Structural Eccentric Mass Vibrator” (SEMV). The test programme was designed by Oxford

University, and site operations were managed by Fugro Structural Monitoring Ltd. The tests were

carried out in December 2003 and January 2004.

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EQUIPMENT AND TESTING PROCEDURES

The caissons were fabricated from mild steel, with the principal dimensions given in Table 1.

The lids of the caissons were stiffened by I-sections. Ports were provided for attaching the pump

and for venting the caissons. An A-frame structure was attached by pins to the 3.0m caisson to

transfer loads from the SEMV or hydraulic jack. Load cells were fitted to measure the axial load in

all four legs of the frame.

Reaction loads for testing were provided by a steel frame, and the layout of the entire testing

assembly is shown in Figure 3. The reaction frame was supported on two square foundations (2m x

2m x 1m deep), which were also installed by the suction method. The frame was installed in a pit of

depth approximately 1.5m at the Bothkennar test site, see Figure 4. Details of the soil properties at

Bothkennar are reported in the collection of papers in Géotechnique, Volume 42, Number 2 (June

1992). The best estimate is that the base of the pit corresponds to a depth of 1.75m in Nash et al.

(1992), so that the undrained shear strength (measured by the undrained triaxial test), taken from the

figures in Nash et al. (1992) is zsu 96.143.11 += where us is in kPa and z is the depth in metres

below the base of the excavation. Salient values are therefore kPa4.11=us at the soil surface,

13.4kPa at the base of the 1.5m caisson and 14.4kPa at the base of the 3.0m caisson. The bulk

density of the clay at relevant depths is estimated as 1680kg/m3. Throughout the testing period the

base of the test pit was covered by about 0.25m of water. The vertical bearing capacities of the

small and large caissons were estimated (using the method of Houlsby and Martin, 2003) as 164kN

and 746kN respectively.

In addition to the clevis pin load cells at the base of the A-frame on the 3.0m caisson, loads

applied by all hydraulic jacks used in the testing were measured by further load cells.

Displacements of the caissons were measured by draw-wire transducers attached to a scaffold frame

based at least 1.5m away from the outside of the caisson. The draw-wires were attached to the

caisson via upstand frames to keep the transducers clear of the water. Six transducers were used to

resolve all six degrees-of-freedom of movement. The co-ordinates of the transducers were first

determined using surveying techniques. During SEMV tests, displacements were also monitored by

means of accelerometers. One pore water pressure transducer was fixed centrally to the underside of

the lid of each caisson, and two more on the inside of the caisson just above the base of the skirt (at

opposite ends of a diameter). The general layout of the instrumentation is shown in Figure 5. All

transducer data were logged at 50Hz for most tests, with averages over 50 readings used for further

analysis, and at 400Hz for the SEMV tests (without data averaging).

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The caissons were installed according to the following procedure. (a) The caisson was

lowered to the soil surface and allowed to penetrate under its own weight with the interior vented to

atmosphere. (b) Water was pumped into the caisson until it was full. (c) The vent was closed and

water pumped out of the caisson to install it to full depth. During phase (c) dead weights were

added to the caisson to correct (as far as possible) any errors in levelling of the caisson. All

subsequent testing (except where noted) was carried out with the caisson vents sealed.

The tests on the 3.0m caisson were relevant to a monopod. Small amplitude cyclic horizontal

loads were applied at the top of the A-frame (4.23m above the lid of the caisson) by means of the

SEMV, operating at 10Hz, at which frequency the applied load was kN0.5± . These loads are

intended to be primarily representative of wave loads experienced by a prototype structure, but

scale to waves of different magnitudes (and return periods) for different sizes of prototype. No

attempt has been made to scale the wave frequency, as undrained conditions are assumed in both

test and prototype, and dynamic effects may be accounted for as discussed below. A fixed bias to

the horizontal loading (representing a wind and/or current loading) was achieved by suspending a

400kg block from a pulley system attached to the top of the A-frame (see Figure 3(b)). The vertical

load on the caisson throughout these tests was augmented by a 2400kg concrete block. During the

first series of SEMV tests an interesting observation was that, whilst the caisson moved only

imperceptibly, vibration at 10Hz transmitted through the ground set up a resonance in the scaffold

reference frame for displacement measurement (thus effectively rendering the displacement

measurements useless). This could not be eliminated satisfactorily by stiffening the frame, and in

subsequent tests displacements were measured by accelerometers attached directly to the caisson.

A second set of tests on the 3.0m caisson involved large amplitude (but low frequency) cyclic

horizontal loading from a hydraulic jack placed approximately horizontally between the top of the

A-frame and the main reaction frame, Figure 3(a). The amplitude of the loads was steadily

increased until large (>200mm) movements of the loading point occurred. These tests were

principally intended to assess the performance of the caisson under extreme conditions.

The tests on the 1.5m caisson were relevant to the tetrapod design. It was first loaded to a

fixed vertical load by means of a hydraulic jack. Cyclic inclined loading was then applied using a

second jack (inclined at 2:1 to the horizontal). Packets of 10 cycles of increasing load amplitude

were applied. The intention was that during these cycles the load in the vertical jack would be held

constant, but the stiffness of the hydraulic system rendered it difficult to control this load, which

therefore showed significant fluctuations. At the end of the testing the inclined jack was

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disconnected, and the caisson pulled out rapidly by the vertical jack to assess the ultimate tensile

capacity.

After the tests were completed the caissons were removed from the ground simply by

reversing the installation process: i.e. by reconnecting the pump and pumping water back into the

caisson. The 1.5m and 3.0m caissons were each installed and tested at two locations in the same test

pit at Bothkennar: a brief summary of the tests completed is given in Table 2.

TEST RESULTS

Installation

Figure 6 shows the records of measured suction against penetration depth for two installations

of the 3.0m caisson and one of the 1.5m caisson. Interruptions to the applied suction have been

removed from the records for clarity. It can be seen that the variation of suction with depth is

reasonably repeatable for the 3.0m caisson. Also shown on the figure are the computed profiles of

suction, using the procedure described by Houlsby and Byrne (2004) (modified slightly to account

for the fact that the caisson is not entirely submerged). The calculations used the strength profile

quoted in the previous section, together with an adhesion factor 5.0=α both inside and outside the

caisson and 9=cN on the rim of the caisson.

It can be seen that in general the computed suction pressure agrees well with the observations

as the full depth of the caisson is approached, but underestimates the suction at shallower depths.

The most probable explanation is that the fitted strength profile from Nash et al. (1992) is

appropriate principally for depths greater than about 3.5m from original ground surface (i.e.

m75.1>z approximately). At shallower depths the evidence from vane tests (Nash et al., 1992) is

that the strength increases significantly, probably due to past desiccation. The strength of the

shallow soil is probably underestimated in the calculation, leading in turn to an underestimation of

the suction.

Vibration tests on 3.0m caisson

Figure 7 shows the record of the applied moment (deduced from the load cells in the legs of

the A-frame) against time for an SEMV test. The test starts at an offset moment of approximately

16.6kNm (from the block and pulley system shown in Figure 3(b)). As the eccentric mass on the

SEMV starts to rotate, it exerts an inertial force at the top of the A-frame which varies with the

square of the angular velocity. The amplitude of loading therefore builds up steadily with the

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frequency, with a minor fluctuation at about 7Hz ( s11≈t in Figure 7) due to a resonance of the

suspended mass providing the offset moment.

Figure 8 shows the resulting moment-rotation response. The initial small amplitude cycles (at

low frequency) plot as the densely packed curves in the centre of the diagram, showing a high

stiffness and relatively little hysteresis. The line through the data shows a rotational stiffness of

225MNm/radian. As the amplitude of moment increases the moment-rotation loop opens up

gradually, until the steady state of 10Hz cycling is reached, at which stage an approximately

elliptical loop (the outer loop in the diagram) is continually retraced. The open loop arises from

damping which has three possible causes (a) viscous material damping, (b) plastic dissipation of

energy in the soil and (c) radiation damping.

The data can be interpreted by first taking the Fast Fourier Transform of both the moment and

rotation to convert to the frequency domain, and then taking the ratio between the two FFT’s to

obtain the complex, frequency-dependent impedance. The real part of the impedance represents the

stiffness and inertial effects, and the imaginary part the damping. The main information on the

effects of frequency on the response is contained within the transients at the beginning and end of

the test, and Figure 9 shows the real and imaginary parts of the transfer function computed for two

20s periods covering these transients. Also shown at 10Hz are the real and imaginary components

computed directly from the steady state response.

The data may be compared with theories for the behaviour of a circular foundation on an

elastic material. Wolf (1994) describes two lumped-parameter models for this case. Wolf presents

models both for surface and embedded footings. Whilst recognising that the caisson is in fact

embedded into the soil, we use here a preliminary analysis based on factors for surface footings.

The analysis of embedded footings requires consideration of cross-coupling between moment and

horizontal loading terms, and these coupling effects differ for the stiffness and damping.

Furthermore the role of the inertial terms in the rocking mode is not fully resolved. The principal

effect of ignoring the footing embedment is that, in the following, the stiffness coefficients may be

underestimated (by a factor which may be in the region of 1.5, but depending on the assumptions

about interactions on the side of the caisson, and the variation of stiffness with depth), and the shear

modulus correspondingly overestimated. Whilst this affects the absolute values of the moduli

discussed below, it does not affect their relative values.

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Wolf’s first model is a 3-parameter model represented conceptually by Figure 10(a). For a

rigid circular footing at the surface of an elastic soil subjected to moment loads, so that MF ≡ (the

applied moment) and θ≡u (the corresponding rotation), the rotational stiffness is ( )ν−=

138 3GRk ,

where G is the shear modulus and ν is Poisson’s ratio. The damping and mass coefficients are

calculated by kvRcs

γ= and kvRm

s2

2µ= , where γ and µ are dimensionless coefficients and

ρ= Gvs is the shear wave velocity ( ρ is the clay density). Note that although the soil is

assumed to be purely elastic, there is loss of energy through “radiation damping”, which accounts

for transmission of energy away from the foundation to an infinitely distant boundary. Adapting the

methods of Wolf (1994) and Das (1993) suggests the values 242.0=γ and 24.0=µ . The real part

of the impedance is mk 2ω− and the imaginary part is cω , and these are shown on Figure 9 for

comparison with the data, computed for MPa5.12=G . Examining first the real part, the theory

provides a reasonable fit to the data at low frequency, but overestimates the stiffness at higher

frequencies: this is because the higher frequencies involved higher amplitude cycling, for which the

secant shear stiffness of the soil would be expected to reduce.

Wolf (1994) suggested the alternative 5-parameter model shown conceptually in Figure 10(b)

as a more accurate representation of the foundation behaviour. The stiffness K is as for the 3-

parameter model, and damping and mass coefficients are defined in a similar way to those in the 3-

parameter model, but with suggested values for 5.0=ν of 00 =γ , 0267.00 =µ , 345.01 =γ ,

29.01 =µ . (Note that for the moment-rotation case the general 5-parameter model thus reduces

effectively to a 4-parameter model). In this case the real and imaginary parts of the impedance can

be calculated as

µ−

γµ+

µ− 0

221

21

21

2

11 a

aak and ( )

γ+

γµ+γ

µ02

121

21

21

2

1 aaak where svRa ω= is

the dimensionless angular velocity. Figure 9 shows that the 5-parameter model gives a very similar

variation of the real part of the impedance to the 3-parameter model. However, within the range of

frequencies tested it gives a lower imaginary part (which represents the damping).

Comparing with the data, it is clear that, in order to fit the real part of the impedance the

stiffness of the foundation would have to be reduced as the frequency increased (and amplitude of

loading increased). Of course this reduction in stiffness is accompanied by an increase in the

material damping, which is not taken into account by this model. One therefore expects the

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damping (imaginary part of the impedance) to be underestimated at higher frequencies by the

simple elastic model. The 5-parameter model therefore provides a better representation of the real

behaviour, since this underestimates the damping. It is worth noting that whilst the rotation of the

caisson during steady state cycling is very closely fitted by a 10Hz sinusoid, the moment contains

significant higher frequencies, with the 20Hz component being about 15% of the amplitude of the

10Hz fundamental. Such a response is indicative of non-linearities (such as plastic dissipation) not

accounted for in the simple 3- or 5-parameter models.

