MODEL TESTING OF FOUNDATIONSFOR OFFSHORE WIND TURBINES

First year report

Model Testing of Foundations for Offshore Wind Turbines

First Year Report

Felipe Villalobos

Keble College

2003 Felipe VillalobosKeble CollegeOctober 2003

Contents

1 INTRODUCTION 3

2 PREVIOUS RESEARCH 62.1 Ultimate bearing capacity of shallow foundations . . . . . . . . . . . . 6

2.1.1 Theory of plasticity and limit equilibrium . . . . . . . . . . . . 62.2 Offshore shallow foundations under combined loading . . . . . . . . . . 7

2.2.1 Flat footings on dry sand . . . . . . . . . . . . . . . . . . . . . 72.2.2 Flat, conical and spudcan footings on saturated sand . . . . . . 82.2.3 Spudcan footings on clay . . . . . . . . . . . . . . . . . . . . . . 9

2.3 Suction Caissons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 92.3.1 Monotonic loading . . . . . . . . . . . . . . . . . . . . . . . . . 92.3.2 Vertical cyclic loading and tensile capacity . . . . . . . . . . . . 102.3.3 Rotational cyclic loading . . . . . . . . . . . . . . . . . . . . . . 10

2.4 Comments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12

3 DESCRIPTION OF THE EQUIPMENT AND SAND SAMPLESUSED 133.1 Performance of the rig . . . . . . . . . . . . . . . . . . . . . . . . . . . 133.2 Geotechnical properties of the white Leighton Buzzard sand and Dogs

Bay sand . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13

4 FIRST RESULTS 174.1 Bearing capacity tests . . . . . . . . . . . . . . . . . . . . . . . . . . . 17

4.1.1 Installation caused by self-weight penetration . . . . . . . . . . 194.1.2 Vertical load-displacement relationship . . . . . . . . . . . . . . 20

4.2 Moment loading tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . 214.2.1 Test descriptions . . . . . . . . . . . . . . . . . . . . . . . . . . 214.2.2 Determination of yield points and displacement vectors . . . . . 254.2.3 Experimental results . . . . . . . . . . . . . . . . . . . . . . . . 264.2.4 Modelling the rotational response . . . . . . . . . . . . . . . . . 29

5 FUTURE RESEARCH PROPOSALS 325.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 325.2 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . . . . . 33

5.2.1 Suction assisted installation . . . . . . . . . . . . . . . . . . . . 335.2.2 Vertical cyclic loading . . . . . . . . . . . . . . . . . . . . . . . 345.2.3 The tensile capacity . . . . . . . . . . . . . . . . . . . . . . . . 355.2.4 The ultimate moment capacity . . . . . . . . . . . . . . . . . . 365.2.5 Moment cyclic loading . . . . . . . . . . . . . . . . . . . . . . . 37

1

5.2.6 Moment loading tests on clay . . . . . . . . . . . . . . . . . . . 385.2.7 Laboratory tests and field tests . . . . . . . . . . . . . . . . . . 385.2.8 Sample preparation . . . . . . . . . . . . . . . . . . . . . . . . . 38

5.3 Timetable . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 395.4 Funding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 39Bibliography . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

2

Chapter 1

INTRODUCTION

The need for increased production of clean and sustainable energy in the near fu-ture has caused the search for alternative sources of energy to usual hydroelectric, fossilfuel or nuclear. Wind energy has proved to be one of the best options for electricitygeneration, with promising forecasts for the future (Greenpeace report, 2000). Thistype of energy is presently the fastest growing energy technology in the world. Offshoregeneration utilises the higher wind speeds offshore. Capital costs (installation, cablecosts and electrical losses) are around 30-50% higher than onshore. The extra revenueof 20% for near shore and 40% far offshore, from extra energy generation justify thedecision to go offshore (Milborrow, 2003). Offshore electricity costs are dropping anddepending on the technological developments could reduce to a third of present levels.Offshore wind power is becoming more competitive with the other power sources, as aresult.

The UK government has decided to develop a plan for implementing this energy policy.Therefore it has been decided to build wind farms offshore of the UK coast during thenext years, with the target to supply the 10% of all the electrical energy requirementsby 2010. Thus, the Crown Estates have released 13 sites around the UK (see Fig-ure 1.1(a)) and as a preliminary step there will be installed 540 turbines, but it hadbeen estimated that about 5000 turbines might be necessary to achieve the 10% target.

In this project the geotechnical engineering has a very important role to play in termsof providing the best technical, environmental and economic solution. The latter ismaybe the key for the success of the project in terms of competitiveness. In the windfarm project the cost of the foundations has been estimated as between 15% and 40%of the total installation cost (Houlsby and Byrne, 2000).

The experience gained in the construction of foundation for wind turbines is withgravity bases and piles. The first option is not suitable for the large towers since thesize and weight required becomes excessive. The height of the hub might be locatedabout 95m above seabed, see figure 1.1(b) (with a shallow water level between 5 and 20m), and the dimension of the blades gives a diameter of 96 m approximately (Byrne,2003). For the second option the restrictions are the high cost and the long time ofinstallation (both of which could increase if the weather condition become harsh). Afeasible solution might be the suction caissons. This kind of foundation has been stud-ied and used in application for the oil and gas industry in the construction of platformsand other offshore facilities. Research in this area has been conducted mainly in the

3

NGI (Norwegian Geotechnical Institute), University of Texas in Austin, COFS (Centrefor Offshore Foundation Systems), Delft University and University of Oxford. However,as figures 1.1(b) and 1.1(c) show, the loading from a wind turbine structure differs fromthat for a oil rig - for the former the moment loads are very large in comparison withthe vertical loads whilst for the latter is not the case (Byrne, 2002 and 2003). One of

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Figure 1.1 Proposed UK offshore wind farm sites. (UK Crown Estates, 2002) Figure 1.3 Three legged Jack up unit (after Byrne et al., 2002b)

Figure 1.2 Wind turbine dimensions and magnitude of loads (after Byrne et al., 2002b). Figure 1.4 Different configurations for offshore wind structures: pile, multi caisson and single caisson structure foundation (after Byrne et al., 2002b)

6MN

4MN

100m

90m

30m

200MN

25MN

Steel pile

Caissons Caisson

Turbine support structure

Seabed

Water surface

NOT TO SCALE

9 FIGURES

(a)

40

Figure 1.1 Proposed UK offshore wind farm sites. (UK Crown Estates, 2002) Figure 1.3 Three legged Jack up unit (after Byrne et al., 2002b)

Figure 1.2 Wind turbine dimensions and magnitude of loads (after Byrne et al., 2002b). Figure 1.4 Different configurations for offshore wind structures: pile, multi caisson and single caisson structure foundation (after Byrne et al., 2002b)

6MN

4MN

100m

90m

30m

200MN

25MN

Steel pile

Caissons Caisson

Turbine support structure

Seabed

Water surface

NOT TO SCALE

9 FIGURES

(b)

40

Figure 1.1 Proposed UK offshore wind farm sites. (UK Crown Estates, 2002) Figure 1.3 Three legged Jack up unit (after Byrne et al., 2002b)

Figure 1.2 Wind turbine dimensions and magnitude of loads (after Byrne et al., 2002b). Figure 1.4 Different configurations for offshore wind structures: pile, multi caisson and single caisson structure foundation (after Byrne et al., 2002b)

6MN

4MN

100m

90m

30m

200MN

25MN

Steel pile

Caissons Caisson

Turbine support structure

Seabed

Water surface

NOT TO SCALE

9 FIGURES

(c)

40

Figure 1.1 Proposed UK offshore wind farm sites. (UK Crown Estates, 2002) Figure 1.3 Three legged Jack up unit (after Byrne et al., 2002b)

Figure 1.2 Wind turbine dimensions and magnitude of loads (after Byrne et al., 2002b). Figure 1.4 Different configurations for offshore wind structures: pile, multi caisson and single caisson structure foundation (after Byrne et al., 2002b)

6MN

4MN

100m

90m

30m

200MN

25MN

Steel pile

Caissons Caisson

Turbine support structure

Seabed

Water surface

NOT TO SCALE

9 FIGURES

(d)

Figure 1.1: (a) Proposed UK offshore wind farm sites (taken from www.bwea.com); (b) Windturbine dimensions and magnitude of loads (after Byrne et al., 2002b); (c) Three leggedjack up unit (after Byrne et al., 2002b), and (d) Different configurations for offshore windstructures: monopile, multi-caissons and single caisson structure foundation (after Byrne etal., 2002b)

the first targets is to achieve the best arrangement for the suction caissons to form thefoundation system for the wind turbine. The options are monopod, tripod or tetrapod

4

- but more than four legs could also be used (see figure 1.1(d)). The structural designapproach must take into account the fact that the most unfavourable conditions areimposed by the transient tensile loads in the upwind leg for multiple caissons, whereasfor monopods the most unfavourable condition pertains to the overturning moment.In both cases the problem is not the ultimate load capacity which will be very large,but the accumulated deformations under cyclic loading (Byrne et al., 2002a).

The challenge in this industry and university research project is to determine an op-timum design procedure for suction caissons for offshore wind farms. The activitiesinvolved in the project cover site investigation, conceptual design, laboratory testing,numerical model, large scale field trial, scour investigation, final design and installationstudy (Byrne et al., 2002b). The present report introduces the laboratory investiga-tions carried out in the first year, and those that will be carried out in the future. Thetests will provide data needed for the implementation of the theoretical framework anddesign methods.

In this report following the literature review, the rig and sand used will be detailed.Subsequently the results of the bearing capacity tests for model caissons of differentlength of skirt and different sand densities will be given and analysed. Then the com-bined load tests carried out with model caissons of diameter 202 mm and 293 mm willbe also shown and examined. Finally future work proposals will be presented.

5

Chapter 2

PREVIOUS RESEARCH

2.1 Ultimate bearing capacity of shallow founda-

tions

2.1.1 Theory of plasticity and limit equilibrium

In 1920, Prandtl made the first attempt at analysing the bearing capacity of a stripmetal tool using the theory of plasticity, based on a material of zero density that pos-sessed cohesion and friction angle. This was followed by inclusion of the surcharge effectby Reissner in 1924 and then extension of these ideas to soil foundations problems byTerzaghi in 1943. By means of limit equilibrium calculations Terzaghi (1943) developeda general bearing capacity equation for strip footings which has been widely used in theengineering practice. This equation has been subject to several adjustments either tomodify the values of the original three bearing capacity factors or to add other factors(as explained in the next paragraph). As a result, many papers have been publishedin this area, including Meyerhof (1951, 1953, 1963), Brinch Hansen (1970) and Vesic(1973, 1975). These papers have extended the ideas of Terzaghi to areas such as theapplication of combined loading. Each paper gives its own semi-empirical expressionsfor the bearing capacity factors (and factors like shape, depth, inclination load and basetilt factors) and many give expressions for the failure surfaces, relying on intuition orin experimental evidences.

