Penetration Resistance of Offshore Skirted Foundationsand Anchors in Dense Sand
Knut H. Andersen1; Hans Petter Jostad2; and Rune Dyvik3
Abstract: Penetration of skirts is an essential design issue for offshore skirted foundations and anchors in sand. Skirts may not penetratefar enough into dense sand by the available submerged weight alone. It may therefore be necessary to apply underpressure inside the skirtcompartment to produce an increased driving force and to reduce the penetration resistance. This paper recommends procedures tocalculate penetration resistance and required underpressure for skirts penetrated in dense sand with and without interbedded clay layers.The recommendations are based on interpretation of skirt penetration data from prototypes, field model tests, and laboratory model testsin dense sand. The paper first presents a model to calculate the penetration resistance of skirts penetrated by weight, or other externalvertical load that does not cause flow of water in the sand. Two models are considered; one based on bearing capacity equations withfriction angles from laboratory tests, and the other one based on empirical correlations with CPT tip resistance. The bearing capacitymodel gives more consistent correlations with the empirical data than the CPT model. Thereafter, a model to account for the effect ofunderpressure applied inside the skirt compartment is proposed. This model is developed based on interpretation of available prototypeand model test data from skirts penetrated by underpressure. The results show that underpressure facilitates skirt penetration in sandconsiderably by providing both an additional penetration force and a reduced penetration resistance. It is also shown that interbedded claylayers can prevent flow of water through the sand and eliminate the beneficial reduction in penetration resistance.
DOI: 10.1061/�ASCE�1090-0241�2008�134:1�106�
CE Database subject headings: Sand; Layered soils; Weight; Model tests; Offshore structures; Foundations; Anchors.
Introduction
Skirted foundations and anchors have proven to be competitivesolutions for various types of floaters, fixed offshore platforms,and subsea systems. More than 485 skirted anchors, also calledsuction anchors, had been installed for anchoring floaters at morethan 50 different sites by year 2004 �Andersen et al. 2005�. Mostof these anchors are in clay, but some are also in sand or layeredsoils. An example of skirted anchors is shown in Fig. 1. Examplesof skirted foundations in sand are the offshore steel platforms atthe Draupner E and Sleipner T sites in the North Sea �Fig. 2��e.g., Tjelta 1995�. Skirted foundations can also be used to in-crease the moment fixity and the bearing capacity of offshorejack-up platforms �Jostad and Andersen 2006� and can be an at-tractive foundation alternative for wind power structures �e.g.,Houlsby et al. 2005�.
One main design challenge for skirted structures in sand is topenetrate the skirts deep enough to obtain the required capacity
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and to keep within acceptable displacements during design load-ing. The high penetration resistance in sand normally requires thatan underpressure is applied within the skirt compartment to pro-duce an increased driving force in addition to the weight. Theunderpressure will also generate hydraulic gradients in the sandoutside and within the skirt compartment. These gradients reducethe penetration resistance at the skirt tip and inside the caisson.Caution is needed, however, not to exceed the critical gradientwhere one may create erosion channels along the skirts, excessiveloosening of the sand within and below the skirts, or groundfailure. If there are clay layers within the sand, the flow of watermay be prevented, and the penetration resistance will not bereduced. In cases with layered sand and clay, very high under-pressures may therefore be needed to penetrate the skirts. Further,the penetration resistance for anchors with stiffeners may bevery sensitive to small changes in stiffener arrangement and soillayering.
This paper first presents a model to calculate the penetrationresistance of skirts penetrated by weight, or other external verticalload that does not cause flow of water in the sand. Two modelsare considered; one based on bearing capacity equations and fric-tion angles from laboratory tests, and the other one based onempirical correlations with CPT tip resistance. Thereafter, amodel to account for the effect of underpressure applied inside theskirt compartment is proposed. The calculation models are devel-oped based on interpretation of available prototype and model testdata from skirted structures.
Most of the work presented herein was done as part of a jointindustry-sponsored project completed in June 2001. Houlsby andByrne �2005� have later published design procedures for installa-tion of skirted foundations in sand. Their procedure is based on a
theoretical framework with several simplifying assumptions, as
no theoretical solution exists to calculate the resistance of a pen-etrating skirt accounting for the effects of the large penetrationdepth over skirt thickness ratio, the increased stresses due to fric-tion along the inside and outside of the skirt and the pore waterflow, and the variation in friction angle in the soil due to largevariations in shear strain and normal stresses around the skirt tip.The stress distribution in the soil due to friction along the outsideskirt, which influences both the resistance along the outside skirtand at the skirt tip, is assumed. The effect of the applied under-pressure inside the skirt compartment is taken into account byusing effective unit weight of the sand corrected for steady statepore water flow condition.
Fig. 1. Suction caissons for the Njord field in the North Sea �Photo:Courtesy of Per Sparrevik, NGI�
Table 1. Measured and Calculated Skirt Penetration Resistance in Proto
CaseD
�m�z
�m�t
�m�
qc
CPTat z
�MPa�
qc
CPTat z /2�MPa�
Dr
�%���
�kN/m3��
�deg� �
Prototype 5 0.75 0.02 8.3 4.15 �100 9.8 44 3
A
Draupner 12 1.8 0.04 55 27.5 �100 10.3 44.5 4
E
Sleipner 14 1.95 0.0375 44 22 �100 10.3 45.5 4
T
Harding 5 7 0.035 30 13c — 10 43.5 3
A7
Harding 5 7 0.035 30 7c — 10 43.5 3
A8aBearing capacity method with triaxial friction angle.bCPT method with tip resistance and wall friction related to CPT.c
Equivalent qc value accounting for variation in qc with depth.
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Simplified equations for the skirt penetration resistance arealso used for the work presented herein, however, the solution iscalibrated by observed skirt penetration data from prototypes,field model tests, and laboratory model tests.