Jacking tests on 3.0m caisson

Following the SEMV tests, the 3.0m caisson was subjected to further cycles of moment, but

under quasi-static conditions, by loading with a hydraulic jack, Figure 3(a). Figure 11(a) shows an

example of the resulting moment-rotation curve for the first few, small amplitude, cycles of test

Jack_3.0_2, showing significant hysteresis even at this stage. Figure 11(b) shows the continuation

of the same test to larger amplitude, showing that not only does hysteresis increase with amplitude,

but also that there is a degradation of the stiffness over several cycles of loading. The degraded

response does, however, appear to be gradually stabilising.

Figure 12 shows very large amplitude cycles from test Jack_3.0_1. These cycles are largely of

curiosity value, since at such large displacements a full-size foundation would have “failed” for all

practical purposes. It is worth noting, however, the characteristic shape of the cycles in which, after

an initially stiff unloading, a very flexible response is observed, followed by a slight stiffening. This

behaviour is typical of a “gapping” response in which the stiffening occurs as a gap (created by the

previous half cycle) is closed. Indeed gaps several tens of millimetres wide and up to 1.02m deep

were measured down the side of the caisson during these cycles.

The SEMV and jacking tests may be compared as follows. At each frequency the real part of

the impedance can be used to deduce a secant shear stiffness of the soil. By making use of the fact

that the SEMV applies a load proportional to the square of the frequency, an amplitude of loading

can also be attributed to each frequency. Dividing the amplitude of loading by the impedance allows

an amplitude of rotation to be determined. Hence the secant stiffness can be expressed as a function

of amplitude of rotation. The results are shown on Figure 13, with the set of points for the initial

ramp up showing slightly higher stiffness than the ramp down, possibly because some degradation

of stiffness is attributable to the cycling. Also shown on the figure is a single point deduced from

the steady state conditions: the confidence attached to this value is much higher than for any of the

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other data, because it is based on many more readings. It is, however, entirely consistent with the

transient data.

Also shown on Figure 13 are the secant stiffness values from jacking test Jack_3.0_2

(conducted immediately after SEMV_2_2), and Jack_3.0_1 (conducted during an earlier installation

of the caisson). These are consistent with the SEMV data in that they show a continuing reduction

of the stiffness with increasing amplitude of cycling, indeed the shape of the ( ) G−θ∆log curve is

very similar to the familiar pattern for variation of stiffness with strain on a ( ) G−γ∆log plot. The

rotation of the caisson is of course approximately proportional to the shear strain amplitude in the

soil. Test Jack_3.0_1 shows a somewhat higher stiffness than Jack_3.0_2, probably because better

control of level was achieved during installation of this caisson, so that a better contact between soil

and the lid of the caisson was probably achieved. (In practice it is likely that any void between the

soil and the lid would be grouted to ensure best performance of the caisson). Finally Figure 13

shows a simple fit to the variation of the secant shear stiffness based on the hyperbolic moment-

rotation relation AMMMM

KM

+−

=θ max

max

0, where 0K is the initial value of the rotational stiffness,

maxM is the maximum moment, and AKK += 2500 where 50K is the secant rotational stiffness

at half the maximum moment. The curve is constructed for MNm/radian2520 =K (corresponding

to MPa140 =G ), kNm135max =M and 8=A .

As mentioned above, a more accurate interpretation of the moment loading tests could be

made by accounting for the embedment of the foundation in the calculation of stiffness factors (see

e.g. Doherty and Deeks (2003)), but this would simply reduce the absolute values of the estimated

stiffness, and not significantly change the relative values or the overall interpretation of the pattern

of response.

Jacking tests on 1.5m caisson

The 1.5m caisson was first loaded to approximately 120kN by the vertical jack, and then

subjected to cyclic loading from the inclined jack. Because of practical difficulties with

simultaneous control of pressures in two hydraulic jacks, the load path was not, however a simple

line in V-H space, but involved a rather complex path as shown in Figure 14 (for clarity in this

figure a considerable amount of data has been removed, and the paths plotted only for a few cycles

near the beginning, middle and end of the test). Importantly, however, the path involved the main

element of loading in the field, in that horizontal loading is accompanied by changes in vertical

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loading too: the path was selected to represent a realistic ratio between these changes. Note,

however, that in the prototype tetrapod the caisson would be restrained against rotation, whereas in

the tests the caisson was free to rotate.

Figure 15(a) shows the horizontal load against displacement for the first few cycles of test

Jack_1.5_2. In spite of the complications caused by scatter in data, the hysteresis loops are clear.

Figure 15(b) shows the continuation of the test to a series of packets of ten cycles at increasing

amplitude of load. For the tests up to kN30±=H there is very little degradation of response, but

the tests at kN40±=H show a clear degradation with cycling. Using the bearing capacity factors

of Houlsby and Martin (2003) the computed bearing capacity for the foundation is 163kN, so this

degradation may well be due to the bearing capacity of the foundation being reached during the

compressive cycles, see Figure 16.

Figure 16 shows the vertical movements throughout test Jack_1.5_2. During the first few

cycles there is very little vertical movement. During the intermediate cycles some vertical

downward movement accumulates during the cycling, but rapidly stabilises, and finally the largest

cycles cause ongoing significant vertical movements.

It appears therefore that for both horizontal load and moment cycling there is a pattern of stiff

response with little hysteresis at very small cyclic loads only. As loads increase the stiffness reduces

and hysteresis increases, but the loops are fairly stable. Eventually a load level is reached at which a

rapid deterioration of performance with number of cycles is observed.

Pullout tests on 1.5m caisson

At the end of the jacking tests on the 1.5m caisson, the caisson was pulled out rapidly by

means of the vertical jack. The results for test Pull_1 are shown in Figure 17. The tensile load

decreases rapidly to about -150kN, and after some minor fluctuations pullout occurs at a relatively

constant load. Figure 18 shows the variation with time of pore pressure measured under the lid of

the caisson and at the tip of the caisson. Also shown is the total vertical load, converted to the

dimensions of pressure by dividing by the area of the caisson. It can be seen that the pore pressure

under the lid is approximately -80kPa (relative to atmospheric pressure) throughout pullout,

indicating that cavitation has probably occurred, with the formation of a void between the caisson

lid and the soil. The tip pore pressure is approximately -100kPa, indicating cavitation beneath the

tip. The relatively small difference between the total load (converted to pressure) and the pressure

beneath the lid represents the friction of the sides of the caisson. The pressure difference of about

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12kPa converts to shear stress of about 2.25kPa on the inside and outside of the caisson, indicating

an α value (shear stress divided by undrained shear strength) of only 0.2. Note also on Figure 17

that once the tensile load exceeds about 15kN (corresponding approximately to the estimated

friction) the stiffness drops significantly below that encountered during compressive loading. (The

initial compressive loading curve is shown on Figure 17 for comparison, reversed and shifted to the

same origin as the pullout). Experience from model testing in sands (Kelly et al. 2004) suggests that

if the friction value is exceeded during cycling, that rapid degradation of the foundation would

occur.

After about 115mm of pullout (at time 16:44) the vent to the caisson was opened, at which

stage the pressure within the caisson rose to near atmospheric and extraction occurred at much

lower loads.

IMPLICATIONS FOR FULL-SCALE FOUNDATIONS

The principal purpose of the tests described here is for calibration of “force resultant”

theoretical models based on work-hardening plasticity to describe the response of caisson

foundations (Houlsby, 2003). However, some simple scaling can be applied to the results of the

tests to make some preliminary estimates of the sizes of caissons that would be needed for full scale

wind turbine installations.

A 3.5MN wind turbine in typical offshore conditions would result in an overturning moment,

in extreme conditions, of approximately 120MNm (Byrne and Houlsby, 2003). If at this stage it is

determined that (say) an acceptable one-way rotation of the foundation is 0.001 radian, then the

results from Figure 11 indicate that for a 2-way rotation of 0.002 radian a typical mobilised usG

value for a soft clay as at Bothkennar would be about 175. Assuming a soil of strength kPa60=us ,

it can be estimated that a caisson of diameter 26m would be required to provide a sufficiently stiff

response. The cyclic nature of the applied loading is principally due to the waves, which may have a

period of about 10s. For this case the dimensionless frequency a would be about 0.34, indicating

that the dynamic effects would be small, and a quasi-static analysis of the foundation would be

justified.

If alternatively an approach based on strength were adopted, then it may be estimated that the

3.0m caisson was able to sustain cyclic moments of about 70kNm without significant degradation

of response. Since the moment capacity scales linearly with the shear strength, and with the cube of

the foundation size, it is concluded that a foundation of 22m diameter would be required in 60kPa

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clay with similar properties to that at Bothkennar. Either a strength or a stiffness criterion therefore

results in a foundation of comparable magnitude, but serviceability considerations (i.e.

deformations) lead to a requirement for a larger foundation.

If a tetrapod were to be designed then first the caisson spacing must be determined. For an

overturning moment of 120MNm and a weight of the structure of say 6MN, then a spacing of 40m

is needed if tension is to be avoided completely. The maximum loading on an individual caisson

would be 3MN, which could be carried in a clay of strength 60kPa with a factor of safety of about

1.5 by a caisson of diameter 4.0m. The estimated shear load of 4MN could also be carried by

foundations of this size. It is difficult, however, to assess the influence such a caisson would have

on the stiffness of the structure without more detailed knowledge of the structure itself.

CONCLUSIONS

A series of field trials of caisson foundations in soft clay are described. The tests are relevant

to both monopod and tetrapod designs for foundations for offshore wind turbines. Installation of the

caissons was achieved by suction. High frequency, low amplitude cyclic moment tests on a 3.0m

caisson showed that the response was affected by stiffness, inertial and damping effects. Low

frequency cyclic moment tests on the 3.0m caisson indicate a stiff response at low amplitude, with a

gradual reduction of stiffness and increase of hysteresis at large amplitude. There was evidence of

gapping at the side of the caisson under very large amplitude cycles. Cyclic inclined loading tests

on a 1.5m diameter caisson also show a reduction of stiffness and increase of hysteresis as load

amplitude increases, with a significant reduction in stiffness after the compression to tension

boundary is crossed and frictional capacity exceeded. Pullout of the 1.5m caisson indicated that

ultimate tensile resistance is governed by cavitation beneath the foundation. The tests contribute to

the development of design procedures for offshore wind turbines founded on caissons.

ACKNOWLEDGEMENTS

This research was sponsored by the DTI and a consortium of companies (Fugro Ltd, SLP

Engineering Ltd, Garrad Hassan, General Electric Wind Ltd, Aerolaminates Ltd and Shell

Renewables Ltd). The authors are very grateful to the Royal Society for the Protection of Birds (and

in particular to Mike Trubridge) for making the site available for this testing. The authors thank

Dr A. Blakeborough for use of the SEMV designed by him, and for advice on interpretation of the

SEMV tests.

13

REFERENCES

Bransby, M.F. and Randolph, M.F, (1998) “Combined loading of skirted foundations”, Géotechnique, Vol. 48, No. 5, 637-655

Bye, A., Erbrich, C., Rognlien, B. and Tjelta, T.I. (1995) “Geotechnical design of bucket foundations”, Proc. Offshore Technology Conference, OTC 7793

Byrne, B.W. and Houlsby, G.T. (2002) “Experimental Investigations of the Response of Suction Caissons to Transient Vertical Loading”, Proc. ASCE, Journal of Geotechnical Engineering, Vol. 128, No. 11, Nov., pp 926-939

Byrne, B.W. and Houlsby, G.T. (2003) “Foundations for Offshore Wind Turbines”, Phil. Trans. of the Royal Society of London, Series A, Vol. 361, December, pp 2909-2930

Byrne, B.W., Houlsby, G.T., Martin, C.M. and Fish, P. (2002) "Suction Caisson Foundations for Offshore Wind Turbines", Wind Engineering, Vol. 26, No. 3, pp 145-155

Byrne, B.W., Villalobos, F. Houlsby, G.T. and Martin, C.M. (2003) “Laboratory Testing of Shallow Skirted Foundations in Sand”, Proc. Int. Conf. on Foundations, Dundee, 2-5 September, Thomas Telford, pp 161-173

Byrne, B.W. and Houlsby, G.T. (2004) “Experimental Investigations of the Response of Suction Caissons to Transient Combined Loading”, Proc. ASCE, Jour. of Geotechnical and Geoenvironmental Engineering, Vol. 130, No. 3, pp 240-253

Das, B.M. (1993) “Principles of Soil Dynamics”, Brooks/Cole, California

Doherty, J.P. and Deeks, A.J. (2003) “Elastic response of circular footings embedded in a non-homogeneous half-space”, Géotechnique, Vol. 53, No. 8, October, pp 703-714

