Shaft capacity of drivenpipe piles in clayby ROBERT M. SEMPLE", BSc, MSc, PhD, CEng, MICE 8W. JOHN RIGDEN>, BSc, MSc, CEng, MICE

IntroductionTHE DEMANDS OF offshore constructioncontinue to focus attention on improvingaxial capacity prediction for driven pipe piles.Considerable research efforts have beenmade over the last 10-15 years involving soilmechanics theories, model and field scalepile testing. Recently, the AmericanPetroleum Institute (API) sponsored a two-year project, undertaken by Olson and hisco-workers at the University of Texas atAustin, to establish a data bank of static pileload test records. Existing data werethoroughly examined and comprehensivelydocumented in that well conceived andexecuted study".

Application of the test data in offshoredesign requires extrapolation to the pile sizesand to some of the soil conditionsencountered. Accordingly, capacity criteriaderived from pile load test results shouldreflect sound physical principles. This Paperpresents new criteria for skin friction in clayinterpreted from the API data base.

Analytical developmentsSimple effective stress analyses are well

represented by the proposals of Burland'ndMeyerhoPI wherein skin friction is related tothe effective overburden pressure, 0„, by aparameter P that incorporates the frictionalcharacteristics of the soil and a coefficient ofearth pressure at failure, K,. As discussed byRandolph4', Meyerhof's recommendation forK„derived from soil mechanics theory andpile load test data, are well supported byresults of recent high quality model tests'4.

More sophisticated effective stressmethods based on critical state concepts andcavity expansion theory"" have contributedsignificantly to understanding while not fullyexplaining test observations (e.g.").Recentanalytical developments'ave includeddetailed consideration of likely strain pathsaround the pile tip during installation, andthis may increase the reliability of theoreticalprediction.

Initial attempts to generalise pile load testinformation used the most obvious soilcharacteristic, undrained shear strength, s„.The a coefficient, defined as the fraction of s„mobilised as skin friction, has generally beencorrelated with undrained shear strength.However, McClelland" noted that a decisionwas taken in early Gulf of Mexico offshorepractice to relate a to the degree ofoverconsolidation of the soil. This decisionrecognised that in existing a—s„relation-

'Director, McClelland Ltd., McClelland House, ChantryPlace, Headstone Lane, Harrow, Middlesex.>Manager, Cwil and Geotechnical Branch, CentralEngineenng Department, BP International Ltd., London.This Paper was presented at the ASCE annual conventionheld in October 1994 in San Francisco. It was included in

an ASCE special technical publication entitled "Analysisand Design of Piled Foundations", published by TheAmencan Society of Cwil Engineers, 345 East 47th St.,New York, New York 1OOI 7-2398.

Symbols used in this Paper

D = pile outside diameter

K,

modified length factor

lateral earth pressure coefficient

k = constant of proportionality

L = pile embedded length

I F = length factor

OCR = overconsolidation ratio

Pl = plasticity index

Qs, = calculated shaft capacity

Q, = measured shaft capacity

s„=undrained shear strength

t = unit skin friction

z = relative soil-pile displacement

a = skin friction coefficient = t/s„

P = skin friction coefficient = t/o„

173 ——pile-soil stiffness ratio

0„=vertical effective stress

= angle of internal friction

ships, which showed a decreasing withincreasing s„the shear strength values had aclose positive correspondence with the over-consolidation ratio, OCR, of the soils. Hencea was taken as unity for normally con-solidated clays, regardless of s„and smallera values were used for overconsolidatedclays as independently suggested byWroth". For many areas of offshoredevelopment outside the Gulf of Mexico, thisconcept was effectively lost in the mid-1970's when the American PetroleumInstitute incorporated an a—s„correlationinto RP2A'.

Semple" estimated OCR values for somepile load test sites and was able to show thatmeasured a values could be related to OCR.Randolph Er Wroth4'onverted Meyerhof'ssemi-empirical effective stress procedure"into curves of a versus the strength ratio s„/0„which is related to OCR but can be deducedmore directly from site investigation data.Comparison of these a—s„/o„curves with piletest data'ndicated reasonable agreementfor relatively short piles but that scale effectsdue to pile length and flexibility" 'houldalso be considered. The relationship of a tos„/o„and pile scale is considered furtherherein.