Prototype and Model Test Data
The prototype and model test case records are summarized inTables 1–3. The tables list the diameter �D�, the skirt wall thick-ness �t�, and the depth �z� that the anchors penetrate to under thegiven weight. All the cases are with steel skirts. The tables alsodefine the sand characteristics, defined by the measured cone pen-etration resistance �qc�CPT��, the relative density �Dr�, and thedrained peak friction angle ���. The friction angle is the drainedtriaxial friction angle recommended in the soil parameter reportsfor the various projects. � is normally determined at an effectivenormal stress representative for the stresses in the soil along a
Fig. 2. Draupner E jacket with skirted foundations �Photo: Courtesyof Tor Inge Tjelta, Statoil�
potential failure surface for anchor capacity calculations �typi-cally about 100–250 kPa�. The measured penetration load�Pf ,measured� is given for all cases. The measured side friction isgiven for one of the model tests, determined as the differencebetween the total resistance to penetrate to a given depth minusthe measured skirt tip resistance at that depth.
Prototype Cases
Two of the prototypes cases in Table 1, Draupner E and SleipnerT, are skirted foundations with painted steel skirts. The otherthree, Prototype A and Harding A7 and A8, are suction anchorswith corroded steel skirts.
Table 2. Measured and Calculated Skirt Penetration Resistance in Labor
CaseD
�mm�z
�mm�t
�mm�
qc
CPTat z
�kPa�
qc
CPTat z /2�kPa�
Dr
%��
�kN/m3��
�deg��
�de
PEN5 557 40 8 65 33 84 9.85 43.2 38
PEN9 557 50 8 80 40 82 9.85 41.6 37
PEN12 557 50 8 80 40 83 9.85 41.5 37
PEN13 557 20 8 33 16 81 9.85 42.5 38
PEN1-3 557 50 8 80 40 82 9.85 42.5 38
100 160 80
150 240 120
200 320 160
250 400 200
280 450 225aBearing capacity method with triaxial friction angle.bCPT method with tip resistance and wall friction related to CPT.
Table 3. Measured and Calculated Skirt Penetration Resistance in Field
CaseD
�m�z
�m�t
�mm�
qc
CPTat z
�MPa�
qc
CPTat z /2�MPa�
Dr
�%���
�kN/m3��
�deg� �
16/11 E1 1.5 0.4 12 5.7 2.85 �100 10.3 44.5 4
16/11 E2 1.5 0.8 12 18 6.9c �100 10.3 44.5 4
16/11 E3 1.5 0.2 12 2.8 1.4 �100 10.3 44.5 4
Sleipner T1 1.5 0.8 12 13.8 5.6c �100 10.3 45.5 4
Sleipner T2 1.5 0.8 12 13.8 5.6c �100 10.3 45.5 4
Sleipner T3 1.5 0.4 12 4.8 2.4 �100 10.3 45.5 4
Sleipner T4 1.5 0.8 12 13.8 5.6c �100 10.3 45.5 4
Sleipner T5 1.5 0.8 12 13.8 5.6c �100 10.3 45.5 4
Øresund 2 0.2 10 0.42 0.21 — 10 41 3
2 0.3 10 1.0 0.5 — 10 41 3
2 0.2 10 0.75 0.3 — 10 37.6 3
2 0.3 10 2.8 0.5 — 10 37.6 3aBearing capacity method with triaxial friction angle.bCPT method with tip resistance and wall friction related to CPT.c
Equivalent qc value accounting for variation in qc with depth.
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At the Harding anchor locations, the soil conditions were quitevariable, with occasional layers of clay and dense sand. There arethus uncertainties about the exact layering and the friction angle.The two anchors in Table 1 �A7 and A8� were penetrated by anunderpressure about six to seven times the underpressure thatwould have given a critical exit gradient if flow had occurred.This means that clay layers have prevented flow to occur, and theskirt penetration is therefore interpreted by assuming that the pen-etration resistance is not influenced by flow of water.
Based on the measured underpressures with depth it is as-sumed that the dense sand starts 4 and 3 m above final skirt tippenetration depth in Anchors A7 and A8, respectively. The side
shear resistance in remolded clay and loose sand above the densesand will be small and within the uncertainties of the penetrationresistance in the dense sand. The penetration resistance above thedense sand is therefore neglected.
Laboratory Model Tests
The laboratory model tests in Table 2 were run by NGI for Statoil.The laboratory model tests had a diameter of 0.557 m and 8 mmthick, 0.32 m long skirts and were run in a test tank with adiameter of 1.6 m �Sparrevik 1995�. The skirt tip resistance wasmeasured by four calibrated strain gauge transducers on beamsintegrated into the skirt at 90° distance. The sand was preparedsubmerged by a combination of hydraulic gradient and vibrationto a height of about 0.85 m. There was some difference in therelative sand density among the tests. The friction angle thereforevaried somewhat from one test to another. Tests PEN5, 9, 12,and 13 were first penetrated by a weight representative for proto-type conditions. Further penetration was achieved by applyingunderpressure inside the caisson. The data in Table 2 are for theend of the weight penetration phase. Test PEN1-3 was penetratedto full depth by applying a higher weight than in the other tests,and gives data for larger penetration depths than the other tests.
Field Model Tests
The field model tests in Table 3 include tests at 16/11E, SleipnerT, and one test in Øresund. The 16/11E and the Sleipner T testswere performed for Statoil by the joint venture NGI/Fugro-McClelland. The loading system consisted of a seabed frame withdrill string jacks. The submerged weight of the test rig �275 kN�and the reaction load from the jacks were transferred to the sea-bed by four legs. The Øresund test was performed for MasterMarine A/S to test the feasibility of using suction anchors formooring heavy offshore lifting barges �Johansson et al., 2003�.The self-weight penetration was not the primary purpose of thetest, and very accurate measurement was not performed of theself-weight penetration depth. There were also uncertainties in theCPT resistance in the upper decimeters.
Penetration Resistance under No Pore Water Flow
Calculation Models
GeneralThe penetration resistance of a skirt without stiffeners, Pf, is thesum of the bearing capacity of the skirt tip, Qtip, and the frictionalong the skirt wall Qside, i.e.