Gourvenec, S. and Randolph, M.F. (2003) “Bearing capacity of a skirted foundation under V,H,M loading”, Proc. 22nd Int. Conf. On Offshore mechanics and Arctic Engineering, 8-13 June, Cancun, Mexico, OMAE2003-37014

Houlsby, G.T. (2003) "Modelling of Shallow Foundations for Offshore Structures", Invited Theme Lecture, Proc. Int. Conf. on Foundations, Dundee, 2-5 September, Thomas Telford, pp 11-26

Houlsby, G.T. and Byrne, B.W. (2000) “Suction Caisson Foundations for Offshore Wind Turbines and Anemometer Masts”, Wind Engineering, Vol. 24, No. 4, pp 249-255

Houlsby, G.T. and Byrne (2004) “Calculation procedures for installation of suction caissons in clay”, submitted to Geotechnical Engineering

Houlsby, G.T and Martin, C.M. (2003) "Undrained Bearing Capacity Factors for Conical Footings on Clay", Géotechnique, Vol. 53, No. 5, June, pp 513-520

Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2003) "Pressure Chamber Testing of Model Caisson Foundations in Sand", Proc. Int. Conf. Foundations, Dundee, 2-5 September, Thomas Telford, pp 421-431

Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2004) "Tensile Loading of Model Caisson Foundations for Structures on Sand", Proc. ISOPE, Toulon, in press

Nash, D.F.T., Powell, J.J.M. and Lloyd, I.M. (1992) “Initial Investigations of the Soft Clay Test Site at Bothkennar”, Géotechnique, Vol. 42, No. 2, pp 163-181

Wolf, J.P. (1994) “Foundation Vibration Analysis Using Simple Physical Models”, Prentice Hall, New Jersey

14

Diameter Skirt length DL ratio Wall thickness Approximate mass (including appurtenances)

1.5m 1.0m 0.67 8mm 670kg

3.0m 1.5m 0.5 8mm 2000kg

Table 1: Details of caisson dimensions

Caisson Installation Test type Code Notes

Installation Inst_1.5_1

Jacking test Jack_1.5_1

1

Pull out Pull_1

Installation Inst_1.5_2 No suction data

Jacking test Jack_1.5_2

1.5m

2

Pull out Pull_2 Inclined jack attached, but not pressurised

Installation Inst_3.0_1

SEMV tests SEMV_1_1

SEMV_1_2

SEMV_1_3

No accelerometer data

No accelerometer data

No accelerometer data

1

Jacking test Jack_3.0_1

Installation Inst_3.0_4

SEMV tests SEMV_2_1

SEMV_2_2

50Hz logging

3.0m

2

Jacking test Jack_3.0_2

Table 2: Outline of caisson tests carried out at Bothkennar

Flow

Pressuredifferential

W

Flow

Figure 1: Installation of a suction caisson

15

CaissonCaissons

Turbinesupportstructure

Seabed

Water surface

h h

D

sD

Figure 2: Possible configurations for suction caisson foundations for wind turbines

4000 4000 6000

30001500

H

H

H B

CC

A

RR

W

V

A

B

C

V

LL

L

L L LL

(a) (b) Figure 3: Outline of field testing equipment, dimensions in mm (water in excavation and displacement reference frames not shown). (a) arrangement for jacking tests on 1.5m and 3.0m 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

16

Figure 4: Test rig showing the 1.5m caisson installed and 3.0m caisson in place for installation

Pressuresensors

Draw-wiredisplacementtransducers

VerticalAccelerometer

3-axisAccelerometer

Figure 5: Outline of instrumentation on caisson (draw wire transducer reference frame not shown)

17

0

200

400

600

800

1000

1200

1400

1600

0 5 10 15 20 25 30 35 40Suction (kPa)

Dis

plac

emen

t (m

m)

1.5m Inst_1.5_11.5m calculated3.0m Inst_3.0_13.0m Inst_3.0_23.0m calculated

Figure 6: Records of suction during penetration

-20

-10

0

10

20

30

40

0 2 4 6 8 10 12 14 16 18 20

Time (s)

Mom

ent M

(kN

m)

Figure 7: Moment v. time for initial phase of test SEMV_2_2

18

-10

0

10

20

30

40

50

-0.0004 -0.0003 -0.0002 -0.0001 0 0.0001 0.0002 0.0003 0.0004

Rotation θ (radians)

Mom

ent M

(kN

m)

Figure 8: Moment-rotation response of caisson in test SEMV_2_2

0

50000

100000

150000

200000

250000

0 1 2 3 4 5 6 7 8 9 10

Frequency (Hz)

Impe

danc

e (k

Nm

/radi

an)

Real, 3-parameterReal, 5-parameterReal SEMV, ramp upReal SEMV, ramp downReal, 10Hz steady stateImaginary, 3-parameterImaginary, 5-parameterImaginary SEMV, ramp upImaginary SEMV, ramp downImaginary, 10Hz steady state

Figure 9: Complex M−θ transfer function for test SEMV_2_2 compared with theoretical expressions for 3- and 5-parameter models

19

m

k c

Fu

m0

k c0

Fu

m1

u1c1

(a) (b)

Figure 10: Conceptual models for stiffness, damping and mass of a foundation (a) 3-parameter model, (b) 5-parameter model

-100

-80

-60

-40

-20

0

20

40

60

80

100

-0.002 -0.001 0 0.001 0.002

Rotation θ (radians)

Mom

ent M

(kN

m)

-100

-80

-60

-40

-20

0

20

40

60

80

100

-0.004 -0.002 0 0.002 0.004 0.006 0.008

Rotation θ (radians)

Mom

ent M

(kN

m)

(a) (b)

Figure 11: Moment-rotation curve for loading of 3.0m caisson, test Jack_3.0_2. (a) detail of small amplitude cycles, (b) medium amplitude cycles

20

-200

-150

-100

-50

0

50

100

150

200

250

-0.04 -0.02 0 0.02 0.04 0.06 0.08 0.1

Rotation θ (radians)

Mom

ent M

(kN

m)

Figure 12: Moment-rotation curve for loading of 3.0m caisson, test Jack_3.0_1: large amplitude cycles

0

2

4

6

8

10

12

14

16

0.00001 0.0001 0.001 0.01 0.1

∆θ (radians)

G (M

Pa)

SEMV_2_2, ramp upSEMV_2_2, ramp downSEMV_2_2, 10Hz steady stateJack_3.0_2Jack_3.0_1Hyperbolic fit

Figure 13: Computed secant shear modulus from test SEMV_2_2 and jacking tests on 3.0m caisson

21

-50

-40

-30

-20

-10

0

10

20

30

40

50

0 50 100 150 200

Vertical Load (kN)

Hor

izon

tal L

oad

(kN

)

Figure 14: Load path in cyclic inclined loading test Jack_1.5_2 on 1.5m caisson

-15

-10

-5

0

5

10

15

0 0.5 1 1.5 2

Horizontal Displacement (mm)

Hor

izon

tal L

oad

(kN

)

-50

-40

-30

-20

-10

0

10

20

30

40

50

-20 -10 0 10 20 30 40 50 60

Horizontal Displacement (mm)

Hor

izon

tal L

oad

(kN

)

(a) (b)

Figure 15: Horizontal movement during inclined loading test Jack_1.5_2 on 1.5m caisson. (a) detail of small amplitude cycles, (b) large amplitude cycles

22

0

20

40

60

80

100

120

140

160

180

0 5 10 15 20 25Vertical Displacement (mm)

Verti

cal L

oad

(kN

)

Figure 16: Vertical movement during inclined cyclic loading test Jack_1.5_2 on 1.5m caisson

-180

-160

-140

-120

-100

-80

-60

-40

-20

0-200 -175 -150 -125 -100 -75 -50 -25 0

Vertical displacement (mm)

Verti

cal l

oad

(kN

)

PulloutLoading curve (shifted and reversed)

Figure 17: Load v. displacement during pullout test Pull_1

23

-120

-100

-80

-60

-40

-20

0

20

16:25 16:30 16:35 16:40 16:45 16:50

Pres

sure

(kPa

)

Lid pressureSkirt pressureTotal load/area

Figure 18: Record of pull out of 1.5m caisson, test Pull_1

1

Field trials of suction caissons in sand for offshore wind turbine foundations

G.T. Houlsby1, R.B. Kelly1, J. Huxtable2 and B.W. Byrne1 Keywords: bearing capacity, sand, dynamic test, foundations, stiffness

ABSTRACT

A programme of testing on suction caisson foundations in an artificially prepared sand test

bed near Luce Bay, in Scotland, is described. The tests are relevant to the design of either monopod

or quadruped foundations for offshore wind turbines. Records are presented for suction installation

of the caissons, cyclic moment loading under both quasi-static and dynamic conditions to simulate

the behaviour of a monopod foundation, and cyclic vertical loading and pullout of caissons to

simulate one footing in a quadruped foundation. Variations of stiffness with loading level of the

foundation are observed, with high initial stiffness followed by hysteretic behaviour at moderate

loads and degradation of response at high loads. Some implications for the design of wind turbine

foundations are briefly discussed.

INTRODUCTION

The offshore wind energy industry is a very rapidly expanding sector of vital economic

importance in the UK, and the foundation represents an important part of the costs of offshore wind

turbine installations (Byrne and Houlsby, 2003). This paper describes a series of tests directed

towards the understanding of the behaviour of suction caisson foundations in sand, as possible

foundations for offshore wind turbines.

An earlier paper, by Houlsby et al. (2004), describes in detail the background motivation to

this testing, and presents equivalent data for tests in clay. The testing programme described here

follows a similar pattern to the tests in clay, and includes tests directed towards the design of both

monopod and quadruped foundations for offshore wind turbines. The most important load case for a

monopod foundation is the applied overturning moment, whilst for a quadruped foundation thee

vertical loading on an individual caissons is most important. Data were obtained from the

installation phases for each caisson. Loading of the caissons was by means of a combination of dead

weights, hydraulic jacks and inertial loading from a “Structural Eccentric Mass Vibrator” (SEMV).

The test programme was designed by Oxford University, and site operations were managed by

Fugro Structural Monitoring Ltd. The tests were carried out in February and March 2004. 1 Department of Engineering Science, Oxford University 2 Fugro Structural Monitoring Ltd.

2

THE LUCE BAY TEST SITE

The tests were carried out in a specially prepared sand bed in an aggregate extraction quarry

near Luce Bay, Dumfries and Galloway, Scotland. The bed was prepared by placing selected fill in

layers approximately 250mm thick and compacting them by multiple passes of a wheeled loader.

The sand bed was approximately 40m x 15m x 3.5m deep. A grading curve for the sand is shown in

Figure 1. Although the sand itself is almost single-sized at about 0.3mm to 0.4mm, it can be seen

that there is a significant (15%) gravel content. The sand bed was also observed to contain a

proportion of small rounded stones up to about 80mm in size, even though every attempt was made

to remove this coarser material in the preparation of the bed. The fines content is negligible. The fill

was placed in an unsaturated condition above the water table in the summer. The entire sand bed

was allowed to flood slowly over a period of several months, and by the time of the testing had been

submerged for four months. During testing there was about 150mm depth of water over the sand

surface.

The bed of sand was characterised principally by means of in situ testing. Three CPT tests

with pore pressure measurement, two cone pressuremeter tests and three seismic cone tests were

conducted. The cone resistance records of three CPT and two cone pressuremeter tests are shown

superimposed in Figure 2, showing (within some variability) a strong increase of cone resistance

with depth, principally due to the increasing stress level. Using standard correlations, the estimated

relative density of the sand bed was 80-85%. In the lower part of the sand layer there appears to be

some looser material, and beneath the sand are much softer deposits that were not investigated in

detail.

Figure 3 shows profiles of shear modulus with depth inferred from the cone pressuremeter

and seismic cone tests. Down to a depth of 2m there is generally good agreement between the

seismic and cone pressuremeter data, with the exception of that from seismic cone test 1. At greater

depths, the cone pressuremeter data suggests a higher shear modulus than the seismic cone data.

The shear stiffness can be characterised by the relationship aa pppG ′= 2500 , where ap is

atmospheric pressure and p’ is the mean effective stress estimated on the basis that the saturated

unit weight was 10.3kN/m3 and estimating K0 as 0.5. This relationship has been fitted to the upper

bound of the test data, as disturbance during testing is expected to lead to underestimates of the

small strain shear modulus. We note, however, that the shear moduli observed at this site are quite

high for sand at shallow depths, and this may be due to the compaction procedures used in

preparation of the bed.

