Calculation Procedures For Installation Of Suction Caissons
Calculation Procedures for Installation of Suction CaissonsbyG.T. Houlsby and B.W. ByrneReport No. OUEL 2268/04University of OxfordDepartment of Engineering ScienceParks Road, Oxford, OX1 3PJ, U.K.Tel. 01865 273162/283300Fax. 01865 283301Email Civil@eng.ox.ac.ukhttp://www-civil.eng.ox.ac.uk/
Design procedures for installation ofsuction caissons in clay and other soilsG.T. Houlsby1 and B.W. Byrne1Keywords: clay, foundations, suction caissons, installationAbstractSuction-installed skirted foundations, often referred to as suction caissons, are being increasinglyused for a variety of offshore applications. In designing a caisson a geotechnical engineer mustconsider the installation process as well as the in-place performance. The purpose of this paper is topresent calculation procedures for the installation of a caisson in clay. For clay sites, the caissonwill often be used as an anchor, with the ratio of the skirt length (L) to the diameter (D) as high as 5.Calculation methods are presented for determining the resistance to penetration of open-endedcylindrical caisson foundations with and without the application of suction inside the caisson.Comparisons of between predictions and case records are made. A companion paper (Houlsby andByrne, 2004) describes the calculation procedure for installation in sand soils. Finally comments aremade here about installation in s variety of soils other than homogeneous deposits of clay or sand.IntroductionA suction caisson is a large cylindrical structure, usually made of steel, open at the base and closedat the top. It might be used either as a shallow foundation or as a short stubby pile (often called asuction anchor). The shallow foundation option is more common at sandy soil sites (e.g. Bye et al.,1995; Hogervost 1980; Tjelta, 1994; Tjelta, 1995; Tjelta et al., 1990) whilst the anchor/pileapplication is commoner in clay or layered soils (e.g. Colliat et al., 1996; Colliat et al., 1998;Erbrich and Hefer, 2002; Lacasse, 1999; Solhjell et al., 1998). Figure 1 shows typical diameter andskirt depths for various projects reported in the literature (the figure is taken from Byrne (2000) andwith further data from Tjelta (2001)). More recently there is an emerging application of caissons asthe foundations for offshore wind turbines (Byrne et al., 2002; Byrne and Houlsby, 2003).This paper addresses installation in clays and other soils whilst a companion paper (Houlsby andByrne, 2004) considers installation in sand. In the anchor application the caisson will be designedso that the skirt length (L) is much greater than the diameter (D) and the ratio L/D might be as largeas 5 (as shown in Figure 1). As oil and gas exploration heads further offshore and into deeper water,it is likely that anchor applications will become more common. There are particular advantages tousing the suction caisson over other anchoring methods (e.g. drag anchors), in that the caisson canbe accurately located, allowing complex mooring line arrangements to be accommodated. Theability to remove a caisson (by simply reversing the installation procedure) allows mooring linearrangements to be altered over the life of a production vessel; and removal at the end of the designlife.After an initial penetration into the seabed caused by self weight, a suction (relative to seabed waterpressure) is applied within the caisson, which forces the remainder of the caisson to embed itself,leaving the top flush with the seabed. The purpose of this paper is to present design calculations forthe installation of the caisson. Separate calculations are of course necessary to assess the capacity ofthe caisson once installed – either as a shallow foundation or as an anchor. Analyses are presentedfor the magnitude of the self-weight penetration, the relationship between suction and further1Department of Engineering Science, Parks Road, Oxford OX1 3PJ1
penetration, and the limits to penetration that can be achieved by suction. The analyses are“classical” in the sense that they make simplifying assumptions, borrowing techniques from bothpile design and bearing capacity theory. More rigorous analyses, using for instance finite elementtechniques, could be used for particular installations. The analyses presented here should, however,provide a reasonable approximation for design purposes. Similar methods (although differing insome details) to those described below have been published e.g. by House et al. (1999), but ourpurpose here is to draw together a comprehensive design method and compare with case recordsfrom several sources.Installation in ClayFigure 2 shows the key variables in the suction caisson problem, so far as the installation isconcerned. For the purposes of the installation calculation the strength of the clay is characterisedby an undrained strength, which is assumed to increases with depth linearly in the formsu suo ρz . The methods described below can readily be adapted to more complex strengthvariations.Self-weight penetrationThe resistance to penetration