A true representation of offshore pileresponse should recognise time varying loadeffects on soil resistance'nd performancecriteria". However, this Paper considers onlythe static capacity of pipe piles as it isassessed for conventional steel framedoffshore structures.

Recent pile testingResearch in Europe and the USA has

tended to focus on cyclic loadresponse"""". The most recent statictests performed on relatively large pilesections, at Empire, La. and Long Beach,Calif.'37, provided important newinformation for data base interpretation. Theindustry-sponsored "ESACC" project""stimulated static testing of instrumentedpiles at a University of California test site",and at model scale at Cambridge

University'ecent

data interpretationsWorking with the data base they compiled

for API, Dennis & Olson'eveloped an a—s,correlation in which a reduced from 1 to 0.3with increasing shear strength. A length-related correction factor was formulated toaccount for the capacity of long piles in (stiff)normally consolidated clays beingunderestimated by this a—sa correlation.

Randolph" found that the a valuescorrelated well with the strength ratio sJI7„ingeneral accordance with the conceptpublished earlier by Randolph Er Wrothx'. Forstrength ratios up to unity, Randolph foundthat a = k(s,/17„)-'ith k = 0.5 provided agood average fit to the test data. Randolphfurther generalised the relationship bydefining k as the square root of the strengthratio for normally consolidated soil. Therelationship was later expanded to includestrength ratios greater than unit44.

Randolph" also noted a possible effect ofpile length, with a tendencyfor longer pi(estohave lower capacities than computed fromthe average a—s,/17„correlation. Herecommended that the pile length effect beassessed separately, considering pilecompressibility and likely residual values ofskin friction, as described by Randolph4'.

New approachIn assessing prospects for improving pile

capacity predictions for long offshore piles,Focht Et Kraft" recommended that theproblem be conceived as having twocomponents. The first component is the loadtransfer response of a soil element ascharacterised by its local peak and residualskin friction and associated displacements,commonly termed t—z curves. The secondcomponent involves the integration of thislocal soil response over the pile length withdue regard to pile compressibility and theresulting variation in load transfer betweenpeak and residual values at various points onthe pile shaft. Focht Er Kraft suggested as anobjective the development of an analysisbased on t—z soil response curves in whichthe value of peak skin friction is predicted byapplying soil mechanics theory to assessingstress changes in soil elements.

Peak skin friction for a soil profile is thevalue that would be deduced from a load teston an incompressible pile. Accordingly, oneapproach to interpreting the API data base is

January 1986 11

first to distinguish peak skin friction asmeasured on relatively rigid piles and then toidentify appropriate adjustments for longer,less stiff piles.

Pile test resultsAlthough the 1982 API data base~

contains results from over 1 000 pile loadtests, only a fraction of these are from drivensteel pipe piles in predominantly cohesivesoil profiles. The data were quality gradedusing a five point scale and the lowest qualitydata excluded from further consideration bythe compilers in developing correlations'.The same approach has been taken in thisstudy.

Table I lists data from 24 sites that we haveselected for analysis. The data are given inorder of increasing strength ratio, sJo„,which are averages for the embedded pilelengths. The identifying Load Test Numbers(LTN) from the API listing are indicated.Several of the data lines in Table I areaverages of more than one test pile at a site.This averaging was performed to avoidclutter on data plots and giving undue weightto essentially repetitive results from a soil-pile condition.

The second from last data line refers toLTN 860 in the API listing. However the piledata on Table I were taken from Rigden etaf"who presented results for a driven pipe pile

which are considered more relevant than thejacked pile data from the same site given inthe API listing. All the piles represented in

Table I were installed by driving. Most avalues are from compression tests afterallowing for end bearing which wascalculated to be small for these friction piles.Unit end bearing was taken as 9 x s„at thepile tip.

Source documents for the data areidentified on Table II. Inspection of thedocuments indicates that a few of the soilprofiles had a relatively thin, surficial sandstratum that would have made a modestcontribution to the measured axial capacity.Unit skin friction in these surficial sands wastaken as 0.4o„, and the pile capacityattributed to cohesive resistance wasadjusted accordingly. Only in one case didthe calculated adjustment approach 10%ofcapacity, adjustments for the few remainingcases being less than 5%.

The pile lengths in Table I are thoseembedded in clay beneath surficial sand,excavation or casing over the upper part ofthe pile where these occurred.