3

EQUIPMENT AND TESTING PROCEDURES

The caissons, loading frame and instrumentation methods used for the tests in sand were

essentially identical to those employed in the previous clay tests at Bothkennar (Houlsby et al.,

2004). A caisson of diameter of 3.0m and a with skirt 1.5m deep was used for moment loading

tests. A second caisson of diameter of 1.5m and with a 1.0m skirt was used for vertical loading

tests. An outline diagram of the test set-up is given in Figure 4, and a photograph of the rig installed

at the Luce Bay testing site shown in Figure 5.

Prior to suction-assisted installation the caissons were allowed to penetrate under their own

weight with the vent to the caisson open to air. In the case of the 3m caisson a 2400kg mass was

added to the caisson to cause a slight further penetration. The caisson vent was then closed and

suction applied by pumping out the trapped air from inside the caisson. The pump used was capable

of pumping both air and water. A flow meter was installed between the caisson and pump.

The testing procedures in sand were very similar to those adopted for the tests in clay

(Houlsby et al, 2004). The principal exception related to the “SEMV” tests using a Structural

Eccentric Mass Vibrator to apply relatively high-frequency cyclic loading. Experience from the

earlier clay tests indicated the value of having a spread of data obtained at different frequencies of

loading. At Luce Bay a series of tests were therefore carried out, in which the frequency was

increased in steps, with approximately 15s of cycling at each frequency. A more reliable definition

of the variation of the complex impedance with frequency was possible using this method.

The 1.5m and 3.0m caissons were each installed and tested at two locations in the test bed at

Luce Bay: a brief summary of the tests completed is given in Table 1.

TEST RESULTS

Installation

Figure 6 shows the records of measured suction against penetration depth for one installation

of the 3.0m caisson and one of the 1.5m caisson. It can be seen that the variation of the required

suction with depth is similar for both caissons. Also shown on the figure are the computed profiles

of suction, using the procedure described by Houlsby and Byrne (2004) (modified slightly to

account for the fact that the caisson is not entirely submerged). The parameters used in the

calculations were a soil friction angle of 45°, Ktanδ = 1, an effective unit weight of 10.3kN/m3, the

ratio of permeability inside the caisson to outside the caisson was 3 and stress distribution factors

‘m’ and ‘n’ were taken as 1. The vertical load, including self weight, applied to the caissons was

4

7kN and 60kN for the 1.5m and 3.0m diameter caissons respectively. Using these values, the final

pressure for installation of the 3m diameter caisson is predicted closely, while the final pressure

required to install the 1.5m diameter caisson is under-predicted. During the actual installations there

were a number of stoppages in pumping for a variety of reasons. It was observed that after each

stoppage movement only occurred once the suction returned to a value similar to that before the

stoppage.

Figure 7 shows the excess pore pressure measured by sensors placed inside the caissons,

50mm above the tip of the skirts, plotted against suction measured immediately beneath the lid of

the caisson. Although there is some scatter, largely related to the stoppages, there is a strong

correlation between the two pressures, as would be expected. Approximately 65% of the suction

pressure is measured near the tip of the skirts, and this fraction appears to remain approximately

constant throughout the installations.

Figure 8 shows a comparison between the volume of air and water removed from the caisson

(by integration of the flow measurement) and the volume computed from the cross sectional area of

the caisson multiplied by the depth of penetration. There is a close relationship between the two. It

can be concluded that during the installation (a) little heave occurred within the caisson and (b) the

volume of water seeping through the sand was small. Either of the above phenomena would have

resulted in significantly larger volumes from the flow measurements.

Vibration tests on 3.0m caisson

Figure 9 shows the record of the applied moment (deduced from the load cells in the legs of

the A-frame) against time for an SEMV test in which the excitation is ramped to a constant

frequency of 8Hz and held for about 18s. As the eccentric mass on the SEMV starts to rotate, it

exerts an inertial force at the top of the A-frame which varies with the square of the angular

velocity. The amplitude of loading therefore builds up parabolically with the frequency.

Figure 10 shows results demonstrating the moment-rotation response, and was compiled by

extracting data from a number of tests, each similar to the one shown in Figure 9, but at different

frequencies. From each test the steady state response has been extracted. The small amplitude

cycles (at frequencies less than 6Hz) are not shown, as the movements induced in the caisson were

too small to be accurately recorded by the accelerometers. As the amplitude of moment increases

the moment-rotation loop opens up gradually, until the steady state of 10Hz cycling is reached, at

which stage an approximately elliptical loop is continually retraced. The open loop arises from

5

damping which has three possible causes (a) viscous material damping, (b) plastic dissipation of

energy in the soil and (c) radiation damping. The minor axes of these ellipses are much smaller than

those recorded at Bothkennar, indicating that much less damping occurs in dense sand than in soft

clay.

The data can be interpreted by first taking the Fast Fourier Transform of both the moment and

rotation signals to convert to the frequency domain, and then taking the ratio between the two FFT’s

to obtain the complex, frequency-dependent “impedance” (a generalised form of the rotational

stiffness). The real part of the impedance represents stiffness and inertial effects, and the imaginary

part the damping. At Bothkennar, the main information on the effects of frequency on the response

was obtained from the transients at the beginning and end of the test, while at Luce Bay the

information was gathered from the steady state of individual tests, stepped at 1Hz intervals from

1Hz to 10Hz. Figure 11 shows the real and imaginary parts of the transfer function computed from

5Hz-10Hz. Also shown as discrete points are the average responses computed at each steady state.

The averaged points can be regarded as much more reliable, as they are derived from much more

data, but it can be seen that in fact the transfer function deduced from the transient response agrees

well with the averaged points.

The data may be compared with theories for the behaviour of a circular foundation on an

elastic material. As discussed by Houlsby et al. (2004), Wolf (1994) describes two lumped-

parameter models for this case, a 3-parameter model and a 5-parameter model. The fitted response

from a 3-parameter model is shown on Figure 11 for comparison with the test data, computed for

MPa85=G and ν = 0.2. Examining first the real part, the stiffness is chosen to provide a reasonable

fit to the data at frequencies in the range 5Hz to 6Hz, but overestimates the stiffness at higher

frequencies. This is of course consistent with the common observation that the stiffness of soils

reduces with strain amplitude, so that a constant stiffness model cannot fit behaviour across a wide

range of strain amplitudes. No fit from a 5-parameter model is shown in Figure 11, as the 5-

parameter model gives a very similar variation of the real part of the impedance to the 3-parameter

model. However, within the range of frequencies tested the 5-parameter model gives a much lower

imaginary part of the impedance (representing the damping). The 3-parameter model appears to be

more satisfactory in this respect.

Jacking tests on 3.0m caisson

Following the SEMV tests, the 3.0m caisson was subjected to further cycles of moment, but

under quasi-static conditions, by loading with a hydraulic jack, Figure 4(a). Figure 12 shows the

6

moment-rotation curve for cycles from test LB_Jack_3.0_1. The moment-rotation response is

initially stiff with little hysteresis. At larger amplitudes of rotation the secant stiffness decreases and

hysteresis increases as the amplitude increases. The cycles at very large amplitude have a

characteristic shape in which, after an initially stiff unloading, a very flexible response is observed,

followed by slight stiffening. This behaviour is typical of a “gapping” response in which the

stiffening occurs as the gap created during the previous half cycle is closed. Gaps were observed

down the side of the caisson during these cycles.

The moment-rotation curve for jacking test LB_Jack_3.0_2 is shown in Figure 13. In this test

packets of 10 cycles were applied to the caisson with amplitudes of ±42kN/m, ±85kN/m, ±169kN/m

and ±254kN/m. As for the data in Figure 12, the displacement amplitude increases with load

amplitude and hysteresis occurs during the larger amplitude cycles. The unload-reload parts of the

curve in Figure 13 are stiffer than those in Figure 12. There is a slight shakedown apparent in the

data at low amplitudes of rotation, with a slight stiffening occurring over several cycles of the same

amplitude. The moment-rotation response during each packet of cycles appears to approach a steady

state by the end of the packet. Byrne and Houlsby (2004) have made similar observations about

hysteretic behaviour during cycling from small-scale moment loading tests on caissons in the

laboratory.

The SEMV and jacking tests may be compared by deducing the equivalent secant shear

modulus from the moment-rotation response of the footing using standard formulae for response of

a surface footing to moment loads. For the SEMV tests the moment range is taken as the difference

between the maximum and minimum values in any given cycle, and the corresponding rotations are

computed. The secant stiffness can then be expressed as a function of amplitude of rotation as

shown in Figure 14. Also shown on Figure 14 are the secant stiffness values from jacking test

LB_Jack_3.0_1. These are consistent with the SEMV data in that they show a continuing reduction

of the stiffness with increasing amplitude of cycling: indeed the shape of the ( ) G−θ∆log curve is

very similar to the familiar pattern for variation of stiffness with strain on a ( ) G−γ∆log plot. The

rotation of the caisson is of course approximately proportional to the shear strain amplitude in the

soil. Finally Figure 14 shows a simple fit to the variation of the secant shear stiffness based on the

hyperbolic moment-rotation relation AMMMM

KM

+−

=θ max

max

0, where 0K is the initial value of the

rotational stiffness, maxM is the maximum moment, and AKK += 2500 . The curve is

7

constructed for MNm/radian11250 =K (corresponding to MPa1000 =G ), kNm450max =M and

4=A .

More accurate interpretation of the moment loading tests could be made by accounting for the

embedment in the calculation of stiffness factors (see e.g. Doherty and Deeks (2003)), but this

would simply reduce the absolute values of the estimated soil stiffness, and not significantly change

the relative values in the overall interpretation of the pattern of response.

Jacking tests on 1.5m diameter caissons

Two jacking tests were conducted on the 1.5m caisson. Test LB_Jack_1.5_1 involved an

application of combined vertical and horizontal loads to the caisson via vertical and inclined jacks.

A vertical jack applied the mean vertical load, while the inclined jack provided a cyclic loading

component. This test was not successful, as it was not possible to control either the mean or cyclic

loads accurately, because the soil/caisson/jack system was sufficiently stiff that the jacks were not

able to operate independently of each other.

Cyclic vertical loading was applied to the 1.5m diameter caisson in Test LB_Jack_1.5_2

about a mean load of 60kN (not including the self-weight of the caisson). Packets of 10 cycles were

applied to the caisson with load amplitudes starting at ±10kN and increasing in steps of ±10kN to

±100kN, and these are shown in Figure 15. The data show that the secant stiffness decreases as the

amplitude of the load increases. Furthermore the displacements increase markedly during load

packets where the caisson was cycled into tension. It is of interest to note that the net displacement

after each cycle in the ±100kN load packet was downward, even though large tensile displacements

had occurred during each cycle. The implication of this is that, as long as the mean vertical load is

compressive, a caisson foundation cycled into tension will ratchet into the sand rather than out of

the sand. However, the data also suggest that caissons should not be loaded in tension for

serviceability reasons, as the stiffness of the caisson reduces to a level where foundation movements

would render a wind turbine inoperable. The data show that the caisson has a significant cyclic

tensile capacity (in excess of -40kN) but extremely large displacements are required to mobilise this

capacity. These data support conclusions made from small-scale laboratory tests (Byrne and

Houlsby 2002, Kelly et al., 2003, 2004), and this comparison was an important objective for the

large-scale testing programme.

8

Pullout tests on 1.5m caisson

At the end of the jacking tests on the 1.5m caisson, the caisson was pulled out rapidly by

means of the vertical jack. The results for test LB_Pull_2 are shown in Figure 16. The tensile load

decreases rapidly to about -120kN, when a vent plug in the lid of the caisson was opened to prevent

damage to the load cell. The tensile load had not reached a maximum by this stage. The proportion

of the tensile load generated by suction pressure inside the caisson is also shown in Figure 16. The

difference between the total load carried by the caisson and the suction load represents the friction

acting on the skirts of the caisson.

The frictional load generated on the skirts of the caisson, as it was pulled out of the sand, is

shown in Figure 17, along with an estimate of the friction load computed using the method

described by Kelly et al., (2003). The calculation was conducted with Ktanδ = 0.75. This

calculation is based on similar principles to that used to predict the suction during installation

shown in Figure 6. The parameters used in the calculations are identical, with the exception that

Ktanδ = 1 was used for installation. The differences between the values of Ktanδ used are

considered to be relatively small, given the approximations involved in modelling caisson

installation/extraction in an idealised soil.

