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Structural Engineering Branch, ArchSD Page 1 of 71 File Code: Friction Piles.doc Information Paper on Small Diameter Frictional Piles CTW/MKL/CYK/KWK/LPL Issue No./Revision No. : 1/ Issue/Revision Date : March 2013 Information Paper Design of Small Diameter Frictional Piles and Cases Study STRUCTURAL ENGINEERING BRANCH ARCHITECTURAL SERVICES DEPARTMENT March 2013

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Page 1: Information Paper on Small Diameter Friction Piles

Structural Engineering Branch, ArchSD Page 1 of 71 File Code: Friction Piles.doc

Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

Information Paper

Design of Small Diameter Frictional Piles

and Cases Study

STRUCTURAL ENGINEERING BRANCH

ARCHITECTURAL SERVICES DEPARTMENT

March 2013

Page 2: Information Paper on Small Diameter Friction Piles

Structural Engineering Branch, ArchSD Page 2 of 71 File Code: Friction Piles.doc

Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

CONTENTS

Content Page

1. Objectives............................................................................................................. 3

2. Background ........................................................................................................ 4

3. Load-Carrying Capacity of Frictional Piles - a Summary ............................. 6

4. In-Situ Measurements in ArchSD Projects..................................................... 30

5. Summary of Findings ....................................................................................... 48

6. Loading Tests ................................................................................................... 50

7. Pile Group Settlement ...................................................................................... 51

8. Method of Procurement ................................................................................... 53

References

Annex A Estimation of the Length of Piles

Annex B Sample Particular Specification for Design and Construction of

Frictional Mini-Piles

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Structural Engineering Branch, ArchSD Page 3 of 71 File Code: Friction Piles.doc

Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

1. Objectives

1.1 Traditional small diameter frictional piles, which are usually drilled and cast in-

situ piles, are especially suited as foundations for sites with difficult access,

congested or where minimal disturbance to the existing structure is required.

Though Clause 5.3.2 of Code of Practice for Foundations 2004 (“Foundations

Code 2004”) issued by Buildings Department states that the allowable bearing

capacity for such non-driven piles may be determined by the allowable bearing

pressure and bond or frictional resistance of the ground, it further states that

unless for piles socketted into rock, the load-carrying capacity of the piles

should not be derived from a combination of the shaft resistance and end

bearing resistance of the piles unless it is justified that the settlements under

working load conditions are acceptable and adequate to mobilise the required

shaft resistance and end bearing resistance of the piles simultaneously. As such,

traditional mini-piles in Hong Kong are designed to be socketted into rock, and

their allowable capacity is derived solely from the average bond strength

between the grout and rock. An average bond strength is usually adopted for

piles socketted into rock, and the study carried out by the University of Hong

Kong (Department of Civil Engineering 2009) confirmed that most of the axial

load transmitted to the rock socket is dissipated at the top portion of the socket.

1.2 This Information Paper, besides reviewing the design and construction of

different types of small diameter frictional piles, introduces a non-traditional

piling system – the “frictional mini-piles”. All such small diameter frictional

piles, unlike traditional mini-piles, derive its load-carrying capacity from shaft

friction from the soil. The frictional mini-piles are constructed with steel I-

section or a group of reinforcement bars in a pre-bored hole with a temporary

steel casing and then injected with cement grout. Similar type of frictional piles,

constructed with steel I-section in a pre-bored hole by continuous flight

augering (CFA) and then injected with cement grout, had successfully been

employed in four ArchSD sites in the 1990s, including a site in Tung Chung, a

site in Yuen Long and two sites in Ma On Shan. A distinctive feature of the

frictional mini-piles is that the pre-bored hole, instead of forming by CFA, is

constructed with Odex or similar method with a temporary steel casing.

Recently, ArchSD has successfully employed such frictional mini-piles as

foundations in two projects – one in Mid-Levels and the other in Central.

Instrumented piles were also installed these two projects to monitor the stress

distribution along the piles. This Information Paper provides:

a) a literature review on the design of piles deriving their load-carrying

capacity from shaft resistance from soil;

b) summary of the results of instrumented piles in some ArchSD projects;

c) the design shaft friction to be adopted for the design of the different types of

small diameter frictional piles (including the frictional mini-piles); and

d) a particular specification on the design and construction of the frictional

mini-piles.

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Structural Engineering Branch, ArchSD Page 4 of 71 File Code: Friction Piles.doc

Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

2. Background

2.1 Typical sub-soil profile in Hong Kong consists of a layer of loose fill overlying

marine deposit and alluvium. Completely weathered soil then follows before

reaching Grade III or better bedrock. For low-rise development, shallow

foundation in the form of pad and/or raft footing is usually adopted. For

medium-rise or high-rise development, founding the building on topmost fill or

alluvium will result in excessive settlement of the building. Deep foundation in

the form of piled foundation is therefore required.

2.2 Approved systems of piles in ArchSD can be classified into replacement piles or

displacement piles. Replacement piles include non-percussion cast in-situ

concrete piles (e.g. PIP piles); large diameter bored piles; pre-bored rock-

socketted steel H-piles, barrette piles; mini-piles founding on bedrock; and

hand-dug caissons (which has been banned) may only be used for public works

under the conditions imposed by Works Branch Technical Circular No. 9/94

(available: www.devb.gov.hk/). Displacement piles include: precast concrete

piles; precast prestressed tubular piles (e.g. Daido, SS piles); driven steel H

piles; and percussion cast in-situ concrete piles (e.g. Frankie piles, Vibro piles).

2.3 Among these approved systems of piles, driven steel H-piles (a small

displacement piling system) are one of the most popular and economical piling

options in Hong Kong due to the quick installation time and tidy site condition.

However, noise and vibration are particular concerns for some sites, and for

some sloping sites the driving operations also require the construction of heavy

temporary platforms. Replacement non-percussion piling systems are then

adopted. Common systems of non-percussion end-bearing piles include: large

diameter bored piles, pre-bored rock-socketted steel H-piles, and mini-piles. In

fact, these piles derive the resistance mainly from their end-bearing on hard

stratum and partly from the shaft friction between the soil and the pile shaft.

However, there are uncertainties on the shaft friction between the soil and the

pile shaft. Furthermore, if the rock end bearing stratum is available at a

reasonable depth, the shaft friction component is small when compared with the

end bearing component in the overall load carrying capacity of a pile. As such,

for traditional replacement non-percussion piles (e.g. pre-bored rock-socketted

steel H-piles, mini-piles, or large diameter bored piles) the shaft friction

component is usually neglected.

