Linkoping, SGI Varia 65

DESIGN OF PILES IN COHESIVE SOIL

Nguyen Truong rr

SGI, Linkoping, .3weaen, Sepi:( Tit•r-..c 1-;, b 1

SGI Varia 65

DESIGN OF PILES IN COHESIVE SOIL

CONTENTS

SUMMARY 1

ACKNOWLEDGEMENTS ₂

INTRODUCTION 3

1.

1. 1 1. 1. 1 1.1.1.1 1.1.1.2 1.1.1.2.1 1 . 1 . 1 • 2 • 2 1.1.1.2.3 1 • 1 • 1 • 2 • 4 1 • 1 • 1 • 2 . 5 1.1.1.2.6 1.1.1.2.7 1 • 1 • 1 • 2 • 8 1 • 1 • 1 • 2 • 9

Bearing capacity of single piles ₅ Methods based on static formulas ₅

Total stress analysis ₅

End bearing ₅

Shaft friction ₈

Canadian Foundation engineering manual ₉ Australian Code

Swedish Code Danish Standard Buildinq Code of

(SAA) 11

(SBN 75) 1 1

12 the Soviet Union _{1 3}

Experience of Thailand _{1 6}

Brom's recommendation 1 6

Vesic's recommendation

The CTH method _{1 8}

1 . 1 . 1 • 2. 1 0 The method of Caquot and Kerisel 1.1. 2

1.1.2.1 1.1.2.2 1.1.2.2.1 1 • 1 • 2 • 2 • 2 1.1.2.2.3 1 • 1 • 2 • 2 • 4 1.1.2.2.5 1 • 1 • 2 • 2 • 6 1 • 1 • 2 • 2 • 7 1.1.2.2.8 1 • 1 • 2 • 2 • 9

Effective stress analysis for bearing

capacity of piles 24

End bearing capacity 27

Shaft friction resistance 29

Burland 30

Canadian Foundation enqineering manual 33

Meyerhof ₃₄

Vesic's recommendation 37

Vijayvergiya and Focht 40

Flatte et al 43

Bozozuk et al ₅

Blanchet et al

Esriq and Kirby 47

18

23

2

1 • 1 • 2 . 2 . 1 0 J anbu

50

1.1.2.2.11 Parry and Swain 52

1 . 1 • 3 1.1.3.1 1.1.3.2 1.1.3.3 1.1.3.4 1.1.3.5

1 . 1 • 4

1 • 2

1 • 2. 1 1.2.1.1 1.2.1.2 1.2.1.3 1.2.1.4 1.2.1.5 1.2.1.6 1 • 2. 2 1.2.2.1 1.2-2.2 1.2.2.3

1. 2. 3 1 • 3 1. 3. 1 1.3.1.1 1.3.1.2 1.3.1.3 1.3.1.4 1.3.2

Discussion

End bearing capacity of a single pile The a method

The B method

Relation between the B method and the A. method

Relation between the a method and the B method

Summary and recommendation for design

Determination of point and skin resistance from field test

Static cone penetration test Vesic

Nottingam ano Schmertmann Broms

Tong et al Sanglerat

Balasubramaniam et al Standard penetration test Meyerhof

David

Relationship between N and the undrained shear strength

Summary

Negative skin friction Basic concept

Causes

Factor that affect the negative skin friction

Neutral point

Fellenius' observation

Design methods for negative skin friction

53 53 54 57

60 61 64

67 67 67 67 69 70 70 72

74 74 74 75 77 78 78 78 78 78 83

83

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1.3.2.1 Canadian Foundation Engineering Manual 1.3.2.2 Bozozuk

1.3.2.3 Broms 1.3.2.4 Fellenius 1.3.2.5 Kezdi 1.3.2.6 Auvinet

1.3.3 Reducing negative 1. 3. 3. 1 Bitumen coating 1. 3. 3. 2 Protection piles 1.3.3.3 Overlapping piles

skin friction

1.3.3.4 Change of the geometry of the pile group

1.3.3.5 Change of the shape of piles 1 .3.3.6 Reduction of point resistance 1.3.4 Summary

2. Settlement analysis of single piles 2.1 Vesic

2.2 Poulos 2.3 Summary 3. Pile grouns

3. 1 Ultimate bearing capacity of pile groups

3.1.1 Introduction 3.1.2 Design methods 3. 1 . 2. 1 Peck et al

3.1.2.2 Canadian Foundation Engineering Manual and Broms

3.1.2.3 Vesic

3. 1 . 2. 4 Morr .house and Sheehan 3.1.2.5 Brand et al

3.1.2.6 Meverhof

3.1.2.7 Australian Code(SAA) 3.2

3.2.1 3.2.2 3.2.3

Settlement of pile groups Terzaghi and Peck

Tomlinson

Morgan and Poulos

83 85 8, 90 90 91 92 9'.?

92 93 93 93 94 95 97 97 100 101 102 102 102 102 102 104 105 106 106 108 109 11 4 1 1 4 1 1 5 11 7

3. 2. 4

3.2.5 3.3

4.