Prandtl and Hill analysed shear failure mechanisms to derive bearing capacity fac-tors (as a function of the angle of friction) for the case of rough and perfectly smoothcontact between the footing base and the soil. The above methods assumed generalshear failure, without analysing the local shear and punching failure mechanism asdefined by Vesic (1973). There is some agreement of the value of the surcharge bearingcapacity factor Nq given by the four authors already mentioned for the strain planecase. However there has been a great discrepancy in the values of the self weight bear-ing capacity factor Nγ. Its exponential growth with the angle of internal friction φcauses any difference to increase or decrease exponentially too. For values of φ above35◦ the difference can become considerable (more than one and a half for φ = 45◦ forexample). In the original work by Terzaghi for the case of plain strains (where φ isobtained by triaxial tests), there is an inconsistency in the framework of the superpo-sition bearing capacity equation.

6

Taking into account the common origin, the papers that are based on Terzaghi showlittle common agreement of which expressions should be used for the bearing factors.Several expressions have been developed for these factors in different countries (Sieffertand Bay-Gress, 2000), factors including size effects (Zhu et al., 2001) and differentareas of geotechnical engineering. Generally, this semi-empirical approach fails to giveinformation on the displacements - for this it is necessary to use elasticity theory orsome methods based in correlations from in situ or laboratory tests.

2.2 Offshore shallow foundations under combined

loading

The main source of research carried out in offshore foundations has been relatedwith the energy industry, for oil and gas exploitation and currently for wind basedelectricity generation. The structures in offshore facilities have to undergo a systemof loading where the horizontal and rotational components become as relevant as thevertical one. Traditionally the offshore structures have been analysed with the elas-ticity theory and with the empirical equations of Meyerhof, Hansen or Vesic, alreadyreferenced, when condition of failure are considered.

Ultimate bearing capacity models for shallow foundations are based on modelling thesoil and the geometry of the foundation. The response of the soil is forced to occurwith a previously defined mode or geometry of the failure surface which is either in-appropriate or a complicated matter to achieve as an integrated theory for a broadvariety of soils and loading states. Other way of deduction treats the soil-foundationinteraction as a unit. Nova and Montrasio (1991) named this analysis macro-element.Byrne (2000) named similarly macro model since the response of the foundations can beincluded within structural analysis. Its origin is in the paper by Roscoe and Schofield(1957) where two envelopes of dimensionless forces V/Vo versus M/Vol (being Vo thegreatest vertical load and l the length of the pad foundation) deducted from two as-sumed bearing pressure distributions under a pier pad foundation in sand are depicted.They also overlapped in a schematic plot of V versus M the envelopes of forces appliedby the steel stanchion of the structure and the resistant forces of the footing.

2.2.1 Flat footings on dry sand

Ticof (1977) took the Roscoe and Schofield idea and carried out a series of load-ing combinations tests (V , H and M) with the aim of causing the bearing capacityfailure of a shallow rough strip footing model resting on sand. From the plotting ofhis results in the V −H plane he adjusted a symmetric parabolic failure envelope anddefined elliptical plastic potentials and associated displacements. He also plotted adimensionless failure envelope in the V/Vmax − H/Vmax − e/B space (where e is theeccentricity and B the width of the footing). Butterfield and Ticof (1979) presented asuggested three dimensional ”cigar shaped” yield surface. Table 2.1 shows the valuesof the parameters derived from the load controlled tests which size such a yield surface.Further investigations have produced more experimental results over a wide range ofloading paths to failure supporting this model (Georgiadis and Butterfield, 1988; Tan,1990; Nova and Montrasio, 1991; Butterfield and Gottardi, 1994; Montrasio and Nova,1997; Gottardi et al., 1999; Cassidy et al., 2002).

7

Table 2.1: Dimensionless peak loads for three depths D (Butterfield and Ticof, 1979)Parameters at failure D = 0 D = B/2 D = BHmax/Vmax (occurs at V ≈ Vmax/2) 0.12 0.18 0.23Mmax/BVmax 0.1 0.14 0.16

Nova and Montrasio (1991) developed a non-associated strain-hardening plasticitymodel with which they modelled their load controlled three degree of freedom tests(V, M/B, H) on a strip footing on loose silica sand. Further tests were done by Mon-trasio and Nova (1997) on dense sand from which they pointed out that the shapeof the yield locus is not affected by foundation shape whilst varying linearly with theembedment foundation width ratio. Gottardi et al. (1999) using the rig designed byMartin (1994) undertook a programme of displacement controlled tests of model circu-lar footings on dense sand. They defined an expression that fit their experimental data(curves V − w) with the intention to interpret the swipe tests afterwards. The swipetests, originally named sideswipe tests by Tan (1990), were done holding constant thevertical displacement w at a certain level, whilst combining different ratios of horizon-tal displacement u to rotational displacements 2Rθ and holding constant that ratio foreach test as well, whence it was possible to map out the shape of the yield surface.

The variation of the yield surface shape has been found to be different for the case ofrigid circular footings on very loose carbonate sand, although the plasticity treatmenthas performed properly. The differences were related with the absence of a general fail-ure shape load-displacement curve and the larger magnitude of vertical displacementcompared with the dense silica sand (Byrne and Houlsby, 2001). Cassidy et al., (2002)modified Model C and applied it to the same experimental data considered by Byrneand Houlsby (2001) finding a very good agreement.

2.2.2 Flat, conical and spudcan footings on saturated sand

The study of jack-up units in deep water depth has drawn an invaluable knowledgein modelling the behaviour of spudcan footings under the application of combined load-ing. Tan (1990) investigated the response in the V −H plane of conical and spudcanfootings on saturated sand from the results of a series of tests in the drum and beamcentrifuges at Cambridge. Tan (1990) found that monotonic loading tests could bepredicted quite well with a single yield locus. For cyclic loads, however, reasonablemodelling of footing behaviour could only be obtained by introducing an inner yieldsurface. Tan proposed a non symmetrical yield locus and plastic potential finding outthat the peak of the horizontal load occurred at about V/Vo ≈ 0.4.

The effects of partially drained loading over a flat circular footing were investigatedat Oxford by Mangal (1999a and b). He tried to find out how the displacement ratemodifies the response under monotonic vertical and combined loading. He suggestedthat under vertical load the virgin penetration response is dependent on the rate ofpenetration and the density of the sand which was expressed in terms of stiffness.

8

2.2.3 Spudcan footings on clay

At the University of Oxford Martin (1994) designed an advanced three degree offreedom (V, w : H, u : M/2R, 2Rθ) rig to test spudcan footings carrying out the samesort of sideswipe test controlling the ratio of rotation to horizontal displacement forthe conditions of increasing strength with depth clayed soils, especially in the case ofembedment ratios h/2R greater than one and when the larger load is V if it is comparedwith H and M . Since Martin found that the shape of the three dimensional yieldsurface remained almost constant whilst it expanded as a consequence of the footingpenetration he established a definition of a single normalised V/Vo : H/Vo : M/RVo

yield surface with Vo being the pure vertical load capacity at any depth of embedment.Additionally, he developed a plasticity work hardening formulation (Model B) whichwas included in a structural analysis program suitable for jack-up units.

2.3 Suction Caissons

The use of skirted foundations, instead of the traditional piles and drag anchors,has been preferred by their save in money and time. The skirted foundations have beenproved in fixed and floating offshore platforms and in a wide range of other oil and gasfacilities (Andersen and Jostad, 1999).

The NGI has performed a series of field model tests (one static and three cyclic) onsoft clay (Dyvik et al., 1993). Andersen et al. (1993) compared the predictions madewith the foundation design for offshore platforms with the results of the testing of thesuction anchors in terms of pull out loads and cyclic displacements. Their predictedfailure loads agreed reasonably with the experimental results.

2.3.1 Monotonic loading

Recently at the University of Oxford as a part of the offshore foundation area ofresearch tests on very dense dry sand (higher than 95% of the relative density) have beencarried out with model caissons of diameters 100 mm and 150 mm. Aspect ratios skirtdepth diameter of 0, 0.166, 0.33 and 0.66 have been used (Byrne and Houlsby, 1999;Byrne, 2000). The tests on very dense oil saturated sand were carried by Byrne in atank with the enough space to test up in nine sites for the same sample avoiding in thatway the difficulty in interpreting the results for different soil conditions that previouslyMangal (1999) faced up to the partially drained tests. The results of load controlledmonotonic and cyclic tests for suction caissons and flat footings were compared. Firstly,for the monotonic combined loading tests and for the ratios of skirt depth diameterabove mentioned, the patterns of behaviour for flat footings also applied to the suctioncaissons. Secondly the expansion of the yield surface due to the increase in the verticalloading ratio (Vo/Vpeak) during the pre-peak bearing capacity produced a change in theshape of the yield surface. As a consequence of this last evidence a new proposal wasdeveloped in which an inner yield surface expands within a fixed outer yield surface(Byrne, 2000).

9

2.3.2 Vertical cyclic loading and tensile capacity

Relating with cyclic loading in suction caissons Byrne (2000) and Byrne and Houlsby(2002) analysed the case of oil saturated dense sand samples making a comparison ofthe pull test in a V − w plot from a load of 200 N conducted immediately after thecyclic vertical loading test on a dense sample under a mean load of 200 N. A clearmatch was found between the pull curve and the extreme points for each cycle. Theresults proved the same agreement for tests conducted before and after densification.Therefore it would be possible to deduce the vertical cyclic response studying only thetensile monotonic response. The tensile response was observed to be softer than thecompression one and the absolute maximum tensile capacity was reached at the cavita-tion limit of the pore fluid. Nevertheless, the most surprising conclusion from Byrne’scyclic tests was that there was no effect of the rate of loading on the response of 100 mmand 150 mm diameter footings which was corroborated by Stratford (2002). Johnson(1999) also found the same conclusion in similar tests of 300 mm diameter footing. Thelack of degradation of the footing behaviour might be explained by the dense samplesused by Johnson (1999) and the very dense samples used by Byrne (2000). However,the use of a loose sample (Rd < 20%) by Stratford would contradict the explanationgiven above. The sources of that contradiction might be found in the variation of thepore pressure and the frequencies of load application. Unfortunately Stratford did notshow any of those.