IMPLICATIONS FOR FULL-SCALE FOUNDATIONS

The principal purpose of the tests described here is for calibration of “force resultant”

theoretical models based on work-hardening plasticity to describe the response of caisson

foundations (Houlsby, 2003). However, some simple scaling can be applied to the results of the

tests to make some preliminary estimates of the sizes of caissons that would be needed for full-scale

wind turbine installations.

A 3.5MN wind turbine in typical offshore conditions would result in an overturning moment,

in extreme conditions, of approximately 120MNm (Byrne and Houlsby, 2003). If at this stage it is

determined that (say) an acceptable two-way rotation of the foundation is 0.002 radian, then the

results from Figure 14 indicate that for this rotation the shear modulus would be about 15MPa. If it

is assumed that the shear modulus scales with the square root of the mean stress, then it can be

estimated that a caisson of diameter 20m would be required to provide a sufficiently stiff response

in sand with a relative density of about 80%. The cyclic nature of the applied loading is principally

due to the waves, which may have a period of about 10s. For this case the dimensionless frequency

a = ωR/cs, (where ω is the frequency, R is the caisson radius and cs is the shear wave velocity =

9

200m/s), would be about 0.03, indicating that the dynamic effects would be minimal (Wolf, 1994),

and a quasi-static analysis of the foundation would be fully justified.

If alternatively an approach based on strength were adopted, then it may be estimated that the

3.0m caisson was able to sustain cyclic moments of about 0.12MNm without significant

degradation of response. Since the moment capacity scales linearly with the effective unit weight,

and with the fourth power of the foundation size, it is concluded that a foundation of 17m diameter

would be required in sand with a unit weight of 10.3kN/m3 with similar properties to that at Luce

Bay. Byrne and Houlsby (2003) proposed Equation 1 as another method for estimating the diameter

of a monopod foundation at low vertical loads:

( )WVM

2RH2RM

3

1

21 fff +

+=−

(1)

where, M is the overturning moment, R is the radius of the caisson, H is the horizontal load, V is the

vertical load and W is the effective weight of the sand inside the caisson. The factors f1, f2 and f3

were obtained from a limited number of small-scale laboratory tests and are equal to 3.26, 1.07 and

0.71 respectively. If the weight of the structure was 6MN and the horizontal load was 4MN then

Equation 1 indicates that the diameter of the monopod foundation would be about 19m if the ratio

of the length of its skirt to its diameter were 0.5.

Either the strength or the stiffness criterion therefore results in a foundation of comparable

size, but the latter, which is related serviceability considerations (i.e. deformations) leads to a

requirement for a slightly larger foundation.

If a quadruped were to be designed then first the caisson spacing must be determined. For an

overturning moment of 120MNm and a weight of the structure of 6MN, then a spacing of 40m is

needed if tension is to be avoided completely. The confirmation on the basis of large scale tests that

tension should be avoided is considered an important, though negative, outcome from this research.

The maximum loading on an individual caisson would be 3MN, which could be carried in sand like

that at Luce Bay with a factor of safety of about 1.5 by a caisson of diameter 3.5m. The estimated

shear load of 4MN can be carried by a caisson with a diameter of 4m and a skirt length of 2.67m. In

this case the shear loading appears to govern the size of the caisson required, but it is considered

likely that this would change when deformations are taken into account. Unfortunately, however, it

is difficult to assess the influence the caisson size would have on the overall stiffness of the

structure, without more detailed knowledge of the structure itself.

10

A comparison of the estimated caisson diameters for a wind turbine foundation in dense sand

with those estimated for soft clay by Houlsby et al., (2004) shows that the size of a mono-caisson

foundation in dense sand would be about 60% of that in soft clay, whereas the sizes of caissons in a

quadruped foundation would be rather similar.

CONCLUSIONS

A series of field trials of caisson foundations in sand are described. The tests are relevant to

the design of both monopod and quadruped foundations for offshore wind turbines. Installation of

the caissons was achieved by suction. High frequency, low amplitude cyclic moment tests on a

3.0m caisson showed that the response was affected by stiffness, inertial and damping effects. Low

frequency cyclic moment tests on the 3.0m caisson indicate a stiff response at low amplitude, with a

gradual reduction of stiffness and increase of hysteresis at large amplitude. There was evidence of

gapping at the side of the caisson under very large amplitude cycles. Cyclic vertical loading tests on

a 1.5m diameter caisson also show a reduction of stiffness and increase of hysteresis as load

amplitude increases, with a significant reduction in stiffness after the compression to tension

boundary is crossed and frictional capacity exceeded. Pullout of the 1.5m caisson indicated that a

sizable ultimate tensile resistance can be generated but is accompanied by extremely large

displacements. The tests contribute to the development of design procedures for offshore wind

turbines founded on caissons.

ACKNOWLEDGEMENTS

This research was sponsored by the DTI and a consortium of companies (Fugro Ltd, SLP

Engineering Ltd, Garrad Hassan, General Electric Wind Ltd, Aerolaminates Ltd and Shell

Renewables Ltd). The authors are very grateful to Mr Adam Macintosh of Luce Bay Plant Hire for

making the site available for this testing. The authors thank Dr A. Blakeborough for use of the

SEMV designed by him, and for advice on interpretation of the SEMV tests. Dr B.W. Byrne

acknowledges the support provided by Magdalen College, Oxford.

REFERENCES

Byrne, B.W. and Houlsby, G.T. (2002) “Experimental investigations of the response of suction caissons to transient vertical loading”, Proc. ASCE, Journal of Geotechnical Engineering, Vol. 128, No. 11, Nov., pp 926-939.

Byrne, B.W. and Houlsby, G.T. (2003) “Foundations for offshore wind turbines”, Phil. Trans. of the Royal Society of London, Series A, Vol. 361, December, pp 2909-2930

11

Byrne, B.W. and Houlsby, G.T. (2004) “Experimental investigations of the response of suction caissons to transient combined loading”, Proc. ASCE, Journal of Geotechnical and Geoenvironmental Engineering 130, No 3, Mar., pp 240-253.

Doherty, J.P. and Deeks, A.J. (2003) “Elastic response of circular footings embedded in a non-homogeneous half-space”, Géotechnique, Vol. 53, No. 8, October, pp 703-714

Houlsby, G.T. (2003) "Modelling of Shallow Foundations for Offshore Structures", Invited Theme Lecture, Proc. Int. Conf. on Foundations, Dundee, 2-5 September, Thomas Telford, pp 11-26

Houlsby, G.T. and Byrne (2004) “Design procedures for installation of suction caissons in sand”, Proc. ICE, Geotechnical Engineering, in press

Houlsby, G.T., Kelly, R.B., Huxtable, J. and Byrne, B.W. (2004) “Field trials of suction caissons in clay for offshore wind turbine foundations”, submitted to Géotechnique,

Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2003) "Pressure Chamber Testing of Model Caisson Foundations in Sand", Proc. Int. Conf. Foundations, Dundee, 2-5 September, Thomas Telford, pp 421-431

Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2004) "Tensile Loading of Model Caisson Foundations for Structures on Sand", Proc. ISOPE, Toulon, Vol. 2, pp 638-641

Wolf, J.P. (1994) “Foundation Vibration Analysis Using Simple Physical Models”, Prentice Hall, New Jersey

Caisson Installation Test type Code Notes

Installation LB_Inst_1.5_1

Jacking test LB_Jack_1.5_1 Combined vertical and horizontal loading: unsuccessful because of control problems

1

Pull out LB_Pull_1

Installation LB_Inst_1.5_2

Jacking test LB_Jack_1.5_2 Vertical loading only

1.5m

2

Pull out LB_Pull_2

Installation LB_Inst_3.0_1

SEMV tests LB_SEMV_1 With offset weight (asymmetric cycling)

1

Jacking test LB_Jack_3.0_1 Increased amplitude cycling

Installation LB_Inst_3.0_2

SEMV tests LB_SEMV_2 Without offset weight (symmetric cycling)

3.0m

2

Jacking test LB_Jack_3.0_2 Multiple cycles at each amplitude

Table 1: Outline of caisson tests carried out at Luce Bay

12

0

10

20

30

40

50

60

70

80

90

100

0.01 0.1 1 10 100Particle Size (mm)

Perc

ent P

assi

ng

Figure 1: Grading curve for test bed at Luce Bay

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

5

0 5 10 15 20 25 30 35 40 45 50Cone Resistance (MPa)

Dep

th (m

)

Figure 2: CPT data at Luce Bay

13

Figure 3: Estimated profile of shear modulus with depth at Luce Bay

4000 4000 6000

30001500

H

H

H B

CC

A

RR

W

V

A

B

C

V

LL

L

L L LL

(a) (b) Figure 4: Outline of field testing equipment, dimensions in mm (water level and displacement reference frames not shown). (a) arrangement for jacking tests on 1.5m and 3.0m 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

0

0.5

1

1.5

2

2.5

3

3.5

4

0 20 40 60 80 100 120 140

Shear modulus G (MPa)

Dep

th z

(m) Fitted curve

Cone pressuremeterSeismic cone 1Seismic cone 2Seismic cone 3

14

Figure 5: Test rig showing the 1.5m caisson installed in foreground and 3.0m caisson in background after a jacking test had been completed

0

200

400

600

800

1000

1200

1400

0 5 10 15 20 25 30 35Suction (kPa)

Dis

plac

emen

t (m

m)

LB_Inst_1.5_2LB_Inst_3.0_21.5m prediction3.0m prediction

Figure 6: Records of suction during penetration

15

-35

-30

-25

-20

-15

-10

-5

0

5

10

15

20-40-30-20-10010

Lid pressure (kPa)

Skirt

tip

pres

sure

(kPa

)

LB_Inst_1.5_2LB_Inst_3.0_2

Figure 7: Relationship between excess pore pressure measured at caisson tip and applied suction to caisson

0

1

2

3

4

5

6

7

8

9

10

0 2 4 6 8 10

Volume pumped (m3)

Area

x v

ertic

al d

ispl

acem

ent (

m3 )

LB_Inst_3.0_11:1 ratio

Figure 8: Relationship between volume pumped from caisson and installed volume.

16

-20

-15

-10

-5

0

5

10

15

20

42 46 50 54 58 62 66 70 74 78 82Time (s)

Mom

ent (

kNm

)

Figure 9: Moment v. time for 8Hz ramp phase of SEMV test LB_SEMV_2

-30

-20

-10

0

10

20

30

-0.00005 -0.000025 0 0.000025 0.00005

Rotation (radians)

Mom

ent (

kNm

)

6Hz

7Hz

8Hz

9Hz

10Hz

Figure 10: Moment-rotation response of caisson in test LB_SEMV_2 at progressively higher amplitudes (and frequencies) of loading

17

0

200000

400000

600000

800000

1000000

1200000

0 2 4 6 8 10Frequency (Hz)

Impe

danc

e (k

Nm

/rad)

Real LB_SEMV_2Imaginary LB_SEMV_2Real average responseImaginary average responseReal, 3-param, constant GImaginary, 3-param, constant G

Figure 11: Complex M−θ transfer function for test LB_SEMV_2

-600

-500

-400

-300

-200

-100

0

100

200

300

400

500

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

Rotation of caisson centre (2R θ) (m)

Mom

ent (

kNm

)

Figure 12: Moment-rotation curve for loading of 3.0m caisson, test LB_Jack_3.0_1

18

-300

-200

-100

0

100

200

300

-0.02 -0.015 -0.01 -0.005 0 0.005 0.01 0.015 0.02 0.025 0.03

Rotation (2Rθ) (m)

Mom

ent (

kNm

)

Figure 13: Moment-rotation curve for loading of 3.0m caisson, test LB_Jack_3.0_2

0

10

20

30

40

50

60

70

80

90

100

0.000001 0.00001 0.0001 0.001 0.01 0.1

∆θ (radians)

G (M

Pa)

JackingSEMVHyperbolic curve fit

Figure 14: Computed secant shear modulus from LB_SEMV_2 and LB_Jack_3.0_1

19

-50

0

50

100

150

200

0 5 10 15 20 25

Vertical displacement (mm)

Verti

cal L

oad

(kN

)

Figure 15: Cyclic vertical loading of 1.5m caisson from LB_Jack_1.5_2

-140

-120

-100

-80

-60

-40

-20

0-100-80-60-40-200

Vertical Displacement (mm)

Verti

cal L

oad

(kN

)

Total loadSuction load

Figure 16: Load v. displacement during pullout test LB_Pull_2

20

0

10

20

30

40

50

60

70

-100-80-60-40-200Displacement (mm)

Tens

ile L

oad

(kN

)

MeasuredEstimated

Figure 17: Measured and computed friction during pullout of 1.5m caisson in test LB_Pull_2

1

A Comparison of Field and Laboratory Tests of Caisson Foundations in Sand and Clay Kelly, R.B1., Houlsby, G.T2 and Byrne, B.W.3 1Research Assistant, Oxford University 2Professor of Civil Engineering, Oxford University 3Departmental Lecturer, Oxford University Contact address: Prof. G.T. Houlsby, Department of Engineering Science Oxford University, UK OX1 3PJ Tel. 01865 273138 Fax. 01865 283301 Email. guy.houlsby@eng.ox.ac.uk

2

ABSTRACT: Laboratory tests applying vertical and moment loads to suction caissons founded in sand and clay

have been conducted to simulate an equivalent series of field tests. The caissons used in the

laboratory were 0.15m, 0.2m and 0.3m in diameter, while those for the field tests were 1.5m and

3.0m diameter. The loads applied to the caissons in the laboratory tests were scaled from those in

the field tests, and the models were loaded in a near identical manner to the field trials. The test

results are presented in non-dimensional form for comparison. The non-dimensional laboratory

moment test data were similar to the field data in most cases. The non-dimensional data from

vertically loaded caisson tests in the laboratory and in the field show some differences, and possible

reasons for these are discussed.