Appendix Apnendix Anpendix

Appendix

Appendix Appendix

4

Mattes and Poulos 1 21

124 Vesic

125 Summary of the design of pile groups

127 Conclusions

A Analysis of point resistance 129

B Typical values of soil pile adhesion 133 C Classification of the clay according

to the consi_stency of the soil 135

D

s

values for piles in t i l l

(moraine clay ) 1 36

13 9 E Values of Young's modulus

142

F References

SGI Varia 65

SUMMARY

This report is made to review various design methods in terms of total and effective stress analysis for determination of the bearing capacity of piles in co­

hesive soils. The relations between different methods are commented and discussed. A summary of general re­

commendations for calculation of the bearing capacity of piles is presented. Also general methods for calcu­

lation of negative skin friction and settlement of a single pile are summarized.

Some papers and current methods for design of pile groups are selected and reviewed. The report contains diagrams.

tables and typical values of parameters that can be used for design purpose or as a guide in a preliminary design of pile foundations in cohesive soil.

2

I

ACKNOWLEDGEMENTS

This report was done at my visit at SGI during 1981 according to SAREC's (SIDA) programme to which appreci­

ation is expressed.

Great thanks to Dr Jan Hartlen, Director of SGI for his recommendation on the work's programme, and his assistance and encouragement.

Grateful thanks to Dr Bo Berggren at SGI for critical reading of the manuscript and invaluable discussions and recommendations.

Gratitude is expressed to Mrs Eva Dyrenas for her expert typing of the manuscript and Mrs Rutgerd Abrink for draw­

ing the figures.

also express my thanks to other members of SGI for their kindness and their assistance during my time at SGI.

Linkoping September 1981

Nguyen Truong Tien

SGI Varia 65

INTRODUCTION

In the last decade, numerous studies have been performed to determine the behaviour of single, axially loaded piles in cohesive soil. There are many factors influencing the behaviour of piles. The most important factors are: soil conditions, pile dimension, installation methods, pile material and stress-strain history of the soil. Therefore, besides existence of a method which takes into account the variety of conditions, the designer must possess a good knowledge of engineering science.

The ultimate bearing capacity of a single pile in cohesive soil is in general limited by the ultimate strength of the surrounding soil. The ultimate bearing capacity of a pile can be evaluated from

a) Calculation methods based on the measured or estimated shear strength of the soil.

b) Static penetration tests where the resistance is measured when a penetrometer is pushed down into the soil at a constant rate.

c) Dynamic penetration tests where the ultimate bearing capacity of a pile is calculated from the number of blows required to drive a penetrometer a given distance into the soil.

d) Pile load test.

The accuracy of different calculation methods depends to a large extent on the measurements of the strength and resistance. The most reliable method to determine the ultimate bearing capacity of single piles is by pile load tests.

Negative friction will be produced where the surrounding soil exhibits a downward movement with respect to the pile shaft, and this effect can cause excessive settlement of the piles with severe damage of the structure. Consequently

4

there is a great interest in practical methods of re- ducing the negative skin friction.

The settlement analysis of pile foundations depends on the position of the load transfer from the pile to the soil; and this is a complicated problem. Therefore, only approximate solutions of this problem are available.

The behaviour of a pile group differs from that of a single pile. The ultimate bearing capacity of a pile group depends on soil type, size of the group, spacing, length of piles and the construction procedures. The evaluation of the ultimate bearing capacity and the

settlement of a pile group is based on empirical methods.

A review of various design methods for determination of the bearing capacity and the settlement of the piles is presented in this report. In the report diagrams, tables, empirical expressions for design purposes have been

collected. Relationships between different methods and recommendations for design have been summarized.

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1 • BEARING CAPACITY OF SINGLE PILES

Vertical axial loads applied on single piles are trans- mitted to the surrounding soil from the pile by skin friction and end-bearing. Qf is defined as the ultimate load where both the total shaft resistance Q s' and the point resistance Qp are mobilized simultaneously.

= A f + A q s s p p

= Qs ⁺ Qp

where A, A are the shaft and point tip area of the pile s p

respectively and f , q are the unit skin friction and s p

unit point resistance. According to the Canadian Foundation Engineering Manual (1978)

if _Cu < ^{100 kPa} _Qf

=

_Qs

and C, _u > ^{1 00 kPa} _Qf

=

_Qs ⁺ _Qp

where C is the cohesion in undrained condition.

u

Two types of approach are currently in use to evaluate Qf: the total stress analysis and the effective stress analysis. Also two methods are generally used in pre- diction of the ultimate bearing capacity of the pile:

the method based on static formulas and the method based om the result of field tests.

1.1 Method based on static formulas

1.1. Total_stress_analysis_for_bearing_caEacity 1.1.1.1 End bearing

The end-bearing Q is related to the undisturbed undrained p

cohesion c of the soil below the pile (Terzaghi, 1943) and is given by the formula

Qp = c N A ⁺ 0 N A - W

U C p V q p p

W is the weight of the pile, but in general W and

p p

6

a A are omitted because the weight of the pile is often

V p

about equal to the displacement soil (N =1 in cohesive q

soil ( c/>= 0) )

A = cross sectional area of the pile tip, m² p

CU = cohesion of the soil, kN/m² N ,N = bearing capacity factors

C q

If soft clays, N is often assigned a value of 9, but i t

C

can vary from 5 in very sensitive,normally consolidated clays (Ladanyi, 1973) to over 10 in overconsolidated clays

(Skemton, 1951). It also varies with the internal angle of friction. (Fig. 1 , Meyerhof, 1976)

Meyerhof (1976) limits the value of Q at the critical p

depth D of the pile.