2.4 However, on some sites in Hong Kong (e.g. in Mid-Levels, Tung Chung, or

some newly reclaimed land) sound bedrock can only be found at very deep

level, e.g. more than 60m from the ground level. In such cases, a cost effective

solution is to adopt replacement piles relying on the shaft friction component.

Shaft friction can be developed after small relative displacements between the

soil and the pile shaft though may reach ultimate shortly after that, and hence

shaft friction component often contributes the bearing capacity in the working

load situations.

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Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

2.5 Besides PIP piles, frictional mini-piles have now been employed commonly in

Hong Kong for those sites with bedrock at great depth. Frictional mini-piles are

a hybrid system combining traditional mini-piles with replacement non-

percussion cast in-situ piles. Traditional mini-piles consist of a steel permanent

casing with internal diameter not greater than 400 mm, with a group of

reinforcement bars in the middle as the load bearing element and the remaining

cavity filled with cement grout. They are required to be socketted into bedrock,

and hence derive their load carrying capacity from end-bearing on the bedrock.

The frictional mini-pile is constructed with steel I-section or a group of

reinforcement bars in a pre-bored hole with a temporary steel casing and then

injected with cement grout. The frictional mini-pile is not required to be

socketted into bedrock, and as such it behaves similar to the other non-

percussion cast in-situ piles deriving their load carrying capacity from the shaft

friction along the length of the piles. Moreover, steel casing is only temporarily

provided for pre-drilling and will be removed during the subsequent grouting

work.

2.6 The advantages of frictional mini-piles are that they are especially suited as

foundations for loading that is not high and for the sites that are with difficult

congested access (Figure 1), or requirements for minimal disturbance to the

existing structure, or bedrock can only be found at very deep level.

Figure 1 Congested site with difficult access

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Information Paper on Small Diameter Frictional

Piles

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Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

3. Load-Carrying Capacity of Frictional Piles –a Summary

3.1 Figure 2(a) shows the forces acting upon an axially loaded pile. Figure 2(b)

shows the typical relationship of shaft resistance Rs and end bearing Rb

components of the pile founded on soil with the settlement of the pile. In theory,

by integrating the mobilised shaft friction fs over the surface of pile can give the

total shaft resistance Rs.

Figure 2(a) Stresses and forces on an axially loaded pile

Figure 2(b) Typical shaft resistance and end bearing versus displacement

in a pile founded on soil

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Information Paper on Small Diameter Frictional

Piles

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Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

Because relatively large displacements are required for piles founded on soil to

mobilise the end bearing capacity in normal range of acceptable settlement

criterion (Figure 2(b)), the ultimate bearing capacity of such a frictional pile

may develop up to 80 – 90% of its capacity through shaft friction (Holt et al

1982, Kwok 1987). However, the design parameters for the shaft resistance

along the length of pile also show great variation. The pile-soil interface shear

friction, besides determined by the stress history of the soil, is affected by the

following key parameters (Brown et al 2007):

a) the construction method;

b) the shear displacement of the soil at the pile-soil interface;

c) the in-situ soil properties (e.g. soil composition, water content, saturation,

stiffness and strength);

d) the concrete/grout properties (e.g. composition, viscosity, pressure, stiffness

and strength).

3.2 Effect of construction methods on shaft friction

3.2.1 Among the parameters listed in the above paragraphs, construction method

affects the shaft friction substantially. That is, despite of the same soil and grout,

piles constructed with the different construction method can have significantly

different shaft friction (Kay and Kalinowski 1997; Lo and Li 2003). The

construction method affects the relative volume of soil displaced in proportion

of the pile volume, the magnitude of the increase in the effective horizontal

stress at the pile-soil interface, the relative roughness of the pile-soil interface

(Figure 3(a)), and the effective diameter of the pile (Brown et al 2007). In

Hong Kong, pile driving by means of hydraulic hammer is commonly adopted

to install displacement piles such as driven steel H-piles. For replacement piles,

the following three methods are commonly employed to form the holes:

a) for large diameter bored piles, various excavating tools such as grabs,

chisels are used to form the holes within a temporary casing;

b) for PIP piles, CFA is used to form the hole;

c) for mini-piles or pre-bored rock-socketted steel H-piles, Odex method is

usually used to form the holes.

(a) smooth interface (b) rough interface

Figure 3(a) Roughness at pile-soil interface

(Source: Rollins et al 2005)

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Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

3.2.2 Besides the difference in the method to form the holes, Armour et al (2000), as

illustrated in Figure 3(b), classifies the different concreting/grouting operations

for replacement piles. In Type A, concrete/grout is placed under gravity head

only. In Type B, grout is placed into the hole under pressure at around 0.5 to

1MPa as the temporary steel drill casing is withdrawn. In Type C, a two-step

process of grouting is employed with cement grout placed under gravity head as

with Type A and prior to hardening of the primary grout (after approximately 15

to 25 minutes), injection of grout via a sleeved grout pipe at a pressure of at

least 1MPa. In Type D, a two-step process of grouting is employed similar to

Type C with grout in the second step injected via a sleeved grout pipe at a

pressure of 2 to 8MPa. A pair of double packers is usually used inside the

sleeved pipe so that specific horizons can be treated several times. Among the

different construction methods, it can be expected that the shaft friction

increases from Type A to Type D construction.

Figure 3(b) Classification of construction method for replacement piles

(Source: modified from Armour et al 2000)

3.3 Shaft friction for different types of frictional piles

3.3.1 Section 3.2 describes the differences in forming the hole and concreting/

grouting operations for installing replacement piles. It should be noted that the

different construction methods affect the effective horizontal stress on the pile-

soil interface, and hence the shaft friction. Figure 4 summarises their load-

settlement behaviours of piles relying on pile-soil friction, including driven steel

H-piles, augered piles and bored piles (O’Neill 2001). Two lines have been

added to Figure 4 to show the expected load-settlement behaviour of frictional

mini-piles with and without post-grouting along the pile shaft.

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Information Paper on Small Diameter Frictional

Piles

CTW/MKL/CYK/KWK/LPL

Issue No./Revision No. : 1/ Issue/Revision Date : March 2013

Figure 4 Load-settlement graphs of common frictional piles

(Source: Modified from O’Neill 2001)

3.3.2 Driven piles and bored piles

As expected, large displacement piles (e.g. driven Daido piles) are of the highest

shaft friction. During driving, the surrounding soil is displaced laterally,

causing an increase in the effective horizontal stress (Brown et al 2007). For

large diameter bored piles, the soil is removed by grabs, and there are temporary

casing throughout the installation. Concreting is of Type A (Figure 3(b)).

Therefore, the effective horizontal stress at the pile-soil interface tends to reduce

or remains unchanged during the construction (Brown et al 2007). Hence shaft

friction should be the lowest. The behaviour of small displacement driven piles

(e.g. driven steel H-piles) will lie between that of bored piles and that of large

displacement piles.