2.3.3 Rotational cyclic loading

Experiments of caisson foundations subjected to moment cycling at different rates (6s, 10 s and 12 s in the Byrne’s tests, 2000) verified the same absence of rate effects for thevertical cycling case alike (Stratford, 2002; Byrne, 2000). In addition to the 100 mm and150 mm diameter caisson footings Stratford (2002) included symmetric and asymmetrictests with a tripod model with 50mm diameter caissons and 50 mm of skirt depth ineach of the legs arranged on a 150 mm diameter circle. The most unfavourable case wasthe asymmetric one, i.e. when the single foot is pushed into the ground and the othertwo are in tension, as intuitively would be expected. Furthermore he conducted testswith and without moment connections (transmission or not of the moment from the legto the footings) obtaining the weakest response when no connections are present. Inthis way each caisson rotates individually instead to do it as whole structure. In generalthe results of the rotational cyclic tests by Byrne (2000) as well as by Stratford (2002)verified the kinematic hardening behaviour described under the Masing definitionsand also covered the two rules incorporated by Pyke (1979). The Masing definitionsare: i) equality in the initial tangential shear modulus with the shear modulus oneach loading reversal and ii) the unloading part of the curve is twice the reloadingone keeping the same shape. The Pyke additional rules or extended Masing rulesare: i) any reloading-unloading curve should suit to the initial backbone loading curvewhen the previous maximum shear strain is exceeded and ii) a subsequent loading orunloading curve follows the previous ones in a stress-strain relationship when they meetin their extremes. Moreover Byrne (2000) concluded that the rotational cyclic loadingbehaviour is dependent on the mean vertical load.

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Table 2.2: Experimental research on suction caissons (after Byrne, 2002)Reference Soil type Footing geometry V:M:H Monotonic/ Test type

2R x L, mm Loading Cyclic

Allersma et al., 1999 sand, clay 60 x 67 H M 150gAllersma et al., 2000 sand, clay 60 x 70 V M, C 150gBrown & Nacci, 1971 sand 254 x 44.5 V M 1g pull-outByrne & Cassidy, 2002; clay 60 x 15,30,60 V:M:H M, C 100gCassidy et al., 2003Byrne & Houlsby, 1999 sand 100 x 0,16,33,66 V:M:H M 1g dryByrne & Houlsby, 2002 sand 150 x 50 V M, C 1g

100 x 33, 66 saturatedByrne & Houlsby, 2003 sand 150 x 50 V:M:H M, C 1g

100 x 33, 66 saturatedCluckey & clay 152.5 x 305 V, V:H M, C 100gMorrison, 1995 PulloutEl-Gharbawy & clay 125 x 250 V, V:H M 1g pull-outOlsen, 1998 100 x 400,600

50 x 600El-Gharbawy & clay 125 x 250 V, V:H C 1g pull-outOlsen, 1999 100 x 400, 600

50 x 600Helfrich et al., 1976 sand 415 x 250 V M 1g pull-outHouse & clay 30 x 120 V M 120gRandolph, 2001Larsen, 1989 sand, clay 104, 204, 305 x 450 H M, C 1g pull-outRandolph et al., 1998 calc. clay 45 x 106 H, V:H M, C 120gRao et al., 1997 clay 75 x 75, 112.5, 150 V M 1g pull-outSteensen-Bach, 1992 sand, clay 48 x 80, 96, 160 V M 1g pull-out

65 x 108, 130, 21680 x 133,160, 266

Wang et al.,1977 sand, silt, 111 x 9.5, 55 V M 1g pull-outclay 140 x 13, 70

200 x 14.5, 82.6337 x 35, 162

Watson, 1999 clay, 60 x 25 V, H, M, C 100g, 150gcalc. silt V:Hcalc. sand

Whittle & clay 50.8 x 51 V M 1g pull-outGermaine, 1998

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2.4 Comments

The bearing capacity theories do not include the case of suction caisson type offoundation and, as it was pointed out, they do not offer any answer to the determi-nation of displacements. The macro model has proved to be an appropriate approachto the problem of combined loading. Now the challenge is to define a yield surfaceusing the plasticity theory for the new conditions of the suction caisson foundationsapplied to the wind turbines. The analysis of the suction caisson behaviour with thework hardening plasticity theory requires validation through the experimental evidence.

Table 2 summarizes the experimental research that has been carried out with suc-tion caissons. The majority of these works have covered ratios V/γ′D3 greater than 1,whereas in the offshore wind problem that ratio is less than 1 (Byrne, 2003). Previoustesting in experimental research considered principally high levels of V compared withH and M , when the opposite is the case in the offshore problem. Finally relativedensities of sand samples tested should be in the range encountered in the seabed soilswhich is around 35% and 75%.

12

Chapter 3

DESCRIPTION OF THEEQUIPMENT AND SANDSAMPLES USED

3.1 Performance of the rig

The loading rig was designed by Martin (1994) to test spudcan footings on clay.Afterwards the rig has been adapted to apply larger loads to test flat and caissonfootings on dense sand (Mangal, 1999 and Byrne, 2000). The development of theprogramme used to control the rig is detailed in Byrne (2000). This programme waswritten in Visual Basic, its function being to control three stepper motors througha data acquisition card connected to the computer. The stepper motors slide thehorizontal and vertical plates in addition to the rotational carriage plate. The recordingof displacements (LVDTs) and loads (Cambridge Load Cell) allows the use of feedbackon either displacements or loads. In the present report, combined load tests werecarried out under the control of V and the M/2RH ratio. In the next section detailsof these tests are given. Figure 3.1 shows the general set up of the rig. The tankused for carrying out the tests has an internal diameter of 1100 mm and a depth of400 mm. This depth of the tank was reached by joining the original tank of 250 mmdepth (used for the bearing capacity tests) to a tank with the same internal diameterand depth 150 mm. It was necessary to incorporate different kind of spacers to lift therig different distances to allow the pushing in of the caissons with skirts of 150 mmand 200 mm, and at the same time to change the direction of the loading acting planeby 90◦. This was done so as to minimise the disturbance among the four testing sites.Thus four symmetrical sites inside the tank were available to carry out tests (see figure3.2).

3.2 Geotechnical properties of the white Leighton

Buzzard sand and Dogs Bay sand

Leighton Buzzard sand has been used in the bearing capacity as well as in combinedload tests. Dogs Bay sand was used for one series of bearing capacity tests on a loosesample. Figure 3.3 depicts the particle size distribution for both sands. From theshape of the curves and the values of the coefficients of uniformity (presented in thetable 3.1), it is clear that both sands are uniformly graded. Values of the specific

13

41

1 Long LVDT vertical displacement 2 Long LVDT horizontal displacement 3 Long LVDT rotational displacement 4 Stepper motor vertical plate 5 Stepper motor horizontal plate 6 Stepper motor rotational plate 7 Shaft with VHM load cell inside 8 Base plate 9 Caisson 10 Reaction frame 11 Sample 12 Tank 13 Spacers 14 I-beams

Figure 3.1 The three degree of freedom loading rig Figure 3.2 Arrangement of the caissons inside the tank and direction of the rotation load to reduce the disturbance of the sample affecting the other test sites.

1

2

3

4

5

6

7

89

10

11

12

13

14

Direction of the rotation load

Figure 3.1: The three degree of freedom loading rig

41

1 Long LVDT vertical displacement 2 Long LVDT horizontal displacement 3 Long LVDT rotational displacement 4 Stepper motor vertical plate 5 Stepper motor horizontal plate 6 Stepper motor rotational plate 7 Shaft wit VHM load cell inside 8 Base plate 9 Caisson 10 Reaction frame 11 Sample 12 Tank 13 Spacers 14 I-beams

Figure 3.1 The three degree of freedom loading rig Figure 3.2 Arrangement of the caissons inside the tank and direction of the rotation load to reduce the disturbance of the sample affecting the other test sites.

1

2

3

4

5

6

7

89

10

11

12

13

14

Direction of the rotation load

Figure 3.2: Arrangement of the caissons inside the tank and direction of the rotation load toreduce the disturbance of the sample affecting the other test sites

14

Table 3.1: Mean grain size and coefficient of uniformitySand D50 (mm) Cu

White Leighton Buzzard (after Schnaid, 1990) 0.8 1.3Dogs Bay (after Nutt, 1993) 0.25 2.66

Table 3.2: Properties of the different sands used during the experimental workSand Mineralogy Gs γmin, KN/m3 γmax, KN/m3

White Leighton Buzzard silica 2.65 14.65 17.58(after Schnaid, 1990)Dogs Bay carbonate 2.75 9.52 13.6(after Nutt, 1993)

gravity of the grains, minimum and maximum unit weights are shown in the table 3.2for both sands. For the silica sand the individual grains are mostly quartz mineralwith subangular to subrounded shapes (Schnaid, 1990). In the case of the carbonatesand the high angularity of the grains, together with the voids presence inside thegrains leads to grain crushability due to stress concentrations (Nutt, 1993). Usingthe semi empirical formulation proposed by Bolton (1986) the variation of the peaktriaxial angle of friction with the relative density was defined assuming the values ofthe critical friction angle φ′

crit of 34.3◦ and 40.3◦ for the silica and carbonate sand,respectively (Schnaid, 1990; Nutt, 1993). Figure 3.4 shows this variation.

41

Figure 3.2 Arrangement of the caissons inside the tank and direction of the rotation load to reduce the disturbance of the sample affecting the other test sites.

0

20

40

60

80

100

0.01 0.1 1 10

Particle size (mm)

Perc

enta

ge p

assi

ng (%

)

Dogs Bay sand White Leighton Buzzard sand

Direction of the rotation load

Figure 3.3: Grading curve for the sands tested

15

42

Figure 3.3 Grading curve for the sands tested.

30

32

34

36

38

40

42

44

46

48

50

0 10 20 30 40 50 60 70 80 90 100

Relative Density (%)

φ'(d

egre

e)

Leighton Buzzard Sand Dogs Bay Sand

Figure 3.4 Evaluation of peak triaxial friction angles for Leighton Buzzard and Dogs Bay sand using Bolton’s stress dilatancy approach.