Soil shear strengthUndrained shear strengths are given in

Table I. Sampling quality was generally goodwith thin-walled pushed samplers beingused at almost all sites.

Table II indicates how the reference soilstrength was measured. The most commontest type was unconfined compression sothis value was usually selected from the APIlisting even if other measurements wereavailable. Exceptions were made for very softand soft soils (s„(40kPa) for which vanestrengths were selected, where available, assuch soils are too weak for reliablecompression testing. Vane strengths werenot available for the first three data lines in

Table I; however the soil testing appears tohave been performed carefully on 125mmdiameter piston samples. After reviewingeach source document, the strength profilegiven therein was sometimes preferred to theinformation in the API listing (LTN 106,487-491, 495). Soil unit weight data for LTN 23were provided by BP for whom the tests wereperformed.

Pile geometryPile embedment length, L, has been

expressed as a multiple of pile diameter, D, in

Table I. The simple aspect ratio L/D cannotrepresent all the variables influencing theeffects that pile geometry and stiffnesscharacteristics may have on axial capacity.The pile scale or "length effect" f2 2'3' isprobably related to pile stiffness as it affectseither lateral whip during driving orcompressibility under axial loading. If pile

TABLE I: SUMMARY OF PILE LOAD TEST RESULTS TABLE II: SOURCES OF PILE TEST DATA

Load D, in

Test L, in milli-

Number metres metres

(f) (2) (3)L(D

(4)

s„, in o„, in

kilo- kilo-

pascals pascals(5) (8)

spa, a(7) (8)

LoadTest

Number

(f)

ShearStrength

Test

(2)Source reference

(3)

86,20

3,7,17478,489491,493854,855

86886987345142

444,450507,508

3015045

844,846848.851

8563256743

443,449368,369435.436437,438

70998106

547,54931,32

829,830495,497

23,24

20.4 762 2721.6 457 4719 2 610 321 5.2 356 4312 2 356 3443.9 305 14496.0 610 15873.8 610 12122 6 767 3066.4 325 20530.5 325 9445.7 325 14229.0 330 8813 7 325 4218 3 325 5748 2 610 79

303131

104162388067

17060455239453364

144147142448718162354273651223153148105112

51152

0.21 0.920.21 0.930.22 0.990.23 1.050.23 1.000.23 0.790.23 0.550.25 0.710.26 1.130.27 0.520.29 0.650.35 0.420.37 1.020.40 0.940.43 0.970.43 0.59

11.6 114 101 21 44 047 064

12.2 168 7214 0 351 4039.6 274 14530 5 610 5022 9 325 7125.9 325 80

1630

165525261

3359

297919199

0.50 0.620.51 0.780.56 0.490.57 0.590.57 0.520.62 0.56

14.9 528 2832.0 274 1 1 71 2.8 325 4016 8 610 281 3.7 325 4213 1 274 4820.4 610 34

9.1 450 2018 3 762 24

5311596

100137110208144335

66141110

8711280

10554

115

0.79 0.520 82 0.520.87 0 571.15 0.551.22 0.471.37 0.491.98 0.442.65 0.512.90 0.46

25 3 274 92 185 244 0 76 0 48

86,20

3,7,1 7487,489491.493854,855

86886987345142

444,450507,508

3015045

844,846848,851

8563256743

443,449368,369435,436437,438

70998106

547,54931,32

829,830495,497

23,24

U

U

U

U

U

V

U

U

U

U

U

U

U

U

U

U

M

FV

U

U

U

U

U

U

M

U

U

0M

M

U

Mansur B Focht"Mansur & Focht"Mansur Ef Focht"Cox, Kraft Ef

Verner'ox,

Kraft EfVerner'onfidential

ConfidentialConfidential

Pelletier Ef Doyle"Peckss

Darragh EfBell'eck"

Raymond"Woodward, Lungren Ef Boitanosx

A.R.E.A s

McCammon Ef Golder

Kirby Ef Rousselia

Confidential

Hutchinson B Jensen"Peck"McCammon Ef GolderPeckPeck"

Peck"

Tog rolesEndley, Ulrich Ef Gray"Stermac, Selby 8 Devatta"U.S.A.C.E.Woodward, Lungren Ef Boitanosx

O'eill, Hawkins Ef MaharHeerema"Rigden er a('"Fox, Sutton Ef Oksuzler"

Note: L = Embedment length;D = Outside diameterlm = 3.28ft; 1mm = 0.039in; 1 kpa = 20.9psf

Note: U = Unconfined compressionQ = Quick tnaxialV = Laboratory miniature vane

FV = Field vaneM = Other test

12 Ground Engineering

whip controls, then the ratio of the travellingstress wave length to pile diameter should beconsidered~. Another reason for the "lengtheffect" may be progressive failure of strain-softening soil as the pile shaft compressesunder axial load. From this viewpoint, it is therelative pile-soil stiffness in axial loading thatis important.