3.3.3 CFA piles

For augered piles using CFA (e.g. PIP piles), the grouting operation is similar to

Type A (Figure 3(b)). Yet, shaft friction for some augered piles (e.g.

displacement CFA piles (or termed “DD piles” in the US)) can achieve the

highest shaft friction among the replacement piles (O’Neill 1994) if suitable

plants are properly used. It is due to its distinct construction method of such

DD piles, where a large diameter CFA with greater torque is used for forming

the hole. The large displacement by the large diameter CFA with greater torque

displaces the soil laterally, and therefore tend to increase the stresses in the

surrounding soil (similar to driven piles). However, in the traditional CFA pile

construction, the shaft friction at the pile-soil interface is far more complicated

than those of driven piles and large diameter bored piles. The continuous

augering operation can maintain the effective horizontal stresses near the value

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Information Paper on Small Diameter Frictional

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that existed before the installation, and may in some cases can achieve higher

effective stresses than the preconstruction state. However, it should be noted

that the rate of auger penetration is extremely important when installing augered

piles (Brown 2005). Prolonged augering at one depth without penetration will

flight the soil surrounding the CFA (a phenomenon called “Archimedes pump”),

resulting in excessive removal of the surrounding soil by loosening and

allowing the adjacent soil to fall into the hole. In such case, the soil surrounding

the pile is decompressed and the effective horizontal stresses are then decreased,

and thus the shaft friction cannot achieve the theoretical values. PSE should

therefore note that though Figure 4 shows that augered piles can achieve the

highest shaft friction among the replacement piles, in reality traditional CFA

piles may not be able to achieve such high shaft friction and may sometimes be

smaller than that for frictional mini-pile with post grouting. For details, PSE

may refer SEB Information Paper Review of PAKT-IN-PLACE Piles

Installation (available: http://asdiis/sebiis/2k/resource_centre/).

3.3.4 Frictional mini-piles

For the frictional mini-piles without post grouting along the pile shaft, the

construction method lies between Type A and Type B (Figure 3(b)). Their

load-settlement behaviour is therefore expected to be better than (though close

to) that of large diameter bored piles, as the grout filling the pre-bored holes is

placed under a small pressure. Improvement of the shaft friction for the

frictional mini-piles can be achieved by post grouting along the pile shaft, and

the construction method will then follow Type C or D (Figure 3(b)). The post

grouting is expected to increase the diameter of the mini-piles and hence the

effective horizontal stress so as to increase the shaft friction. However, it should

be noted that pre-bored holes for frictional mini-piles in Hong Kong are usually

formed by Odex method, and hence the densification effect is not predominant.

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Information Paper on Small Diameter Frictional

Piles

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3.4 Idealised pile-soil interface shear friction

3.4.1 Numerous studies (e.g. Misra and Chen 2004; Frizzi and Meyer 2000) have

been carried out to study the relationship of shear friction with shear movement

for the soil at pile-soil interface. Figure 5 shows one of such studies plotted in

dimensionless axes of developed shear to ultimate shear resistance (f/fmax)

against the shear displacement to diameter of the pile (W/D).

Figure 5 Relationship of shear resistance with shear displacement at pile-

soil interface (Source: modified from Frizzi and Meyer 2000)

3.4.2 Theoretical model of development of shaft friction with displacement

Although the actual load-transfer mechanisms developed along the pile-soil

interface are highly complicated (Mayne and Harris 1993; Paik et al 2003; Yang

et al 2006), an idealised model for the soil at the pile-soil interface is usually

adopted (Misra and Chen 2004). Figure 6(a) plots the relationship between the

idealised pile-soil interface shear friction and shear movement u. In the

idealised model, is assumed to vary linearly with u in the elastic zone, and is

then assumed to be constant once the soil movement exceeds critical shear

displacement uc. Mirsa and Chen (2004) derive the following governing ODE

for the shear displacement u(x) at a distance x from the pile tip in the pile of

diameter D and length l:

eud

ud

0for 0)(

)( 2

2

2

1for 0)(

and 2

2

2

ecu

d

ud

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Information Paper on Small Diameter Frictional

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where k = stiffness factor (Figure 6(a)),

λ =

mk

kl ,

km = axial stiffness of the pile,

= non-dimensionless length = x/l,

and e = transition from elastic to plastic zone = le/l (Figure 7)

Misra and Chen (2004) then give the general solutions to this governing ODE

with the following shear displacement u()along the pile shaft with an applied

axial load of P :

elastic zone: 10for sinh

cosh

uP

Pu

elasto-plastic zone: e

uP

Pu

0for

sinh

cosh

1for 1

12

1e2

22

e

u

eP

Pu

and plastic zone: 10for 1

12

12

2

uP

Pu

where = Pul/km,

and Pu = ultimate load carrying capacity of the pile = Dlc.

The location of transition from elastic to plastic zones e at a given load P can

be solved from the following equation:

0cosh

1tanh1

eu

e

eP

P

For the value of uc, Figure 6(b) shows the summary of Luo et al (2000) of the

test results by Cartier and Gigan (1983), Lim and Tan (1983), Murray et al

(1980), Billam (1972), Chang et al (1977), and Taylor (1948). Their summary

indicates that, for silty soil, uc lies between 0.8mm to 5.6mm, and 2.5-5.6mm is

the mode. Thus, a shear movement of around 3 to 6mm may already cause full

mobilisation of the pile-soil interface friction.

Figure 6(a) Idealised pile-soil interface

shear friction with shear displacement

(Source: Misra and Chen 2004)

Figure 6(b) Frequency distribution table

of the test data of uc

(Source: Luo et al 2000)

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3.4.3 Zoning in a frictional pile

When a frictional pile is subjected to vertical load, the movement at the pile

head will be larger and the movement decreases along the length of the pile.

Figure 7 shows the variation of shaft friction along a pile in a homogenous

uniform soil medium using the idealised model (Wong 2003; Misra et al 2004).

When the pile is in elasto-plastic stage, the shaft friction along the length may

be idealised into two zones: plastic and elastic (Figure 7) (Misra and Chen

2004), depending the shear movement at the pile-soil interface. Elastic zone is

at the lower portion of the pile where the shear displacement at the interface is

still less than uc. Plastic zone occurs at the top portion of the piles where the

shear displacement at the interface exceeds uc. Unlike an end-bearing pile, not

the whole length of the pile (especially the portion near the pile tip) for a

frictional pile will be mobilised.