Figure 3.4: Grading curve for the sands tested

16

Chapter 4

FIRST RESULTS

4.1 Bearing capacity tests

The main purpose of the bearing capacity tests was to define the load-displacementbehaviour of a skirt model footing, for different length of skirt. Five samples wereprepared and in each one nine tests were conducted. All the 45 test results are publishedin Villalobos et al. (2002). In the present report, representative information will bereproduced. The geometry of a suction caisson is given in the figure 4.1a and a pictureof the seven caisson scale model footings tested is shown in the figure 4.1b. Figure 4.2depicts two general curves V − w for the same caisson footing, these curves belong totests with different densities and different kind of sands. The upper one corresponds toa dense sample of Leighton Buzzard sand and the lower one to a loose sample of DogsBay sand. Two common parts can be described in the loading-displacement V − wcurves. The first one (or A part) characterises the installation of the caisson whilstonly the forces over the tip and the friction over the external and internal surface wallsof the caisson are involved. The second part or B corresponds to the bearing capacitywhen, as well as the forces over the tip and over the caisson walls are been applied,the force over the internal upper circular area of the caisson become predominant. Theshape of the A part are similar for both, whilst for the B part they have significantdifferences. The jump from A part to B part occurs when the internal top plate of thecaisson touches the soil. For the case of dense sand this jump begins before the 51 mm(length of the internal skirt for this example) which reveals the dilation of the soil pluginside the caisson. Afterwards the vertical load increases until a peak is reached, in thecase of dense sand or the curve slope reduces, in the case of loose sample. Figures 4.3and 4.4 show the complete series of tests from where belong the two curves mentionedabove. The seven length of skirt L used were: 0, 12.75, 25.5, 38.25, 51, 76.5 and 102mm. The diameter 2R = 51 mm and the wall thickness t = 1.6 mm were the same inall test. The trend is clear for the two series of tests (figures 4.3 and 4.4) and it canbe seen that the shape of the curves are similar. In both figures the friction part or Apart of the curves are coincident for all the tests, whilst for the B part there is almosta linear trend if the peak points were picked up in the silica sand and the inflexionpoints in the carbonate sand. For the two densest series of tests (Rd = 74% and 88%)appeared a visible heave around the footing when the soft response started to occurdue to the dilation of the sand. The velocity of the vertical displacement was 0.05mm/s for all these series of tests.

17

43

Figure 4.1a Outline of suction caisson

Figure 4.1 Caissons used b) in the bearing capacity tests and c) in the moment load tests.

V’

2Ri

2Ro

t h

mudline

L

b)

c)

(a)

(b)

44

Figure 4.1c Caissons used in the moment loading tests.

(c)

Figure 4.1: (a) Outline of a suction caisson; caissons tested for (b) bearing capacity , and (c)moment loading

0

50

100

150

200

250

300

350

400

450

0 20 40 60 80 100 120

Vertical Displacement, w (mm)

Ver

tical

Loa

d, V

(N

)

A

B

Figure 4.2: Load displacement curves for tests FV55 and FV65. The same skirt depth forboth L = 51 mm and L/2R = 1. Rd = 88% for the silica sand and Rd = 26% for thecarbonate sand, respectively

18

4.1.1 Installation caused by self-weight penetration

The self weight penetration of the caisson into the ground can be described as inpile design practice. That is determining the friction forces in the internal and externalwalls of the caisson plus the end bearing capacity in the perimeter tip. The frictioncan be expressed by calculating the vertical effective stress next to the caisson, thenmultiplying it by the lateral pressure coefficient K to evaluate the horizontal effectivestress and finally determining the mobilised angle of friction between the caisson walland the soil. The end bearing can be defined as usual as the superposition of thesurcharge and self weight parts, where it is assumed a plain strain condition for astrip footing of width t for the caisson rim. The following expression can describe thisapproach,

V ′ = γ′h2

2(Ktanδ)o2πRo + γ′h

2

2(Ktanδ)i2πRi + γ′hNq2πRt +

12γ′Nγ2πRt2 (4.1)

friction on outside friction on inside end bearing on annulus

Where h is the penetration of the caisson due to the load V ′. Figure 4.3 and 4.4show the curves (given in equation 4.1) that fit the data. These curves are lowest inboth cases. The reason for this conservative approach is that the above equation (4.1)does not take into account the enhancement of vertical stresses close to the pile dueto frictional forces further up the caisson, whereupon it underestimates the force forfull penetration (Houlsby, 2002). The procedure developed by Houlsby (2002) allowsexpressing the vertical load in the same way as (4.1) but with the difference of includ-ing the constant stress distribution of the frictional contribution outside and inside thecaisson wall by means of the factors Zo = f(Ro, z, fo) and Zi = f(Ri, z, fi).

V ′ =Z2i,o

{exp

(h

Zi,o

)− 1− h

Zi,o

}γ′(Ktanδ)i,o2πRi,o

+ Zi

{exp

(h

Zi

)− 1

}γ′Nq2πRt +

12γ′Nγ2πRt2

(4.2)

where for the soil within the caisson,

Zi =

Ri

[1−

(1− fih

Ri

)2]

2(Ktanδ)i

(4.3)

0

100

200

300

400

500

600

700

800

900

0 20 40 60 80 100 120 140 160Vertical Displacement, w (mm)

Ver

tical

Loa

d, V

(N

)

eq. 4.2

eq. 4.1

Figure 4.3: Load-displacement curves for Rd = 88% in Leighton Buzzard sand

19

and for the soil outside the caisson,

Zo =

Ro

[(1 + foh

Ro

)2

− 1

]2(Ktanδ)o

(4.4)

Figures 4.3 and 4.4 also show the fit of the equation (4.2) to the data. From thesefigures the upper fit curves which include the friction mass of soil around outside andinside the caisson prove to be a better fit.

The average diameter of the sand grains is 0.8 mm for the Leighton Buzzard sandand 0.25 mm for the Dogs Bay sand (see table 3.1) and the thickness of the caissonwall is 1.6 mm. Thus there would be a scale effect in the calculation of the plain strainbearing capacity on the caisson rim. Nevertheless it believes that the scale effect is lessinfluent in the estimations of the friction force as the other parameters according to fitcurves obtained and plotted in figures 4.3 and 4.4.

4.1.2 Vertical load-displacement relationship

The data recorded in this series of tests will be used in modelling the vertical load-displacement relationship for suction caisson foundations. The curves obtained providethe hardening rule for plastic vertical displacement, and also the elastic stiffness onunloading. Also it is possible to get a relationship between the size of the yield surfaceand the plastic penetration (Gottardi et al., 1999). The fit curve to the data shouldinclude the friction and bearing capacity parts as was shown in figure 4.2 to 4.4. It isfelt that equations like 4.2 to 4.4 are a good choice for the first part of the curve andan exponential equation would suit properly the second one.

0

50

100

150

200

250

300

350

400

0 20 40 60 80 100 120 140 160

Vertical Displacement, w (mm)

Ver

tical

Loa

d, V

(N

)

eq. 4.2

eq. 4.1

Figure 4.4: Load-displacement curves for Rd = 26% in Dogs Bay sand

20

4.2 Moment loading tests

4.2.1 Test descriptions

The moment loading tests are intended to represent loading state on the caissonfor which the vertical load is held constant, reproducing the same condition as for theself weight of the complete turbine structure, whilst a rotation is applied. The positiveconstant values of V used were 0, 20, 50 and 100 N. In addition negative constantvalues of V were used too in a few tests, for V of -10, -20, -30, -40 and -50 N. However,the value of V undergone by the soil-caisson system during the installation stage washigher due to the friction in the internal and external caisson walls and the bearingcapacity tip. The force measured for both caissons tested was as much as 600 and 700N when the top inside the caisson touched the soil plug. When the skirt penetrationbecame close to the completely embedded condition V grew sharply as can be seenin the figure 4.5h. Figure 4.5 (h, n and o) depicts the test FV58 10 which was doneexclusively to show what happened under the application of V beyond the transitionpoint (around 600 or 700 N). Due to the load cell capacity restrain the rig allowedreaching a vertical load of 2400 N. In the previous description and analysis of the bear-ing capacity tests one can examine that subsequent part of the soil-footing response(see figures 4.2 to 4.4). From the figure 4.5h the tension capacity of the caisson canalso be determined. The value of the maximum tension load was -59 N (see figure4.5o). The next step was to reduce the magnitude of V to the defined level between0 to 100 N for the positive loads case. After that using a subroutine in the same Vi-sual Basic program implemented by Byrne (2000) the rig was controlled giving it theinstruction of how many moves will be required in the sequence. In all the cases theobject of the first move was to reach the target constant load V. For the second moveit was to hold that target load within a certain tolerance. And for the third and sub-sequent moves (in the case of cyclic loads) it was input the displacement, velocity andacceleration for each direction (w, w, w; 2Rθ, 2Rθ, 2Rθ; u, u, u). In all the tests only therotation component was running with values different from zero. The combination ofhorizontal and rotational loads, H and M (the environmental loading state) has beentreated as a constant ratio during all the series of combined loading tests. The ratiohas been expressed as M/2RH and the values adopted during the testing have been0.5, 1 and 2. Thus it covers the variation range of interest for the wind turbine problem.

The dimensions of the two aluminium model scale caisson footings tested were: i)diameter 2R = 293 mm, skirt length L = 150 mm and wall thickness t = 3.4 mm; andii) diameter 2R = 202 mm, skirt length L = 200 mm and wall thickness t = 3.4 mm.Figure 4.1a shows an outline of a suction caisson and figure 4.1c the caisson footingstested. Table 4.1 summarises the main features of the tests given in the present report.A first series of trial tests were carried out. In those the effects of the boundary condi-tions, displacements, velocities and sample conditions were investigated. The velocityof the tests (ranging between 0.005 mm/s to 0.01 mm/s) proved not to be importantin the results of load displacements curves obtained. The combined loading tests re-trieved a lot of information which was possible to record with the data acquisitionsoftware used. The plots reviewed after each test were the forces (V, M/2R,H) andthe displacements (δw, 2Rδθ, δu) versus time as well as a combination of loads versusdisplacements (figure 4.5 shows the most important plots).

21

0

100

200

300

400

500

600

700

800

0 500 1000 1500 2000Time (sec)

Ver

tical

Loa

d, V

(N

)

(a)

-100

-80

-60

-40

-20

0

20

40

60

80

100

0 500 1000 1500 2000

Hor

izon

tal L

oad,

H (

N)

(b)

-60

-40

-20

0

20

40

60

0 500 1000 1500 2000

Mom

ent L

oad,

M/2

R (

N)

(c)

-150

-100

-50

0

50

100

150

0 500 1000 1500 2000

Ver

tical

dis

plac

emen

t, δw

(m

m)

(d)

-3

-2.5

-2

-1.5

-1

-0.5

0

0.5

1

1.5

0 500 1000 1500 2000

Hor

izon

tal d

ispl

acem

ent,

δu (

mm

)

(e)

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

0 500 1000 1500 2000

Time (sec)

Rot

atio

nal d

ispl

acem

ent,

2Rδθ

(m

m)

(f)

M/2R = 0.5655H + 0.0096R2 = 0.9994

-60

-40

-20

0

20

40

60

-100 -80 -60 -40 -20 0 20 40 60 80 100Horizontal Load, H (N)

Mom

ent L

oad,

M/2

R (

N)

(g)

-100

400

900

1400

1900

2400

0 20 40 60 80 100 120 140Vertical displacement, δ w (mm)

Ver

tical

load

, V (

N)

(h)

22

-50

-40

-30

-20

-10

0

10

20

30

40

50

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2Rotational Displacement, 2R δθ (mm)

Mom

ent L

oad,

M/2

R (

N)

(i)