Pf = Qtip + Qside �1�
The bearing capacity and the wall friction have been calculatedby two different models, as explained in the following. The firstmodel uses bearing capacity theory with friction angle from labo-ratory tests. Both triaxial and estimated plane strain frictionangles are applied. The second model uses an empirical modelwhere the skirt tip penetration resistance is related to the cone tipresistance measured in CPT tests. The skirt wall friction is relatedboth to the CPT tip resistance and to effective stresses and friction
angle.
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Bearing Capacity Theory with Triaxial Friction AngleIn the bearing capacity model the penetration resistance is calcu-lated as
Pf = Qtip + Qside = qtipAtip + fs,avAwall �2�
where qtip=0.5��tN�+qNq. fs,av=0.5 K �� z tan � �average fric-tion over skirt length�. Atip=�Dt. Awall=2�Dz �sum of insideand outside�; Nq=e� tan � tan2�45+� /2�. N�=1.5�Nq−1� tan �.q=effective overburden pressure outside skirt at skirt tip level;��=effective unit weight of sand; �=peak drained friction angleof sand; r=roughness factor between skirt wall and sand;�=friction angle between sand and skirt wall ��=r��; K=ratiobetween horizontal and initial vertical effective stresses at skirtwall; t=skirt wall thickness �assumed to be small compared to thediameter�; D=diameter of foundation or anchor �external D canbe used since t is small�; and z=depth below sand surface.
The expressions assume homogeneous sand. In case of layer-ing, the skin friction, fs,avAwall, can be calculated as the sum of thecontributions from the individual layers.
Normally, the effective overburden is assumed to be q=��z,but in this case with long skirts relative to the skirt wall thickness,the skirt wall friction may give an important additional verticalnormal stress outside the skirt wall �e.g., Clausen, 1998�. Thisadditional vertical stress may cause an increased skirt tip resis-tance, which instead of using a depth factor, is added to the ef-fective overburden as follows:
q = ��z + � f f s,tip = ��z�1 + � fK tan �� �Clausen 1998� �3�
where � f =ratio between vertical normal stress increase and skirtwall friction at skirt tip level and fs,tip=skirt wall friction at skirttip level. Stiffeners will give an increased resistance, which maybe calculated in the same way as the tip resistance.
Bearing Capacity Theory with Plane Strain Friction AngleThe equations are the same as described earlier, but uses the planestrain friction angle instead of the triaxial friction angle.
The argument for using plane strain instead of triaxial frictionangle is that the displacement pattern around the skirt tip may becloser to plane strain than to axisymmetry as the skirt wall is thincompared to the skirt diameter. However, plane strain tests arenormally not included in conventional soil investigations andmeasured plane strain friction angles have not been available forthe case records in Tables 1–3. The plane strain friction angles forbackcalculation have therefore been estimated, as explained later.
Empirical Model Related to CPT ResistanceIn the empirical model the penetration resistance is calculated as
where qc=cone tip resistance measured in CPT tests;ktip=empirical constant relating skirt tip resistance to qc; andkside=empirical constant relating skirt wall friction to qc.
An alternative model that has also been used is to relate theskirt tip resistance to the CPT resistance, as above, but to calcu-late the skirt wall resistance from effective stresses and frictionangle, i.e.
Bearing Capacity Theory with Triaxial Friction AngleThe calculations are based on the drained triaxial friction angle,�, determined at an effective normal stress representative for thestresses in the soil along a potential failure surface for capacitycalculations �typically about 100–250 kPa�. An interface frictionangle between skirt wall and sand of �=0.9� was used, corre-sponding roughly to the ratio between the recommended values of� and � for piles in dense sands in API �1993�. A ratio betweenvertical normal stress increase and skirt wall friction of � f =1 wasapplied, based on trial and error backcalculation of the data. Thelateral earth pressure coefficient, K, is discussed in the following.The same K value was used along inside and outside skirt, even ifK along the inside and outside wall may be different. The result ofthe backcalculation is presented in Tables 1–3 and commentedupon in the following.
A lateral earth pressure coefficient of K=0.8 was used for theprototype cases. The value of K=0.8 is the same as recommendedby API �1993� for calculation of drained shaft friction of open-ended unplugged piles. Draupner E and Sleipner T had paintedskirts. Paint has been shown to give a significant reduction in skirtfriction in clay, both in the prototype and in ring shear tests withpainted interface �Dendani and Colliat 2002; Andersen et al.2005�. However, a series of ring shear tests on sand at NGIgave the same interface strength with and without paint. The in-terface friction angle was thus not reduced in the cases with paint.Table 1 shows that these assumptions give good correlation be-tween calculated and measured penetration resistances for the fiveprototype cases.
In the laboratory model tests �Table 2�, the density of the sandvaried somewhat among the tests, giving some variation in theapplied friction angle from one test to another. The lateral earthpressure coefficient was increased to K=1.1 to obtain good agree-ment with the measured resistance. An increased K seems reason-able, as the radial wall of the sand bed container will limit thelateral displacements of the sand outside the skirt and causehigher lateral reaction stresses against the skirt wall during skirtpenetration than in a free field situation.
The results for Test PEN1-3, which was penetrated to a largedepth/diameter ratio by weight, show that the calculation modelgives good agreement with the measured skirt penetration resis-tance at all depth/diameter ratios.
The tip resistance was measured separately in Test PEN1-3.This made it possible to check the calculation of each of the twocomponents, i.e., tip resistance and wall resistance. The result isincluded in Table 2, showing that the agreement is good not onlyfor the calculated and measured total resistance, but also forthe calculated and measured skirt wall resistance throughout thepenetration depth. The data in Table 2 also show that the directcontribution of the skirt wall resistance to the total resistance isnegligible at small penetration depths and generally contributesless than 15%, even at large penetration depths. Indirectly, how-ever, the skirt wall resistance still contributes significantly to thetip resistance. According to the calculations, the interaction fromthe skirt wall resistance contributes about 60% of the calculatedtip resistance at full penetration depth. For a prototype founda-tion, like the Draupner E, the same calculations show that skirtwall resistance at end of weight penetration �1.8 m� is about 13%and contributes about 35% to the skirt tip resistance.