Figure 7 Zoning of soil along a frictional pile

(Source: modified from Misra et al 2004)

Figure 8 predicts the load-settlement curve in a black line of a frictional pile

obtained from the above general solution against the actual load-settlement of

three mini-piles measured by Misra and Chen (2004). The elastic, elasto-plastic

and plastic behaviour at increasing shear displacement are clearly shown.

Figure 8 Calculated shear displacement against measured displacement

(Source: Misra and Chen 2004)

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3.5 Methods to calculate shaft friction

The mobilised unit shaft friction (fs) along the pile shaft is theoretically

determined by the sum of pile to soil cohesion and friction components in the

following equation:

fs = ca + σ’h tans

where ca and s are respectively the adhesion and friction parameters between

the soil and the pile shaft, and σ’h is the effective horizontal stress due to

overburden. Numerous theoretical methods (e.g. Nordlund method (1963), α-

method (Tomlinson 1971), -method (Burland 1973; Fellenius 1991),

Nottingham and Schmertmann CPT method (1975, 1978), method based on

SPT-N values (Meyerhof 1976)) have then been developed to compute the shaft

resistance along a pile shaft.

3.6 α (total stress) method

α (total stress) method suggests that the ultimate capacity of the pile is be

determined from the undrained shear strength (cu) of the cohesive soil

(Tomlinson 1971). This method further assumes that the shaft resistance is

independent of the effective overburden pressure, and the unit shaft resistance fs

is therefore given by the following equation:

fs = ca = αcu

where α is an empirical adhesion factor to relate the average undrained shear

strength along the pile length. The factor α depends on the nature and strength

of the cohesive soil, pile dimensions, method of installation, and time effects.

Typical values of α range from 1.0 for soft clay to 0.30 for very stiff clays

(Kulhawy and Jackson 1989). α method, however, assumes slow dissipation of

water and is therefore not applicable in most of the soils in HK. Fellenius (2011)

further commented that the load transfer between a pile and the soil is governed

by effective stress rather than the undrained shear strength. In Hong Kong,

GEO Publication No. 1/2006 - Foundation Design and Construction (GEO

2006) published by the Geotechnical Engineering Office therefore recommends

the use of either -method or method based on SPT-N values (“N-value

method”), which are applicable to both cohesive and cohesionless soil.

3.7 (effective stress) method

3.7.1 (effective stress) method models the long-term drained shear strength

conditions of piles using the effective stress, and the ultimate unit shaft

resistance fs is calculated by Coulomb’s friction law using the following

equation:

fs = σ’v

where (a dimensionless coefficient) = Kstan,

σ’v = average effective overburden pressure along the pile shaft,

Ks = lateral earth pressure coefficient,

and = friction angle between the soil and the pile shaft.

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The ultimate unit shaft resistance fs is limited to 200kPa (O’Neill and Reese

1999), whilst GEO (2006) recommends a limit of 150kPa unless a higher value

is substantiated by site specified tests.

3.7.2 Table 1(a) summarises the typical values of obtained from back analysis of

field test data for different methods of installation and types of pile given in

GEO (2006). The range of given is similar to those specified in overseas

research, which gives a minimum of 0.15 (Fellenius 2011) and a maximum

value of 1.20 (O’Neill and Reese 1999; Caltrans 2008). Besides summarising

the back analysis of field test data, GEO (2006) also gives the field data of each

site in its Appendix, and it should, however, be noted that the test data (Figure

9) show very high variability. Besides the construction method, also depends

on the types of soil. Fellenius (2011) gives approximate range of values of

(Table 1(b)) for different types of soil; but cautions that the values of can

“deviate significantly from [those] values.”

Table 1(a) Typical values of for different types of piles

Pile Type Soil Type

Small displacement driven

piles (e.g. driven steel H-

pile)

CWG 0.1 - 0.4

Loose to medium dense sand 0.1 - 0.5

Large displacement driven

piles (e.g. driven Daido SS

piles)

CWG 0.8 - 1.2

Loose to medium dense sand 0.2 - 1.5

Large diameter bored piles CWG 0.1 - 0.6

Loose to medium dense sand 0.2 - 0.6

Shaft-grouted bored piles CWG 0.2 – 1.2

(Source: GEO 2006)

Table 1(b) Typical values of for different types of soil

Soil Type

Clay (25o 30

o) 0.15 - 0.35

Silt (28o 34

o) 0.25 - 0.50

Sand (32o 40

o) 0.30 - 0.90

Gravel (35o 45

o) 0.35 - 0.80

(Source: Fellenius 2011)

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(a) Replacement piles without shaft grouting

(b) Replacement piles with shaft grouting

(c) Displacement piles

Figure 9 values for piles installed in saprolites1 in Hong Kong

(Source: GEO 2006)

1 GEO (2006) defines “saprolite” as “mass that retains the original texture, fabric and structure of the

parent rock”. In Hong Kong, saprolite may be used interchangeably with “decomposed granite” (Lo

and Li 2003).

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3.7.3 The following two approaches are usually used to evaluate the value of β

adopted in Drilled Shafts: Construction Procedures and LRFD Design Methods

issued by the US Federal Highway Administration (Brown et al 2010):

1) the depth-dependent β method, which establishes an empirical relationship

of β versus depth is determined from field load tests; and

2) the rational method based on soil mechanics theory, which expresses β in

terms of Ks and δ.

3.7.4 Depth-dependent β method

The depth-dependent β method was first introduced in 1978, and was claimed to

provide conservative estimates of side resistance given the uncertainties

associated with construction effects (O’Neill and Reese 1978; Brown et al

2010). Since then, there have been a lot of field tests confirming its applicability.

Kulhaway and Chen (2007) calculated the values of from the available field

data for replacement piles in gravels and cobble soils. Their results are shown

in Figure 10. Figure 11(a) shows the variation of summarised by Rollins et

al (2005) together with two lines inserted by Fellenius (2011) for piles in sand

the values recommended by CFEM (1992) and GEO (2006). All results show

that peak value of occurs at the pile head decreasing with depth,

corresponding to the larger shear displacement at the top portion of the pile as

predicted in Section 3.4.

