NEGATIVE SKIN FRICTION ON PILES IN LAYERED SOIL DEPOSITS

By K. S. Wong" Member, ASCE, and C. I. Teh2

ABSTRACT: A simplified numerical procedure for the analysis of negative skin friction on piles in a layeredsoil deposit is proposed. Pile-soil interface behavior is modeled by nonlinear soil springs. A framework fordetermining the model parameters from conventional soil tests data has been established. This procedure isused in the back-analysis of seven well-documented test piles in different soil deposits. The good agreementbetween the computed and the measured values for all seven piles confirmed the validity of the proposedapproach.

METHOD OF ANALYSIS

(1)

(b)

,.~

.,-w.,-w,.,

I..~~t%r'~ ....

-La,... ID

Layer I

(a)

...(.)

where [Kp ] = pile stiffness matrix; {wp } = nodal displacementvector; and {P} = nodal load vector. For piles with cross­section geometry other than circular, an equivalent circularpile that gives the same perimeter area is used. The compres­sion stiffness of the actualyile is preserved by using an equiv­alent Young's modulus, Ep ,

EpA c = EpAp (2)

where Ep = Young's modulus of the actual pile; A p and A c

= actual and equivalent circular pile cross-sectional areas,

homogeneity and anisotropy, which are common features ofnatural soils, cannot be adequately modeled by such ideal­ized solutions. To account for soil inhomogeneity, Poulos andDavis (1980) recommended the use of a mean soil modulusin Mindlin's solution. More significantly, elastic continuumbased approaches do not model the important mechanism ofpile-soil slip. When slip is considered, it is usually achievedby curtailing the interface stresses to some limiting values(Chow et al. 1990).

This paper describes a simplified procedure for the analysisof negative skin friction on pile in a stratified soil deposit.The pile-soil interface is modeled by hyperbolic soil springsin a manner similar to the load-transfer method. The actualpile-soil slippage is not expicitly modeled. The main intentionis to capture the complex phenomenon of pile-soil interactionwith a simple hyperbolic model. Attention is focused on howto determine the relevant soil parameters for the analysis.The validity of this approach is subsequently verified by ex­tensive comparisons with well documented case histories.

The problem of a pile located in a layered soil undergoingconsolidation settlement is depicted in Fig. 1. The pile isdiscretized into a finite number of cylindrical bar elements.The load deformation behavior of the pile can then be ex­pressed as

FIG. 1. Downdrag on Pile In Layered Soli Deposit: (a) Pile In Lay­ered Soli Undergoing settlement; (b) Discretization of Problem

Negative skin friction will develop whenever the adjacentsoil settles more than the pile. The magnitude of downdragcan be quite significant. Bjerrum et al. (1969) reported fieldmeasurements in which the downdrag forces exceeded thestructural capacity of the piles. Very often the downdrag con­tinues to increase even at relatively large soil settlement. Caserecords reported by Indraratna et al. (1992), Fukuya et al.(1982), and Clemente (1984) revealed that the measureddowndrag continued to increase with increasing soil settle­ment well beyond 200 mm. For the case reported by Lee andLumb (1982), the maximum downdrag was not achieved untilthe settlement reached about 400 mm.

Based on the available field data, it is evident that largesettlements may be needed to mobilize the full negative skinfriction on a pile. However, this does not necessarily implythat large soil movements are required to initiate slippage atthe pile-soil interface. A downdrag problem is basically asettlement problem. As the excess pore pressure dissipateswith time, there is a corresponding increase in soil settlementand effective stress. This in turn causes an increase of shearstrength at the pile-soil interface and a corresponding increasein the downdrag. This process continues until the soil becomesfully consolidated. As for pile-soil slippage, it may have startedat a very small settlement and continue to slip along with theincrease in settlement and effective stresses.

In a conventional analysis, downdrag is usually estimatedusing closed-form equations proposed by different research­ers (Terzaghi and Peck 1967; Fellenius 1972). These methodsoften assumed that the negative skin friction is fully mobilizedabove the neutral point regardless of the magnitude of soilsettlement. The location of the neutral point is determinedeither empirically or by some iterative schemes involving thebalances of forces acting on the pile (Fellenius 1972).

More recently, several continuum approaches have beensuggested for the analysis of end-bearing piles (Poulos andMattes 1969; Poulos and Davis 1980; Teh and Wong, in press,1995). The method proposed by Poulos and coworkers wasbased on the solution of a point load in an elastic half-space(Mindlin's solution) and modified empirically to account forthe presence of the rigid bearing layer. Ng et al. (1976) andKog et al. (1986) applied the solution of a point load in alayered elastic half-space (Chan et al. 1974) to study the effectof a compressible bearing layer on negative skin friction. In-

INTRODUCTION

'Assoc. Prof., School of Civ. and Struct. Engrg., Nanyang Tech.Univ., Singapore 2263.

'Sr. Lect., School of Civ. and Struct. Engrg., Nanyang Tech. Univ.,Singapore 2263.

Note. Discussion open until November 1,1995. To extend the closingdate one month, a written request must be filed with the ASCE Managerof Journals. The manuscript for this paper was submitted for review andpossible publication on April 19, 1994. This paper is part of the JournalofGeotechnical Engineering, Vol. 121, No.6, June, 1995. ©ASCE, ISSN0733-941O/95/()()()6-0457-0465/$2.00 + $.25 per page. Paper No. 8291.

JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995/457

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

The stiffness of the spring at any given load level is evaluatedas the tangent to the hyperbolic curve, which may be writtenas

where the subscript i denotes the node number; Pi == nodalforce; kSi = initial tangent of the hyperbolic curve; Wi ==relative pile-soil settlement; Rf == a hyperbolic constant; andPHi = maximum allowable nodal load given by

1. At the beginning of the nth increment, compute theelements of [Ks ] using (7).

2. Solve (13) to obtain {Il.Wp }n'3. Compute a new matrix of [Ks ] based on the tangent

modulus at the displacement {wI' },,-I + T){Il.Wp }", where0<T)<1.