In the 16/11E and Sleipner T field model tests �Table 3�, thelateral earth pressure coefficient also had to be increased to a
value higher than K=0.8 to get good agreement. This can be
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explained by the loads from the weight of the test rig and thereaction loads from the jacks, which were transferred to the sea-bed by four pods located only 0.65 m from the anchors.
Agreement between calculated and measured penetration re-sistance of the Øresund field test could be obtained with param-eters within the possible parameter range. In this case K=0.8 wasused, as there were no reaction pods. However, the uncertainty inthe measured penetration depth and the CPT resistance in theupper decimeters were too large to draw definite conclusions forthis test. Calculations with various possible parameter choices areshown in Table 3.
In summary, the bearing capacity calculation model, wherethe increase in skirt tip resistance from interaction with the skirtwall friction is taken into account, gives good agreement with themeasured penetration resistance when the triaxial friction angle isused. The most important contribution of the skirt friction isthrough its influence on the skirt tip resistance.
For prototypes in dense sand with no interaction from neigh-boring structures, good agreement is obtained with an earth pres-sure coefficient against the skirt wall of K=0.8, a friction anglebetween skirt wall and sand of �=0.9�, and a ratio between ver-tical normal stress increase and skirt wall friction at skirt tiplevel of � f =1. A higher value of K must be considered if thereare other structures nearby. Since these parameters are empirical,they should be used together with Nq and N� values determinedby the above-given expressions and a peak triaxial effective fric-tion angle determined at effective stresses representative forcapacity analyses �i.e., at vertical effective consolidation stress of�vc� =100–250 kPa�, even if the effective stresses may be highernear the skirt tip.
Bearing Capacity Theory with Plane Strain Friction AnglePlane strain friction angles were not available for the cases thatwere backcalculated �Tables 1–3� and had to be assumed. Twoseries of calculations were made; one where the plane strain fric-tion angle was assumed to be 1.1 times the triaxial friction angle,and one where the plane strain friction angle that gave best agree-ment with the measured penetration resistance was searched for.The empirical factors K and � f found when the triaxial frictionangle was used may include a plane strain effect, and thereforenew sets of K and � f were allowed. The interface friction anglebetween skirt wall and sand was kept as �=0.9�.
Equally consistent agreement with the measured penetrationresistances was not obtained by using estimated plane strain in-stead of triaxial friction angles. Use of plane strain friction anglealso has the disadvantage that plane strain �biaxial� tests are moreexpensive to perform than triaxial tests, and normally not per-formed as part of a typical offshore soil investigation.
Empirical Model Related to CPT ResistanceAs mentioned, there are two versions of this empirical model.They both relate the skirt tip resistance to the CPT resistance, butone version relates the wall resistance to the CPT resistance,whereas the other version calculates the wall resistance from theeffective stresses and the friction angle.
The result of the backcalculation where both skirt tip and wallresistances are related to the CPT tip resistance is presented inTables 1–3. Different values of kside have been assumed, and thevalue of ktip that gives best agreement with the measured penetra-tion resistance has been determined for each of them.
In the prototypes, Table 1 shows that the backcalculatedvalue of ktip varies between 0.01 and 0.55 when a value of
kside=0.0015 is used. If the side shear factor is reduced to
kside=0.0010, the ktip range increases to 0.03 and 0.6. The ktip isthus not very sensitive to moderate variations in kside. If kside isincreased to kside=0.00175 all the measured resistance is taken asside shear in the Draupner E case and ktip becomes 0.0, which isnot reasonable.
The lowest ktip values for the prototypes in Table 1 occur forDraupner E and Sleipner T. One possible reason is that the sideshear is low due to the painted walls in these cases, even if this isnot supported by the ring shear tests mentioned previously. If kside
is set to zero, ktip increases to 0.08 for Draupner E and to 0.13 forSleipner T.
These results indicate that it is difficult to find one set of ktip
and kside that fits all the measured prototype skirt penetration re-sistances. It may be worth noting however, that the high factorsoccur for the Harding anchors, which were penetrated to largerdepth/diameter ratios than the three other prototypes. Thus theremay be a depth/diameter effect on the ktip and kside factors, but thisis not supported by the laboratory model test as discussed in thefollowing.
In the laboratory model tests �Table 2�, backcalculation wasdone by first finding the kside factor that gave best agreement withmeasured wall resistance with depth in Test PEN1-3. This gave akside=0.0053. Good agreement with measured total penetrationresistance was found at all depths with ktip=1.03–1.19 �Table 1�.With kside=0.0053, the best agreement with the measured totalskirt penetration resistance in the five laboratory model tests wasfound with ktip in the relatively narrow range of 0.93–1.24. How-ever, in the laboratory model tests, the CPT equipment was notidentical to the field cone, and the CPT was run after the modeltesting. The measured CPT resistance is therefore uncertain, andthe laboratory model tests are not so well suited to develop em-pirical side and tip factors for use in prototype cases.
In the field model tests �Table 3�, backcalculation of the16/11E and Sleipner T tests with kside=0.0015 gave ktip in therange of 0.09–0.24, and backcalculation with kside=0.0010 gavektip in the range of 0.12–0.25 �Table 3�. This is within the rangefound for the prototypes. As for the prototypes, ktip is not sosensitive to moderate variations in kside.
In the Øresund field tests, the data in Table 3 show that agree-ment between the calculated and measured penetration resistancecan be obtained for combinations of kside and ktip within the rangeof kside and ktip in the prototypes. The uncertainty in the measuredpenetration depth and the CPT resistance in the upper decimeterswere too large to draw more definite conclusions from these tests.
In summary, the empirical model which relates the skirt tipand wall penetration resistance to the CPT resistance gives a largerange in the backcalculated ktip. The parameter range was found tobe
ktip = 0.01 – 0.55 when kside = 0.0015 and
ktip = 0.03 – 0.60 when kside = 0.0010.