Figure 10 Variation of for replacement piles in cohesionless soil

(Source: Kulhaway and Chen 2007)

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Figure 11(a) Variation of for replacement piles in sand

(Source: Fellenius 2011)

Design and Construction of Continuous Flight Auger (CFA) Piles published by

the US Federal Highway Administration (Brown et al 2007) recommends the

use of the depth-dependent β method of either O’Neill and Reese (1999) method

or Coleman and Arcement (2002) method. O’Neill and Reese (1999) method

was based on a design trend line for related to the depth of the soil layer by

data fitting. As also depends on the types of soil, the method tries to relate the

types of soil to their standard penetration test blowcount (SPT-N values) by

scaling down by the ratio of N/15 for loose sand layers with N15. The

following equations for calculating the value of at depth z (in m) from the

ground level are derived:

sand with N15: = 1.5 – 0.245 × z0.5

soil with N15: = 15

N×{1.5 – 0.245× z

0.5

Coleman and Arcement (2002) method is derived from loading tests on a

number of CFA piles in mixed soil conditions consisting of alluvial, loessial

deposits and inter-bedded sands and clays in Mississippi and Louisiana, the US,

and the following set of equations for for calculating the value of at depth z

(in m) from the ground level as follows:

silty soils: = 2.27× z-0.67

sandy soils: = 10.72 × z-1.3

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Figure 11(b) shows the plot by Coleman and Arcement (2002) comparing their

derived equations against that by O’Neill and Reese (1999) for sand with their

data for silts and clays. The main difference between those of O’Neill and Reese

(1999) and Coleman and Arcement (2002) is that in the latter equations has

larger value at the pile head, which decreases rapidly with depth.

Figure 11(b) Comparison of O’Neill and Reese (1999) method with

Coleman and Arcement (2000) method

(Source: Coleman and Arcement 2002)

Using these equations, Figure 12 shows the variation of with depth for

homogenous soil profiles with N15. It shows that will be much greater for

the top soil, and will approach zero at depth around 30m. Brown et al (2010)

account for the variation of with depth by arguing that the values of Ks are

higher near the surface, where many soil deposits are overconsolidated as a

result of burial, erosion, fluctuations in the water table, capillary rise,

desiccation, etc. Brown et al (2010) argue that the effect of preconsolidation is

to increase the in-situ horizontal stress and, therefore, β, and with increasing

depth, most soil deposits trend toward a normally consolidated state, a lower

value of Ks and therefore a lower value of β. However, the validity of such

argument is to be further substantiated, especially the fact that the topmost soil

in Hong Kong is relatively loose and is seldom overconsolidated, and that

though there is preconsolidation effect due to fluctuation in the water table etc,

the increase in effective vertical stress would not be significant. On the other

hand, the variation of along the length of the pile tallies with the zoning soil

profile in Figure 7. Hence, this Information Paper suggests that the variation of

with depth may be due to the different relative pile movement at the pile-soil

interface along the length of the pile. That is, soil at the top portion of the pile

will become fully mobilised with larger relative pile movement, and it is

difficult to mobilise soil fully at lower portion.

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Figure 12 Variation of with depth in a homogenous uniform soil with

N15

3.7.5 Rational method

, being the lumped shaft friction parameter including Ks and , is a function of

the soil strength, soil stress state and its change, and the soil-shaft interface

characteristics (Kulhawy and Chen 2007). The rational method aims at

evaluating separately the parameters that are lumped into β using theory of soil

mechanics. Firstly, the friction angle between the soil and the pile shaft is

related to the friction angle of the soil. Kulhawy (1991) found that for cast in-

situ concrete piles with good construction techniques, a rough interface can

develop, giving / equal to 1.0; but with poor slurry construction this ratio can

be 0.8 or lower.

Secondly, Ks is found to depend on soil displacement, pile installation method,

pile geometry, and the stress changes caused by construction, loading, and

desiccation. Analysis of field load tests has shown that Ks can range from about

0.1 to over 5 (Kulhawy and Mayne 1990). Kulhawy (1991) reported his study of

the effect of installation methods on a number of piles, and found that the values

of Ks related their values to the effect with respect to Rankine at rest earth

pressure Ko within the ratio of 0.67 and 1. Their findings are as summarized in

Table 2. The range shows that the shaft friction developed in large

displacement driven piles is the highest, followed by small displacement driven

piles, and the lowest shaft friction should be that of large diameter bored piles.

Their findings tally with the load-settlement behaviours for these types of piles

in Figure 4. The results further shows that the soil at the pile-soil interface for

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large displacement driven piles behaves in passive stage and Ks approaches

Rankine passive earth pressure coefficient Kp, and that for large diameter bored

piles behaves in active stage and Ks approaches Rankine active earth pressure

coefficient Ka (Vesic 1977; Rollins et al 2005).

Table 2 Ks/Ko for different types of piles

Pile Type Ks/ Ko

Small displacement (e.g. driven H-pile) 0.7-1.2

Large displacement (e.g. driven Daido SS piles) 1.0-2.0

Drilled shafted piles (e.g. large diameter bored piles) 0.67-1.0

(Source: modified from Kulhawy 1991)

It was also in the past limited by the lack of data from load tests to correlate β

with parameters, such as Ks and δ. Given the uncertainties and difficulties in

relating Ks and with Ko and respectively, recent studies as published in

FHWA (Brown et al 2010) and Kulhawy and Chen (2007) have tried to relate

them with SPT-N value of the soil. The correlation of and N value is a

common practice in Hong Kong and presumably the US as well. The following

set of equations is then proposed to calculate the value of β (Brown et al 2010):

= 27.5 + 9.2 log N60

tanKtan

sin

σ'

σ')sin (1β p

v

p

where ’p is the effective vertical preconsolidation stress = 0.47×pa× N600.8

(for

silty sands) or 0.15×pa× N60 (for gravelly soils); N60 is the SPT-N value corrected

for field procedures and apparatus;2 and pa = atmospheric pressure. Brown et al

(2010) reported that in-situ tests have shown generally good agreement with

these correlations. However, factors such as cementation, aging, structuring,

desiccation, etc that may affect Ks and δ have not been accounted in the above

equation.

The rational method, though is better than the depth-dependent β method from a

soil mechanics perspective, has not addressed the relative pile movement

between pile and soil.

3.7.6 Applicability of method

Section 3.1 has identified the key parameters affecting the shaft friction, which

include construction method (including the workmanship), the shear

displacement at the pile-soil interface, site specific soil properties and the

2 N60 assumes that the SPT hammer has about 60% efficiency, and to convert the measured blowcount

SPT-N value to N60, the equation N60 = N×CN×CE×CB×CR×CS×CA×CBF×CC, where CN = overburden

correction factor; CE = energy correction factor; CB = borehole diameter correction factor; CR = rod

length correction factor; CS = sampling method correction factor; CA = anvil correction factor; CBF =

blow count frequency correction factor; and CC = hammer cushion correction factor. For details, see

Skempton (1986).