C

For bored piles, the Canadian manual recommends

= N* c A

C U p

where = ultimate point load, kN

= cross sectional of the pile point, m

= minimum undrained shear strength of the clay at the pile point level: kPa

N* = bearing capacity factor, which is a

C function of the pile point diameter.

Point diameter N*_{_Q_}

less than 0.5 9

0.5-1.0 7

greater than 1.0 6

In very stiff clays and t i l l , cu can be measured by pressure- meter. From experience, the Danish standard recommends that

Q fp = 18 CU Ap

The Australian Code (SAA 1978) subjects according to Whitaker & Cooke (1968) that the value of N varies with

C

SGI Varia 65

L/B, where Lis the length of the pile and Bis the diameter.

If L/B > ^{4 N} _C

=

⁹

and L/B < ^{4 N} _C

=

^5.6

Broms (1972) has pointed out that Qfp = 9 cuAp' where u is obtained from fall cone tests, and the value of Qfp

generally corresponds to 10 = 20% of the total pile bearing capacity.

1000

~r,, L

,Jj

^/

^~ ;/,

J

/

v/

100

~.,,, ~

^,,✓

~

...

^{/ .} / / . / /

"'.,^C: /

//4' /

/,"7 r/

.. _, ),

^{" /}

^,

z

_,,/

.. /

~ /.

N

.,,,,

~ /

u

.,,,,,,, .,,,,,,,,I 1/ ^/

8. ,,. ^/ 20

u 00 ^C: Ne'"""., 1 / /

..

_{/ /} -::>-Nq

V

~

"' i6

10 ....

-

^_,,,.

^.,,,.,

_,,,,

.. [j....-,.,,,

/ /

V

;) ^/

/

V

₁

~.,... ... / /

/

/

~

~ -<,,,

^.,,

' /

:::,,..D,

v-- -

~

--- _-;}

1/

/

⁸

_,,,./

L ^/

l 1

0 10 15 20 25 30 35 40 45

Angle of internal friction, ,; , in degrees

Fig. 1 Bearing capacity factors and critical depth ratios for driven piles. (After Meyerhof, 1976)

3

According to Caquot and Kerisel (1956) the point resistance is given by the formulas

= C N + yD N

. U C q

N and N are functions of the angle of internal friction

C q

and are given in TABLE 1.

YD = overburden pressure at the level of the pile tip.

TABLE 1. Bearing capacity factors of point resistance (After Caquot and Kerisel, 1956)

<P o _{Ne Nq max Nq min}

0 7 1 . 0 1 . 0

5 9.66 1 . 84 1 . 5

1 0 13.82 3.44 2.5

1 5 6.22 4.0

20 12.78 6.67

25 2 6. 1 6 11 . 41

30 56.95 20.37

1.1.1.2 Shaft friction

The shaft resistance is related to the average undisturbed undrained shear strength of the soil

c

_u along the pile.

Q = ac ,1 s u s

Where A = area of the pile shaft in contact with the soil.

s

The adhesion factor a is not constant but varies with soil t ype, shear strength, time, method of installation of the pile, dimensions and shape of the pile and other factors.

It varies from about 1 for piles driven in soft clays to l e ss than 0.25 for piles driven in stiff clay. An upper l imit of 100 kPa is assigned to the shaft adhesion a c

(Bozozuk, 1979). Some criteria on value of a is summarized below.

SGI Varia 65

1.1.1.2.1 Canadian Foundation Engineering Manual

Q = a c A

s u s

The Canadian manual recommends a value of a according to Tomlinson ( 1971) (see Fig. 2 and Table 2) . The value of a is empirical, therefore the bearing capacity of piles resulting from the above formula should be con- firmed by load tests. For the case of bored piles in

clay, where c > 100 kPa, the shaft adhesion is calculated u

by

Q = ac A

s ua s

where c = ultimate adhesion kPa ua

Experience shows that C =(0.3-0.4)C

ua u

c is greatly affected by the excavation process. It is ua

recommended that c is determined from the minimum un- ua

drained shear strength c and that it is limited to a u

maximum of 100 kPa. The ultimate load should be confirmed by load tests.

TABLE 2. Design values of adhesion factors a for piles driven into stiff to very stiff cohesive soil (Tomlinson, 1971).

Case Soil condition Penetration Adhesion

ratio factor

Sand or gravel, over- 20 1 . 25 1 lying stiff to very

stiff cohesive soil >20 see Fig.2 Soft clay or silt over- 8-20 0.4 2 lying stiff to very

stiff cohesive soil >20 1. 07 Stiff to very stiff 8-20 0.4 3 cohesive soil without

overlying strata >20 see Fig.2

- -

1 0

Penetration ratio: Depth of penetration into stiff to very stiff soil/Diameter of pile (relation between L/B in Fig.2).

Undrained Shear Strength (cul lb/ft²

0 1000 3000 4000 5000

l.00 tj

0.75 0 0

~ 0.50

-~ C _.,

.c 0.25

-0 <t 0.00

0 50 100 150 200

Undrained Shcor Strength (cul kPa

(al PILES DRIVEN THROUGH OVERLYING SANDS OR SANDY GRAVELS

Undrained Shear Strength (eul lb/ft²

1000 2000 3000 4000 5000

1.00°

tj _0.75 ...