Relative axial pile-soil stiffness has beentermed rrs and defined by Murff" as the ratioof pile elastic compression, acting as a free-standing column, to the local soildisplacement required to mobilise peak skin

friction. If the displacement corresponding tothe limit of elastic soil response is related tothe pile diameter, then rr, for an open pipepile is a function of the aspect ratio L/D, peakskin friction, and the pile area ratio (steelsection as a fraction of gross sectional area).Although pile area ratios are known for mostof the data in Table I, being on average about10%, there is uncertainty over which of thepipe piles were filled with concrete prior totesting thereby altering their axial stiffness.Further, the relevant peak skin friction valuesare unknown so rr, values cannot be

confidently determined.The ratio LJD only approximately

represents pile geometry and stiffnesscharacteristics. It is an expedient that can beconsidered as replacing more complexexpressions for flexural and axial pilestiffness while reflecting more of thesefactors than does pile length alone.

Conventional a correlationThe data from Table I are plotted as a

versus sv on Fig. 1. in which the L/D value isgiven for each data point. Criteria

1.2

c

8 1.0c08

0.6

0.4

32 O88%%5747'4227

~43

40O

~121

$01

7250 O79 4028 o80 ~ oo28

710O i 158 O117205 ~

48~142

~LID

~30

~34

92

20~ ~145

~API24

34 S =335U

0.4. O142

1.6+c

43 308 32~ o

88o42 g LID

0.8 0144 ~40~121

~94 101O 72 50~158 79 ~ ~~ ~ 40

~205 ~48145 O92 42O~20

~34 24

0.2.

00 50 100 150 200 250

Undrained shear strength, IrPa

Fig. 1 (aboveJ. Alpha vs. undrained shear strength

Fig. 2 (right, topj. Alpha vs. strength ratio (all dataJ

csl0

0.20.2

1.6

0'5 08. 27

0.4

F057

0.8 1.6 3.2Soil strength ratio,S la

Fig. 3 (right, central. Alpha vs. strength ratio (L/DC6DJ

Fig. 4 (right, bottomJ. Length factor vs. aspect ratio0.4-

48

42 34o

20O

24O

Fig. 5 fbelovvj. Comparison of measured and calculated capacities

g0

10-8

0.1-

0.20.2

3.20

E

u.

8 1.6-

Cl

I.0~ 0.8-

E00

0.4-

0.4 0.8 1.6 3.2Soil strength ratio, S„/6„

C = Oversized dosure plate

R = Redriven before test

~ oo~yg ff~ ~' ~

w~ Sc.~Oi 55Oc w +c

RO-

R~ OCR

CR

0.1 10Measured capacity, Q5~IMN

0.210 20 40 160 320

Pile aspect ratio. LID

January 1986 13

F 1.0-

<rmax

0.2

Fig. G. Theoretical length effect

log rrs

recommended by API're represented by thesolid curve. The data conform quite well tothe API criteria with the exception of thethree points for stiff clays having a values of1 or more. These data represent the sites atEmpire and Long Beach"'here pilesections were installed through casings setsufficiently deeply into normallyconsolidated clays that the soils tested wereof a stiff consistency.

The Empire and Long Beach tests wereperformed after the API a—s„criteria wereformulated. Their inclusion on Fig. 1

indicates that s, is not the variable controllinga. Further, this difference in a values for stiff,normally consolidated and stiff, overcon-solidated clays underlay the introduction of apositive length-related adjustment factor in

the a—s„ formulation of Dennis Et Olson'.

Peak a valuesThe pile test data are presented on Fig. 2 as

a versus the strength ratio spa„, usinglogarithmic scales, and again with values ofL/D indicated. Inspection discloses a patternin the L/D values with the relatively longerpiles generally having smaller a values for agiven strength ratio.