4. Repeat the steps 2-4 until changes in {Il.Wp }" are small.5. Update the nodal displacement vector using (14).6. Repeat steps 1-5 for the next load increment.

Initial Spring Stiffness at Pile Toe

The soil behavior at the pile toe is also described by (1),with Pi representing the maximum end bearing force and W;

the relative pile toe-soil displacement. The initial tangent,k,;, is correlated with elastic soil parameters based on thesolution of a rigid punch on a semi-infinite elastic half-space(Randolph and Wroth 1978)

k'i = 4G;r)(1 - v;) (11)

NUMERICAL PROCEDURE

Since the soil spring modeling the interface is nonlinear,the equation governing the development of negative skin fric­tion has to be solved incrementally. At each step, the incre­mental nodal displacement is computed from the equation

[Kp]{llwpt == [K,]{llw, - Ilw,,}n (12)

where {IlWc} == a vector of incremental free field soil settle­ment; and {IlWp} == a vector of incremental pile node dis­placement. The subscript n denotes the increment number.The soil spring stiffness matrix [Ks ] is computed at the be­ginning of the increment using the known pile node displace­ment. Rearranging (12) leads to

[Kp + K,]{llwpt == [K,]{llwc}" (13)

By applying a known value of {Il.Wc}n, (13) can be solved for{Il.Wp }. At the end of each increment, the pile nodal displace­ment is updated as

{wpt = {Wp}"_1 + {Ilwp}" (14)

Although a reasonably accurate approximation to the hy­perbolic load-displacement curve can be obtained by usingsmall incremental steps, a hybrid incremental-iterative pro­cedure has been adopted instead. This method allows the useof larger incremental step size without compromising on so­lution accuracy. The procedure involved in the proposed nu­merical scheme is outlined here:

(4)

(5)

(7)

(8)

(9)

(6a-c)

Wi ~ P;I(1 - P.)

r", = 2p(1 - v;)L

k'i = 2'ITG;l;lln(r",/r,,)

dp/dw = k,)(1 + WY

where

where Is; == limiting shaft friction; and As; shaft area as­sociated with node i. The constant Rf is assumed to be unityin this study.

By defining two dimensionless parameters Pi and Wi' (2)may be written in a more compact, dimensionless form as

respectively. In the absence of any externally applied load,{P} is a consequence of the relative displacement between thepile and surrounding soil. The soil spring at the pile shaft isgoverned by the hyperbolic equation

Pi = Wi/(~ + Rf Wi.) (3)K,. PU/

Initial Soil Spring Stiffness at Pile Shaft

At low stress level, the interface behavior is essentiallyelastic and the initial spring stiffness, k,;, may be correlatedto the elastic small strain soil parameters. For the pile shaft,the general form of the relationship derived by Randolph andWroth (1978) for a pile in an isotropic, homogeneous, elasticsoil medium has been adopted.

where Gi == soil shear modulus; I == pile segment length; rm

== limit radius; and ro == pile radius. For a homogeneous soilor a Gibson's soil in which the stiffness increases with depth,Randolph and Wroth proposed the following expressionfor r",

(10)

where L == pile length; v; == Poisson's ratio; and p == ratioGLlZ/GL' The terms G LIZ and GLare the soil shear moduliat middepth and at the pile toe, respectively. In the presentstudy, which deals with layered soils, the inhomogeneity fac­tor p proposed by Lee (1993) is adopted

p = JGiz dz/GLL

Other empirical correlations for rm in layered soils havealso been suggested (Lim et al. 1993). A survey of theseproposed relationships shows that rm generally ranges fromO.5L to 2.5L for typical Lid values of 25 to 200. Moreover,k'i as expressed by (8) is proportional to G i and inverselyproportional to In(rn;lro ). Hence, the effect on ks; due to achange in rm will be significantly smaller than that caused bythe same percentage change in Gi • Despite the greater impactof G; on k,;, relatively little study has been carried out on theappropriate method for determining G; for downdrag anal­ysis. A detailed discussion on the correlation of G i with con­ventional soil data is included in this paper.

A computer program names NSFpile (Teh 1994) has beencoded using the proposed method. The case studies describedin this paper were conducted using this program.

EVALUATION OF SOIL PARAMETERS

Eqs. (4), (8), and (11) show that the hyperbolic soil springconstants are functions of the soil stiffness and the interfaceshear strength. The required soil parameters include unit skinfriction (Is) at the interface, soil shear modulus (Gi ), Pois­son's ratio (v;) and ultimate end bearing capacity (qn)' Theseparameters are often not directly available in the publishedcase studies. A framework for determining these parametersfrom the available test data is established to enable a rationalback-analysis to be carried out.

Unit Skin Friction, fs

At the pile shaft, the limiting spring force is related to theunit skin friction Us), The ~-method is adopted in this studyand Is is expressed as

458/ JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

TABLE 2. Recommended Input Parameters for Sand, InterfaceFriction Angle

(15)

where a~ = effective overburden pressure adjacent to thepile shaft at the time when the downdrag is to be computed.The manner in which a ~ is estimated will be described inmore detail in a later section.

The coefficient ~ is a function of the soil type, pile materialand the method of pile installation. For a saturated clay, ~can be determined using an empirical correlation similar tothat proposed by Flaate and Seines (1977)

Pile material(1 )

Rough concreteSmooth concreteSteelTimber

aKulhawy (1984).