If it is assumed that the painted Draupner E skirt is completelysmooth �i.e., kside=0.000�, the low value increases to ktip=0.08.
The reason for the large variation in the backcalculated ktip
value may be due to the difference in the geometry and the wallroughness between the skirt and the CPT. If the wall frictioninteracts with the tip resistance, the interaction will be a functionof the geometry and the wall roughness. The interaction effect fora skirt may then not be properly modeled by experience from theCPT. In addition, the tip penetration resistance and thereby the
effective stresses will be different for the skirt and the CPT. As
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the friction angle is a function of the effective stresses, the frictionangle may thus be different around the skirt tip and the CPT.
When the wall resistance was based on effective stresses andfriction angle, the same values of � and K were used as in theeffective stress backcalculation presented earlier. Detailed resultsare not included in Tables 1–3, but a parametric study showed thatthe penetration resistance was only marginally influenced byvarying K within a range of 0.8–1.85. The side shear has lessinfluence on the backcalculated skirt tip parameter than when thebearing capacity theory with triaxial friction angle was used, be-cause there is no interaction between wall resistance and skirt tipresistance in this model. For the prototypes, the backcalculatedvalue of ktip was found to vary between 0.07 and 0.47. The rangeof the backcalculated ktip value is somewhat smaller than andwithin the range found when the wall resistance was related to theCPT resistance, but the range in ktip is still large.
Penetration Resistance under Combined Weightand Underpressure
Effect of Underpressure
By applying underpressure beneath the top plate inside the skirtcompartment skirts can be penetrated deeper than they will pen-etrate under self-weight. This underpressure will have the follow-ing two effects:• The pressure difference between the top and bottom of the
top plate will give an increased driving force equal to thepressure difference times the area of the top plate insidethe compartment.
• If the soil profile consists of sand without clay layers, theunderpressure will cause flow of water from the sea outsidethe skirt compartment through the sand to the top of the sandplug inside the skirt compartment. This flow will change thepore pressure field and the effective stresses in the sand. Theeffective stresses will increase along the outside of the skirts,where there are downward gradients. Inside the skirts, how-ever, the effective stresses will decrease. Beneath the skirt tip,there will be horizontal gradients, which will reduce the skirttip resistance. For a circular skirt compartment, the gradientsare smaller outside than inside the skirt. Thus the net effect ofthe increased resistance outside the skirt and the reduced re-sistances inside the skirt and beneath the skirt tip is a reducedpenetration resistance.If there are clay layers in the sand, the water flow will be
prevented. The underpressure will then increase the penetrationforce, but it will not reduce the penetration resistance.
Equipotential lines developed from finite element flow analy-ses for steady state flow into a circular skirt compartment in sandwithout clay layers are shown in Fig. 3 for two cases; one wherethe permeability is the same inside �kinside� and outside the skirt�koutside�, and the other one where the permeability inside the skirtis 2.5 times higher than the permeability outside the skirts. As theunderpressure is increased, the gradients will increase and reducethe effective stresses inside the skirt toward zero. The gradientwhen the vertical effective stress is reduced to zero is termed thecritical gradient, icr. The critical gradient in soil with stresses fromthe soil only, will be icr=�� /�w. The stresses inside the skirt will,however, be influenced by the skirt friction. The magnitude of theskirt friction, together with permeability variations in the soil andthe skirt depth to diameter ratio, will therefore govern where the
By comparing the two plots in Fig. 3, it can be seen that thecase with higher inside permeability has higher gradients alongthe outside of the skirts and smaller gradients inside the skirtcompartment than the case with uniform permeability. The formerreduces the possibility of hydraulic failure inside the skirt com-partment when the effective stresses are reduced toward zero.
Diagram for Critical Underpressure
Erbrich and Tjelta �1999� used finite element steady state flowanalyses to develop diagrams showing the underpressure thatwould give a critical gradient in the sand as function of the ratioof skirt penetration depth to diameter, z /D. They expressed theunderpressure, �u, in normalized form as a “Suction Number,
Fig. 3. Equipotential lines for cylindrical anchors with z /D=0.45from finite-element analyses; �a� kinside /koutside=1.0; �b� kinside /koutside=2.5.
SN,” defined as
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SN = �H/z���w/��� = �u/�z��� �6�
where H=differential water head between outside and insideskirt; ��=submerged unit weight of sand; and �w=unit weight ofwater. The suction number for the differential water head thatcauses a critical gradient in the sand will be called the “criticalsuction number” and denoted SN,cr in the following.
Erbrich and Tjelta �1999� developed diagrams of critical suc-tion number as function of the ratio between skirt penetrationdepth and diameter, z /D, where the critical suction was definedby either �1� the average gradient between the skirt tip and the topof the sand plug being equal to the critical gradient; or �2� the exitgradient at the sand surface being equal to the critical gradient.The diagrams were established for cases with increased perme-ability in the entire sand plug inside the skirt and for cases withincreased permeability in a thin zone of 0.07 times the diameteradjacent to the inside skirt.
Several factors will influence where the critical gradient isreached first. It is believed, however, that the skirt friction willcause the critical gradient to first occur at the sand surface, closeto the skirt. A diagram of �SN,cr-z /D� for this condition was de-veloped from steady state finite element flow analyses for thecase with increased permeability inside the entire sand plug, aspresented in Fig. 4. The diagram is similar to the diagram pro-duced by Erbrich and Tjelta �1999� for the same condition, butextended to larger depth to diameter ratios to cover a wider rangeof geometries.
Maximum Allowable Underpressure
When the critical gradient is reached and the vertical effectivestress inside the skirt is reduced to zero, there is a potentialfor progressive development of piping and channeling to occur.Piping and channeling would hinder an increase in underpressureand inhibit further skirt penetration. There is also a potential forloosening of the soil, which in turn could reduce the density andincrease the permeability of the sand inside the skirt compart-ment. However, experiences from installation of skirts in proto-type and model tests have shown that gradients close to criticalcan be applied without significant detrimental consequences. Apossible reason is that the skirts may penetrate further and tend to
Fig. 4. Critical suction number, SN,cr, as function of depth to widthratio, z /D, and permeability ratio, kinside /koutside
reduce the gradients when critical gradients are approached.