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concrete/grout properties. Two methods – depth-dependent β and rational –

have been used to evaluate the values of β. The equations based on the depth-

dependent β method only take into account for its variation with the depth (the

variation presumably due to the different shear displacement along the pile

shaft). The effect of other site specific soil properties (including soil types and

degree of consolidation) is only partially considered by classifying them into

sand, gravelly sand and soil with N15. Brown et al (2010) tried to provide the

rationale behind on why load test results indicate β is depth-dependent by

saying that the forming of the hole disturbs the soil, reducing its density and

allowing relaxation of horizontal stress. They viewed that “detailed evaluations

of in-situ strength and state of stress are not warranted because the in-situ

properties are changed by construction and the changes cannot be predicted

reliably”. With this limitation, the depth-dependent β method assumes that the

soil disturbance can reduce the soil friction angle to a lower-bound value

corresponding to the critical state void ratio. The practice of lumping Ks and δ

into a single parameter (β) and then evaluating β solely as a function of depth

therefore neglects the influence of geology, material type, and stress history. Its

use is therefore restricted to site-specific ground conditions (Brown et al 2007).

Should such relationship in one specific site be applied to other sites, the pile

length for different sites with the same load carrying capacity will be the same,

as the variation of with depth will be the same. Kulhaway and Chen (2007)

remarked that they “do not believe that a mean line should be drawn through

these data, which then would be used as a ‘design line’.”

In order to include the effect of site specific soil properties, the rational method

is an improvement of the depth-dependent β method by relating the value of β to

the site specific SPT-N values of the soil. However, the typical soil profile in

Hong Kong shows increasing SPT-N values with depth, and using the above

equations based on the rational method will give increasing values of β with

depth and this does not match β diminishing with depth as shown by load test

results. This Information Paper suggests that, as discussed earlier, relative pile-

soil movement should be considered.

Hence, a practical way recommended by this Information Paper is to consider

the site specific SPT-N values of the soil, which can take into account of the

combined effect of site specific soil properties and the overburden effective

pressure, and this together with considerations of pile-soil movement will be

discussed in the subsequent paragraphs.

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3.8 N-value method

3.8.1 N-value method (Meyerhof 1976) is more commonly adopted in Hong Kong to

calculate the unit shaft friction and base resistance of piles. It estimates pile

capacity based on semi-empirical correlation between SPT-N values results and

static pile load tests. The method correlates SPT-N value directly to shaft

friction, with different coefficients depending on whether the foundation is a

replacement or driven pile. As SPT-N values have been included in the

calculation, the type of soil (including the degree of consolidation and its

strength properties) and overburden pressure have indirectly been incorporated

into the method. As the SPT-N values are readily available in every project,

this method is very easy to use and provides a quick way to calculate the shaft

friction. The following paragraphs will discuss the semi-empirical correlation

values to be used for different construction methods.

3.8.2 Shaft friction for piles without post-grouting

3.8.2.1 Most of the test data (Figure 13) available to date in Hong Kong have been

summarised in GEO Publication No. 1/2006. These data were obtained by

load-test frictional piles to calculate the fmax/N ratio. These data give a mean

value of 0.89 for replacement piles, which is lower than values reported

elsewhere (e.g. Chang and Wong 1995; Tan et al 1998). Moreover, like the

data for , they have a high variability with a coefficient of variation (COV) of

61%.

(a) Replacement piles without shaft grouting

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(b) Displacement piles

Figure 13 fmax/N values for piles installed in saprolites in Hong Kong

(Source: GEO 2006)

3.8.2.2 Meyerhof (1976) provides the correlation factor of the average maximum

mobilised shaft friction fmax (in kPa) with SPT-N values to be 2 for driven pile

and 1 for bored piles. For augered piles, the earlier paragraphs have pointed

out that their shaft friction lies between driven piles and bored piles. In the

design of PIP piles which was a proprietary piling system patented by

Intrusion Prepakt before the 21st century in Hong Kong, owing to their special

construction method, the average maximum shaft friction (in kPa) had

traditionally been taken as fmax = 4.8×N with limiting value for SPT-N values

at about 40. The proprietary method is in fact very similar to CFA pile.

However, recent studies found that such high correlation factor (4.8N) is only

recommended for relatively large displacement DD piles, and for conventional

CFA piles the correlation factor may not be able to achieve such high values

(Brown et al 2007). In view of the experience over the years (summarised in

SEB Information Paper Review of PAKT-IN-PLACE Piles Installation

(available: http://asdiis/sebiis/2k/resource_centre/), ArchSD SEI 04/2010:

Particular Specification for Non-Percussion Cast In-situ Concrete Piles

(available:

http://asdiis/sebiis/2k/MAIN%20TOPIC/technical%20paper/frame.htm) now

specifies that the design shaft friction (in kPa) with a FOS of about 3 for non-

percussion cast in-situ concrete pile (including PIP Piles) is now taken as

varying from a maximum of 1.6×N for CFA piles to 0.7×N for piles formed by

boring with an auger and temporary casing with limiting value for SPT-N

values at about 40, and specifies further that the adopted design values have to

be further verified by trial piles before construction.

3.8.2.3 For other frictional piles, GEO (1996, 2006) provides the correlation factor of

the average maximum shaft friction fmax in (kPa) with SPT-N values for

different types of piles and methods of construction as follows:

fmax= Fgeo×N (kPa)

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where Fgeo is taken as 1.5 to 2 for small-displacement piles (e.g. driven H-piles)

for N up to about 80, and 4.5 for large-displacement driven piles (e.g. Daido

SS piles) with a limiting average shaft resistance of 250 kPa. For replacement

piles formed by boring, Fgeo is taken as 0.8 to 1.4 for N up to 200. GEO

(1996, 2006) further recommends the base resistance to be ignored in

calculating the load carrying capacity of the pile.

3.8.3 Shaft friction for piles with post-grouting

3.8.3.1 Shaft friction of a cast in-situ pile can further be increased by post-grouting

along the pile shaft. Post-grouting is a pressurised process that injects cement

grout to the interface between the surface of the installed pile and soil through

tubes embedded within the pile. The installed pile surface must be cracked

open by injecting either water or grout under high pressure. This is done after

the concrete or grout of the pile has set, but before it has gained significant

strength. An early application of the shaft grouting method overseas was in

1975, where loading tests on six shaft grouted 660mm bored piles showed

that there was an increase in shaft friction of 2.5 times that of piles without

post-grouting (Gouvenot and Gabaix 1975). With regard to the long term

durability of the effect of the shaft grouting, it was reported in a Bangkok site

that there was no loss of shaft resistance for two shaft grouted piles in alluvial

sand and clay when reloaded one year after the first load test (Littlechild et al

1998).