8a

(degrees)(2)

<1>'0.8<1>' to <1>'0.5<1>' to 0.9<1>'0.8<1>' to 0.9<1>'

(16)

Shear Modulus, G,

When soil is modeled as an elastic material, the theory ofelasticity yielded the following equation:

where OCR = overconsolidation ratio; and the subscript NCdenotes normally consolidated clay. If field vane data areavailable, ~NC can be estimated as follows: v'as

(2)

0.3 to 0.40.2 to 0.30.1 to 0.2

C--f/'

V

VJ

II

Consistency(1 )

Soft nonconsolidated clayFirm clayStiff over consolidated clay

apoulos and Davis (1980).

FIG. 3. Correlation between Relative Density and SPT Blow CountCorrected for Overburden Pressure (Holtz and Gibbs 1979)

TABLE 3. Recommended Input Parameters for Clay, Poisson'sRatio

~ 100

c:1 80

~~ 80

40

~~ 20

~ 0o 10 20 30 40 50

CORRECTED SPT BLOWCOUNT. N_

(18)

(17)

/3 = Ks tan B

Different values of K s and /) have been reported in the lit­erature (Poulos and Davis 1980; Tomlinson 1986; Kulhawy1984). The recommended values for large-displacement piles(NAVFAC 1982) are summarized in Tables 1-3.

where I.l. = Bjerrum's correction factor (Bjerrum 1972) forvane shear strength. The I.l.-curve is reproduced in Fig. 2. Inthe absence of test data, ~NC can be taken as 0.22.

For sand, ~ is a function of the lateral earth pressure coef­ficient, K., and the interface friction angle, /)

...\~1\

-...t---~\~ E,Jc, (fGr OCR < ])

I-

where E; and Ej = drained and undrained initial tangentmoduli; and v; and Vs = drained and undrained Poisson'sratios, respectively.

The stress-strain curve of a saturated clay determined fromundrained compression test can be idealized as a perfecthyperbola. Based on the formulation proposed by Duncanet a1. (1980), Ej can be related to the secant deviatoric mod­ulus at 50% failure stress level (E50) as

1200

1000

0 800

~ 600".gIi 400w

200

oo 20 40 60

1.2

~

ri0.8 0

~0.6 u..

wz0.4 c(

>0

0.2 ..Jwu:::

o80 100

E'G i = 2(1 ; v;)

Ei = 2E50

2(1 + vs )(19a,b)

(20)

Taking V s to be 0.5 for the undrained condition, the initialshear modulus may be expressed as

PLASTICITY INDEX, PI

RG. 2. Variations of E..,/cu for Clay with OCR < 3 (Duncan andBuchlgnanl1976) and Correction Factor (....) for Vane Shear Strength(Blerrum 1972) with Plasticity Index Gi = (2/3)E50 (21)

TABLE 1. Recommended Input Parameters for sand

D:Compactness (%) Kb Kc nC

s

(1 ) (2) (3) (4) (5)

Very loose o to 15 0.6 to 1.0 250 0.7Loose 15 to 35 1.0 to 1.4 500 0.7Medium dense 35 to 65 1.4 to 1.6 1,000 0.7Dense 65 to 85 1.6 to 2.0 1,500 0.7Very dense 85 to 100 2.0 to 2.4 2,000 0.7

aNAVFAC (1982).·Poulos and Davis (1980), NAVFAC (1982), and Tomlinson (1986).<Wong and Duncan (1974).

For clays with an OCR less than 3, E50 may be estimatedfrom the curve shown in Fig. 2. This curve is derived fromthe correlation proposed by Duncan and Buchignani (1976).

For sand, E,~ is determined from the following (Duncanet a1. 1980)

(22)

where K = modulus number; n = modulus exponent; andPa = atmospheric pressure. Based on a large collection oftest data, Wong and Duncan (1974) had derived typical valuesof K and n (Tables 1-3). The coefficient of earth pressure atrest, K o ' is computed from the empirical equation proposedby Mayne and Kulhawy (1982)

JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995/459

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

Ko = (1 - sin <1>')OCR,;n4>'

where <1>' = effective friction angle in degrees.

(23) (27)

where N60 = blow count corrected for field procedure.

Poisson's Ratio Effective Overburden Stresses

For clays, Poulos and Davis (1980) had recommended a rangeof values based on soil consistency (Tables 1-3).

where Nc and Nq = bearing capacity factors. The valuesrecommended in NAVFAC (1982) have been adopted in thisstudy.

The Poisson's ratio for soil is frequently estimated fromempirical relationships. For sand, the following equation basedon the theory of elasticity can be used

RETROSPECTIVE ANALYSIS OF CASE RECORDS

The parameters i, and G; are assumed to be functions ofthe effective stress. In a downdrag problem the effective stressesare not constant but change with time as excess pore pressuredissipates. In the evaluation of in the appropriate effectivestresses are those prevalent at the time when downdrag is tobe computed. For most practical problems, it corresponds tothe final effective stresses at the end of primary consolidation.On the other hand, the vafues of G; should reflect the changesin soil properties that occur during the period of consolidationand an average G; should therefore be used. Hence, G; isevaluated based on the average of the initial and "final" ef­fective stresses.

For sand, due to the rapid dissipation of pore pressure, i,and G; are computed from the final effective stresses assumingno excess pore pressure.

(24)

(25)

Ultimate End Bearing Pressure

The ultimate bearing pressure at the pile tip is determinedusing the following equation:

Correlation of Dr and ~' with Standard PenetrationTest (SPT) Data

The determination of the sand parameters (K" Km K, n,and v;) requires either the input of relative density (Dr) orfriction angle (<1>'). These data are not given in the case rec­ords reviewed in this study. Only the blow counts from stan­dard penetration tests (SPT) are available. Hence, D, and <1>'are estimated from the blow counts using published corre­lations.