-100

-80

-60

-40

-20

0

20

40

60

80

100

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2Horizontal Displacement, δ u (mm)

Hor

izon

tal L

oad,

H (

N)

(j)

δ u = 1.0145(2R δθ ) + 0.0005

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

-2 -1.25 -0.5 0.25 1 1.75Rotational Displacement, 2R δθ (mm)

Hor

izon

tal D

ispl

acem

ent,

δu (

mm

)

(k)

δw = -0.3949δu + 114.68

112

113

114

115

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2 2.5Horizontal displacement, δ u (mm)

Ver

tical

dis

plac

emen

t, δ

w (

mm

)

(l)

δ w = -0.3739(2R δθ ) + 114.67

112

113

114

115

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2Rotational displacement, 2R δθ (mm)

Ver

tical

dis

plac

emet

, δw

(m

m)

(m)

650

700

750

800

850

900

950

1000

135 137 139 141 143 145Vertical displacement, δ w (mm)

Ver

tical

load

, V (

N)

(n)

-80

-60

-40

-20

0

20

136 138 140 142 144 146 148Vertical displacement, δ w (mm)

Ver

tical

load

, V (

N)

(o)

Figure 4.5: Examples of recordings from tests carried out. From (a) to (g) and (i) to (k)correspond to test FV47 7 1; (h) to test FV58 10; and (l) to (m) to test FV45 7 1; (h) isa expanded view of figure (h) where is observed a clear change of the curve slope when theinternal top of the caisson touches and then compresses the plug soil, and (o) shows thecaisson pull-out

23

Table 4.1: Summary of the moment loading tests

38

Table 4.1 Summary of the key moment loading tests

Test Test Name Date V M/2R H Ratio γγγγd Rd L L/2R Observations (N) (N) (N) M/2RH (KN/m3) (%) (mm)

1 FV1_1 19/12/2002 50 39 38 1 15.16 20 150 0.5 10 FV15_2 21/01/2003 0 26 12 2 15.05 16 150 0.5 Velocity

26 FV26_3_1 10/02/2003 0 18 35 0.5 15.37 28 150 0.5 0.005mm/s

30 FV27_3_1 11/02/2003 0 23 24 1 15.37 28 150 0.5 " "

34 FV29_3_1 11/02/2003 100 55.8 28 2 15.48 32 150 0.5

38 FV30_4_1 13/02/2003 100 41 80 0.5 15.24 23 150 0.5 42 FV31_4_1 13/02/2003 100 51 51 1 15.24 23 150 0.5 46 FV32_4_1 18/02/2003 100 54.5 27 2 15.24 23 150 0.5 50 FV33_4_1 18/02/2003 100 37 74 0.5 15.24 23 150 0.5 54 FV34_5_1 21/02/2003 50 30.5 60 0.5 15.18 21 150 0.5 58 FV35_5_1 24/02/2003 20 24.5 47 0.5 15.18 21 150 0.5 62 FV36_5_1 24/02/2003 20 35.7 18 2 15.18 21 150 0.5 66 FV37_5_1 25/02/2003 50 45 23 2 15.18 21 150 0.5 79 FV45_7_1 17/03/2003 0 19.5 39 0.5 15.23 23 200 1 cyclic and

83 FV46_7_1 18/03/2003 20 24.5 47 0.5 15.23 23 200 1 Velocity

87 FV47_7_1 19/03/2003 50 27.5 53 0.5 15.23 23 200 1 0.01mm/s

91 FV48_7_1 20/03/2003 100 33 60 0.5 15.23 23 200 1 " "

95 FV49_8_1 25/03/2003 0 31 30 1 15.21 22 200 1 99 FV50_8_1 26/03/2003 20 35 34 1 15.21 22 200 1

103 FV51_8_1 27/03/2003 50 40 39 1 15.21 22 200 1 107 FV52_8_1 27/03/2003 100 44.5 44 1 15.21 22 200 1 111 FV53_9_1 03/04/2003 0 36 18 2 15.46 31 200 1

115 FV54_9_1 03/04/2003 20 42 21 2 15.46 31 200 1 119 FV55_9_1 04/04/2003 50 42 20 2 15.46 31 200 1

123 FV56_9_1 04/04/2003 100 52 26 2 15.46 31 200 1 133 FV58_10 23/04/2003 2400 1 15.3 31 150 0.5 Vertical load

134 FV60_10 23/04/2003 -40 12 11 1 15.3 31 150 0.5

136 FV61_10 28/04/2003 -30 17.8 18 1 15.3 31 150 0.5

137 FV62_11 30/04/2003 -20 26.3 25 1 15.54 34 150 0.5

138 FV63_11 08/05/2003 -10 26.8 26 1 15.54 34 150 0.5

139 FV64_11 08/05/2003 -20 24.9 25 1 15.77 43 150 0.5

140 FV65_11 09/05/2003 -50 14.3 13 1 15.77 43 150 0.5

24

4.2.2 Determination of yield points and displacement vectors

Yielding is defined in solid mechanics as the state when irreversible plastic strainingcommences even if there is no linearity. However, soil is not a continuum but a partic-ulate material. Conceptually in soil mechanics is understood yielding as the transitionfrom a state with defined behaviour characteristics to another one different in terms ofvariation of stresses and strains. The procedures to be used to define yielding shouldbe rational and repeatable, minimizing personal influences. There are some proceduresto define the yield point. For example it could be possible to define the start of yieldingas the point when the experimental curve leaves the initial almost linear part as it iswell established for some clay and some sands. It can also be predefined some valueof displacement (0.3 mm of rotational displacement for example) for which the corre-sponding value of load is identified as the yield point. Other methods could be usedto determine the initiation of yielding in a similar manner as the empirical proceduresuggested by Casagrande for obtaining the preload of a sample of overconsolidatedclay (Poorooshasb et al., 1967). The yield point was initially defined as the point ofminimum radius of curvature following in some way the above criteria, but the resultsobtained were not so consistent as expected. Finally the yield point was determined asthe intersection of two straight lines which fit the experimental curve at the beginningand at the end of it (Graham et al., 1982). This method was used due to its simplicityand the consistency of the results.

It can be seen in the figure 4.6, corresponding to a general test, that there was in-cluded the case for the initial yield point according to the criteria of the initial linearpart of the curve. That would give a value of 18 N as yield point. If the criteria of apredefined displacement as 0.3 mm of rotational displacement is chosen the yield pointwould be 29 N and if 0.25 mm is chosen the yield point would be 28 N and so on (de-pending of the choice of the displacement), only the first case was depicted in the plot.With the criteria similar to the determination of the overconsolidated load the yieldpoint would be equal to 23 N, which is close to the value obtained by the maximumcurvature criteria, that is 24 N. And the procedure chosen gives as a result the valueof 31 N. Table 4.1 and figures 4.7 to 4.10 show the values so determined for the testsanalysed. Figure 4.5 (k, l and m) shows the evolution of the displacements during therotation displacements applied over the caisson. These data have been recorded andprocessed for all the tests carried out. Figure 4.5 (l and m) depicts the test FV45 7 1which is a two cycles test since with a constant vertical load V = 0 N. The caissondiameter was 202 mm and M/2RH = 0.5. The equations that appear in the plots arethe best linear fit to the experimental curve and represent: i) variation of the horizontaldisplacement u as function of the rotational displacement 2Rθ (figure 4.5k). ii) initialvariation of the vertical displacement w as a function of the horizontal displacement u(figure 4.5l) and iii) initial variation of the vertical displacement w as function of the ro-tational displacement 2Rθ (figure 4.5m). The linearity of these relationships might beexplained bearing in mind the linearity among the three different forces (V, M/2R,H).Figure 4.5g shows the change of M/2R as function of H. Although that plot belongsto another test (FV47 7 1) that relationship was always matched for all the tests andM/2RH ratios (0.5, 1 and 2). For the reason that the vertical load was held constantduring all the moment tests there was also a linear relationship among V and H andM/2R.

25

From the slopes of the straight lines in the figure 4.5(k, l and m) the displacementvectors were determined. Figures 4.7 to 4.10 depict the displacement vectors for bothcaissons tested. From the displacement vectors experimentally determined is possibleto define a plastic potential. This will be developed in the theoretical section of theproject.

4.2.3 Experimental results

In the first series of tests the caisson of diameter 293 mm was tested under amonotonic rotation in only one direction to obtain the load displacement response. Inthe second series of tests the caisson of diameter 202 mm was used and tests with twocycles were implemented. Wind, currents and waves loads are applied over the turbinestructure dynamically. Thus cyclic tests provide a better understanding of the responsecaused by the loading conditions found offshore. In general M/2R − 2Rθ and H − ucurves showed an initial approximately linear part with a steep slope (see initial partof curves in figure 4.5i and j). When the slope of the curve starts to change gently toa softer response, a final approximately linear curve is reached, as before. The samecurve shape was found in all the rotational plots. This pattern suited very well with thedetermination of the yield point already explained and with the fit of a mathematicalexpression to the data. It is proposed that a hyperbolic expression would fit well tothe shape of the experimental curves.

A useful way to interpret the results from the moment loading tests is on plots ofinteraction diagrams. Hence the yield points are plotted in M/2R−H and M/2R−Vdiagrams. Figures 4.7 to 4.10 depict those points which are in the table 4.1 as well.Despite differences in density (Rd between 16% and 43% with an average value of 27%)for each sample tested the results are consistent and it is clear to see a trend in termsof the constant load V and from the M/2RH ratio. The response increases with largerV , as well as with larger M/2RH ratio. For the case of low values of V , less than20 N and even negative values, the soil-footing system still resist which reveals thefrictional capacity developed. The lowest value of the negative constant vertical loadapplied was V = -50 N and the tension capacity was -65 N in that case. The changein -6 N from the -59 N pointed out previously is due to a slight increase of the density(compare values of relative density in table 4.1 for tests FV58 10 and FV45 7 1). Thefact that the soil-footing system has moment resistance even under such unfavourableconditions should be incorporated in the modelling of the suction caissons. In figures4.8 and 4.10 appear one point in the V - axis. They represent the weight of the soilplug inside the caisson: 160 N and 95 N, respectively. These values might be used asa parameter of the proportion of the tension capacity.