is calculated as the sum of adhesion on the outside and inside of thecaisson, and the end bearing on the annular rim. The adhesion terms are calculated, following usualpractice in pile design, by applying a factor α to the value of the undrained strength. The endbearing is calculated, again following standard bearing capacity analyses, as the sum of an N q andan N c term. The result is:V ′ hα o su1 (πDo )adhesion on outside hα i su1 (πDi )adhesion on inside (γ ′h su 2 N c )(πDt ) (1)end bearing on annuluswhere su1 suo ρ h 2 is the average undrained shear strength between mudline and depth h,su 2 suo ρh is the undrained shear strength at depth h, α o and αi are adhesion factors on theoutside and inside of the caisson (as used in undrained pile design) and N c is an appropriatebearing capacity factor for a deep strip footing in clay (typically a value of about 9 might beadopted). For undrained analysis N q 1 .Suction-assisted penetrationOnce the self-weight penetration phase has been completed, so that a seal is formed around the edgeof the caisson, it will be possible to commence the suction installation phase. The applied suction inthe caisson is s relative to seabed water pressure, i.e. the absolute pressure inside the caisson isp a γ w hw s . There are a number of practical limits to the maximum attainable value of s.Amongst these are (a) the absolute pressure at which the water cavitates (usually a small fraction ofatmospheric pressure), (b) the minimum absolute pressure that can be achieved by the given pumpdesign, (c) the minimum relative pressure that can be achieved by the pump.The suction causes a pressure differential across the top plate of the caisson, which resultseffectively in an additional vertical load equal to the suction times the plan area of the caisson. Thecapacity is again calculated as the sum of the external and internal friction, and end bearing term.Note that the overburden term is reduced in the end bearing calculation by the suction pressure,assuming that the flow of soil under the rim occurs entirely inwards. The result is:2
()V ′ s π Di2 4 hα o su1 (πDo ) hα i su1 (πDi ) (γ ′h s su 2 N c )(πDt ) (2)which is readily rearranged to:()V ′ s π Do2 4 hα o su1 (πDo ) hα i su1 (πDi ) (γ ′h su 2 N c )(πDt ) (2a)Note that if the variation of soil strength is not simply linear, all that is necessary is to replace su1with the average strength from mudline to depth h, and su 2 with the strength at depth h. Equation(2) gives a simple relationship between suction and depth. For constant V ′ and a linear increase ofstrength with depth (so that su1 and su 2 are linear functions of h), s is a quadratic function of h .Limits to suction assisted penetrationIn addition to the limit imposed by the maximum available suction, there is a limit to the depth ofpenetration that can be achieved by the action of suction. If the difference between the verticalstress inside and outside the caisson, at the level of the caisson tip, exceeds a certain amount, thenlocal plastic failure may occur, and further penetration may not be possible. The mechanism may bethought of as a “reverse” bearing capacity problem, in which the soil flows into the caisson.The average vertical stress (relative to local hydrostatic) inside the caisson at tip level is relativelyπDi hα i su1. The third term in this expression arises fromstraightforward to estimate as s γ ′h πDi2 4the downward friction inside the caisson, and here it is assumed (for simplicity) that this results in auniform increase of vertical stress at all radii in the caisson. Note that the assumption of a uniformincrease in vertical stress within the caisson is clearly unreasonable at small values of h D , but itwill be seen below that this calculation is only needed at h D values greater than about 2, forwhich the uniform increase is a reasonable approximation.The relevant stress outside the caisson is much harder to estimate, since the downward load fromadhesion on the outside of the caisson will enhance the stress in the vicinity of the caisson, but thisenhancement is difficult to calculate. However, making the simplifying assumption that thedownward load from the adhesion is carried by a constant stress over an annulus with inner andouter diameters Do and Dm , this stress (again relative to local hydrostatic) may be calculated asπDo hα o su1γ ′h . Thus the “reverse bearing capacity” failure would occur when2π Dm Do2 4πDi hα i su1πDo hα o su1 s γ ′h γ ′h N c* su 2 , where N c* is a bearing capacity factor222πDi 4π Dm Do 4appropriate for uplift of a buried circular footing. Substituting the solution for s into equation (2)and simplifying gives:()(V ′ N c* su 2) πDi2Di2 hα o su1 (πDo ) 1 2 4Dm Do2 () (γ ′h s N )(πDt )u2 c (3)which can be solved for h. Note, however, that although the above equation appears linear in h, infact su1 and su 2 are themselves linear functions of h, so that the solution again involves solving aquadratic. Furthermore it would be rational to assume that Dm increases with penetration, forinstance in the form Dm Do 2 f o h , where f o is a constant “loadspread” factor. A further3