The relative density is estimated from SPT blow countscorrected for overburden pressure (Nco,,) using the correla­tions proposed by Holtz and Gibbs (1979). The proposedrelationship has been simplified and plotted in Fig. 3. Theinternal friction angle is estimated from D, using the corre­lation by Meyerhof (1956):

<1>' = 28 + 0.15D, (26)

where <1>' is in degrees and D, is expressed as a percentage.The correction of blow count for overburden pressure is

determined using the correlation proposed by Liao and Whit­man (1986)

Case records of seven test piles have been reviewed. Therequired input parameters are determined using the proce­dures described in the preceding section. The procedure hasbeen followed consistently in all cases in order to verify theapplicability of the proposed method. Details of each casestudy are described in the following sections.

Case Study 1: Tokyo Bay, Japan

Fukuya et al. (1982) reported the measurement of down­drag on a test pile located at a recently reclaimed land calledOhgishima, off the coast of Kanagawa Prefecture, in TokyoBay. The reclamation work took 3 years from November 1971to December 1974. A steel pipe pile was driven closed-endedwith a diesel hammer during October-November 1973. Thepile was 37.5 m long, with an outside diameter of 610 mmand a wall thickness of 9.5 mm. The equivalent pile moduluswas 12.7 GPa. The pile was driven through 7 m of sand filland 30.5 m of silt, clay and gravel into a bearing stratum ofdense sand. The soil profile and the Cu data from unconfinedcompression tests are shown in Fig. 4. The water table wasassumed to be at 2 m below the ground surface. The pile wasmonitored for about 41 months, until March 1976. The ground

AXIAL FORCE IN PILE (kN)Settlement (mm) Cu (kPa) f. (kPa) ;; '" '"G, (MPa) 0 0

;; '" '" .... ;; ;; u: 0 0 0

00 0 0

'" '" ;; '" '" .... '" 0DepdI{-) 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0• 0

EUedqS..dfll

'-20kNIm' ....14

S."ySih • 10Y "'16kM/w1J 10PI-oIO 10 so •. Gn~ I gY"lOkNfni N-U

I:r :rSUlyOar r 20 Ii: 20

Y-l6ld'lfa1' PI-4OIOSO [l.W "'. WCl Cl

C1ayt1 SII1 30 • 30T -1I'r>WaJ ••PI-1O

37.'DeueS.1UI N.,. 4040

(0) (b) (e) (eI) (e)

I_ Measured - Computed I(f)

FIG. 4. Case Study 1: Tokyo Bay, Japan (Fukuya et al. 1982): (a) Soil Profile; (b) Settlement Profile; (c) Variation of Measured cu ; (d)Variation of Computed fs; (e) Variation of Computed G,; and (f) Comparison of Measured and Computed Axial Forces in Pile

460/ JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

TABLE 4 Determination of f and G, for Sand Fill, Gravel, and Dense Sand Layers for Case 1.0, <1>' I) K n t. G,

Depth 'Ymoist 'Ysaturated No,," (%) (degrees) Ko v' K. (degrees) Table Table (kPa) (MPa).Soil type (m) (kN/m3 ) (kN/m3 ) Eq. (27)

, Fig. 3 Eq. (26) Eq. (23) Eq. (24) Table 1 Table 2 1 1 Eq.15) Eq. (19)IT.(1 ) (2) (3) (4) (5) (6) (7) (8) (9) (10) (11 ) (12) (13) (14) (15) (16)

Sand fill 0 18" 20" 11 0 52 35.8 0.42 0.3 1.53 25 1,090 0.7 0 07 - - - 86 - - - - - - - - 61.4 21.49

Gravel 13.5 - 20" 13 125 57 36.5 0.41 0.29 1.6 25 1,190 0.7 94 2916.5 - - - 155 - - - - - - - - 116 33.71

Dense sand 37.5 - 20" >50 296 90- 41.5 0.34 0.25 2.15 25 1,900 0.7 - 76.64

Note: The ultimate end bearing capacity in the dense sand computed using Eq. (25) is 58.3 MPa."Assumed value.

TABLE 5. Determination of f. and G, for Silt and Clay Layers for Case 1

Initial Conditions Before Reclamation Final Conditions in March 1976 Initial Shear Modulus

p~ = t. G,Plasticity CU IT: OCR OCWIT:

,OCR = Cu [3 (kPa) Average Eso/cu (MPa)IT.

Depth index (kPa) (kPa) Eq. (28) (kPa) (kPa) p~/rr: (kPa) Eq. (16) Eq. (15) Cu (kPa) Fig. 2 Eq. (21)(m) (2) (3) (4) (5) (6) (7) (8) (9) (10) (11 ) (12) (13) (14)

8 40-50 28 7 38 263 92 2.86 47 0.37 34 37 350 8.6413 40-50 21 40 2.97 119 122 1 26.8 0.22 26.8 24 350 5.5817.5 -50 37 75 2.74 205 160 1.29 43 0.25 40 40 300 8.0023 -50 50 110 2.48 273 194 1.41 56.2 0.26 50.4 53 300 10.6225 -50 70 122 3.32 405 205 1.98 77.9 0.31 63.6 74 300 14.8032.5 -25 65 170 1.98 337 255 1.32 70 0.25 63.8 67.5 700 31.50

AXIAL FORCE IN PILE (kN)

PILE COMPR':SS~N~m) 0 § § §o en 0010

If)

30

40"'-------

I• Measured ... No dlaslpatlon ... Full dissipation I

10

~:I:

ai 20c

70

70

(e)

oS.(nm)

o

(d)

OJ (MPI)

'----......49.1

(c)