Figure 4.11 shows the effect of the increase of V from 0 to 100 N in the two cyclerotational tests for fresh samples and figure 4.12 shows the results with different valuesof V obtained in the same site of a sample. In both figures a consistent location ofthe curves can be seen, that curves with higher V pass above (or below dependent onthe part of the cycle) those with lower values of V . That comparison works very wellbecause the tests have been done in the four sites of a fresh sample that has the samedensity. For the tests carried out in the same site the variation of the density aftereach test it seems to be not very significant for the case of the loose samples tested.However those results are not taken into account for the plotting of the interaction

26

0

10

20

30

40

50

-1 -0.5 0 0.5 1Rotational displacement, 2R δθ (mm)

Mom

ent l

oad,

M/2

R (

N)

0.3 mm

Figure 4.6: Five criteria for the determination of the yield point

0

20

40

60

0 20 40 60 80Horizontal load, H (N) - δ u

Mom

ent l

oad,

M/2

R (

N)

- 2R

δθ

V = 0 N V = 20 N V = 50 NV = 100 N V = -20 N V = -30 NV = -40 N V = -10 N V = -50 N

Figure 4.7: Moment load versus horizontal load and displacement vectors for the 293 mmdiameter caisson

27

0

20

40

60

80

100

-160 -140 -120 -100 -80 -60 -40 -20 0 20 40 60 80 100Vertical load, V (N) - δ w

Mom

ent l

oad,

M/2

R (

N)

- 2

Rδθ

M/2RH = 0.5 M/2RH = 1 M/2RH = 2

Figure 4.8: Moment load versus vertical load and displacement vectors for the 293 mmdiameter caisson

0

10

20

30

40

50

60

70

0 10 20 30 40 50 60 70

Horizontal load, H (N) - δ u

Mom

ent l

oad,

M/2

R (

N) -

2Rδθ

V = 0 N V = 20 N V = 50 N V = 100 N

Figure 4.9: Moment load versus horizontal load and displacement vectors for the 202 mmdiameter caisson

28

0

20

40

60

80

100

-100 -80 -60 -40 -20 0 20 40 60 80 100Vertical load, V (N) - δw

Mom

ent l

oad,

M/2R

(N) -

2Rδθ

M/2RH = 0.5 M/2RH = 1 M/2RH = 2

Figure 4.10: Moment load versus vertical load and displacement vectors for the 202 mmdiameter caisson

diagrams. Only test on fresh samples have been considered. In figure 4.13 one can ob-serve the agreement with the second Masing’s rule. To prove the first rule is necessaryto reload during a loading reversal. As a part of the future research proposal that willbe included.

-60

-40

-20

0

20

40

60

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2Rotational displacement, 2R δθ (mm)

Mom

ent l

oad,

M/2R

(N)

FV48_7_2; V = 0N FV48_7_3; V = 20N FV48_7_4; V = 50N FV48_7_1; V = 100N

Figure 4.11: Comparison among tests of similar densities under increasing constant verticalload on fresh samples

4.2.4 Modelling the rotational response

The data retrieved in the tests carried out will be used in modelling the rotationalbehaviour of suction caisson foundations. The elastic behaviour, yield surface andplastic potential can be defined from the results obtained in the moment tests forlow vertical loads. Work hardening plasticity theory has been applied successfully in

29

-60

-40

-20

0

20

40

60

-2 -1.5 -1 -0.5 0 0.5 1 1.5 2Rotational displacement, 2R δθ (mm)

Mom

ent l

oad,

M/2R

(N)

FV45_7_1; V = 0N FV46_7_1; V = 20N FV47_7_1; V = 50N FV48_7_1; V = 100N

Figure 4.12: Comparison among tests carried out in the same site under different constantvertical load

0

20

40

60

80

-0.5 0.5 1.5 2.5 3.5Rotational displacement, 2R δθ

Mom

ent l

oad,

M/2R

initial loading (doubled)

first reversal relocated

reloading relocated

second reversal relocated

Figure 4.13: Prove of the second Masing’s rule. The unloading or reversal is twice thereloading (initial loading). For the second cycle there is agreement too, but without thedoubled of the reloading

30

the simulation of combined loading for the case of flat and spudcan footings and forclay and dense silica sand and loose carbonate sand under high vertical loads (Martin,1994; Houlsby and Cassidy, 2002; Cassidy et al., 2002). Following these research thetheoretical part of the project will develop a new framework for the wind turbinesproblem. The aim of the new model will be integrate the different loading conditionsfrom which the moment loading will allow modelling the rotational response of suctioncaisson foundations.

31

Chapter 5

FUTURE RESEARCHPROPOSALS

5.1 Introduction

The literature review presented at the beginning has drawn the conclusion that theprevious research on suction caisson foundations that has been focused mainly to theoil and gas industry applications. On the other hand, for the wind turbine problemthere is little data to show where it is necessary to expand the investigation to setthese new problem conditions. Therefore the following testing program will point inthat direction to acquire the appropriate data to validate theoretical models. Two plas-ticity models have been developed at Oxford to analyse spudcan footings: Model B andModel C. The first one is for modelling the response in clay and was calibrated againsttests carried out by Martin (1994). The second one is for modelling the response insand. Model C was calibrated with the data provided by Gottardi et al. (1999) andhas been validated by Houlsby and Cassidy (2002). It has been also validated withsome modifications for loose carbonate sand by Cassidy et al. (2002). Both modelsare limited for the case of spudcan and flat footings. Besides they can resolve correctlythe monotonic loading case, but it is not the same for elastic and cyclic, where theirperformance is limited.

As part of the foundation for offshore wind farms project the theoretical investiga-tion will improve these problems. The suction caisson foundations are going to beanalysed taking into account their behaviour mainly under working loads rather thanonly under ultimate or failure loads because the design is conditioned by the accumu-lation of plastic deformations during the operation life of the wind turbines. Thus theunderstanding of the dynamic response must be investigated. The new model will bedeveloped based on the hyperplasticity theory. It will need the input of experimentaldata which cover the behaviour of the foundation under the monotonic and dynamicloading condition. It is assumed that normally the suction caisson foundations willundergo loads under yielding. Consequently, testing will be useful in modelling theresponse inside the yield surface for monotonic as well as for cyclic loading.

The diagram in figure 5.1 shows the interaction with the research project parts. Twokinds of foundation structure arrangements will be considered: i) tripod or tetrapodand ii) monopod (see figure 1.4). For the first two the tensile response controls thedesign of the foundation. For the third case, the overturning moment becomes the de-

32

52

Figure 5.1 - Interaction amid the parts that constitute the research project.

0

100

200

300

400

500

600

700

0 20 40 60 80 100 120 140

Vertical Displacement, w (mm)

Ver

tica

l Loa

d, V

(N

)

Figure 5.2 Three possible load – displacement relationships under different suction pressures before and after the top – soil plug contact. A self weight load of V = 40N is assumed for the caisson of diameter 293mm.

Real Foundation

Theoretical Model

Field Tests

Chamber Tests

Laboratory Tests

Fatigue Tests

Installation without suction Increasing

constant suction

Figure 5.1: Interaction among the parts that constitute the research project

sign key. Dissimilar loading regimes will be applied to the model caissons where cyclicloading is going to be an essential one. The effect of the suction pressure applied duringthe installation will be assessed. Combinations of different geometries of the caissonfootings are required to find out aspect ratio and size effects. It is also necessary torecreate the type of foundation soil in the offshore seabed. Tests will be carried outfor a range of relative density for sandy soil and for a range of different strength forclay soil determined previously in the field by means of site investigation and geologicalsurvey. Finally laboratory tests will be considered shadowing the field test.

5.2 Experimental Program

5.2.1 Suction assisted installation

As a construction condition, the foundations of wind turbines will be set up onthe seabed under the self weight initially and assisted by suction pressure to reachthe desired level of final penetration. The tests that have been already completedrelated with the installation have included the situation of self weight penetrationon dry samples (see previous section). It is certain that the suction component willhave a key role in the installation works. Thus experiments will be conducted insaturated samples using both model footings of diameter 202 mm and 293 mm. Fromthe bearing capacity results and from the moment test were obtained the results ofthe vertical load-displacement response in the installation or ’frictional’ region. In thefield it is expected to install the caissons until a certain level (much lower than thehalf of the caisson skirt length even for loose sand or soft clay). At that level suctionis applied to push in fully the skirt length to the required length, depending on theproject design. This situation needs to be explored letting the model caisson footingsink under its own weight, say 20 N or 50 N and measuring the respective verticalpenetration. Afterwards, the penetration of the caisson will be due to the suctionpressure applied. The level of suction will be controlled to see and evaluate changes

33

in the gradient when the top inside the caisson touches the ground. Subsequently, thesuction pressure will keep applying to determine at least the initial steep slope of thecurve. Figure 5.2 depicts shapes of the curves envisaged. Attention will be paid to the

52

Figure 5.1 - Interaction amid the parts that constitute the research project.

0

100

200

300

400

500

600

700

0 20 40 60 80 100 120 140

Vertical Displacement, w (mm)

Ver

tica

l Loa

d, V

(N

)

Figure 5.2 Three possible load – displacement relationships under different suction pressures before and after the top – soil plug contact. A self weight load of V = 40N is assumed for the caisson of diameter 293mm.

Real Foundation

Theoretical Model

Field Tests

Chamber Tests

Laboratory Tests

Fatigue Tests

Installation without suction Increasing

constant suction

Figure 5.2: Three possible load - displacement relationships under different suction pressuresbefore and after the top - soil plug contact. A self weight load of V = 40 N is assumed forthe caisson of diameter 293 mm

fluid flow inside the sand around the caisson. The suction applied should not increase toa level which could produce piping failure. The following information will be collected:rate of installation, the suction pressure and the flow rate. Sanham (2003) pointed outthat the self weight has a negligible effect in the rate of installation once suction isapplied. On the other hand he found increasing (close to linear) relationships betweenthe installation rate versus pressure applied and the installation rate versus flow rate.Two aspect ratios will be used: 0.5 and 1 for the model caissons of diameter 202 mmand 293 mm. In addition, three levels of suction pressure are going to be applied. Inaddition to two different initial vertical load or self weight will be considered plus twodifferent densities. The total number of tests to be conducted is 24.

5.2.2 Vertical cyclic loading

The transition between compression to tension (and vice versa) will be investigatedthrough vertical cyclic testing. The change of stiffness in each cycle is one of the pa-rameters to be assessed. Representative vertical loads and soil characteristics will beused to adjust the wind turbine conditions. Byrne (2000) and Byrne and Houlsby(2002) analysed such tests for the case of very dense sand saturated by oil. Frequenciesbetween 0.1 Hz and 1 Hz will be considered. For this range of frequencies consolidationof sand samples is expected to occur in the first five cycles if loads applied during thesefirst cycles represent extreme events. Extreme events are understood as representingof storms that the structure will undergo.

The use of extreme events during cycling has proved that the lower amplitude cy-cles do not cause too much degradation compared to the extreme events. Byrne (2000)

34

found almost the same response when were applied different loading time histories dur-ing the testing at essentially the same period (10 s). The load time histories were: i)constrained new wave ii) modulated sine wave and iii) stepped sine wave. From the2000 cycles test for a very dense sand sample it can be concluded that no more than100 cycles were enough to analyse the response of the caisson footing since the restfollowed the same trend (Byrne, 2000). Another conclusion can be drawn from theresults achieved by Byrne (2000) and Byrne and Houlsby (2002); the number of cyclicloading tests can be reduced, using the results from the monotonic tests. Plots of ex-treme points belonging to the vertical cyclic loading compared with plots of pull testsmatched almost perfectly even in the case of before and after densifying the sample.