Introduction Suction caissons have been used to provide innovative solutions to deep-water foundation and

anchoring problems during the last 15 to 20 years. For a number of these projects, programmes of

model testing designed to simulate prototype caisson behaviour have been conducted. Laboratory

studies of small-scale model caissons at 1g include Larsen (1989), Steensen-Bach (1992) and El-

Gharbawy and Olson (1998); and centrifuge studies were carried out by Watson et al. (2000),

House and Randolph (2001), Bang et al. (2001) and Clukey et al. (2003). Published data from field

trials using larger-scale caissons are less common, but include Hogervorst (1980), Tjelta (1986),

Dyvik et al. (1993), and Cho et al. (2003). In most of these studies the data are used to calibrate

theoretical models which are then used to predict prototype behaviour. No attempt has been made to

compare directly the load-displacement response of caissons at different scales. Such comparisons

are, however, vital if model testing is to be used to predict field behaviour.

Recently, a wide ranging programme of 1g laboratory tests, and larger scale field trials investigating

suction caissons as foundations for offshore wind turbines has been conducted at Oxford University

and reported by Byrne (2000), Byrne et al. (2002), Byrne et al. (2003), Kelly et al. (2003, 2004)

and Houlsby et al. (2005a, 2005b). Suction caissons for offshore wind turbines will initially be

founded in relatively shallow water. Compared to applications in the oil and gas industry they will

have lower vertical loads, but proportionately higher moment and horizontal loads. As foundations

for offshore wind turbines, caissons may be arranged to form a tetrapod or a monopod structure.

Caissons in the tetrapod arrangement resist the overturning moment through opposing vertical

reactions, while a mono-caisson resists the applied moment directly. As in previous studies, the

laboratory tests were designed to calibrate a theoretical model of caisson behaviour, and data from

field tests were used to validate the model. However, unlike previous studies, a series of laboratory

tests have been performed to shadow the field trials, in order to directly investigate effects of scale.

3

Vertical load tests in sand have been conducted using caissons with diameters of 0.15m and 0.2m in

the laboratory to compare with a 1.5m diameter caisson in the field. Moment load tests in sand were

conducted using 0.2m and 0.3m diameter caissons in the laboratory and 3.0m in the field. Moment

tests in clay used 0.2m diameter caissons in the laboratory and 3.0m in the field. Cyclic loading

tests involving purely vertical loading were not carried out as part of the field trials in clay.

Dimensionless equations are described below to compare the laboratory and the field test data.

Dimensionless equations for comparison of laboratory and field tests Laboratory and field tests are compared on the basis that, in dimensionless terms, both the stiffness

and the strength should be similar in equivalent tests. Consider first the strength, which is assumed

to be governed by a simple bearing capacity approach. In this case, a simple dimensional analysis

suggests that in clay vertical and horizontal loads will scale by 2Rsu , and moments by 3Rsu ,

where R is the caisson radius and us a representative undrained strength. In drained sand loads

scale by 3Rγ′ and moments by 4Rγ′ , where γ′ is the unit weight. This form of scaling applies

when the densities in laboratory and field are such that the bearing capacity factors are similar for

the two cases. This will require the laboratory tests to be at a lower relative density than the field

tests, to account for the reduction of friction angle with stress level. The tests reported here do not

precisely satisfy this condition. However, the applied vertical loads were only a small fraction of the

compressive bearing capacity, and in this case the issue of density appears to be less important, as

evidenced by the successful use of the scaling approach described below.

The consideration of stiffness is a little more complex. The elastic stiffness matrix for a caisson

subjected to vertical, moment and horizontal loads V, M, and H loading may be written as:

=

uRθw

kkkk

kRG

HRM

V2

00

0022

24

43

1 (1)

where G is the shear modulus of the soil, w, 2Rθ and u are vertical, rotational and horizontal

displacements and ki are dimensionless elastic constants

Consider the moment-rotation relationship, which can be obtained from Eq. (1) as:

( ) ( )

−

−=

MRHkk

kkkGRθM

22

42

2432

3

(2)

The shear modulus in sand is known to be related to the mean confining stress. One of the principal

objections to modelling structures at small scale in the laboratory, at 1g, is that the stresses are

4

much smaller than in the field. We attempt to account for the stress level by expressing the shear

modulus as:

n

a

v

a pσ

ApG

′= (3)

where pa is atmospheric pressure (used as a reference pressure), vσ′ is a representative effective

vertical stress, A is a dimensionless constant and n is the pressure exponent. Wroth et al. (1979) and

Coop and Jovicic (1999) have reported that the value of n is not unique, but depends on the strain

level. Wroth et al. (1979) suggest that n ranges from 0.435 at very small strain to 0.765 at large

strain For Ham River silica sand, Coop and Jovicic (1999) suggest that n = 0.59 for very small

strains. Wroth and Houlsby (1985) suggest that a value of 0.5 would capture most of the important

features of the increase in stiffness with pressure, and this value is used here to compare the

laboratory and field data.

Both stress magnitude and strain amplitude will vary around the caisson, but a representative stress

may be estimated (neglecting load spreading effects) at some depth aR below the caisson as:

( )aRhγRV

v +′+π

=σ′2

(4)

where, V is the applied vertical load, h is the length of the caisson skirt and a is a dimensionless

constant.

Substituting Eq. (4) into Eq. (3) gives:

( )

++

γ′πγ′= −

22)2(42 3

)1( aRh

RVRApG nn

a (5)

Which in turn may be substituted into (2) to give:

( )

( ) ( )( )

nna a

Rh

RγV

πMRHkk

kkkAγR

p

RγM

++

′

−

−

′θ=

′

−

2224

222 342

2432

1

4 (6)

If tests at different scale are conducted keeping M/2RH, ( )32RV γ′ and h/2R constant, and taking n

to be 0.5, then the moments and rotations in these tests will be related by:

( ) ( )

γ′

′=

′ Rh

RV

RHMf

γRp

θRγ

M a2

,2

,222 31

5.0

4 (7)

5

where 1f is a dimensionless function. This suggests that satisfactory comparison of both stiffness

and strength can be achieved by plotting ( )42RγM

′ against

5.0

2

′γRpθ a .

A similar process can be conducted for vertical loading in sand, giving:

( ) ( )

′

′=

′ Rh,

Rγ

Vf

γRp

Rw

RγV ma

22222 32

5.0

3 (8)

Where mV has been introduced as a mean vertical load during the test which may be used to

estimate an appropriate shear modulus. Eq. (8) suggests that vertical loading data can be compared

by plotting ( )32RγV

′ against

5.0

22

′γRp

Rw a .

A similar analysis can be applied to moment-rotation loading in clay, except that the shear modulus

is taken to be proportional to the undrained shear strength of the clay, assuming the OCRs for the

clays are similar in the laboratory and in the field (e.g. Wroth and Houlsby, 1985). The results is:

( ) ( )

= ,OCR

Rh

RsV,

RHMθf

RsM

uu 2,

222 233 (9)

Suggesting that ( )32RsM

u may be plotted directly against θ .

The dimensionless mean vertical loads applied to the caissons must be similar for the above

procedures to be used to compare the test data. Butterfield and Gottardi (1994), and others, have

shown that the rotational and translational capacity of a shallow footing depends on the applied

mean vertical load as a proportion of the ultimate vertical capacity. This dependence is taken into

account by requiring similarity of 3)2( Rγ

Vm′

or of 2)2( RsV

u

m as well as of density or OCR.

Field trials The field trials of suction caissons in sand and clay were conducted from December 2003 to March

2004 in Scotland. The clay site was at Bothkennar, near Stirling, and a full description of the field

tests there is given by Houlsby et al. (2005a). Bothkennar was the EPSRC geotechnical test site,

and Nash et al. (1992) have comprehensively described the properties of the high plasticity silty

clay at that site. A 20m long by 10m wide by 1.75m deep test pit was dug. The base of the pit lay

6

just beneath a shelly layer described by Nash et al. (1992). The undrained shear strength of the silty

clay at this level, as inferred from Nash et al. (1992), was 11.4kPa, increasing linearly by

1.96kPa/m. The pit was flooded with water to a depth of 200mm. A 3m diameter caisson having a

skirt 1.5m long and a wall thickness of 8mm was installed at two locations within the pit using

suction. A 1m3 concrete block was placed on the lid of the caisson to provide a constant vertical

load in addition to the self-weight of the caisson. Cyclic horizontal loads were applied 4.23m above

the lid of the caisson using a hydraulic jack mounted between a separate reaction frame and a

double A-frame structure fixed to the caisson. The loads applied to the caisson were measured by

shear pin load cells attached to the hydraulic jack. Displacements were measured using draw-wire

sensors attached to a reference frame installed around the caisson. Cyclic horizontal loads, with the

load amplitude increasing by 5kN per cycle, were applied to the caisson at its first location. Similar

tests were conducted at the second location, but a packet of 10 cycles with horizontal load

amplitude of ±20kN was included within the test programme.

The sand site was at Luce Bay, near Stranraer, and a full description of the field tests there is given

by Houlsby et al. (2005b). A purpose-built sand embankment was constructed from spoil material

within a working sand and gravel quarry at Luce Bay. The spoil material consisted of 85% fine to

medium grained silica sand, 15% fine to coarse grained gravel with the occasional cobble and peat

inclusion. The embankment was constructed in a worked-out area of the quarry during the summer

of 2003, when the water table was low. The embankment was constructed in 250mm thick layers,

which were compacted using the tracks of the construction machinery. The embankment was about

40m long by 10m wide and 3.5m deep. Cone penetrometer tests and maximum and minimum

density tests were performed by Fugro Ltd to determine the properties of the sand. The maximum

and minimum dry unit weights of the sand were 17.8kN/m3 and 13.7kN/m3 respectively. The

relative density of the compacted sand was inferred from the cone penetrometer data to be 80-85%.

The submerged unit weight at a relative density of 80% was determined to be 10.3kN/m3. The field

tests were conducted in the winter, when the water table had risen above the top of the

embankment.

Two caissons were tested at Luce Bay: a 1.5m diameter caisson, which had a skirt length of 1m and

a thickness of 8mm, and the 3m diameter caisson previously used at Bothkennar. Both caissons

were installed using suction. The 1.5m caisson was loaded using a hydraulic jack attached to the

caisson and a reaction frame. Packets of 10 cycles with amplitudes increasing by ±10kN per packet

up to ±100kN were applied to the caisson about a mean load of 60kN. The loads were again

measured using the clevis pin load cells, and the displacements by draw-wire sensors fixed to an

independent reference frame. The 3.0m diameter caisson was again installed at two locations. In the

7

first test it was loaded in a similar manner to the test Bothkennar. Packets of 10 cycles were applied

in the second test at horizontal load amplitudes of ±10kN, ±20kN, ±40kN and ±60kN.

Laboratory tests The laboratory tests were conducted using test beds constructed from Redhill 110 sand and

Speswhite kaolin clay. Redhill 110 is a poorly graded fine-grained silica sand. It has maximum and

minimum dry unit weights of 16.8kN/m3 and 12.8kN/m3 respectively. Saturated test beds 350mm

deep were constructed in a container 1.1m in diameter. The samples were densified using vibration.