'----...... 73.7

(b)

30

15

'-'"':"-----' 373

Exx:es! Pore-lkh)

Fill pl•••dbcfon> 1900

(a)

Dcplb(m)

o2~

Clay

13 ----111-----

40 _---JJ- _

20-

27-

FIG. 5. Case Study 2: Sorenga, Norway (Bjerrum et al. 1969): (a) Soil Profile; (b) Overburden Pressure and Excess Pore-Pressure Distri­butIons; (c) VarIation of Computed f.; (d) VariatIon of Computed G,; (e) Settlement Profile; and (f) Comparison of Measured and ComputedPile Compression and Axial Forces In Pile

settled approximately 290 mm during this period. The settle­ment profile obtained in March 1976 is shown in Fig. 4(b).Primary consolidation had nearly completed. The maximumdowndrag was measured at about 2,390 kN. The neutral pointwas located at approximately 82% of the pile length be­lowground.

In the computation, the parameters for the sand fill, gravel,and the bottom dense sand layers were based on average SPTN-values. The stress history of these deposits was not known.The OCR was assumed to be 1.0 in the computation of Ko .

The detailed data are summarized in Table 4.For the silt and clay layers, ~NC and OCR were not known.

They were estimated from the correlation proposed by Mesri(1975) and Lacasse et aI. (1978)

(Cjp')YANE = (0.22 ± 0.04) x OCR082 (28)

where (cu/P')YANE = normalized and corrected field vanestrength. Since no field vane tests were conducted, the cor­rected field vane strength was assumed to be the same as theCu from unconfined compression tests (Nakase et aI. 1972).

The coefficient ~NC was taken to be 0.22 and the OCR wasestimated from (28). Table 5 summarized the actual param­eters used to obtain is and G;. The computed variations of isand G i with depth was shown in Figs. 4(d) and 4(e), respec­tively.

The computed and measured variation of downdrag withdepth are shown in Fig. 4(f). Generally good agreement hasbeen obtained. At a surface settlement of 290 mm, the max­imum computed downdrag of 2,335 kN compares reasonablywith the measured value of 2,390 kN.

Case Study 2: Sorenga, Norway

Bjerrum et aI. (1969) reported the monitoring of downdragin two test piles at Sorenga, in the harbor of Oslo. Two pileswere installed in the fall of 1966. Only the pile without anenlarged base (denoted as pile G in the original paper) isreviewed in the present study. The steel pipe pile was drivento rock at 40 m. The pile had an external diameter of 500mm and a wall thickness of 8 mm. The equivalent pile mod-

JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995/461

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

22------

UI

AXIAL FORCE IN PILE (kN)... N (,01 ..

o 8 8 8 8o 1';-----~

30"----__------'

20

~ 10g

tUJo

\.......-Sal1 --8012 --Sol 3

(el(dl

• Mu. PMtI

•

Ie)

30L----...J

lbl

o +--+--+-"rl

SetUemenl (mm)I'> c.>

o g g g

20

~ 10oS

two

SonOal'Y-ISkNfml

PI-70

(a)

SaDd

6m 16mI I I

01:

Mod. SIIlf Oa)'y-17_ PI-SO

slidaa)'29 y -19_ PI-SO

26

DcpIh(m)

o ----:-N,...._'"'m=:--­2 --~:.::::..;:.=...--

-io- FBI '" Wcalbered Oa)'6 y -17_S

FIG. 6. Case Study 3: Bangkok, Thailand (Indraratna et al. 1992): (a) Soli Profile; (b) settlement Profile; (c) Effective Stress Distributions;(d) Variation of Computed f.; (e) Variation of Computed G,; and (f) Comparison of Measured and Computed Axial Forces In Pile

I- Measured --ComputedI

AXIAL FORCE IN PILE (kN)-. -. N I'\)

en 9: UI 0 ttlo g 0 g g 8

o

30

10

60ra 140m+--+---+

Of:::SetUemenl (mm) '. (kPa) Gf (MPa)

8~N '"Depth(-)

HewfUl 0 '" w ... g;- g; g; 8 g; r:l .. g:0 0 0 0 0 0 0

• 02 1

_JI1IIY-17kN1.-3 N--14

Meci. Plae S._d

y 0'9"""'" N-ol4to3

10IIInD SI.,. 0.,

T • u.s tNhra' '1-50 :[:I:l-ll.l!:l 20

:.u~:::"

~7.s

DeaR S.ad II: Cnvel

., -20kNIm' N-1I1050 30

" c.) Cb) Ce) Cd)(e)

FIG. 7. Case Study 4: Melbourne, Australia (Walker and DarvaIl1973): (a) Soli Profile; (b) Settlement Profile; (c) Variation of Computedf.; (d) Variation of Computed G,; and (e) Comparison of Measured and Computed Axial Forces In Pile

ulus was 13 GPa. The site was mantled by 13 m of fill followedby 27 m of soft clay overlying bedrock [Fig. 5(a)]. The fillwas placed at the end of last century and consisted mainly ofsawdust and sandy organic mud. The water table was locatedat 2 m belowground. Small excess pore pressures were re­corded before piling [Fig. 5(b)]. Two years after pile instal­lation, the pile head and the surrounding ground were foundto have settled about 13.8 mm and 70 mm, respectively. Themaximum downdrag was 2,500 kN.