Peak skin friction is measured by stiff pilesso the data for the relatively shorter piles,including those tested at Empire and LongBeach, are plotted separately on Fig. 3. Thescatter is small, and curves were readily fittedto the data by separate regression analysesfor strength ratios greater than unity and inthe range 0.35to 1.For strength ratios belowabout 0.35, a value of unity was selected fora as these soils are essentially normallyconsolidated, s„ increases with depth, andthe strength ratio tends to be constant overthe pile length. In soil profiles having greateraverage strength ratios, the s„profile is moreuniform, and the strength ratio varies overthe pile length with relatively large valuesoccurring near the ground surface.

The curve on Fig. 3 is labelled aato suggestthat the relationship is for peak skin friction.The curve actually represents piles havingL/D = 40+20, and so is likely to be a lowerbound to peak a values.

In this formulation, the strength ratio is ameasure of soil overconsolidation. Noattempt has been made to normalise thestrength ratio to account for possible effectsof variation in basic soil plasticitycharacteristics as expressed by plasticityindex, Pl, or the associated angle of internalfriction, gi. One possibility is to relate eachstrength ratio to the value for the same soil in

a normally consolidated condition, (s„/o„)„,.There is, however, conflicting information onhow this value varies with Pl or lg" e, whilethe normally consolidated strength ratioobtained from compression test data can beindependent of Pl". Given this uncertainty,the term s„/o„has been taken to representsoil overconsolidation without furthergeneralisation.

The correlation between a and strength

ratio on Fig. 3 bears some resemblance inform to the existing API a—s, criteria. Itreflects conventional wisdom that a is unityfor normally consolidated clays and 0.5 forheavily overconsolidated clays.

Pile length effectHaving established a correlation for peak

skin friction from the relatively shorter piles,the effect of greater pile embedments hasbeen evaluated borrowing from thetechnique used by Dennis Et Olson'. For eachstrength ratio in Table I, a value of a, wasobtained from the correlation on Fig. 3, andthis was used to normalise thecorresponding measured value a . Theresulting ratios of a„/a are shown on Fig. 4as a function of the pile aspect ratio, L/D.Approximate bounds to the data areindicated together with an average lineobtained by regression analyses.

The average value of unity for a /a~ for therelatively shorter piles is of courseconditioned by the fact that these piles(average L/D = 40) were used to define ap.Values of ct /a for relatively longer piles fallprogressively below unity. The data scatter isdue partly to the limitations of L/D in

representing pile stiffness, and to otherfactors that indicate a "length effect" forindividual test piles. These factors arediscussed subsequently.

Adopting the term LF, or length factor, fora /a~, the solid line between the data boundson Fig. 4 provides a pile embedment factor tobe applied to the a, values from Fig. 3. Theaverage unit skin friction for a pile is thenLFa~„where s„ is the average value over thepile length.

EvaluationComputed values of pile shaft capacity,

Qac, based on LFa~„are compared with themeasured values, Qa„, on Fig. 5. Agreementis good, the data falling within a fairly tightscatter band. The average ratio of measuredto computed pile capacity is 1.01 with astandard deviation of 0.15.

The results on Fig. 5 indicate that thea—s,/rT„—L/D correlations provide a good fit tothe pile test data which supports anapproach based on peak skin frictionmodified, where appropriate, to account forpile scale. Examination of the pile load testdocuments does, however, indicate certaindeficiencies in execution of some tests thatwill have biased the data towards low avalues. These effects were discussed by Kraftet al" who adjusted their data interpretationto minimise this bias.

Length factor adjustmentOlson Er Dennis" note that piles redriven

and reloaded to failure exhibited capacitieslower than on initial loading. They alsorecognised the harmful effect on skin frictionof an oversized closure plate at the pile tip.The data to which these considerations applyare identified on Fig. 4. With one exceptionthey are all relatively long piles so the bias isincorporated into the pile length factor, LF. Ofthe six LF values less than 0.7, all but one ofthe piles were either redriven or had anoversized closure plate, or both as in the caseof the two smallest values. As evidence of LFless than 0.7 derives from particularlyadverse circumstances, we consider an LFvalue of 0.7 to be a reasonable lower limit. Itshould also be remembered that the ci,correlation on Fig. 3 was derived from pileswith L/D = 40+20, thereby incorporatingsome scale effect and providing a lowerbound to peak a values.