_

- - - c L = G r e a t than 208 0 ~

0

~

0.50

,..._ -- _---

·;;; - - - .,.,-L=I08

" ^0.25

.c -0 0

C

--- _---

<t 0.000

50 100 150 200

Undroined Shear Strength (cul kPa (b) PILES DRIVEN THROUGH OVERLYING SOFT CLAY

Strength (cul lb/ft²

0 3000 4000 5000

1.00

Sand or Sandy Gravels

tj 0 0.75

~ 0 0.50

C

·;;., 0 .c 0.25

-0 <t

0.000

(Cl PILES

...

, ',

'\..

'

'<L=Great than 408

'

_'\.

_{,_ '}

...

_____

'

^...

\_--;::~s--

50 100 150 200

Undrained Shear Strength (cul kPa WITHOUT DIFFERENT OVERLYING STRATA

Note:Curves not applicable to Hor cruciform section or to bored or driven and cast in place piles

Safety factor should not be less than 2.5 except for design based on adequate loading testdata

Fig. 2 Adhesion factor vs. shear strength for different penetrations into stiff clay. (After Tomlinson, 1970)

SGI Varia 65

1.1.1.2.2 Australian Code (SAA)

ct cA

s

The values of a are presented in Fig. 3.

1.0

o.e \

l;j

~

0

\

1-

u

0. 6

~

\.

'\

z

~

0 ~

~

u

^0.4

-...

--- -

:::>

0 w

~ 0.2

0 0 100 200

AVERAGE UNDRAINED SHEAR STRENGTH, Cu,kPa

Fig. 3 Reduction factor a vs. undrained shear strength for piles in clay. (After Australian Code, 1978)

1 • 1 . 1 • 2. 3 SBN 7 5 ( 1 9 7 5)

The bearing capacity is calculated by

= a o u P... ,...;::,

where ^{A =} shaft area of the pile

s ₌ shear strenght determined by fall cone

CU test or vane test

a = 0.5, 0.8 and 1.0 for steel, concrete and timber piles respectively according to Broms' recommendation (1972).

1 2

For a pile in tension the maximum load is 40 kN.

- SBN 75 does not allow the use of the upper 20% (or at least 3 m when calculating the bearing capacity of a floating pile).

- If a load test has been carried out at a constant rate of penetration (CRP), the maximum allowable load corre- sponds to 2/5 of the ultimate load with respect to soil failure (FS=2.5).

- When the bearing capacity is calculated from the un- drained shear strength the safety factor is equal to 3 • 0 .

A load test is carried out if the pile Class A is used.

(Q >600 kN). The procedure for pile driving tests is a 11ow-

explained in Report No 59 of the Commission on Pile Research.

1.1.1.2.4 Danish Standard

The value of a varies with the pile material as follows:

Q = ac A s u s

Timber pile a= 0.4 Concrete pile a= 0.32 Steel pile a= 0.28

The Danish standard requires that partial coefficients in failure analysis is used. With a normal combination of the loads (dead load+ live load+ snow or dead load + live load) the following partial safety factor should be applied.

tan<jld

=

_~ 1 tan<jlf (tan<jld

=

use in design) c·u

C

=

_{1. 75}

SGI Varia 65

and the bearing ca~acity of the pile:

FS = 2 withou~ :toad testing__-- FS = 1 ,.6 with load testing

1.1.1.2.5 Building Code of the Soviet Union (Luga, 1965) see Kezdi (1975)

The maximum allowable load of the pile is calculated by:

P = nm (0Eaf. 1. ⁺ A q)

max is i pp

where n

=

coefficient reflecting the scatter of the physical characteristics, usually n

=

^0.7

m

=

¹ for buildings, for bridges and hydraulic structures, see Table 3

0

=

perimeter of the pile

a

=

factor of safety (see Table 4) f. lS

=

specific value of mantle

(see Table 5)

friction (Mp/m^{2 )} 1. l

=

thickness of i:th l~yer

A p

=

cross section of the pile tip qp

=

ultimate value of

(Table 6)

specific tip resistance

d

C=1rd A = 1rcl' /4

Fig. 4 Data for the pile formula published in the Soviet Building Code.

Table 3 Values of coefficient m

Structure

High piling Low piling

Table 4.

1-5 0.80 0.85

Number of Piles

6-10 0.85 0.90

0.90 1.00

31 1.00 1.00

Values of the coefficient a.