Cyclic loading tests result in a large amount of data. One way to reduce it is de-creasing the logging rate. In the case of long tests with a lot of cycles (more than 20or 50) the reduction of experimental data can be done normalising the extreme valuesof loads for each cycle.

The first tests will be carried out applying small load amplitudes following increas-ing ones until displacements start to increase significantly. Therefore a sequence offour cyclic loads will be applied starting from 100 N ± 10 N, 100 N ± 20 N, and so onuntil to find the amplitude from which the displacements become significant. For threedifferent ranges of cyclic loads, two aspect ratios (L/2R = 1 and 0.5 with 2R = 202mm and 293 mm), two frequencies (1 Hz and 0.1 Hz) and, dry and saturated samplesthe total amount of tests will be 48, if loose and medium dense samples are included.

Vertical fatigue tests will be carried out in parallel as part of the research projectand they will give information about the caisson response under long term loadingconditions. Tests with more than one hundred thousand long-term cycling loading willbe undertaken on a footing under certain levels of vertical load. Three drained testswill be conducted to replicate the long term tests applying no more than 20 cycles. Dueto the rig load rate constrains, the three forecaster tests will consider load amplitudein the rage of -10 N to 230 N (with the average around 110 N which scales the deadload of the wind turbine) and load rates around 1 second. Results of these forecastertests will be taken into account in the planning of the long term tests.

Furthermore tests in a pressurised chamber will be carried out vertical cyclic teststo replicate the water pressure (Kelly et al., 2003). Water pressure up to 200 kPa willbe applied and the caisson footing used in the testing will have diameter and skirtlength of 280 mm and 180 mm, respectively. The caisson footing response will beexplored with the use of an actuator that can apply larger and faster vertical loadson water saturated sand. With the actuator a range of loads between 10 and 100kN can be applied and a more wide range of frequencies can be used (0.25 Hz to 10Hz). The change of the stiffness from compression to tension (or the contrary) will bespecially analysed. The effect of cavitation of the pore fluid will be also investigated.Comparison of results will be only possible for the low frequencies and dense samples.

5.2.3 The tensile capacity

The response of the tripod/tetrapod arrangement is limited by the upwind founda-tion in tension (Byrne et al., 2002b; Stratford, 2002). There is evidence that the tensile

35

capacity of the suction caisson foundations is non-zero for the case of very dense sand(Byrne and Houlsby, 2002). In undrained conditions cavitation of the pore fluid hasbeen the limit for the high tensile capacity achieved in Byrne’s (2000) pull out tests.From the present report can be seen that the same first conclusion for the loose sandsamples used and for the drained tests carried out.

Nevertheless, it is necessary to prove other densities and loading conditions. A se-ries of monotonic tension loading tests aimed to consider this condition will be carriedout at rates, varying from drained to partially drained response. Although previousresearch has not found a significant influence of the loading rate in the response ofthe caisson footings (Mangal, 1999; Johnson, 1999; and Byrne, 2000), the variation ofthe loading rate and pore pressures will be important to study under the wind turbineproblem conditions. Then tensile capacity tests will be considered under saturatedconditions. They will be carried out after some of vertical cyclic tests have been done.Thus comparison of tensile responses will be considered for three different cyclic load-ing amplitude histories and three different rates (1 s, 6 s and 10 s). The number oftension tests - including loose and medium dense samples will be 18.

The pressurised chamber tests which replicate the water pressure will explore the sametensile response with the use of an actuator that can apply larger and faster verticalloads on water saturated sand. Comparison will be made as explained before.

5.2.4 The ultimate moment capacity

The response of the monopod structure is conditioned by the overturning momenton the foundation. Rotational displacement tests under vertical constant load havebeen already completed. The results have been presented in the previous section. Inaddition swipe tests will complete the investigation of ultimate monotonic momentcapacity. Vertical constant load and swipe tests allow to define the yield surface andthe direction of the displacement vectors. Low vertical load will be considered.

In a swipe test footings are pushed vertically into the ground until a certain levelof load is reached. At this point the vertical displacement is kept constant. Subse-quently the footing is displaced laterally by means of either horizontal movement orrotational movement. The load path traced in that way will allow to define the yieldsurface as has been done amongst others by Tan (1990), Martin (1994) and Gottardiet al. (1999). Those research considered mostly the ’normally consolidated’ case whichinvolves relatively high levels of vertical loads during testing. To represent the case ofthe wind turbine problem the first step of the swipe test description above that is, theapplication of the vertical load, must consider the ’overconsolidated’ case. Once thecertain level of load is reached for the fully embedded caisson will follow a verticallyunloading until a certain low load. Special interest will be given to the negative startingvalues of the vertical load with the idea to discover if the yield surface obtained wouldmatch with plots obtained in the moment loading tests with constant vertical load (seefigures 4.8 and 4.10). The vertical displacement is held, whilst the test is conducted.Figure 5.3 depicts how curves of swipe tests might become. In order to compare theprobable curves shown in the figure 5.3 with the plots of the figures 4.8 and 4.10 acontrol of the load ratio M/2RH will be conducted without holding V constant duringthe progress of the test, w will be kept constant instead. Dry swipe tests will consider

36

54

Figure 5.3 – Plausible load path obtained doing a swipe test starting with negative values of vertical load and holding the vertical displacement for the caisson diameter 293mm.

0

20

40

60

80

100

-160 -140 -120 -100 -80 -60 -40 -20 0 20 40 60 80 100Vertical load, V (N) - δ w

Mom

ent l

oad,

M/2

R (

N)

- 2

Rδθ

M/2RH = 0.5 M/2RH = 1 M/2RH = 2

Figure 5.3: Plausible load path obtained doing a swipe test starting with negative values ofvertical load and holding the vertical displacement for the caisson diameter 293 mm

three initial vertical loads of -20 N, -40 N and -60 N plus two aspect ratio L/2R. Inaddition to the three ratios of M/2RH (= 0.5, 1 and 2) which will be optional sinceit is believed that with the ratio of 1 would be adequate. Loose and medium densesamples will be used. Finally, 12 swipe tests will be carried out.

5.2.5 Moment cyclic loading

The continuous hyperplasticity cyclic loading model needs to be validated with ex-perimental results. Consequently tests will be carried out to assess the response ofthe suction caisson foundations. Byrne (2000) used the information obtained in themonotonic loading tests to study the cyclic moment tests. The number of cyclic testscan be reduced in the same way as Byrne and Houlsby (2002) suggest for the verti-cal loading case. Transient tests carried out by Byrne (2000) showed that there is nosignificant difference between tests conducted at different rates (6 s, 10 s and 12 s).Monotonic tests with two cycles have already carried out and are presented in the previ-ous section. These tests were done under drained conditions on loose dry sand samples.

Transient tests of no more than 20 cycles with a period of 10s will be conducted.Reloading during some of the loading reversal cycles will be done to prove the firstMasing’s rule. Constant vertical loads with values of -20 N, 50 N and 100 N will beconsidered. Mean and load amplitude values of M/2R and H will become: 0 ± 50 Nand 0 ± 100 N. Using loose and medium dense samples the total number of tests is 24.Four saturated tests will be undertaken under a constant vertical load of 100 N withM/2R and H equal to 0 ± 100 N.

Moment fatigue tests will be carried out as part of the research project and theywill give information about the caisson response under long term loading conditions.Tests with more than one hundred thousands long-term cycling loading will be under-taken on a footing under certain levels of moment load with a fixed M/2RH ratio. Toforecast what would happen three drained tests will be conducted to replicate the longterm tests just applying 20 cycles. Due to rig constrains (high frequencies) the threeforecaster tests will consider a mean and load amplitude values of M/2R equal to 0 ±10 N with load rates around 0.1 Hz. The vertical load will be held at 100 N during the

37

tests. Results of these forecaster tests will be useful to make decisions in the planningof the long term tests.

5.2.6 Moment loading tests on clay

The soil foundation mainly found in the projected wind farm sites is sand. However,it has been prospected in some of the sites clayed subsoil that underlies the sand layer.Thus the response of the foundation on clay must be taken into account. Monotonicand cyclic moment loading tests will be undertaken with similar characteristics in termsof rate load and magnitude of loads. A total of six tests will be conducted.

5.2.7 Laboratory tests and field tests

Field tests will be undertaken with caissons of diameter 1.5 m and 3 m (Kelly, 2003).In these field tests loads will be applied in such a way to replicate the tetrapod/tripodand monopod cases. As already has been explained that in the first case the responsein compression and tension of the caisson will be tested and for the second case themoment loading response will be tested. Because the tests in the laboratory willlook for the same kinds of response it will be possible compare them if certain mainparameters are reasonably similar. Nonetheless, depending on how the field tests aregoing to be conducted, some of the laboratory tests will be more comparable thanothers. Therefore laboratory tests shadowing the tests carried out at the clay and sandsite will be conducted to analyse among others size effect. Thus it will be possible toscale the small model results to the prototype size for those particular comparativetests. Furthermore that scaling will be extended to the other tests results.

5.2.8 Sample preparation

The programme of testing described will cover the necessary information required todevelop and validate the hyper-plasticity model implemented in the theoretical sectionof the wind farm project. As a consequence, it will be possible the design of the windturbines foundations under the conditions encountered in the offshore of the UK. Theexperimental work will include predominantly sandy soil although clayed soil will betested as well. Dry Leighton Buzzard sand samples prepared with relative densitybetween 35% and 75% will be generated by means of gently pouring the loosest onesand by means of pouring in layers and subsequent vibration the densest ones. In thecase of partially drained conditions, the sand samples used will be finer than for thedrained case (silica sand). Baskarp cyclone sand has been used in previous laboratorystudies by Mangal (1999) and Byrne (2000). That sand can be saturated with oil ofviscosity of 100 centistokes to capture the transient behaviour properly. The use ofpore pressure transducers will allow recording the changes in pore pressure in the plugsoil, inside the caisson. Therefore the effective stresses will be determined and it willbe possible to define a range of parameter values for which weak responses would bereached. Those parameters to be controlled during the tests will be the rate of loading,level of loads and density of the sample. Preparation of clay samples can be doneusing Speswhite kaolin and the consolidation system available in the soil mechanicslaboratory at Oxford. The slurry preparation and consolidation allow good qualityhomogeneous clay specimens to be obtained (Martin, 1994).