3.8.3.2 Shaft-grouted technique has been employed in Hong Kong for mini-piles since

the early 1990s (Lui et al 1993) and for barrettes and large diameter bored

piles since the late 1990s. In Hong Kong, post-grouting is carried out using

tube-a-manchette in stages after casting the piles. The system adopted for shaft

grouting consists of 50mm diameter mild steel tube-a-manchette pipes, with

manchettes spaced at about 1m intervals along the pipes (Figure 14). The

tube-a-manchette pipes are fixed, using normal tie wire, to the outside of the

reinforcement cage and within the zone of the concrete cover for large

diameter bored piles, or to the steel H-piles. Within 24 hours after

concreting/grouting, water is injected through the tube-a-manchette pipes at

the perimeter of the shaft with a pair of double packers in stages to crack the

green grout/concrete (Figure 15), which can then be followed by shaft

grouting operation.

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(a) Tube-a-manchette pipe without

manchette

(b) Tube-a-manchette pipe with

manchette

(c) Double packers and pumps (d) Inflated packer inserted in the

Tube-a-machette pipe

(e) Tube-a-manchette pipe installed

adjacent to the flange of steel section

(f) Pressure gauge for post-grouting

Figure 14 Tube-a-Manchette pipe

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(a) Arrangement of Tube-a-manchette pipe along perimeter of shaft

(b) Post-grouting by means of tube-a-manchette

Figure 15 Cracking of green grout and post-grouting

3.8.3.3 For shaft-grouted mini-piles, the horizontal soil stress around the pile

perimeter will increase due to the soil modification resulting from pressure

grouting and this effect may be considered as similar to a pile without post-

grouting but with an enlarged pile diameter. In-situ measurements have been

carried out to correlate the shaft friction at the pile-soil interface after the shaft

grouting operation. Littlechild et al (1998) reported that the fmax/N value was

5 with a maximum value of 260kPa and 200kPa for shaft-grouted piles

respectively in sand and clay in Bangkok, whilst Stocker (1983) reported

maximum values of 250 to 400kPa in sands and 200 to 270kPa in clays. In

Hong Kong, Chan et al (2004) used fmax/N of 2.85 for a frictional mini-pile

with post-grouting construction method in a project for the former Kowloon-

Canton Railway Corporation (the “KCRC”) in Tuen Mun, Hong Kong, and

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found satisfactory performance in the subsequent loading test to twice the

design working load. Littlechild et al (2000), based from a number of loading

test results on the foundations in a KCRC project, reported that the fmax/N

values range from 1.3 to 3.6. Their study further noted that shaft grouting is

more effective for completely weathered materials with SPT-N values less

than 60, and that for soil with SPT-N values greater than 60, the percentage of

increase in shaft friction is less marked.

3.8.3.4 GEO (2006) summarized the data of loading tested on shafted grouted

replacement piles in Hong Kong (Figure 16), and found that the fmax/N values

can range from 1.4 to 5.5. The highest value was reported by Lui et al (1993)

in a project in Mid-Levels, at which an average mobilized shaft friction fmax of

5.5N was recorded in an instrumented pile in completely weathered granite

with a maximum value of 270kPa. GEO (2006), unlike that for piles without

post-grouting, does not recommend the range of fmax/N for piles with post-

grouting, and only suggest the fmax/N values should be limited to 5 with a

limiting N-value of 100.

Figure 16 fmax/N values for replacement piles with shaft-grouted installed

in saprolites in Hong Kong

(Source: GEO 2006)

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3.9 Section Summary

3.9.1 In Section 3, three of the key parameters affecting the shaft friction along the

length of the pile have been discussed, including:

a) the construction methods,

b) the shear displacement at the pile-soil interface, and

c) the site specific soil properties.

This Information Paper then summarises and N-value methods in estimating

the shaft friction along the length of a pile. The depth-dependent method, has

the advantage of relating the shaft friction with the depth, with greater values at

the pile head tallying with the greater shear displacement at the pile-soil

interface at such region. However, it does not correlate with the construction

methods and soil properties, e.g. soil strength and degree of consolidation. The

rational method method has included the effect of soil properties of a specific

site by relying to the SPT-N values of the soil. Though it is better than the

depth-dependent β method from a soil mechanics perspective, it has not

addressed the relative pile movement between pile and soil.

N-value method, on the other hand, has the advantage of relating the shaft

friction with the SPT-N values, and hence incorporates indirectly the soil

properties in its calculation. Moreover, the effect of overburden pressure has

also been indirectly taken into account in the SPT-N values. It further takes into

account of the various construction methods by using different correlation factor

for different types of piles. This is why N-value method is widely used in

calculating the shaft friction for a frictional pile. To incorporate the variation of

shear displacement at the pile-soil interface along the pile length, an average

correlation factor Fgeo is usually adopted for calculating the pile capacity.

3.9.2 Upper bounds of shaft friction

In Section 3, available literature has been reviewed, and it was noted that there

has been a consensus that there are upper bounds of the shaft friction. For

example, GEO (1996, 2006) gives an upper bound for N to about 80 for small-

displacement piles and limits the shaft friction for large-displacement piles to

270kPa. Similarly, for shaft-grouted piles Littlechild et al (1998) reported that

the maximum shaft friction in sands lies between 250 to 400kPa, and further

reported that shaft grouting is more effective for soil with N less than 60.

Micropile Design and Construction Guidelines published by the US Federal

Highway Administration (Armour et al 2000) summarises the values of typical

shaft friction with reference to the different construction methods in described

in Section 3.2 and different types of soil. Table 3 is an extract of the summary

for sandy soil, and it can be seen that the shaft friction increases from Type A to

D construction methods and also increases from with increasing percentage of

gravel. Furthermore, the ranges of shaft friction are in line with those observed

by literature. Post grouting along the shaft (i.e. Type C or D) is very effective

for soil with weak soil; but for good soil, the percentage of the increase in the

shaft friction is less obvious. This observation tallies with the observation of

Littlechild et al (1998).

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Table 3 Typical range of shaft friction

Soil types

Typical range of shaft friction (kPa) for different

construction methods

Type A Type B Type C Type D

Sand with silt 70 – 145 70 – 190 95 – 190 95 – 240

Sand with silt and

gravel 95 – 215 120 – 360 145 – 360 145 – 385

Gravel with sand 95 – 265 120 -360 145 – 360 145 - 355

(Source: Micropile Design and Construction Guidelines (2000))

4. In-Situ Measurements in ArchSD Projects

4.1 In-situ measurements in the 1990s

Numerous in-situ measurements have been carried out since the 1990s on the

shaft friction along the length of frictional piles in both private and public sector

projects. These in-situ measurements have then been published (e.g. Ng and Lei

2003; Li 2000; Ng et al 2001; Yau 2000; Lei and Ng 2007; Littlechild et al

2000). In ArchSD, Dr H Y WONG (our ex-SGE/NP) carried out in-situ

measurements on eight instrumented piles of the four sites (two for each site)

during the 1990s (Wong 2003). All piles on the four sites were installed by

Intrusion Prepakt using PIP pile system. On three out the four sites, steel

sections have been inserted so that the load carrying capacity of the piles was in

the range of 2200kN to 2700kN, and on the remaining site, the load carrying

capacity of the piles was 1461kN (the typical load-carrying capacity of 610mm

PIP pile). These eight instrumented piles were loaded to twice their working

capacity, and the pile-head settlement and strains along the length of piles were

measured. Raw data and more details of his instrumentation works can be

found in Wong (2003).