Vibrated Pile

Type of Pile Driven Pile Sand Coarse Silt Silt Clay

Sol id Pile 1 1.1 0.9 0. 7 0.6

Pipe pile 0.9 1.0 0.9 0.7 0.0

Table 5 Maximum unit values of mantle friction,

Sand, Fine Sand Silts and Clay Consistency Index le Average

Depth Coarse

of Layer, to Rock

m Medium Fine Flour 0.8 0.7 0.6 0.5 0.4 0.3

1 3.5 2.3 1.5 3.5 2.3 1.5 1.2 0.5 0.2

2 4.2 3.0 2.0 4.2 3.0 2.0 0.7 0.7 0.3

3 4.8 3.5 2.5 4.8 3.5 2.5 2.0 0.8 0.4

4 5.3 3.8 2.7 5.3 3.8 2.7 2.2 0.9 0.5

5 5.6 4.0 2.9 5.6 4.0 2.9 2.4 1.0 0.6

7 6.0 4.3 3.2 6.0 4.3 3.2 2.5 1.1 0.7

10 6.5 4.6 3.4 6.5 4.6 3.4 2.6 1.2 0.8

15 7.2 5.1 3.8 7.2 5.1 3.8 2.8 1.4 1.0

20 7.9 5.6 4.1 7.9 5.6 4.1 3.0 1.6 1.2

25 8.6 6.1 4.4 8.6 6.1 4.4 3.2 1.8

30 9.3 6.6 4.7 9.3 6.6 4.7 3.4 2.0

35 10.0 7.0 5.0 10.0 7.1 5.0 3.6 ^2.2

' 2

ton/m

Screw and Bored

Piles 0.8 1.1 1.3 1.4 1.5 1.6 1.7 1.8 2.0 2.2 2.4 2.6

14

If the pile has an enlarged base, qp has to be multiplied by the factor given in Table 7.

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Table 6 Ultimate value of specific tip resistance,ton/m²

Depth of Pile Tip, m

4 5 7 10 15 20 25 30 35

Granular Soils Gravel Coarse Sand Medium Sand Fine Sand Cohesive Soils

le 1.0 0.9 0.8 0.7 0.6

820 530 380 280 180

880 560 400 300 190

950 600 430 320 210

1050 680 490 350 :::40

1170 750 560 400 380

1250 820 620 450 310

1340 880 680 500 340

1420 940 740 550 370

1500 1000 800 600 400

Table 7 Reduction coefficients for enlarged pile bases

Soil Type Beneath Base Ratio of

Base and Shaft Lean Clay Clay

Diameters Sand Coarse Silt le~ 0.5 le~ 0.5

1.0 1.00 1.00 1.00 1.00

1.5 0.95 0.85 0.75 0.70

2.0 0.90 0.80 0.65 0.50

2.5 0.85 0.75 0.50 0.40

3.0 0.80 0.60 0.40 0.30

Coarse Silt

0.5 120 130 140 150 160 170 180 190 200

The following simple formula estimates the bearing capacity of traditional piles under usual conditions:

in plastic clay Qf

=

^{3 A} _s

in mixed soil Qf

=

^{6 A} _s

in sand and gravel Qf

=

^{1 0 A} _s

where A is

s the mantel surface in m2 and Qf the failure load in ton.

1 • 1 • 1 • 2 • 6 Experiences from Thailand

Based on the test loading of piles in Bangkok clay:

Holmberg (1970) has obtained a relationship between the adhesion factor a and the undrained shear strength c as shown in Fig. 5. According to a new study of

u

Balasubramaniam et al(1981)this relationship is still _I recommended for practical purpose.

l-1 0

.j.J

u rd 4-l

Fig. 5

, 2 1 - - - - . - - - - . . . - - - - , , - - - , - - - - , - - , - - - ,

o 0 25 - 30 cm. wooden piles.

1 0 1 - - - - . - - - + - - - - ' - - - - f - - - - f - - - - l - - - l

)(!Concrete

T'

^{o 22 l 22} cm, prestressed concrete

pile.

09

o.a l - - - + - ' . f - - - - J - - - . , 1 - - - . , 1 - - - . , - - - - 1

\

~✓Wooden piles 01

o Octagonal ( 0 58 cm} conc,ele pile.

a 35 x 35 cm reinforced concrele pile

( ) lndicoles lime interval in weeks between piling and load teslinQ, (4)1\~2

)

0 6 1 - - - t t - < . - - - i - - - + - - - - 1 - - - 1 - - - 1

~-~---

04

(3)

•uJ 03

Upper and lower (:(¥values of wooden piles correspond 10 a shape factor of s = I O and I I :espe..:w,f!ly.

xs-10),. ( 2 ) - = = (2).;.. _ _ (1)(9) - -

(I)

0 2 1 - - - - - ----+----+---+----+---+---+----+---+---t

0 I -

4 8 10 12 14 16 18 20 22 24 26 28 30 32 34 36 38

( I I J)

Average c (field vane tests) u

Relationship between adhesion factor and shear strength of the clay. (After Holberg, 1970).

1.1.1.2.7 Broms' recommendation (1972)

The skin friction resistance is calculated by

Q = ac A

s u s

The adhesion f s = ac u is evaluated from the standard test method and th_e adhesion factor according to Table

..t •

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1 6

TABLE Sa. Adhesion factor

a) C-u < 50 kPa Adhesion factor ^(a)

Steel piles 0.5

Concrete piles 0.8 Timber piles 1. 0 b) ^C > 50 kPa C

u

Steel piles 1 0 kPa Concrete piles 30 kPa Timber piles 50 kPa

ExperieBces in Sweden indicate that the undrained shear strength of a clay is often overestimated by the standard test methods (fall cone tests, vane tests, or unconfined compression tests) when the liquid limit or the fineness number exceeds 80 (LL ~ fineness number, Karlsson 1961).