Modified relationshipsOn Fig. 3, the best fit line to the heavily

overconsolidated clays has a slight negativeslope that is largely conditioned by the twosmallest a values in Table I. In both cases thepiles were tested shortly after installation.Further, the result from the most heavilyoverconsolidated soil in the data base wasobtained after several cycles of redriving andtesting. Given these unfavourablecircumstances, the adoption of a minimumvalue of 0.5 for a~ appears reasonable. Themaximum value of unity for a~ applies toessentially normally consolidated clayswhich may be defined as those having astrength ratio less than 0.35.

For predicting pile capacity, we believe thatthe best-fit a—s,/rT,—L/D correlations in Figs.

o 1.6

c6

o 0.8.c05

V

0.4.CL

(1.0,0.35)

(0.5, 0.8I

0.20.2 0.4 0.8 1.6 3.2

Soil strength ratio, S„lo„

1.6

05

S 0.8-cI0.7, 120)

0.420 40 80 160 320

Pile aspect ratio LID

Fig. 7. Criteria for capacity prediction

If the empirical evidence for the pile lengthfactor is attributed to relative pile-soilstiffness effects in axial loading, then theexpected shape of the relationship betweenLF and L/D can be inferred from theory. AsL/D is related to the pile-soil stiffness factortt3 the characteristic shape of the theoreticalLF—rr, relationship shown on Fig. 6 wouldsuggest a lower limit for LF corresponding tothe ratio of residual to peak skin friction.From such considerations,

Randolph'uggestedthat the measured capacity of ahighly compressible pile could be 70%of thecapacity based on peak skin friction.

It is not known, however, that the "lengtheffect" is entirely due to progressive soilfailure in axial loading. Pile whip duringdriving may cause irrepairable damage to thesoil, particularly in overconsolidated claysthat cannot flow back to establish a pressureon the pile wall. This effect may partly explainthe reduction in a, with increasing strengthratio. In addition, the nature and magnitudeof lateral pile oscillations during drivingdepend on, among other factors, the lengthsof the stress wave and the pile. Someprogress has been made in quantifying thisbehaviour . However, an analytical modelproviding insight into pile whip effects,analogous to that for axial pile-soft stiff-ness", remains to be developed. Pile andstress wave lengths may combine to providea skin friction reduction effect related to aflexural stiffness ratio for which the pileaspect ratio, L/D, is an approximation.

14 Ground Engineering

00

c0

c~ 25

50 100Undrained shear strength, kPa

150 200 250

c0Es

ecIllc.

Ultimate pile capacity. MN

0 10 20 30 40 0 20 40 60 600

25-

50.

75.75-

100.100.

125125-

150

150

Fig. 8. Strength profiles for normally consolidated clays offshore Fig. 9. Compressive pile capacities for normally consolidated claysLouisiana offshore Louisiana

3 and 4 should be modified to the formsshown on Fig. 7. The main alteration isconstraining the length factor, nowdesignated F, to have a minimum value of0.7. This modification avoids theunderestimation of capacity for very long,slender piles that otherwise would result ifthe a values measured for such piles underunfavourable circumstances were taken atface value. Calculated shaft capacities basedon unit skin friction equal to Fa~„with Fand

a~ from Fig. 7, still correlate well with themeasured values.

Application to designThe relationships given on Fig. 7 may be

used in predicting pile capacity. Due to thenature of the information available, the skinfriction correlation and scale factor werederived using average results for each piletest. An incremental pile capacity method, inwhich the soil profile is considered layer bylayer, is ultimately preferable but itsformulation must await the development ofan appropriate data base. Layer-by-layerapplication of the criteria on Fig. 7 wouldrepresent a misapplication of the existingtest data, and would generally invoke errorson the conservative side.

Standard offshore practice in many areasis to measure s, by quick triaxial tests onsamples recovered by pushing a thin-walledsampler. The criteria developed herein arebased on high quality pushed samples, butthe reference undrained shear strength wasmost commonly measured by the unconfinedcompression test. However, this test is notappropriate to stiff, structured clays forwhich quick triaxial tests are necessary. In

other clays, unconfined and quick triaxialcompression tests give comparable resultswhere sampling disturbance is minimised.Accordingly, we recommend that thereference shear strength be taken as that of apushed sample tested in triaxial compressionat natural water content. Where othersampling and testing techniques are applied,equivalent soil strengths can be obtainedusing adjustment factors where these areavailable from correlation studies of similartype soils (e.g.')