The maximum driving force is obtained when the maximum un-derpressure is applied, that is the underpressure, ucr, that gives thecritical exit gradient. The maximum driving force is then
W� + ucrAtop = W� + SN,cr��z�D2/4 �7�
where Atop=area of the top plate inside the skirt compartment. Asthe wall thickness normally is small compared to the diameter, itis assumed that the inner and outer diameters are equal.
The penetration resistance for an underpressure giving a criti-cal gradient can be estimated by assuming that the inside skirtfriction and the tip resistance are reduced to zero when the criticalgradient is reached. With this assumption the penetration resis-tance is given by the friction along the outside skirt wall, which is
where ��v�=increase in vertical effective stress along the outsidewall due to the water flow. The increase in vertical effective stressat the outside skirt tip will be equal to the reduction in porepressure at that point, i.e., ��v�−�utip. Assuming that the porepressure change at the skirt tip outside and inside the skirt is thesame, the pore pressure at skirt tip is then given by
Assuming icr=1.0 and no head loss beneath the skirt tip maygive a penetration resistance on the outside skirt on the high side.On the other hand, the assumption of full loss of resistance alongthe inside skirt and at the skirt tip may give a resistance on thelow side. The two assumptions may therefore counteract eachother.
The penetration resistance then becomes
Foutside = 0.5K��z2�DSN,cr tan � �10�
and the skirt cannot be penetrated further when
Foutside W� + SN,cr��z�D2/4 �11�
i.e., when
W� �2Kz/D tan � − 1�SN,cr��zAtop
where
Atop = �D2/4 �12�
This equation can be used to check whether the skirt penetrationto a given depth is possible. The maximum possible penetrationdepth can be found by applying the equation for various depths todetermine at what depth the weight is no longer greater than�2Kz /D tan �−1�SN,cr��zAtop.
Interpretation of Data from Prototypes and ModelTests
In order to determine the effect of underpressure on skirt penetra-tion resistance, the penetration resistance measured at variousskirt penetration depths in prototypes and model tests penetratedby underpressure are plotted in Fig. 5. The diagram showsthe normalized skirt penetration resistance as function of the nor-malized applied suction number. The measured skirt penetration
resistance, Pf ,Flow, is normalized with respect to the calculated
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skirt penetration resistance for no pore water flow, Pf ,NoFlow. Thelabels at the data points indicate the ratio between skirt penetra-tion depth and skirt wall thickness, z / t.
The value of Pf ,NoFlow was determined by the bearing capacitycalculation model described earlier, using triaxial friction angle.The suction number, SN, is normalized with respect to the suctionnumber for the underpressure that gives critical gradient at the topof the sand plug inside the skirt, SN,cr. The critical suction numberwas determined from the diagram in Fig. 4 for given values of theratio between skirt penetration depth and diameter, z /D, and theratio between the permeability inside and outside the skirt,kinside /koutside. The difference in permeability was determined byassuming that the sand inside the skirt is loosened due to dila-tancy occurring in the sand plug due to shear strains caused bysand displaced by the penetrating skirts. The density reductionwas calculated with dilatancy parameters from drained triaxialtests, assuming all displaced soil moving into the skirt compart-ment and uniform shear strain in the sand inside the skirt. Theeffect of density on permeability was determined from Fig. 6.More details are given in the next section.
The diagram in Fig. 5 shows that the normalized penetrationresistance decreases with:• Increasing normalized suction number, i.e., with increasing un-
derpressure.• Decreasing aspect ratio �ratio between penetration depth and
skirt thickness, z / t�. Only Test PEN5 deviated somewhat fromthis trend. The aspect ratio plays a role because the relativecontributions of inside skirt friction, outside skirt friction, andtip resistance depend on the aspect ratio, and because thesecomponents will also be influenced differently by a givenunderpressure.The data in Fig. 5 are used as the basis to draw the empirical
contours presented in Fig. 7. These contours can be used tocalculate the penetration resistance under combined weight and
Fig. 5. Normalized penetration resistance as function ofunderpressure �normalized suction number� measured in prototypesand model tests
Interpretation of the prototype and model test data also sug-gested that:• The reduction in normalized penetration resistance as function
of normalized suction number is defined by one curve for eachz / t ratio, independent of the normalized weight.
• There are some tendencies for reduction in normalized pen-etration resistance with reduction in penetration rate until thepenetration rate is slow enough for steady state pore pressuresto develop.
• The density of the sand inside the skirt compartment is re-duced by the skirt penetration. This was found from soil heavemeasurements inside the skirts and, in the model tests, by per-meability measurements and cone penetration tests.
• Underpressures giving close to critical exit gradients wereapplied in the prototype cases. In the model tests the exitgradients were generally somewhat smaller than the critical
Fig. 6. Permeability measured in hydraulic permeability laboratorytests on two sands prepared at various densities
Fig. 7. Recommended diagram for penetration resistance as functionof suction number and aspect ratio
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value. Even if gradients close to critical were applied, exces-sive loosening of the sand or other detrimental consequenceswere not observed.
Calculation Procedure
Based on the above-described interpretation, the penetration re-sistance and the required underpressure to reach a given penetra-tion depth can be calculated as illustrated in Fig. 8 and describedin the following.1. Calculate the penetration resistance under no pore water
flow, Pf ,NoFlow. In principle, any reliable method can be used.2. Determine the critical suction number, SN,cr, by calculating:
• The volume change due to dilatancy caused by shearstrains from soil displaced by the penetrating skirts as�vol=−4�r sin � / �1−sin ��, where �r=−2t /D. The dila-tancy parameter can be determined from volumetric andaxial strains measured in appropriately consolidateddrained triaxial tests, preferably sheared in extension, assin �=−d�vol / �2d�a−d�vol�. sin � should be the secantvalue from initial conditions to a radial strain equal to theskirt compartment due to skirt penetration, �r=−2t /D. Asan approximation, one may assume �=32.80+0.66� basedon experience from in-house data at NGI.