The kaolin was consolidated from slurry in two 450mm diameter containers to create 430mm high

test beds. It was consolidated in stages over a period of one month, and was subject to a maximum

pressure of 200kPa. The undrained shear strengths of the clay, inferred from shear vane tests, were

10.8kPa and 9.0kPa at the surface of the two containers increasing to 12.4kPa at a depth of 100mm

in both containers.

Four caissons were used in the laboratory tests, two with aspect ratios (length to diameter) of 0.66

used to model the 1.5m field caisson and two with an aspect ratio of 0.5 to model the 3.0m field

caisson. The caisson geometry, test bed properties and mean vertical loads applied to the caissons

are given in Tables 1 to 4 for the vertical load tests in sand, moment load tests in sand and in clay.

The mean vertical loads applied to the field caissons in Tables 1 to 4 include the self-weight of the

caissons.

In the laboratory the caissons were installed either by suction or by pushing in sand, and just by

pushing in the clay. The loads were applied using a three-degree-of-freedom test rig designed by

Martin (1994). The rig can control vertical, horizontal and moment loads, or displacements,

independently to follow a desired load path (Byrne, 2000). The loads were recorded using a

Cambridge-type load cell (Bransby, 1973) and vertical, horizontal and rotational displacements

were recorded using LVDTs.

In the laboratory tests the cyclic vertical loads were load-controlled. In contrast, the tests applying

overturning moments were controlled by applying fixed rotations scaled from the field trial values.

Comparison of laboratory and field test data in sand Vertical load tests

Data from laboratory tests on 0.15m diameter caissons, where one caisson was installed by pushing

and one by suction are compared with data from the vertical field test in Figures 1(a) and 1(b). In

each test packets of cycles of increasing amplitude were applied. There are two key features of the

8

behaviour during the cycling: (a) the stiffness in a given cycle and (b) the accumulation of

deformation over several cycles. It is immediately apparent from Figure 1(a) that larger

accumulated deformation occurred in the laboratory test in which suction was used for installation

than with pushed installation. This is attributed to the probable loosening of the sand adjacent to the

caisson during the suction installation. By comparison the accumulated deformations in the field

test (also suction installed, but a caisson 10 times larger), Figure 1(b), fall between the two

laboratory tests in terms of the dimensionless plot. The implication is that disturbance due to

suction installation is relatively less important for the larger caisson, which seems entirely plausible.

Suction installation may create a localized zone of disturbance adjacent to a caisson, which does not

increase in proportion to caisson diameter

Figures 2(a) and 2(b) show two laboratory tests, of slightly different sizes, both with pushed

installation. The larger test shows less accumulated deformation, supporting the hypothesis that

accumulated deformation reduces with scale. Significantly more accumulated deformation occurs in

suction-installed tests than with pushed installation. These trends are shown in Figure 3 in which

the accumulated deformations are plotted against cycle number. (Note that the cyclic load amplitude

increases every 10 cycles). In most cases the accumulated deformations were downwards, except

for the smaller caisson installed by pushing, which initially moved slightly upwards as the cyclic

loads were applied, and only moved downwards when the minimum vertical loads became tensile.

Whilst clear trends can be discerned, it is difficult without more test data to quantify the

accumulated deformations as a function of size, loading amplitude and installation method.

Differences between the tests may also be attributable to minor variations in relative density.

Turning now to stiffness, Figure 4 shows the dimensionless stiffness plotted as a function of

dimensionless cyclic rotation amplitude (on a logarithmic scale). The stiffness is in each case

defined as the secant stiffness over the unloading branch of a cycle (which differs slightly from the

secant stiffness on loading because of the influence of the accumulated deformation). This plot may

be compared to the shape of a typical plot of shear modulus against shear strain amplitude. It is

clear that the stiffness drops rapidly with cyclic amplitude, but both laboratory and field tests show

remarkably similar trends, indicating that the method for normalising the data is highly successful

in achieving comparability of stiffness in any one cycle. Such comparability is in fact apparent from

the shapes of the larger cycles that can be seen in Figures 1 and 2: in each case the shape and size of

each individual cycle is closely comparable, with the differences between tests being mainly

attributable to the accumulated deformations. This point is emphasised in Figure 5, in which

individual cycles at small, medium and large amplitudes from the different tests have been

superimposed at normalised scale.

9

Note that the data fall onto two well-defined curves in Figure 4. The left hand curve relates to the

first set of cycles in each test, and the right hand curve the remaining cycles in each test. The

apparent separation in the data is to a certain extent an artefact of the requirement that, for each load

amplitude, the points fall on a single curve.

The variation of stiffness during packets of cycles is explored in more detail in Figure 6, which

shows the unloading stiffness plotted against cycle number. Note that after each 10 cycles the load

amplitude is increased, resulting in a lower stiffness. During each packet of cycling, however, there

is no discernable trend of variation of stiffness. Figure 6 also highlights differences of stiffness

between the tests that are more apparent here than in Figure 4. It should be emphasised though that

the normalisation process involves a factor of about 33 between the stiffness in the largest and

smallest tests: the remaining differences which fall within a factor of about two should be viewed

within this context. Note that on the normalised scale the field tests at small amplitude appear rather

stiffer than the laboratory test, whilst at larger amplitude they are less stiff. This systematic

variation may be attributable to the fact that different values of the exponent n in Eq. (3) would be

applicable at different amplitudes of rotation.

Note that as cyclic amplitude increases the hysteresis loops become more open (damping increases).

Once cycles become sufficiently large that the minimum vertical load becomes tensile, a transition

to a much more flexible response occurs and the hysteresis loops acquire a characteristic “banana”

shape (see Figure 5), similar to that reported by Byrne and Houlsby (2002). For practical foundation

designs it seems prudent to avoid this more flexible behaviour, so that caisson foundations should

be designed to avoid tensile loading. Note that, paradoxically, the rate of accumulated downward

movement increases once tensile loading is reached.

Moment load tests

Data from the moment loading laboratory and field tests, where the load amplitude increased with

each cycle, are compared in Figures 7(a) and (b). Scaled rotations were applied to the laboratory

tests to allow the resulting moment loads to be compared.

The first few cycles during the laboratory test are shown in Figure 8(a). The scatter in the data

reflects noise affecting the transducer readings at the very small displacements applied to the

caisson. The shape of the normalized moment-rotation curves in the laboratory test with pushed

installation and in the field test were similar at small rotations, see Figures 8(a) and (b), but

diverged as the magnitude of the rotations increased (Figure 7). The characteristic shape of the

hysteresis loops at large rotations in the field trial was due to gapping between the caisson and the

sand (which was observed visually). Gapping was not observed in the laboratory test. This may be

10

because the absolute (as opposed to normalised) magnitude of the rotations in the laboratory tests

was much smaller than that in the field test. The process of normalisation of loads and stiffness is

therefore unable to account entirely for all phenomena observed, although it seems very satisfactory

at small rotations. The fact that the magnitudes and shapes of the curves in Figures 8(a) and (b) are

similar is a very significant result. Foundations for offshore wind turbines are likely to be designed

to operate in the range of normalized rotations shown in Figures 8(a) and (b) rather than at larger

deformations, as shown in Figure 7, where the normalized responses of the field and laboratory

caissons differ. The maximum normalized moment continued to increase with rotation amplitude in

the field trial, whereas it seemed to approach a limit in the laboratory test. The normalized moment

loads were significantly smaller in the laboratory test where the caisson was installed by suction,

again indicating a detrimental effect of sand disturbance during installation at small scale.

The normalized moment-rotation responses in laboratory and field tests in sand where packets of 10

cycles were applied to the caissons are shown in Figure 9(a) and (b). The normalised behaviour of

the caissons is again similar at small rotations, and becomes increasingly different as the

deformations increase. Although the shapes of the loops are dissimilar, in fact the loads achieved at

comparable normalised rotations were very similar.

Normalised unloading stiffness data for all of the moment loading tests are presented in Figure 10.

The unloading stiffness is again defined as the peak-to-peak normalised moment divided by the

normalised rotation during the unloading half cycle. The normalised unloading stiffnesses are

remarkably similar for all of the tests, taking into consideration that the scaling incorporates a factor

in excess of 13600, which results from incorporating a stress-dependent stiffness into the non-

dimensional equations as well as from the direct scaling factor. The difference between the

normalised moment at larger rotations in the field and laboratory test data is probably due to the

non-dimensional equations not fully taking into account changes with stress level and density.

Comparison of laboratory and field data in clay Data from the first field test at Bothkennar are compared with laboratory test data in Figures 11(a)

and (b). The normalised strength and stiffness are (as for the sand) more similar at low rotations

(Figure 12) than at larger rotations (Figure 11). Data from the initial few cycles in these tests in

Figures 11(a) and (b) shows a good correlation between the normalised behaviour to rotations up to

about 0.0015 radians, although the caisson in the laboratory test showed a slightly stiffer response.

The ultimate capacity of the caisson in the laboratory test was less than in the field test by about

30%, and the peak-to-peak stiffness was also less in the laboratory than the field at large rotations.

11

Data from the moment-rotation tests, where a packet of 10 cycles was applied to the caissons, are

presented in Figures 13(a) and (b). The normalised strength and stiffness of the field caisson was

significantly lower than the laboratory caisson at low to medium rotations, particularly in the

positive quadrant. The normalised strength of the field caisson only exceeded that of the laboratory

caisson at large displacements. The stiffness of the field caisson in this test was less than in the first

test, Figure 11. This may be due to softening of the Bothkennar silty clay over time, given that the

second test was conducted 28 days after the first test, and 44 days after the test pit was excavated.

The lower strength and stiffness of the field caisson in the positive quadrant of Figure 13(b) might

have been caused by an uneven installation. At the end of installation there was a 90mm difference

of level across the caisson. One side of the caisson might have therefore have been in better contact

with the clay than the other side.

The normalised unloading stiffness data for the clay tests are shown in Figure 14. The normalised

stiffnesses are very similar, providing a very satisfactory verification of the normalisation procedure

involved, bearing in mind that before normalisation the stiffnesses differ by a factor of more than

3900.

Concluding remarks

Normalisation procedures have been proposed to allow laboratory tests and field trials of

foundations in sand and clay to be compared in terms of both stiffness and capacity. These have

been applied to a series of laboratory tests specially conducted to mimic a series of field trials. The

tests concentrated on cyclic loading conditions relevant to the design of offshore wind turbine

foundations.

In sand the normalisation procedures take account of differing stiffness at different stress levels.

Comparisons were made for both cyclic vertical and cyclic moment loading. In both cases the

normalisation procedures were found to achieve generally satisfactory comparison between tests at

very different scales. For vertical loading the normalised stiffness within any given cycle was

highly repeatable at different scales, but the accumulation of displacement with cycles could not be

compared quantitatively. Qualitative comparisons indicated less accumulation of displacement for

larger caissons, and (at least at small scale) much more for suction as opposed to pushed installation

of caissons. Moment loading tests in sand indicated an excellent comparison at low rotation

amplitude, with more divergence at higher amplitudes.

12

Moment loading tests in clay led to similar conclusions as for sand: again the comparison of

normalised performance was much closer at small rotation than at large rotation. The former are,

however, more relevant to realistic design cases.

It should be borne in mind that the scaling relationships proposed here address stiffness and

capacity values which, in absolute terms, differ by several orders of magnitude. In that context, the

quality of agreement between laboratory and field results is highly satisfactory, even when some

differences, highlighted above, remain. These comparisons lend confidence to the use of such

scaling relationships in the design of full-scale structures. Data from loading of such structures,

currently not available, would of course be essential for confirmation of the applicability of these

procedures across a wider range of scales.

Acknowledgements

The authors are grateful to the DTI and EPSRC for the funding of this research. The authors would

also like to acknowledge the industrial participants to this research project: SLP Engineering Ltd,

Fugro Ltd, Garrad Hassan, GE Wind, NEG Micon, Shell Renewables Ltd and HR Wallingford.