If sand is present in a soil profile, the mostrational approach is to apply the criteria on

Fig. 7 using average values of s„and a,for theclays, and to select the scale factor using thetotal embedded pile length. This gives theaverage skin friction in clay from which theshaft capacity in clay can be determined. Theshaft capacity in sand can be separatelyassessed in the usual way and added to givethe total shaft capacity. At present, pilelength effects in sand are broadly accountedfor in offshore practice by the use of limitingskin friction values.

The prospects for improving pile design byappropriate site-specific testing areincreased by finding that an empiricalapproach based on peak skin frictionadjusted for pile scale effects accuratelyrepresents existing pile capacitymeasurements. Peak skin friction can bemeasured either by testing a pile segment or,more economically, by the use of precisionin-situ instruments4 calibrated against piletest measurements. If site-specific data areavailable then analyses incorporating pilestiffness characteristics can be used in anindependent assessment of likely scaleeffects.

Comparison withcurrent practice

The criteria on Fig. 7 have been applied tothree representative offshore sites forcomparison with capacity predictionsobtained from the current API methods. TheAPI methods used are those given in

Paragraph 2.6.4b of the 1982 edition ofRP2A'. Method 1, commonly termed "f=

d'or

skin friction equals undrained shearstrength), is restricted to highly plastic clayssuch as are found in the Gulf of Mexico.Method 2, for other clays, corresponds to theline designated API on Fig. 1.Normally consolidated clay

Site investigation practice in the Gulf ofMexico has involved percussive samplingwith miniature vane or unconfinedcompression testing to determine shearstrength of clays. The criteria on Fig. 7 areintended for use with quick triaxial strengthsfrom pushed samples. Quiros et al'avepresented a comprehensive set of soilstrength data to about 150m penetration in anormally consolidated clay at a locationoffshore Louisiana which they designate

"Site 2". Undrained shear strength profilesappropriate to the API method and to thecriteria on Fig. 7 are shown on Fig. 8. Thequick triaxial strength profile on Fig. 8 isessentially the same as the interpreted shearstrength profile for the site given in thesource publication.

Compressive capacity curves developedusing API Method 1,the criteria on Fig. 7, andthe appropriate shear strength profiles onFig. 8 are presented for two different pilediameters on Fig. 9. Unit end-bearing wastaken as 9 x s„at the pile tip. The curve for the1.83m OD pile attributed to the API methodagrees with the corresponding pile capacitycurve given by Quiros et al".

The a—s„/o„-L/D criteria produce thegreater capacities on Fig. 9. In percentageterms, the differences are greater for thelarger diameter pile which is less affected bythe length factor in the new procedure. Atpenetrations of interest for the 1.22m ODpile, the proposed method gives predictedcapacities about 10% greater than the APImethod, producing a small reduction inrequired pile length of the order of 5%. Thecorresponding increase in predicted capacityfor the 1.83m OD pile is about 25% whichreduces the required pile penetration byabout 10%.Heavily overconsolidated clay

Fig. 10 gives a shear strength profile forthe Heather Field located about 100km eastof the Shetland Islands in the North Sea.Quick triaxial shear strength and verticaleffective stress profiles were obtained fromDurning & Rennie". The other curve shownon Fig. 10 represents the average value s,/o„to a given depth. These silty clays are heavilyoverconsolidated and particularly strong, theaverage s,/a, and s„values for the profilebeing about 2 and 500kPa, respectively.

For the 1.83m (60in) OD piles installed atHeather Field, the criteria on Fig. 7 yield an avalue of 0.5 which is also the value given byAPI Method 2 commonly used for North Seainstallations. In hard, heavily over-consolidated silty clays the a—s,/o„—L/Dcriteria proposed herein will generally givethe same pile capacity as the existing APImethod.Overconsolidated clay

North Sea cohesive soil profiles are

January 1986 15

commonly less heavily overconsolidatedthan at Heather Field. Fig. 11 gives quicktriaxial shear strength profiles for twopushed sample borings at Magnus Field4',which is located about 200km northeast ofthe Shetland Islands. A thin sand stratumoccurred in the upper part of the profile atthis site. Also indicated on Fig. 11 are curvesof the average s,/o„values that reduce toabout 0.5 to 0.6at the depths of interest. TheMagnus piles were 2.13m OD and the targetpenetration was about 80m4'.