development would be to allow the enhancement of the stress to vary (say linearly) from zero atDm to a maximum at the caisson surface.It is worth, however, considering some approximate solutions for the maximum penetration. Formany cases the final term (the end bearing) is small. We consider also the case where the appliedload V ′ is small, and make the approximation Do D Di . If we write Dm mD , then equation(3) leads to the following result for this simplified case:N* s 1 h c u 2 1 D 4α o su1 m 2 (4)The factor N c* 4α o is likely to be in the region of about 3, although it could vary considerably.The factor su 2 su1 would be 1.0 for a homogeneous soil, and 2.0 for the extreme of a soil with astrength increasing linearly with depth from a value of zero at the surface. The final factor variesfrom 1.0 if m is assumed to be very large, to 0.75 if say m 2 . The overall result is that thecalculated maximum attainable value of h D is likely to be from about 2.5 for stiff clays (withstrengths approximately uniform with depth) to 5 for soft normally consolidated clays (withstrengths approximately proportional to depth). The effect of accounting for the external load V ′would be to increase these values. Equation (4), however, provides a useful estimate of themaximum h D ratio of a suction-installed caisson that could be reliably installed in clay. Ifdifferent assumptions are made about the way the external adhesion load enhances the verticalstress are made, the same broad conclusions arise, although the precise figures will vary.It should be noted that some measured values of installations indicated that higher h D ratios thanimplied by the above calculation may be achievable. The above may therefore be treated as aconservative calculation.Note also that the end bearing calculation in equations (1) and (2) does not take into account anyenhancement of the stress level inside or outside the caisson due to the frictional terms. This followsconventional piling design calculations, in which no such correction is usually included. If thiseffect was to be taken into account, the factor γ ′h in equation (1) would be replaced by whicheverπDo hα o su1πDi hα i su1or γ ′h (almost invariably the former). Onceis the smaller of γ ′h 2π Dm Do2 4πDi2 4()suction is started, γ ′h s in equation 2 is replaced by the smaller of γ ′h s γ ′h πDi hα i su1πDo hα o su1()2π Dm Do2 4or(usually the latter except at very small suction). In practice these changes makeπDi2 4very small differences to the calculation.The effect of internal stiffenersMost suction caissons include some internal structure, usually consisting of either vertical plates orannular plates, to provide strength and stiffness to the cylindrical shell, either to suppress bucklingduring suction-assisted penetration, or (in the case of a caisson anchor) to reinforce the caisson atthe pad-eye connection. The analysis for the case of annular stiffeners is not considered here, butthe use of vertical stiffeners results in only a small change in the calculation.4
In principle stiffeners could be located on the outside of the caisson, but this option does not usuallyseem to be adopted. The additional resistance offered by the stiffeners can be taken into account byan adhesion term of the form hαsu1l , where l is the perimeter length of the stiffeners (usuallyapproximately twice the plate length for thin plate stiffeners), and an end bearing term of the form(γ ′h su2 N c )A , where A is the end area of the stiffeners. The area on which the suction acts (on theleft side of Equation 2) should also be reduced by A, although this correction will usually be tiny.Note that if the stiffeners do not extend the full depth of the caisson, appropriate corrections arerequired for the value of h used in the contribution from the stiffeners, and in the appropriate su1and su 2 values.In the calculation of the maximum attainable depth using suction, note that the terms involvingadhesion on the inside of the caisson cancel, and have no overall effect on the calculation. The sameis true for terms resulting from the resistance from internal (but not external) stiffeners, so forinternal stiffeners only Equation (4) can still be used.Example 1Consider a suction caisson of outside diameter 12m, wall thickness 45mm and depth 5m. Such acaisson might be considered as a foundation for an offshore structure. The caisson is stiffened by 30plates 25mm thick and 200mm deep welded as radial fins on the inside of the caisson, andextending for the top 4m of the caisson only. The soil profile is idealised as a layer 2m thick ofconstant strength 20kPa, with below that a linear increase of strength from 25kPa at 2m at a rate2.5kPa/m. The buoyant unit weight is taken as 6kN/m3. The end bearing factor N c is taken as 9,and the adhesion factor α as 0.6 for the outside of the caisson and 0.5 for inside and for thestiffeners. The maximum applied vertical load (including the weight of the caisson and buoyancyeffects) is 1000kN and the water depth is 50m.The calculations