The unit weights of the fill and clay were not available.They were back-calculated from the overburden pressure dis­tribution reported by Bjerrum et at. (1969). The moist andsaturated unit weight of the fill were estimated to be 16 and18.5 kN/m3 , respectively. The saturated unit weight of thesoft clay was 19 kN/m3 . The only known information on thefill was its composition. The following parameters were as­sumed: ~ = 0.3; Eso/cu = 500; and v; = 0.35. The soft claylayer was assumed to be normally consolidated with a ~ of0.22 and v; of 0.35. A plasticity index of 20, typical for claysin that region, was assumed. The corresponding Eso/cu ratiowas deduced to be 1,000 from Fig. 2. The excess pore pres­sures at the end of the test were not reported. Since the fillwas placed more than 70 years ago, the change in excess porepressure during the 2-year test period was likely to be small.

462/ JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995

Nevertheless, two analyses were conducted. The first as­sumed no dissipation of excess pore pressure; and the secondassumed full dissipation. The computed variations of Is andG; with depth are shown in Figs. 5(c) and 5(d). The bedrockwas assumed to be rigid and nonyielding. The profile of soilsettlement was not known. The settlement profile is assumedto be constant in the fill layer and linearly decreasing in theunderlying clay as shown in Fig. 5(e).

The computed downdrag and pile compression are shownin Fig. 5(f) together with the measured data. Generally goodagreement has been obtained between the measured and com­puted values. It was also noted that the small excess porepressure has very little effect on the final results. The maxi­mum computed downdrag assuming full dissipation of excesspore pressure is 2,380 kN, which compares well with themeasured value of 2,500 kN. The neutral point was locatedat the pile tip as expected for an end bearing pile.

Case Study 3: Bangkok, Thailand

Indraratna et at. (1992) reported a field study of downdragon driven piles at a site 10 km east of Bangkok. The hollowprestressed concrete pile was 27 m long with an external di­ameter of 400 mm and a wall thickness of 75 mm. The equiv­alent pile modulus was 30.5 GPa. The pile was driven through

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

Gj (MPa)N .. CD CD

00000

fs (kPa)..... CD f\) m

00000

...........,..••

Effective Stresa (lePe)

8 0 8 8 ~ 8Cu (kPa)

o g: 8 §

t\.\'.~.\0'\ ..,.. 'a.'\

".\ .'. \

10

40

o

Settlement (mm)

o ~ ~ ~ g

50 ........------'

g 20

~0­Wo 30

50----_ Sa_

IludySUI1_11_3

P1-2035----

SudJSl1t1_19_3

431_.:.;",--.::;15__

9

IJ

25----

DopIb(m)

o ,11I2 SUI1 Sa...

PI-.1-20~

~

(a) (b) (e) (4) (e) (f)

50 -'------------'

•

AXIAL FORCE IN PILE (kN)...... ...... fI,,) N

o g g 8 § 8

30•

AXIAL FORCE IN PILE (kN)...... f\) c,.) ..

§ § § §oo

10

40

g 20

b:wo 30

AXIAL FORCE IN PILE (kN)~ I\) § §

0 § §0

10

g 20J:I-0-w0 30 •

•40

50

(gl

FIG. 8. Case Study 5: Tokyo, Japan (Endo et al. 1969): (a) 5011 Profile; (b) Settlement Profile; (c) Variation of Measured cu; (d) EffectiveStress Distributions; (e) Variation of Computed f.; (f) Variation of Computed G,; and (g) Comparison of Measured and Computed AxialForces In Plies

fill and soft clay into a stiff clay stratum. The soil profileincluding the new fill is shown in Fig. 6(a). A test embank­ment (12 m x 22 m x 2 m high) was constructed to induceground settlement. The pile was monitored for 9 months. Theground settled about 270 mm during this period. The settle­ment profile is shown in Fig. 6(b). Piezometer readings in­dicated that the primary consolidation had completed. Themaximum downdrag reached 305 kN. The neutral point islocated at approximately 85% of the pile length belowground.

Indraratna et al. reported a value 5 MPa for E' and a v;of 0.2 for the fill material, which is equivalent to a Gi of 2.08MPa. The Gi and v; values of the weathered clay are 1.7 MPaand 0.33, respectively. The average unit skin friction, Is, ofthese two soils were assumed to be 10 kPa, which corre­sponded approximately to a 13 of 0.3. The stress increase dueto the surcharge was computed using Boussinesq's solutionfor finite areal load. The profiles of preconsolidation pres­sure, effective overburden pressure before and after fill place­ment are shown in Fig. 6(c). The available data on the bearingstratum do not permit a direct estimate of the pile toe bearingcapacity. Three sets of c' and <1>' values were tested: (1) c'= 50 kPa and <1>' = 23°; (2) c' = 20 kPa and <1>' = 26°; and(3) c' = 0 kPa and <1>' = 28°. The corresponding qb-valuesare 1.75 MPa, 2.47 MPa, and 3.15 MPa, respectively. Thecomputed profiles of f, and Gi are shown in Figs. 6(d) and6(e).

The computed and measured downdrag profiles are shownin Fig. 6(f). The effect of different end-bearing pressures on

the computed downdrag was insignificant. The computeddowndrag are in close agreement with the measured data.

Case Study 4: Melbourne, Australia

Walker and Darvall (1973) reported the downdrag of asteel pipe pile driven closed-ended through a highly stratifiedsoil [Fig. 7(a)]. The pile was 34 m long, with an externaldiameter of 760 mm and a wall thickness of 11 mm. Theequivalent pile modulus was 12 GPa. After the pile was driven,a test embankment (100 m x 200 m x 3 m high) was con­structed. The pile was monitored for about 8 months. Primaryconsolidation was completed within the first 4 months andthe final vertical stresses did not exceed the preconsolidationpressures. The ground settled about 29 mm after 4 months[Fig. 7(b)]. The maximum downdrag was measured to be 1.8MN.