As shown on Fig. 12, pile capacity curvesfor the profiles derived using API Method 2and Fig. 7 are the same to about 60mpenetration, both methods giving a = 0.5 tothis depth. At 80m penetration, the capacitypredicted by the a—s,/B,-L/D method isabout 30% greater than the API value whichagain corresponds to a =0.5.The associatedreduction in required pile penetration isabout 10m, or 12% of the embedmentlength required by the API method.

Although similar shear strength profileswere obtained from the two borings, thereare differences which are reflected in the pilecapacity predictions for each method.However, the range in pile capacity isreduced using the new criteria. Of the twoprofiles, Boring 11 gave the greater averageshear strength. This is compensated by alower a value, as the sJa„ratio is also greaterfor Boring 11.The Magnus example demon-strates the reduced sensitivity to undrainedshear strength, and hence to measurementerrors, in stiff overconsolidated clay. This canbe deduced from inspection of the a—sJa„relationship on Fig. 7.

ConclusionsMeasured shaft capacities of driven pipe

piles in clay represented in the 1982API database can be reproduced quite accurately bycorrelations based on the concept of a peakskin friction reduced to account for pile scaleeffects. Peak skin friction is obtained from acorrelation between a and the strength ratios,/a„which indicates the degree of soiloverconsolidation. The reduction factordepends on pile aspect ratio, L/D. It

represents the influence of pile stiffness in

Fig. 10 (below, left).Soilinformation forHeather Field, NorthSea

00

c0

BI

20-

100

Bl

Fig. 11 (right). Soilinformation forMagnus Field, NorthSea

40-

Fig. 12 (below,right). Compressivepile capacities forMagnus Field, NorthSea

controlling the degree to which eitherprogressive failure can develop in staticloading or the soil is damaged by pile whipduring driving, or both.

Interpreting pile test data within aframework of soil overconsolidation ratio andpile stiffness is believed to reflect soundphysical principles that increases confidencein extrapolating the pile test information tolarge offshore foundations. Criteriasuggested for use in pile design are given onFig. 7.

When compared with current practice, asembodied in the 1982 edition of API RP2A,the a—s,/Tr„—L/D criteria generally producemore optimistic results. For typical piles andnormally consolidated highly plastic clays in

the Gulf of Mexico the required pile lengthsare 5 to 10</o less. At the other extreme ofhard, very heavily overconsolidated siltyclays there is no difference in the predictedcapacity of offshore piles. Pile lengths are

reduced by 10—15% in clays that areintermediate in terms of strength and degreeof overconsolidation.

Consideration should also be given toapplying the underlying concepts on a site-specific basis as a means of improving piledesign for offshore structures. Site-specifically, peak skin friction can beeconomically measured and the scaleadjustment assessed analytically.

200Undrained shear strength, kPa

300 400 500 600

AcknowledgementsThe study reported herein was initially part

of a state-of-practice review performed inconnection with BP's planning of a pile loadtest programme in the UK. We are grateful toBP International Ltd. for permission topublish this Paper.

The work was further stimulated bypublication of the API data base compiled byR.E. Olson and his co-workers at theUniversity of Texas at Austin, and by

0c

c0+

10-

20.

30.

40-

250

r~r

JrI

IIIJJ

I/I

IIII

/

IIrIIlrrrI

s„/o„

Undrained shear strength, kPa

750 1 000

c0+EO

Clce

1000

00

20.

40-

60.

60

0.5

25

1.0

50

1.5 2.0 2.5 3.0Soil strength ratio, Se lay

Ultimate pile capacity, IirIN

75 100 125

500 10 15 20

Soil strength ratio, S„lo„10016 Ground Engineering

deliberations of the API RP2A Geotechnica(Work Group on possible new methods ofpredicting skin friction.

We particularly wish to thank J.A. Focht ofMcClelland Engineers, Inc., and M.F.Randolph of Cambridge University who havethrough their own work, stimulatingdiscussion, and constructive criticism,directly contributed to the ideas given in thisPaper. The contents are, of course, entirelyour responsibility.

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January 1986 19