• The relative density, Dr, inside the skirt compartment fromDr= �emax−e� / �emax−emin�, where e=e0+�vol�1+e0� voidratio and emax and emin maximum and minimum void ra-tios, respectively.
• The ratio of permeability inside and outside the skirt,kinside /koutside, by for instance entering a diagram as in Fig. 6with the calculated relative density inside the skirt compart-ment and the relative density before skirt penetration.
The critical suction number, SN,cr, can then be determined by en-tering the diagram in Fig. 4 with kinside /koutside from the last bulletin Step 2 and the depth to diameter ratio of the skirt, z /D, for the
Fig. 8. Illustration of procedure to determine normalized penetrationresistance and suction number
3. Determine whether the specified penetration depth can bereached, by checking that
W� �2Kz/D tan � − 1�SN,cr��zAtop, where Atop = �D2/4
�13�
4. Calculate:• The ratio between the submerged weight and the penetra-
tion resistance under no pore water flow at the given pen-etration depth, W� / Pf ,NoFlow, and
• The ratio Wsoil plug� SN,cr / Pf ,NoFlow=0.25�D2z��SN,cr /Pf ,NoFlow. The ratio represents the slope of the line that willbe followed in the diagram in Fig. 8 when an underpressureis applied inside the skirt.
5. Determine Pf / Pf ,NoFlow and SN /SN,cr by plotting a line withan intersection of W� / Pf ,NoFlow at the vertical axis and a slopeof Wsoil plug� SN,cr / Pf ,NoFlow in the diagram in Fig. 7, as shownin Fig. 8. This line represents the driving forces, andPf / Pf ,NoFlow and SN /SN,cr are defined by the intersection ofthis line and the curve with the z / t representative for thespecified penetration depth. With this information, the pen-etration resistance can be determined by multiplyingPf / Pf ,NoFlow with Pf ,NoFlow from Step 1, and the required un-derpressure can be determined by multiplying SN /SN,cr withz��SN,cr, using SN,cr calculated in Step 2.
The penetration resistance and the required underpressure canbe determined as function of depth by repeating the calculationsfor other penetration depths. The influence of the submergedweight on the penetration resistance and the required underpres-sure can be determined by repeating the calculations with differ-ent submerged weights.
Penetration Resistance in Sand with InterbeddedClay Layer
If there are continuous clay layers interbedded in the sand, thewater flow will be prevented and no reduction in penetration re-sistance will be achieved by applying underpressure. The effect ofthe underpressure will then be an increased penetration force.This increase in penetration force can be calculated as �uAtop.
Four examples where underpressure has provided an increasedpenetration force, with clay layers having prevented flow ofwater, are presented in Table 4. Table 4 gives the measured pen-etration resistance, Pf, and the measured underpressure �suctionnumber, SN�, normalized to the calculated penetration resistanceunder no flow, Pf ,NoFlow, and the calculated critical suction num-ber, SN,cr, respectively. Pf ,NoFlow is calculated by the bearing ca-pacity model, using the triaxial friction angle. There are someuncertainties about the exact layering and the friction angle in
Table 4. Normalized Penetration Resistance and Underpressure �Suction
Casez
�m�D
�m�t
�m���
�kN/m3��
�deg��
�deg�
Prototypes
Harding A7 7 5 0.035 10 44 39
Harding A8 7 5 0.035 10 44 39
Field model tests
Sleipner B2 1.2 1.5 0.012 11.2 42.5 38
Sleipner B4 1.0 1.5 0.012 11.2 45.5 41
these cases, but the data are believed to be representative.
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The underpressure required to penetrate the skirts is veryhigh �SN /SN,cr=5.7–8.7� and exceeds significantly the underpres-sure that will give critical gradients in the soil plug in sand with-out interbedded clay layers. Table 4 also shows that the weightcontributes very little to the penetration force �W� / Pf ,NoFlow
= �0.0–0.06�. These results confirm that clay layers may preventflow of water due to applied underpressure. They also show thatunderpressures significantly in excess of the underpressure thatwould cause critical gradient in the soil can be applied, providinga penetration force significantly greater than the self-weight. Incases of sand with interbedded clay layers, it may be difficult topenetrate the skirts through dense sand layers without providingpenetration force by means of underpressure, especially for largedepth to diameter ratios. It should be noted that the four cases inTable 4 did not have inside stiffeners. If the skirts are equippedwith inside stiffeners, it is uncertain whether drainage channelsmay form above the inside stiffeners, and how these may influ-ence the penetration resistance and the maximum underpressurethat can be applied.
Conclusions
Penetration Resistance under Self-Weight
Interpretation of skirt penetration data from prototypes and modeltests shows that the penetration resistance under self-weight orother external load can be calculated by a bearing capacity calcu-lation model using the drained triaxial friction angle and takinginto account the increase in skirt tip resistance from interactionwith the skirt wall friction. Equally good agreement with the mea-sured penetration resistances was not achieved when using bear-ing capacity formulation with estimated plane strain frictionangles. An empirical model, which relates the skirt penetrationresistance to the CPT resistance, gives a large range in the back-calculated skirt tip bearing capacity factor, ktip.