For the analysis, the properties of the fill and sand werederived from the SPT N-values. The moist and saturated unitweights were assumed to be 17 kN/m3 and 19 kN/m3, re­spectively. The water table was deemed to be located at 1.5m below the original ground surface. The stress increase dueto the surcharge was determined using Boussinesq's solution.No data on the silt layer were available, so it was assumedto have properties of a sand with a relative density of 25%.For the silty clay, I3Nc was assumed to be 0.22. Consolidationtest yielded an OCR of 1.57 at a depth of 18.5 m. This OCRvalue was assumed to be applicable to the entire clay layer.The average plasticity index of the silty clay was about 65.

JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995/463

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

Based on these parameters, is and G; at various depths wereobtained [Figs. 7(c and 7(d)]. Based on a corrected N-valueof 22, the sand at the pile toe was estimated to have a Dr of74% and 4>' of 39°. This resulted in a qb of 33 MPa.

The computed and measured downdrag forces are shownin Fig. 7(e). Good agreement has again been obtained.

Case Study 5: Tokyo, Japan

Endo et al. (1969) reported the measurement of downdragon three vertical piles and two batter piles at a site locatedabout 3 km west of downtown Tokyo. Only the vertical pileswere analyzed. The piles had an outside diameter of 609.6mm and a thickness of 9.5 mm. The equivalent pile moduluswas 12.7 GPa. Two of the piles were founded on a loose siltysand stratum at 43 m depth. One was driven closed-ended(cE43) and the other was open-ended (oE43). The third pile,31 m long, was a friction pile. It was driven closed-ended withits toe ending in the silt deposit (cF31). The piles were drivenin June 1964. Measurements were monitored until October 1%7.It was reported that settlement in the area was caused by pump­ing of ground water for industrial purposes. The soil profileconsisted of 2 m of fill, 37 m of silt, 6 m of loose diluvial sandfollowed by a layer of dense sand [Fig. 8(a)]. The ground-watertable was at 1.5 m below the surface. The settlement profileobtained in April 1966 is shown in Fig. 8(b). The ground surfacesettled 131 mm. The variation of undrained shear strength withdepth based on the unconfined compression tests are shown inFig. 8(c). The effective vertical stresses and the maximum pastpressures are shown in Fig. 8(d).

For the analysis, the fill was assumed to have a Dr of 50%and a moist and saturated unit weight of 18 kN/m3 and 20kN/m\ respectively. The computed t and G; at 2 m depthare 22.4 kPa and 9.95 MPa, respectively. The diluvial sandat 43 m has an average SPT N-value of 20. The corrected N­value was 8, which gives rise to a Dr of 42% and a 4>' of 34°.The computed q/, and G; were 30.55 and 70.38 MPa, respec­tively. For pile cF31, the end bearing capacity in the silt wascomputed based on an assumed friction angle of 30°, whichresults in a qb of 3.93 MPa. The computed f, and G; profilesare shown in Figs. 8(e) and 8(f), respectively.

The computed and measured downdrag of the three pilesare shown in Fig. 8(g). The computed values for piles oE43and cF31 are in good agreement with experimental data. Largerscatter was observed in the computed and measured down­drag of cE43. One possible explanation could be attributedto local variation of soil properties. This can be seen in thesudden increase of measured downdrag between the depthsof 20 m to 25 m for cE43, which was not observed in theother two piles.

CONCLUSIONS

A simplified load-transfer method was proposed for theanalysis of downdrag on single piles in layered soil. A con­sistent procedure has been established for the evaluation ofthe input parameters from conventional soil test data. Thecase records of seven test piles consisting of various pile typesin different soil deposits around the world were analyzed.Results of the analysis have confirmed that the proposed ap­proach is capable of predicting with reasonable accuracy thedowndrag on single piles.

APPENDIX. REFERENCES

Bjerrum. L. (1972). "Embankments on soft ground." Proc., Spec. Conf.on Perf. of Earth and Earth-Supported Struct., ASCE, New York,N.Y .. Vol. 2. I-54.

Bjerrum. L.. Johannessen, 1. J., and Eide, O. (1969). "Reduction of

464/ JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995

skin friction on steel piles to rock." Proc., 7th Int. Conf. in Soil Mech.& Found. Engrg., Mexico, Vol. 2,27-34.

Chan, K. S., Karasudhi, P., and Lee, S. L. (1974). "Force at a point inthe interior of a layered elastic half-space." Int. J. ofSolids and Struct.,10,1179-1199.

Chow, Y. K., Chin, J. T., and Lee, S. L. (1990). "Negative skin frictionon pile groups." Int. 1. Numerical & Analytic Methods in Geomech ..14(1),75-91.

Clemente, F. M. Jr. (1984). "Downdrag, negative skin friction and hi­tumen coatings on prestress concrete piles," PhD thesis, Tulane Univ.,New Oleans, La.

Duncan, J. M., and Buchignani, A. L. (1976). "An engineering manualfor settlement studies." Geotech. Engrg. Rep., Dept. of Civ. Engrg..Univ. of California, Berkeley.

Duncan, J. M., Byrne, P., Wong, K. S., and Mabry, P. (1980). "Strength.stress-strain and bulk modulus parameters for finite element analysesof stresses and movements in soil masses." Geotech. Engrg. Rep. USCIGT/80-01, Dept. of Civ. Engrg., Univ. of California, Berkeley.

Endo, M., Minou, A., Kawasaki, K., and Shibata, T. (1969). "Negativeskin friction acting on steel pipe piles in clay." Proc., 7th Int. Cantin Soil Mech. & Found. Engrg., Mexico, Vol. 2, 93-98.

Fellenius, B. H. (1972). "Downdrag on long piles in clay due to negativeskin friction." Can. Geotech. 1.,9(4),323-337.

FIaate, K., and Seines, P. (1977). "Side friction of piles in clay." Proc.,9th Int. Conf. in Soil Mech. & Found. Engrg., Tokyo, Japan Vol. 1,517-522.