38

5.3 Timetable

The following timetable is proposed:

Table 5.1: Timetable

32

Table 5.1 Timetable 2 0 0 3 2 0 0 4 2 0 0 5Experiment No Aug Sep Oct Nov Dec Jan Feb Mar Apr May Jun Jul Aug Sep Oct Nov DecJan Feb Mar Apr May Jun Jul Aug

5.2.1 245.2.2 515.2.3 185.2.4 125.2.5 315.2.6 65.2.7 4Data analysis

Thesis writing up

Total No of tests 146

Experiment:

5.2.1 Suction assisted installation 5.2.2 Vertical cyclic loading 5.2.3 The tensile capacity 5.2.4 The ultimate moment capacity (Swipe tests) 5.2.5 Moment cyclic loading 5.2.6 Moment loading tests on clay 5.2.7 Laboratory tests and field tests

Experiment:5.2.1 Suction assisted installation5.2.2 Vertical cyclic loading5.2.3 The tensile capacity5.2.4 The ultimate moment capacity (Swipe tests)5.2.5 Moment cyclic loading5.2.6 Moment loading tests on clay5.2.7 Laboratory tests and field tests

5.4 Funding

A DTI grant of £478k has been awarded to the Civil Engineering Group for theresearch project: Suction caisson foundations for offshore wind farms. The part relatedwith laboratory testing has £212k for 24 month. A percentage of that amount has beenused during the last year in laboratory staff, equipment and computing. The Groupwas awarded by EPSRC £221k for the completing of the third year research programwhich will cover laboratory testing as well as theoretical work.

39

Bibliography

[1] Andersen, K. and Jostad, H. P. (1999). Foundation design of skirted foundations andanchors in clay. Offshore Technology Conference, Houston, Paper 10824

[2] Andersen, K., Dyvik, R., Schrder, K., Hansteen, O. and Bysveen, S. (1993). Field testsof anchors in clay II: Predictions and interpretation. ASCE Journal of GeotechnicalEngineering, 119, No 10, pp. 1532-1549

[3] Bolton, M.D. (1986). The strength and dilatancy of sands. Geotechnique 36, No. 1, pp.65-78 and discussion 37, No. 2, pp. 219-226

[4] Brinch Hansen, J. (1970). A revised and extended formula for bearing capacity. BulletinNo 28, Danish Geotechnical Institute, Copenhagen, 5-11

[5] Butterfield, R. and Ticof, J. (1979). Design parameters for granular soils (discussioncontribution). Proc 7th Int Conf Soil Mech Fndn Engng, Brighton, 4, pp. 259-261

[6] Butterfield, R. and Gottardi, G. (1994). A complete tree-dimensional failure envelopefor shallow footings on sand. Geotechnique 44, No. 1, pp. 181-184

[7] Butterfield, R., Houlsby, G.T. and Gottardi, G. (1997). Standardized sign conventionsand notation for generally loaded foundations. Geotechnique 47, No. 5, pp. 1051-1054

[8] Byrne, B.W. and Houlsby, G.T. (1999). Drained behaviour of suction caisson founda-tions on very dense sand. Offshore Technology Conference, Houston, Paper 10994

[9] Byrne, B.W. (2000). Investigations of suction caissons in dense sand. DPhil thesis, TheUniversity of Oxford

[10] Byrne, B.W. and Houlsby, G.T. (2001). Observations of footing behaviour on loosecarbonate sands. Geotechnique 51, No. 5, pp. 463-466

[11] Byrne, B.W. (2002). ”Investigating novel foundations for offshore wind turbines”.Proc. Int. Conf. Offshore Mechanic and Artic Engineering, Oslo, Norway. Paper NoOMAE2002-28423

[12] Byrne, B.W. and Houlsby, G.T. (2002). Experimental investigations of the responseof suction caisson to transient vertical loading. ASCE Journal of Geotechnical andGeoenvironmental Engineering 128, No. 11, pp. 926-939

[13] Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2002a). Cyclic loading of shallowoffshore foundations on sand. Proc. Int. Conf. on Physical Modelling in Geotechnics(ICPMG), St John’s, Newfoundland, Canada

[14] Byrne, B.W., Houlsby, G.T., Martin, C.M. and Fish, P.M. (2002b). Suction caissonfoundations for offshore wind turbines. Jour. of Wind Eng. 26 No 3, pp 145-155

[15] Byrne, B.W. (2003). Proposal for laboratory testing of caisson foundations for offshorewind turbines. Report FOT009, Department of Engineering Science. The University ofOxford

[16] Cassidy, M.J., Byrne, B.W. and Houlsby G.T. (2002). Modelling the behaviour ofcircular footings under combined loading on loose carbonate sand. Geotechnique 52,No. 10, pp. 705-712

[17] Dyvik, R., Andersen, K., Hansen, S.B. and Christophersen, H.P. (1993). Field tests ofanchors in clay I: Description. ASCE Journal of Geotechnical Engineering, 119, No 10,pp. 1515-1531

40

[18] Georgiadis, M. and Butterfield, R. (1988). Displacements of footings on sand undereccentric and inclined loads. Canadian Geotechnical Journal, 23, pp. 199-212

[19] Gottardi, G., Houlsby, G.T. and Butterfield, R. (1999). The plastic response of cir-cular footings on sand under general planar loading. Geotechnique 49, No. 4, pp. 453-470

[20] Graham, J., Pinkney, R.B., Lew, K.V. and Trainor, P.G.S. (1982). Curve-fitting andlaboratory data. Canadian Geotechnical Journal, 19, pp. 201-205

[21] Greenpeace Report (2000). North Sea Wind - A powerhouse for Europe technicalpossibilities and ecological considerations. http://www.owen.eru.rl.ac.uk

[22] Houlsby, G.T. and Byrne, B.W. (2000). Suction caisson foundations for offshore windturbines and anemometer masts. Journal of Wind Energy 24, No. 4, pp. 249-255

[23] Houlsby, G.T. (2002). Design calculations for installation of suction foundations.Internal report. Department of Engineering Science, The University of Oxford

[24] Houlsby, G.T. and Cassidy, M.J. (2002). A plasticity model for the behaviour offootings on sand under combined loading. Geotechnique 52, No. 2, pp. 117-129

[25] Johnson, K. (1999). Partially drained loading of shallow foundations. Fourth yearproject. Department of Engineering Science, The University of Oxford

[26] Kelly, R.B. (2003). Proposed field tests of caisson foundations for offshore wind turbines.Report FOT008, Department of Engineering Science. The University of Oxford

[27] Kelly, R.B., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2003). Pressure chambertesting of model caisson foundations in sand. ICOF 2003

[28] Mangal, J.K. and Houlsby G.T. (1999). Partially-drained loading of shallow foundationson sand. Offshore Technology Conference, Houston, Paper 10991

[29] Mangal, J.K. (1999). Partially drained loading of shallow foundations on sand. DPhilthesis, The University of Oxford

[30] Martin, C.M. (1991). The behaviour of jack-up rig foundations on clay. First yearreport. Department of Engineering Science, The Oxford University

[31] Martin, C.M. (1994). Physical and Numerical Modelling of Offshore Foundations UnderCombined Loads. DPhil thesis, The University of Oxford

[32] Meyerhof, G.G. (1951). The ultimate bearing capacity of foundations. Geotechnique 2,No 4, pp. 301-332

[33] Meyerhof, G.G. (1953). The ultimate bearing capacity of foundations under eccentricand inclined loads. Proc. 3rd ICSMFE 1, Zurich, pp. 440-445

[34] Meyerhof, G.G. (1963). Some recent research on the bearing capacity of foundations.Canadian Geotechnical Journal, 1, pp. 16-26

[35] Milborrow, D. (2003). ”Offshore wind rises to the challenge”. Windpower Monthly, 19,No.4, pp. 51-56

[36] Montrasio, L. and Nova, R. (1997). Settlements of shallow foundations on sand:geometrical effects. Geotechnique 47, No 1, pp. 49-60

[37] Nova, R. and Montrasio, L. (1991). Settlements of shallow foundations on sand.Geotechnique 41, No 2, pp. 243-256

[38] Nutt, N.R.F. (1993). Development of the cone pressuremeter. D.Phil. thesis, TheUniversity of Oxford

[39] Pyke, R. (1979). Non linear soil models for irregular cyclic loadings. ASCE Journal ofthe Geotechnical Engineering Division 105 GT6, pp. 715-726

[40] Poorooshasb, H.B., Holubec, I. and Sherbourne, A.N. (1967). Yielding and flow of sandin triaxial compression: Part I. Canadian Geotechnical Journal, 4, No 4, pp. 376-388

[41] Roscoe, K.H. and Schofield, A.N. (1957). The stability of short pier foundations insand. Discussion. British Welding Journal, pp. 12-18

41

[42] Sanham, S.C. (2003). Investigations into the installation of suction assisted caissonfoundations. Fourth year project. Department of Engineering Science, The Universityof Oxford

[43] Schnaid, F. (1990). A study of the cone pressuremeter test in sand. DPhil thesis, TheUniversity of Oxford

[44] Sieffert, J.G. and Bay-Gress, Ch. (2000). Comparison of European bearing capacitycalculation methods for shallow foundations. Proc. Instn Civ. Engrs Geotech Engng,143, pp. 65-74 and discussion 149, pp. 63-64

[45] Stratford, O. (2002). Cyclic loading of caisson foundations in sand. Fourth year project.Department of Engineering Science, The University of Oxford

[46] Tan, F.S.C. (1990). Centrifuge and numerical modelling of conical footings on sand.PhD thesis, Cambridge University

[47] Terzaghi, K. (1943). Theoretical soil mechanics. New York, John Wiley & Sons[48] Ticof, J. (1977). Surface footing on sand under general planar loads. PhD thesis,

Southampton University[49] Vesic, A. S. (1973). Analysis of ultimate loads of shallow foundations. ASCE Journal of

the Soil Mechanics and Foundation Division, 99, No 1, pp. 45-73[50] Vesic, A. S. (1975). Bearing capacity of shallow foundations. In Foundation Engineering

Handbook (ed. Winterkorn, H.F. and Fang, H.Y.), Van Nostrand, New York, pp. 121-147[51] Villalobos, F.A., Byrne, B.W., Houlsby, G.T. and Martin, C.M. (2003). Bearing

capacity tests of scale suction caisson footings on sand: Experimental data. ReportFOT0005/1, Department of Engineering Science, The University of Oxford

[52] Wang, M. C., Demars, K.R. and Nacci, V.A. (1977). Breakout capacity of modelsuction anchors in soil. Canadian Geotechnical Journal, 14, pp. 246-257

[53] Zhu, F., Clark, J. and Phillips, R. (2001). Scale effect of strip and circular footings rest-ing on dense sand. ASCE Journal of Geotechnical and Geoenvironmental Engineering127, No. 7, pp. 613-621

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