The undrained shear strength for clay is generally re- duced in Sweden as follows:

TABLE 8b. Reduction coefficient

Fineness number Reduction coefficient (Approx.equal to LL)

80-100 0.9

100-120 0.8

120-150 0.7

150-180 0.6

>180 0.5

The bearing capacity of a pile which has been driven into a normally consolidated clay is approximately equal to the calculated, when the undrained shear strength has been evaluated by fall cone tests.

1

e

If the undrained shear strength is determined by uncon-

fined compression tests the critical load is underestimated.

The table Ba shows that the upper limit of the unit skin friction resistance is equal to 50 kPa. When c < 50 kPa,

u

the unit skin friction resistance is approximately equal to the undrained shear strength according to Broms (1972) and Tomlinson (1957).

1.1.1.2.8 Vesic (1977)

For overconsolidated clay Vesic has recommended a = 0.45

1.1.1.2.9 The CTH method (1979)

Bengtsson et al (1979) state that the shear strength should be determined by vane tests because

- lower cost than fall cone test

- more reliable value of the shear strength than the fall cone test at greater depth.

No consideration should be taken to the pile material because results of Torstensson (1973) show that piles of different materials but with the same shaft area, shape and dimensions have approximately the same bearing capacity.

The surface of failure occurs at a small distance from the pile in clay.

The shear strength of soft, highly plastic clay is de- pendent on the rate of deformation. The shaft resistance for a cylindrical pile one month after installation was equal to 0.9 times the undrained shear strength determined in a field test with the same time to failure.

SGI Varia 65

Torstensson (1973) showed that the shear strength deter- mined with the field vane test varies with the time to failure:

where

T /T = 1.21. (tto )-0.053 er o

T = critical shear strength er

t₀ = time to failure in a standard test (1 min) t = any time to failure if t = 3 h, T = 0.9T

er o

The displacement modulus can be calculated by Butterfield et al (1971) and Torstensson (1973) for the case that the pile base is negligible. The ratio of the stiffness of the pile and soil is calculated and used in Fig. 6a.

0.7

! I ^l

I! I

1L/d ^{I l}

^{: I}

^I

0.6 ^{I l ! I} : I ^l !

I! I

I

^{2oi ! r}

Ii

0.5 ^{I I! 1 i I ' I ! i}

14o,5

d

^I I ¹ I t l ^{I ! I}

0:4 _{• MOOI} ^{i 60! ! , : I}

I :i~

l;

l

! i '! i

0.3 ^I

I I/ ^{1 f} I ^I _{I : .} I'

(!)(/) Q2 i I

i:

_' ^{: ] l 1}

I l ^I I I

...____,

-a 0.1 _I

' Ii! 1

(/) '

~ ^{I I !} I

0 ^{l ' '}

2 2 5 2 5

10³ 104 105 ₁₀⁶

Ep/Gs

Fig. 6a Diagram for determination of the initial dis- placement modulus.

Ks to be used for calculation of load/displacement curve.

d = equivalent pile diameter G = shear modulus of the clay

Es= equivalent Young's modulus for the pile op= Tshaft/Ks

(After Bengtsson et al, 1979)

20

From results of tension tests on floating piles, Torstens- son (1973) presented a normalized shaft shear stress dis- placement curve (Fig. 6b).

1.0

Q85

0.5

---- -7---1--- A

-

--j

_1---I ^! _{I I}

I ~---==-====.r---J B

i r---1

^C

I I

CU () '-..

. I

I I

I I

I- 0 O ^{025 C.6} _{1 2 3 4} 5

6/8f

Fig. 6b Idealized relationship between shear stress ratio

.h/c ) along the pile shaft surface and relative

a

displacement (o/of) of the pile with respect to the surrounding soil.

= friction resistance

of= relative displacement at failure Ca

A = curve representing conditions at a low rate of displacement

B,C= curve representing conditions at a high rate of displacement.

(After Bengtsson et al, 1979)

The procedure of calculation is:

1. Determine with help from Fig. 6athe initial displace- ment modulus K from given data

s

a) the length of the pile L b) the diameter of the piled

c) Young's modulus of the pile material E

If the pile is nonhomogeneous E = crois sectional axial stiffness divided by nd²

/i

d) Shear modulus of the soil G

s

For normally consolidated, soft, highly plastic clays in Sweden Gs = 15 0 c u

SGI Varia 65

2. The complete relation between the shaft shear stress and the displacement can be calculated by Fig. 6b.

Note:

a) The failure load is calculated by

Q = f · 0L f s

f s 0L

0.9

vane

= 0.9 Tfu tf

= shaft area of the pile

= the mean value of failure shear strength from vane test

= factor due to time to failure (=0.9 if time to failure= 3 h)

= due to the time of installation ⁽¹ month after installation)

b) The value of the displacement o can be evaluated from K

o

s = Tshaft

where T is the shaft friction, assuming that the shaft

displacement o a t a load equal to½ bearing capacity is calculated by Ts/2 where Ts is the failure unit skin friction of the pile. The initial modulus cannot be used to directly calculate the value of ofailure.

c) To account the variation of if the shear modulus and the undrained shear strength are not constant a finite difference program can be used.

3. Simplified calculation method

As results from load tests show the displacement of the tip is less than 25% of the displacement of the pile head for a load in the permissible range. There- fore, i t can be assumed that the pile tip does not move or the axial deformation of the pile is equal to

the pile displacement. The axial force can be expected to fall between the type of stress distribution 7a and 7b.