Fukuya, T., Todoroki, T., and Kasuga, M. (1982). "Reduction of neg­ative skin friction with steel tube NF pile." Proc., 7th Southeast AsianGeotech. Conf., Hong Kong, Vol. 1,333-347.

Holtz, W. G., and Gibbs, H. J. (1979). "Discussion of 'SPT and relativedensity in coarse sand.''' 1. Geotech. Engrg., ASCE, 105(3),439-441.

Indraratna, B., Balasubramaniam, A. S., Phamvan, P., and Wong, Y.K. (1992). "Development of negative skin friction on driven piles insoft Bangkok clay." Can. Geotech. 1.,29(3),393-404.

Johannessen, 1. J., and Bjerrum, I. (1965). "Measurement of thecompression of a steel pile to rock due to settlement of the surroundingclay." Proc., 6th Int. Conf. in Soil Mech. & Found. Engrg., Montreal,Canada, Vol. 2, 261-264.

Kog, Y. c., Karunaratne, G. and Lee, S. L. (1986). "Effects of NSFon piles in layered soils." Geotech. Engrg., Bankok. Thailand, 17(2),211-234.

Kulhawy, F. H. (1984). "Limiting tip and side resistance." Proc., Symp.on Design and Analysis of Pile Found., ASCE, New York, N.Y., 80­98.

Lacasse, S. M., Ladd, C. c., and Baligh, M. M. (1978). "Evaluation offield vane, Dutch cone penetrometer and piezometer testing devices."Res. Rep., Massachusetts Inst. of Technol. (MIT), Cambridge, Mass.

Lee, C. Y. (1993). "Pile groups under negative skin friction." 1. Geotech.Engrg., ASCE, 119(10), 1587-1600.

Lee, P. K. K., and Lumb, P. (1982). "Field measurements of negativeskin friction on steel tube piles in Hong Kong." Proc., 7th SouthwestAsian Geotech. Conf., Hong Kong, 363-374.

Liao, S. S. C., and Whitman, R. V. (1986). "Overburden correctionfactors for SPT in sand." J. Geotech. Engrg., ASCE, 112(3),373-377.

Lim, C. H., Chow, Y. K., and Karunaratne, G. P. (1993). "Negativeskin friction on single piles in a layered half-space." Int. J. for Numer.& Analyt. Methods in Geomech., Vol. 17,625-645.

Mayne, P. W., and Kulhawy, F. H. (1982). "K,,-OCR relationships insoil." J. Geotech. Engrg., ASCE, 108(6),851-872.

Mersi, G. (1975). "Discussion of new design procedure for stability ofsoft clays." J. Geotech. Engrg., ASCE, 101(4),409-412.

Meyerhof, G. G. (1956). "Penetration tests and bearing capacity ofcohesionless soils." J. Soil Mech. and Found. Div., ASCE, 82(1), 1­19.

Nakase, A., Katsuno, K., and Kobayashi, M. (1972). "Unconfinedcompression strength of cohesive soils with high sand content." Rep.of Port and Harb. Res. Inst., 11(4), 83-102.

NAVFAC. (1982). "Foundations and earth structures." Design manual7.2, U.S. Department of the Navy, Alexandria, Va.

Ng, H. K., Karasudhi, P., and Lee, S. L. (1976). "Prediction of negativeskin friction and settlement in piles due to fill surcharge." Geotech.Engrg., 7(1), 25-45.

Poulos, H. G., and Davis, E. H. (1980). Pile foundation analysis anddesign., John Wiley, New York, N.Y.

Poulos, H. G., and Mattes, N. S. (1969). "The analysis of downdrag inend-bearing piles." Proc., 7th Int. Conf. in Soil Mech. & Found. Engrg.,Mexico, Vol. 2, 203-208.

Randolph, M. F., and Wroth, C. P. (1978). "Analysis of deformation ofvertically loaded piles." J. Geotech. Engrg., ASCE, 104(2), 1465-1488.

Teh, C. I. (1994). "NSFpile-a computer program for the analysis of

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.

downdrag on single piles." Geotech. Res. Rep. GT/94/0I, School ofCiv. and Struct. Engrg., Nanyang Tech. Univ., Singapore.

Terzaghi, K., and Peck, R. B. (1967). Soil mechanics in engineeringpractice, 2nd Ed., John Wiley & Sons Inc., New York, N.Y.

Tomlinson, M. J. (1986). Foundation design and construction, 5th Ed.,Pitman Books Ltd., London, England.

Walker, L. K., and Darvall, P. (1973). "Downdrag on coated and un­coated piles." Proc., 8th Int. Conf. in Soil Mech. & Found. Engrg.,

Moscow, U.S.S.R., Vol. 2.1,257-262.Wong, K. S., and Duncan, J. M. (1974). "Hyperbolic stress-strain pa­

rameters for nonlinear finite element analyses of stresses and move­ments in soil masses." Geotech. Engrg. Rep., Dept. of Civ. Engrg.,Univ. of California, Berkeley.

Wong, K. S., and Teh, C. I. (1994). "Downdrag on single piles." Geo­tech. Res. Rep. GT/94/02, School of Civ. and Struct. Engrg., NanyangTech. Univ., Singapore.

JOURNAL OF GEOTECHNICAL ENGINEERING / JUNE 1995/465

J. Geotech. Engrg., 1995, 121(6): 457-465

Dow

nloa

ded

from

asc

elib

rary

.org

by

Eng

enha

ria

Bib

liote

ca o

n 07

/11/

16. C

opyr

ight

ASC

E. F

or p

erso

nal u

se o

nly;

all

righ

ts r

eser

ved.