References

Bang, S., Cho. Y. and Kim, S.S. (2001), “Calibration of suction pile installation design with centrifuge model tests”, 20th Int. Conf. on Offshore Mechanics and Arctic Engineering, paper OMAE/OSU-5023

Bransby, P.L. (1973), “Cambridge contact stress transducers”, CUED/C-SOILS/LN2, Cambridge University

Butterfield, R. and Gottardi, G. (1994), “A complete three-dimensional failure envelope for shallow footings on sand”, Geotechnique, Vol. 44, No.1, pp 181-184

Byrne, B.W. (2000) “Investigations of suction caissons in dense sand”, DPhil thesis, University of Oxford

Byrne, B.W. and Houlsby, G.T.(2002), “Experimental investigations of response of suction caissons to transient vertical loading”, ASCE, J. of Geotech. and Geoenv. Eng., Vol. 128, No. 11, pp 926-939

Byrne, B.W., Houlsby, G.T., Martin, C.M. and Fish, P. (2002)"Suction Caisson Foundations for Offshore Wind Turbines", Wind Engineering, Vol. 26, No. 3, pp 145-155

Byrne, B.W., Villalobos, F., Houlsby, G.T. and Martin, C.M. (2003), “Laboratory Testing of Shallow Skirted Foundations in Sand”, Proc BGA International Conference on Foundations, pp 161-167

Cho, Y., Lee, T.H., Chung, E.S. and Bang, S. (2003), “Field tests on pullout loading capacity of suction piles in clay”, 22nd Int. Conf. On Offshore Mechanics and Arctic Engineering, paper OMAE-2003-37016

Clukey, E.C., Aubeny, C.P. and Murff, J.D. (2003), “Comparison of analytical and centrifuge model tests for suction caissons subjected to combined loads”, 22nd Int. Cont. on Offshore Mechanics and Arctic Engineering, paper OMAE2003-37503

Coop, M.R. and Jovicic, V. (1999), “The influence of state on the very small strain stiffness of sands”, Proc. conf. on Pre-failure deformation characteristics of geomaterials, Torino, pp 175-181

13

El-Gharbawy, S. and Olson, R. (1998), “Laboratory modelling of suction caisson foundations”, Proc. 10th Int. Conf. on Offshore and Polar Engineering, Vol. 1, pp 537-541

Dyvik, R., Andersen, K.H., Hansen, S.B. and Christophersen, H.P. (1993), “Field tests of anchors in clay. I: Description”, ASCE Journal of Geotechnical Engineering, Vol. 119, No. 10, pp 1515-1531

Hogervorst, J.R. (1980), “Field trials with large diameter suction piles”, Offshore Technology Conference, Houston, Texas, paper 3817

Houlsby, G.T., Kelly, R.B., Huxtable, J. and Byrne, B.W. (2005a), “Field trials of suction caissons in clay for offshore wind turbine foundations”, Geotechnique, in press

Houlsby, G.T., Kelly, R.B., Huxtable, J. and Byrne, B.W. (2005b), “Field trials of suction caissons in sand for offshore wind turbine foundations”, Geotechnique, submitted

House, A.R. and Randolph, M.F. (2001), “Installation and pull-out capacity of stiffened suction caissons in cohesive sediments”, proc. 11th Int. Conf. on Offshore and Polar Engineering, Vol.2, pp 574-580

Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2003), “Pressure chamber testing of model caisson foundations in sand”, Proc BGA International Conference on Foundations, pp 421-432

Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2004) "Tensile Loading of Model Caisson Foundations for Structures on Sand", Proc. 12th Int. Conf. on Offshore and Polar Engineering, Toulon, Vol. 2, pp 638-641

Larsen, P. (1989), “Suction anchors as an anchoring system for floating offshore constructions”, Offshore Technology Conference, paper 6029

Martin, C.M. (1994), “Physical and numerical modelling of spudcan behaviour”, DPhil thesis, Oxford University

Nash, D.F.T., Powell, J.J.M. and Lloyd, I.M. (1992), “Initial investigations of the soft clay test site at Bothkennar”, Geotechnique, Vol. 42, No. 2, pp 163-181

Steensen-Bach, J.O. (1992), “Recent model tests with suction piles in clay and sand”, Offshore Technology Conference, Houston, Texas, paper 6844

Tjelta, T.I. (1986), “Large-scale penetration test at a deep water site”, Offshore Technology Conference, Houston, Texas, paper 5103

Watson, P.G., Randolph, M.F. and Bransby, M.F. (2000), “Combined lateral and vertical loading of caisson foundations”, Offshore Technology Conference, Houston, Texas, paper 12195

Wroth, C.P., Randolph, M.F., Houlsby, G.T. and Fahey, M. (1979), “A review of the engineering properties of soils with particular reference to the shear modulus”, University of Cambridge report CUED/D-SOILS TR75

Wroth, C.P. and Houlsby, G.T. (1985), “Soil mechanics – property characterisation and analysis procedures”, Proc. 11th Int. Conf. on Soil Mechanics and Foundation Engineering, San Francisco, Vol. 1, pp 1-50

14

Table 1 Caisson geometry and soil properties for vertical load tests in sand Diameter

(m) Installation

Method Unit Weight

(kN/m3) Relative Density

Mean Load(kN)

Aspect ratio (h/D)

(t/D)

1.5 Suction 10.3 0.80 66 0.66 0.0053 0.2 Pushing 10.0 0.84 0.152 0.66 0.0165 0.15 Suction 9.5 0.69 0.065 0.66 0.0067 0.15 Pushing 9.7 0.77 0.062 0.66 0.0067

Table 2 Caisson geometry and soil properties for moment load tests in sand with increasing

cyclic amplitude Diameter

(m) Installation

Method Unit Weight

(kN/m3) Relative Density

Mean Load(kN)

Aspect ratio (h/D)

(t/D)

3.0 Suction 10.3 0.80 42.4 0.5 0.0027 0.3 Pushing 10.0 0.84 0.041 0.5 0.0117 0.2 Suction 9.5 0.69 0.012 0.5 0.0050 0.2 Pushing 9.7 0.77 0.012 0.5 0.0050

Table 3 Caisson geometry and soil properties for moment load tests in sand with packets of

cycles Diameter

(m) Installation

Method Unit Weight

(kN/m3) Relative Density

Mean Load(kN)

Aspect ratio (h/D)

(t/D)

3.0 Suction 10.3 0.80 42.4 0.5 0.0027 0.3 Pushing 10.0 0.84 0.041 0.5 0.0117 0.2 Pushing 10.0 0.84 0.012 0.5 0.0050

Table 4 Caisson geometry and soil properties for all moment load tests in clay Diameter

(m) Installation

Method su

(kN/m2) Mean Load

(kN) Aspect ratio

(h/D) (t/D)

3.0 Suction 14.4 42.4 0.5 0.0027 0.2 Pushing 12.4 0.162 0.5 0.0050

15

-2

-1

0

1

2

3

4

5

6

-0.01 0.01 0.03 0.05 0.07 0.09

[w/(2R)][pa/(2Rγ')]1/2

V/[ γ'

(2R

)3 ]

Suction installationPushed installation

Figure 1(a): Vertical load: Laboratory test data using

0.15m diameter caissons

-2

-1

0

1

2

3

4

5

-0.005 0 0.005 0.01 0.015

[w/(2R)][pa/(2Rγ')]1/2

V/[ γ'

(2R

)3 ]

Figure 2(a): Laboratory test: 0.15m diameter caisson:

Pushed installation

-2

-1

0

1

2

3

4

5

6

-0.01 0.01 0.03 0.05 0.07 0.09

[w/(2R)][pa/(γ'2R)]1/2

V/[ γ'

(2R

)3 ]

Figure 1(b): Vertical load: Field test data: 1.5m

diameter caisson

-2

-1

0

1

2

3

4

5

-0.005 0 0.005 0.01 0.015

[w/(2R)][pa/(2Rγ')]1/2

V/[ γ

'(2R

)3 ]

Figure 2(b): Laboratory test: 0.2m diameter caisson:

Pushed installation

16

-0.01

0

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0.09

0 10 20 30 40 50 60 70 80 90

Number of cycles

Nor

mal

ised

cum

ulat

ive

disp

lace

men

t [w

/(2R

)][p a

/( γ'2

R)]1/

2

Field test0.15m Suction installed0.15m Pushed0.2m Pushed

Figure 3: Cumulative displacements during vertical loading as a function of cycle number

0

10000

20000

30000

40000

50000

60000

0.00001 0.0001 0.001 0.01

∆w/(2R)[pa/(γ'2R)]1/2

K v/[(

γ'pa)

1/2 (2

R)3/

2 ]

Field test0.15m Sucked0.15m Pushed0.2m Pushed

Figure 4: Normalised vertical unloading stiffness plotted against normalised displacement

17

-1

0

1

2

3

4

5

0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04 0.045 0.05

[w/(2R)][pa/(2Rγ')]1/2

V/[ γ

'(2R

)3 ]1.5m Field0.15m Suction installed0.2m Pushed0.15m Pushed

Figure 5: Superimposed small, medium and large cycles of vertical displacement from different

tests

0

1000

2000

3000

4000

5000

6000

7000

8000

0 10 20 30 40 50 60 70 80 90Number of cycles

Scal

ed u

nloa

ding

stif

fnes

sK v

/[(γ'p

a)1/

2 (2R

)3/2 ]

Field test0.15m Suction installed0.15m Pushed0.2m Pushed

Figure 6: variation of normalised stiffness with cycle number

18

-0.6

-0.4

-0.2

0.0

0.2

0.4

0.6

-0.05 -0.03 -0.01 0.01 0.03 0.05

θ[pa/(2Rγ')]1/2

M/[ γ

'(2R

)4 ]

Lab. suckedLab. pushed

Figure 7(a): Moment loading: Laboratory test data:

0.2m diameter caisson

-0.30-0.25

-0.20-0.15-0.10

-0.050.000.05

0.100.150.200.25

-0.01 -0.005 0 0.005 0.01

θ[pa/(γ'2R)]1/2

M/[ γ

'(2R

)4 ]

Figure 8(a): Small deformation moment loading:

0.2m diameter caisson

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

-0.02 -0.015 -0.01 -0.005 0 0.005 0.01

θ[pa/(γ'2R)]1/2

M/[ γ

'(2R

)4 ]

Figure 9(a): Multiple cycle moment loading:

Laboratory test data: 0.2m diameter caisson

-0.6

-0.4

-0.2

0.0

0.2

0.4

0.6

-0.05 -0.03 -0.01 0.01 0.03 0.05

θ[pa/(γ'2R)]1/2

M/[ γ

'(2R

)4 ]

Figure 7(b): Moment loading: Field test data: 3.0m

diameter caisson

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0.10

0.15

0.20

0.25

-0.01 -0.005 0 0.005 0.01

θ[pa/(γ'2R)]1/2

M/[ γ

'(2R

)4 ]

Figure 8(b): Small deformation moment loading:

3.0m diameter caisson

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

-0.02 -0.015 -0.01 -0.005 0 0.005 0.01

θ[pa/(γ'2R)]1/2

M/[ γ

'(2R

)4 ]

Figure 9(b): Multiple cycle moment loading: Field

test data: 3.0m diameter caisson

19

0

50

100

150

200

250

300

350

400

450

500

0.0001 0.001 0.01 0.1

∆θ[pa/(γ'2R)1/2]

K m/[(

γ'pa)

1/2 (2

R)7/

2 ]

3.0m sand_13.0m sand_20.2m sand_10.2m sand_20.3m sand_2

Figure 10: Normalised unloading stiffness data for moment load tests in sand

20

-0.5

-0.4

-0.3

-0.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

-0.015 -0.01 -0.005 0 0.005 0.01 0.015 0.02

θ

M/[s

u(2R

)3 ]

Figure 11(a): Moment loading: Laboratory test data in

clay: 0.2m diameter caisson

-0.3

-0.2

-0.1

0

0.1

0.2

0.3

-0.002 -0.001 0 0.001 0.002 0.003

θ

M/[s

u(2R

)3 ]

Figure 12(a): Small deformation moment loading:

Laboratory test data in clay

-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

θ

M/[s

u(2R

)3 ]

Figure 13(a): Multi-cycle moment loading:

Laboratory test data: 0.2m diameter caisson

-0.5

-0.4

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

0.4

0.5

-0.015 -0.01 -0.005 0 0.005 0.01 0.015 0.02

θ

M/[s

u(2R

)3 ]

Figure 11(b): Moment loading: Field test data: 3.0m

diameter caisson

-0.3

-0.2

-0.1

0.0

0.1

0.2

0.3

-0.002 -0.001 0 0.001 0.002 0.003

θ

M/[s

u(2R

)3 ]

Figure 12(b) Small deformation moment loading:

Field test data in clay

-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

θ

M/[s

u(2R

)3 ]

Figure 13(b): Multi-cycle moment loading: Field test

data: 3.0m diameter caisson

21

0

50

100

150

200

250

300

350

400

450

500

0.0001 0.001 0.01 0.1

∆θ

K m/[s

u(2R

)3 ]3.0m clay_13.0m clay_20.2m clay_10.2m clay_2

Figure 14: Summary of normalised unloading stiffness data for moment loading in clay