Fig. 7

where

z ^{, r}

. . _{. .}

. . . . . . _{. .}

. .

T

~

. . . _. . _. . . . . . . _. . . _.

T

.··

z

z ~,

Paxial

. .

..· .

. . . . . . . .

. .

. .

Paxial

. .

.

. . . . . . . . . .

Distribution of shear stress and axial load for piles where

22

a) the bearing capacity of the pile tip is neglected.

The dashed lines represent typical shear stress distribution and distribution of axial force obtained during load tests.

b) The bearing capacity of the shaft area is neglected. (After Bengtsson et al, 1979) Pl 0.5 c5 Pl

EA < elast < EA. ^{1 •} 0

p ₌ axial load on the pile head 1 = length of the pile

E = Young's modulus for the pile material A = cross-sectional area of the pile

SGI Varia 65

Experience from behaviour of piles in soft, highly plastic clays in Sweden shows that the point bearing capacity of those piles is less than 10% of the total bearing capacity.

Bozozuk (1979) has recommended that this method is use- ful for primary design, for detail design i t is necess- ary to carry out load tests.

1.1.1.2.10 The method of Caquot and Kerisel (1956)

= A f s s

A s = shaft area of the pile f ⁼ unit skin friction

s

f ⁼ T in clay for ^{cp =} 0 s max

T ^{= C} + 100 c² max 100+7c² and for ^cp > 0

f = T +T¹ s max max

where ,, , _" max ₌ c ( 1 +sin'!' e . ^{,i., ) (} TI/ 2 + cp ) tan cp

The relation of T /c is presented in Table 9.

max TABLE 9.

cp ⁰ 1 0 1 5 20 25 30 35 40

Relationship between cp and T /c.

max

Tmax/c 1. 06 2.06 2.70 3.62 5.01 7.27 10.30

1 . 1 • 2 Effective stress analysis for bearing capacity of iles

It is recommended that the bearing capacity of piles is calculated by effective stress analysis because:

.2 4

- skin resistance of piles is governed by the effective stress conditions around the shaft, the increase in bearing capacity of the friction piles in clay is essen- tially a phenomenon'of radial consolidation of the clay.

The gain in resistance with time should be controlled by the time factor Th defined by

in which eh is the coefficient of radial consolidation and t is the elapsed time since pile driving and B the diameter. Available field data on the subject are

assembled in Fig. 8 after Ve sic ( 19 7 7) .

The method of installing the pile and the sequence of strata through which a pile penetrates has an important effect of the relationship between available shear re- sistance and undrained shear strength. A larger amount of scatter about the average values is shown in Fig.10 after Platte et al (1977). Fig.9 after Vesic (1977) also shows that no correlation exists between skin resistance and undrained shear strength.

The variation of skin resistance of piles in clay could be better understood if test results are interpreted in terms of effective stress and the equation

f ⁼ K tano'a'

S S V

The main difficulty in applying the effective stress approach is to estimate the radial effective stress on the pile at failure and the evaluation of Ks in the above formula, or

SGI Varia 65

u. 0

Fig. 8

Length Type Dia. ft. Soil

type

- - - -

~} steel H 14 {191}

219 D. steel pipe 6 A steel pipe 12

@l precas t 14

@( concrete :ts~el

of

Plpe 24

22

60

{m

1242} ³¹⁶

300

silt soft clay soft clay soft boulder

clay soft to

stiff clay

TIME , SINCE DRIVING (days)

Location Source

Tappan Zee, N.Y. Yang 1956 San Francisco

Michigan Horten Quay

Eugene Island

/

Seed & Rees~, 1957 Housel 1958 Bjerrum et al., 1958

}Mcclelland, 1969 Stevens, 1974 - - - - (theoretical prediction)

Field data on increase of bearing capacity with time for friction piles in clay. (After Vesic, 1 977)

the state of stress around the pile and in the pile itself.

In the effective stress analysis, the end-bearing capacity is related to the effective friction angle of the soil and the vertical effective stress in the ground at tip of the pile. The skin friction is related to the coefficient of friction between the pile and the soil and to the normal horizontal effective stress. The ultimate load is defined by

Q

=

A f + A q

=

Q +Q

f s s p"p s p

where f , q are unit shaft resistance and point resistance s p

respectively, evaluated from effective stress criterion.

N .µ

....____ 4-l i::

I.S _I

SOURCE OF DATA:

S.. S...l.h{l95n B • a; .. ,_ (19Sll E • E;.i. ... ¹ (1961!

G • Gol.., (1913) Gol.., l L'°"°'d {1914) L .L,&5-,«(1~) w • w..,.1,o1 (19SJ) R • Rot.lift & TOMliRSOft (19SJ)

26

I

SYMl!Cl.l:

0 CAST-IN SITU CONCRETE PILE 0 DRIVEN COHCRETE PLE 0 STEEL PLE 0 TIMBER PILE

.. ..

.µ 0 1.0 ~ s

. s.--

⁽¹⁹⁵⁹⁾

^..

(I)

0 i:: m

.µ Ul

·r-l Ul 0.1 (I)

H i::

·r-l

~ Ul

Fig. 9

Fig. 1 0

T • TOfllllU\MWI (1953)

u • U.S.,-, Watorwoy, E,... Ito. (1950)

..

w • Woocfwcnf et al (1961)

..