described above have been implemented in a spreadsheet-based program “SCIP”(Suction Caisson Installation Prediction). Figure 3 shows the calculated loads required to install thecaisson in the absence of suction. Figure 4 shows the predictions from SCIP of the variation ofsuction with depth required for installation, and in this case the maximum suction required is 49kPa.Example 2: Predicted of installation pressures compared to centrifuge testsHouse and Randolph (2001) conducted a series of tests on the centrifuge at the University ofWestern Australia, investigating the installation of suction caissons in normally consolidated clay.The experiments were carried out at 120g. The strength profile of the clay could be idealised as zeroat the surface increasing with depth at a gradient of 144kPa/m to a depth of 67mm then at 204kPa/m(at prototype scale these represent rates of increase of 1.2kPa/m and 1.7kPa/m). The effective unitweight of the soil (accounting for the 120g acceleration) was determined to be 792kN/m3. Thedimensions of the caisson were 30mm diameter, 0.5mm wall thickness and 120mm skirt length.(Equivalent prototype dimensions 3.6m diameter, 60mm wall thickness, 14.4m skirt length). Aneffective vertical load of 15.3N was applied to the caisson.Figure 5 shows the penetration resistance for the caisson without using suction, showing that mostof the resistance is in the skirt friction. Figure 6 shows an estimated suction penetration curve,which shows good agreement with the experimental data reported by House and Randolph (2001).The self-weight penetration amounts to 41mm and the maximum suction pressure required is143.9kPa. An adhesion factor of 0.5 was used for both internal and external walls.5
Example 3: Prediction of plug failureA series of tests were conducted by House et al. (1999) on the laboratory floor to investigate plugfailure during installation of suction caissons in normally consolidated clay. They investigated threecaissons with diameters 10.4mm, 15.9mm and 37.2mm. All caissons had a wall thickness of 0.4mm and an L/D ratio of 8.In the following a comparison is made for the 15.9mm diameter caisson. The soil strength profilewas estimated by House et al. (1999) to be 75kPa/m and the effective unit weight to be 5.9kN/m3.The caissons were initially pushed into the clay to a penetration of approximately one diameterbefore the suction was applied. Assuming a circular end bearing capacity factor of 8.5 (Houlsby andMartin, 2003), the maximum calculated penetration by SCIP that is possible before a plug failure isexpected is 83mm or an h/D 5.2. This can be compared to the conclusions drawn by House et al.(1999). They compare the volume of water withdrawn from the caisson cavity during installation tothe displaced volume within the caisson (assuming heave has not occurred). When more water isevacuated than can be accounted for by the installed portion of the caisson, they infer that plugheave has occurred. Figure 8 shows, for two installations of the 15.9mm diameter caisson, theexcess volume of water removed, plotted against normalised penetration. For the cases shown,House et al. (1999) deduced that plug failure occurs at an L/D ratio between 4 and 5, which agreeswith the prediction given above. Again an adhesion factor of 0.5 was used. Note that although plugfailure occurred it was still possible to install the caisson further. Installation continues until allwater has been withdrawn from the internal cavity. The consequence of plug failure is that thecaisson cannot be installed to its full design depth.Example 4: Nkossa Field InstallationThis calculation involves some modification to the basic procedures described above to account forthe geometry of the caissons used in the Nkossa field off the coast of West Africa (Colliat et al.,1996; Colliat et al., 1998). Two different anchor sizes were used depending on the loadingconditions. We will only consider the installation of the smaller of the two, defined by Colliat et al.(1998) as a Type I anchor. The geometry of the caissons is unusual, as they have a step change indiameter part way down the caisson. The bottom section is 4m in diameter and extends for 4.8mwhilst the top section is 4.5m in diameter and is 7.5m long. The anchor chain lug is located at thechange in caisson diameter. The wall thickness for the pipe sections was 15mm and the designpenetration was 11.8m. The larger top section was to accommodate any soil heave that occurredduring installation. Internal stiffening plates are also believed to have been used. However, these areomitted in the calculation here, as there is insuffi
outside and inside of the caisson (as used in undrained pile design) and Nc is an appropriate bearing capacity factor for a deep strip footing in clay (typically a value of about 9 might be adopted). For undrained analysis Nq 1. Suction-assisted penetration Once the self-weight penetration phase has been completed, so that a seal is formed around the edge of the caisson, it will be possible .
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