The main reason why the bearing capacity model gives moreconsistent agreement with the measured penetration resistancethan the empirical model does may be because the bearing capac-ity model includes interaction between the wall and tip resis-tances. The interaction may not be properly modeled in theempirical model because of differences in geometry and wallroughness between the CPT and the skirt, and different effectivestresses at failure beneath a skirt and the CPT. However, the bear-ing capacity model is sensitive to the friction angle and requiresdetermination of the relative density through CPT testing and em-pirical diagrams that correlate relative density and CPT resis-tance. Since none of the models are straightforward and withoutuncertainties, it may be worth applying both models in designanalyses.
er� in Sand with Clay Layers
� f Nq N� W� / Pf ,NoFlow Pf / Pf ,NoFlow SN /SN,cr
1 115 165 0.03 0.99 6.86
1 115 165 0.03 0.99 6.86
1 91.7 125 �0.0 1.01 5.71
1 146 221 0.06 1.02 8.66
Numb
K
0.8
0.8
1.1
1.6
The bearing capacity model is calibrated based on a friction
angle between skirt wall and sand of �=0.9� and a ratio betweenvertical normal stress increase and skirt wall friction at skirt tiplevel of � f =1. For situations with no interaction from neighboringstructures, an earth pressure coefficient against the skirt wall ofK=0.8 should be used. A higher value of K must be considered ifthere are other structures or constraints nearby or if the sand isoverconsolidated. As these parameters are empirical, they shouldbe used together with Nq and N� values determined by the expres-sions given in this paper and a triaxial friction angle determined ateffective stresses representative for capacity analyses �i.e., at avertical effective consolidation stress of 100–250 kPa�.
For the empirical model, the parameter ranges and combina-tions are:
ktip = 0.01 – 0.55 when kside = 0.0015 and
ktip = 0.03 – 0.60 when kside = 0.0010
when the wall resistance is related to CPT resistance. When thewall resistance is based on effective stresses and friction angle,the parameter range is ktip=0.07–0.47.
The models are calibrated by backcalculation of prototypesand model tests in dense sand and should be used with caution ifthe sand is not dense.
Penetration under Combined Weightand Underpressure
The data from skirts penetrated by combination of weight andunderpressure show that the reduction in penetration resistance ismore significant with increasing underpressure and decreasing as-pect ratio �ratio between skirt penetration depth and skirt thick-ness�. The penetration resistance as function of underpressure isdefined by one curve for each aspect ratio, independent of nor-malized weight.
The data indicate that skirt penetration causes a reduction inthe density of the sand inside the skirt compartment. This gives ahigher permeability in the sand inside the skirt compartment thanoutside. The density reduction is mainly due to dilatancy causedby shear strains in the sand plug from the sand displaced by theskirts. The density reduction can be calculated with dilatancy pa-rameters from triaxial tests.
Underpressures giving close to critical exit gradients wereapplied to reach target penetration depth in the prototypecases. Even if gradients close to critical were applied, excessiveloosening of the sand or other detrimental consequences were notobserved.
In sand with interbedded clay layers, the interpretation showsthat underpressure can be used to provide significant drivingforces, but in this case reduction in penetration resistance cannotbe relied upon.
Future Studies
The procedures to calculate the penetration resistance under self-
weight and underpressure are both based on empirical data and
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may be improved and extended to sand of lower relative densityas more case histories become available. The effect of skirt wallroughness may also deserve further study.
Acknowledgments
The writers would like to thank Statoil, Chevron Petroleum Tech-nology Company, Conoco Inc., Petrobras, and Saudi Aramco,which sponsored the joint industry project within which thework described herein was performed. Statoil and BP are alsoacknowledged for their release of skirt penetration data. Contri-butions from Tor Inge Tjelta of Statoil in his active involvementto initiate the project and as chairman of the steering committee,Per Sparrevik of NGI regarding information and discussions ofthe model tests at NGI, and Niels Mortensen of NGI regardingdiscussions of calculation procedures and ring shear tests to de-termine interface friction between skirt and sand are also greatlyappreciated.
References
American Petroleum Institute �API�. �1993�. “Recommended practice forplanning, designing and constructing fixed offshore platforms.” APIRP 2A, American Petroleum Institute, Washington, D.C.
Andersen, K. H., Murff, J. D., Randolph, M. F., Clukey, E., Erbrich, C.,Jostad, H. P., Hansen, B., Aubeny, C., Sharma, P., and Supachawarote,C. �2005�. “Suction anchors for deepwater applications.” Proc., Int.Symp. on Frontiers in Offshore Geotechniques (ISFOG), KeynoteLecture, Perth, Western Australia, 3–30.
Clausen, C. J. F. �1998�. “Fundamentering av plattformer; Observasjonerog refleksjoner.” 16th Laurits Bjerrum Memorial Lecture, NorwegianGeotechnical Institute, Oslo.
Dendani, H., and Colliat, J. L. �2002�. “Girassol: Design analysis andinstallation of the suction anchors.” Proc., Offshore Technology Conf.,Houston, Paper No. 14209.
Erbrich, C. T., and Tjelta, T. I. �1999�. “Installation of bucket foundationsand suction caissons in sand—Geotechnical performance.” Proc., Off-shore Technology Conf., Houston, Paper No. 10990.
Houlsby, G., and Byrne, B. W. �2005�. “Design procedures for installa-tion of suction caissons in sand.” Geotech. Eng., Proc., Institute ofCivil Engineering, London, 158�GE3�, 135–144.
Houlsby, G., Ibsen, L. B., and Byrne, B. W. �2005�. “Suction caissons forwind turbines.” Keynote lecture, Proc., Int. Symp. on Frontiers inOffshore Geotechniques, Perth, Western Australia, 75–94.
Johansson, P., Aas, P. M., and Hansen, S. B. �2003�. “Field model testsfor a novel suction anchor application.” Proc., 6th Int. Symp. on FieldMeasurements in Geomechanics, Oslo, Norway.
Jostad, H. P., and Andersen, K. H. �2006�. “Potential benefits of usingskirted foundations for jack-up platforms.” Proc., Offshore Technol-ogy Conf., Houston, Paper No. 18016.
Sparrevik, P. �1995�. “Development of new platform foundation conceptthrough instrumentation.” Proc., 4th Int. Symp. on Field Measure-ments in Geomechanics, Bergamo, Italy, Paper, 137–144.
Tjelta, T. I. �1995�. “Geotechnical experience from the installation of theEuropipe jacket with bucket foundations.” Proc., Offshore Technology