^h

.. ..

.. _{.. ..} .. .. _.. .. _..

,:

.. ..

.. ^.. .. ..

- .. .~ ... _...

''al

.. ^{.. ..}

a, ••

^.. ...

r ...

.. .. .. - ^{.. . ..} .. ^. . . _. . ..

..

=..- ... : ... 1

.

^{., g:} _••.• a:..

..

II a: ⁸¹

-·

- ^..

^,:

..

^,,,,

_{. ..} ..

.. ..

^{.. :r:r.::}

-

,

., ..

... -a.

.. .. .. . ^.

1 _z

"' z 0

5 ~

w .;; 0 w "

<(

0: w

>

<(

0.1

undrained

1.0 I.I 2.0

shear strength(ton/ft) 2

2.S

Comparison between skin resistance of pile in clay and undrained strength (After Vesic, 1977)

35 •

30 •

•

25

.

_0,. ^/:,~

20

•

•

15

10

(FLATTE AND SELNES, .)

10 15 20 30 35

MEAN UNDRAINED SHEAR STRENGTH, KN/m2

Observed side friction versus undrained shear strength.

SGI Varia 65

1.1.2.1

where

Fig. 11

End-bearing capacity

a' V =

= N o'A q ^V p

effective vertical stress in the soil at the tip of the pile

N q ⁼ bearing capacity factor (Berezantzev et a 1 1 9 6 1 , Ve s i c ( 1 9 6 3 ) ) ( s e e F i g.. 1 1 )

z

tJ' H 0 +l u

~ m

:>-i +l

·r-1

A p ⁼ cross-sectional area of the tip of the pile

10.000,- - -- - - -- --r--- - ~ -- - ~ - ~ - ~

1000

De Beer.

J;iky Meyerhof

Beresan tsev.

v~siC

U 100

m P.,

m u

b,

~

·r-1 H m

Q)

P'.l

TL"rzaghi

10~ - - -~ - -- - < -- - - - ' - - - - -- - ' - - - '

25 30 .15 40 45 50

Angle of internal friction,degrees

Bearing capacity factors vs. angle of internal friction, according to various authors.

(After Kezdi, 1975)

In compressible silty clay, the bearing capacity factor has a value of about 10 (Blanchet, 1979) when the pile ends in a saturated clay the above equation gives a

reasonable estimate of the point resistance of the pile, (Bozozuk, 1979).

28

Vesic (1975₁ 1977) has been working on the expansion theory, and has recommended the following formula for calculation of the point resistance

Qp = (cN* + a N*)A

C O q p

in which N* and N* are appropriate factors, related to

C q

each other by

N~ = (Nq-1) cot<jl

The value of N* and N* are functions of the angle <jJ and

C q

the rigidity index (I ) (See Appendix A) rr

a : means normal soil stress

0

a 0

= 1+2K_{0 0 ,}

3 V

a': the effective vertical stress in the soil at the

V

foundation level.

SGI Varia 65

1.1.2.2 Shaft friction resistance

The effective unit shaft resistance fs on a pile in homogeneous clay is given by

where

f = c' + K 0 ' tan

o '

S a ^{S V}

c' = unit adhesion between the soil and the pile, a which is independent of the normal stress

K s ⁼ earth pressure coefficient on the pile shaft

() I = effective angle of friction between the pile

and the soil

Clark and Meyerhof (1972) measured the friction between the soil and a steel plate and showed that as the shear rate was reduced, c' decreased, and in a drained test i t _a became equal to cero (Fig.12). Bozozuk (1979) came to the

0.,

.

5 . . . - - - ,

4

3

z

(o} • Und<aint<I Tut Normal Su<fact SIMI

I

llHt:AA RATE •.0-44

;q

!

aHU!l RAT!:• .016

,.,...1

H s, .. 1

rd

..c: (1) rJl r-1

rd ~ o~i_-1.._L-...L-L_j_.J.._-'---'-.,__.___,__-'---I

:s

(d) Orolnod Tut

rJl (1)

~

z

I Z 3 4 5

Normal pressure p.s.i.

Fig. 12 Residual strength results of skin friction test.

a,b,c: undrained test; normal surface steel.

d: drained test. (After Clark and Meyerhof, 1972)

30

same conclusion in a soil pile friction test. Meyerhof (1976) and Vesic (1977) suggest that c~ can be neglected and

f = K 0¹ tano'

S S V

or f = S0'

S V

Some criteria on the calculation o f f are summarized below. s

1.1.2.2.1 Burland

Burland (1973) followed Chandler's (1968) approach and suggested the equation:

Q = A f sf s s

where f

=

K 0¹ tan~d

=

S0'

S S V V

K = earth pressure coefficient.

s

For driven piles, K

s

K = K (safe side),

S 0

clay:

K = 1-sin~'d

0

> K, so it is assumed that

0

and for normally consolidated

the effective overburden pressure

the remoulded drained angle of friction of the soil

(According to Tomlinson (1971), i t is assumed that the failure takes place in the remoulded soil close to the shaft surface so o~ ~d).The reduction factor Scan be written:

S = (1-sin~d)tan~d

Sis not very sensitive to clay type.

For normally consolidated clays S = 0.24-0.29 (~=20-30°).

SGI Varia 65