4.2 In-situ measurements in recent ArchSD projects

4.2.1 Two ArchSD projects have recently employed the frictional mini-piles as the

foundations, and instrumented piles (details in Figure 17(a)) were installed in

order to verify the shaft friction along the pile length. The first project is “A

permanent planning and infrastructure exhibition gallery at City Hall Annex”

(Inform no. 7195U) (the “City Hall Annex Project”). 14 nos. of the frictional

mini-piles were installed to a new annex, and each of the frictional mini-piles

consisted of a Grade S355JR 152×152×37 kg/m UC installed in a pre-bored

hole formed into soil with a temporary steel casing with internal diameter of 305

mm and then injected with cement grout followed by extraction of temporary

steel casing before the setting of grout. The load carrying capacity of the

frictional mini-pile was 580kN. During the installation, the casing of the upper

portion was accidentally left in to a depth of 22m below the ground level, due to

the early setting of the grout before extraction, whilst the temporary casing of

the remaining piles was withdrawn successfully. As such, negligible friction

was expected from the upper portion of the pile, where the temporary casing

remains.

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4.2.2 The second project is “Transformation of the former police married quarters site

on Hollywood Road into a creative industries landmark” (Inform no. 7955V)

(the “Hollywood Road Project”). 34 nos. of the frictional mini-piles were

installed to a new annex, and each of the frictional mini-piles consists of a

152×152 built-up I-section from Grade S355JR 20mm thick steel plates

installed in a pre-bored hole formed into soil with a temporary steel casing with

internal diameter of 305 mm and then shaft grouted with cement grout followed

by extraction of temporary steel casing before the setting of grout. Post-grouting

using tube-a-manchette in stages after casting the piles was carried out to

increase the shaft friction along the length of the pile. The load carrying

capacity of the frictional mini-pile was 1300kN. Typical details of these

frictional mini-piles are shown in Figure 17(b).

Figure 17(a) Details of Instrumented Piles at City Hall Annex Project

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Figure 17(b) Typical details of the frictional mini-pile at City Hall Annex

Project

4.3 Shaft friction for PIP Piles

4.3.1 From the paper reported by Wong (2003) on the 4 ArchSD sites using PIP piles,

this Information Paper carry out an analysis of the instrumentation data. Figure

18 plots the axial force along the depth of the instrumented piles. In the

following paragraphs, both and N-value methods will be used to calibrate the

relationship among the measured shaft friction, SPT-N values and for these

instrumented piles.

(a) Yuen Long Site (b) Tung Chung Site

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(c) Ma On Shan Site A (d) Ma On Shan Site B

Figure 18 Axial load along the length of instrumented PIP piles in four ArchSD

projects (Source: Wong 2003)

4.3.2 method

In Section 3, the following equations for calculating the value of at depth z

from the ground level have been quoted

O’Neill and Reese (1999):

sand with N15: = 1.5 – 0.245× z0.5

soil with N15: = 15

N×{1.5 – 0.245× z

0.5

Coleman and Arcement (2002):

silty soils: = 2.27× z-0.67

sandy soils: = 10.72 × z-1.3

Figure 19 shows the relationship of the value of at depth z from the ground

level for all the eight instrumented piles, with the best-fit trend line in black

shown alongside with the equations of O’Neill and Reese (1999) and Coleman

and Arcement (2002). It correlates with the above equations with the greatest

value at pile head decreasing with depth z. The data of all sites show similar

trend and values, indicating that in these sites generally follows the same

trend with depth z. As the construction method of all these PIP was being the

same, it is reasonable to deduce that a generalised equation can be derived. The

best-fit line is thus:

= 7.5× z-1.2

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It can further be seen that the best-fit line lies close to that equation of Coleman

and Arcement (2002) for sandy soil, indicating the soil types for these eight

sites match closely with the soil type used in that equation.

Figure 19 Variation of with depth z for PIP piles

To investigate the value of with the shear displacement of the soil at the pile-

soil interface, a plot of against the measured displacement at the interface is

shown in Figure 20 and with the trend line, though not a good correlation,

added. The plot confirms that soil behaves elastically with small shear

displacement; but will then behave elasto-plastically with the increase in the

shear displacement exceeding about 4mm. The soil will behave in plastic stage

with a shear displacement exceeding 10mm.

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Figure 20 Varation of with shear displacement movement at pile-soil

interface

Though for the eight instrumented piles the similar values of with depth z,

Section 3 has already pointed out that such correlation does not consider the

soil types (including its properties and the degree of consolidation) at different

sites.

4.3.3 N-value method

Figure 21 plots the relationship of fmax/N and shear displacement at pile-soil

interface with the depth along the instrumented piles for the four selected sites

in the study of Wong (2003). The maximum mobilised fmax/N range from 7 to

25. It was further noted that most of the shaft friction was developed in the top

10m, and the shaft friction at deeper depth was not fully mobilised. The

observations by Wong (2003) tally with the theoretical load profile (Figure 7)

of Misra and Chen (2004) and Misra et al (2004). Plastic zone can be seen with

a shear displacement exceeding 6mm. Wong (2003) accounted for the variation

along the length of the pile by dividing the maximum friction fmax = 4.8 N with

different FOSs as follows:

FOS = 1 for top one-third (i.e. at depth 0 to H/3)

FOS = 2 for middle one-third (i.e. at depth H/3 to 2H/3)

FOS = 3 for bottom one-third (i.e. at depth 2H/3 to H).

However, as per the discussion in Section 3, the pile head moves more than the

pile tip, and the soil behaves plastically as the shear displacement exceeds uc.

Therefore the shear resistance at that portion can be fully mobilised, and the

FOS of 1 proposed by Wong (2003) represents the shaft friction of the soil in

the plastic zone. The degree of mobilisation will then be decreased as shear

displacement at the pile-soil interface decreases, and Wong (2003) used FOS of

2 or 3 to represent the decrease, as it is difficult to have the pile-soil interface to

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behave fully plastically unless there is a plunge of the pile into the ground such

that mobilisation of the friction of the whole pile length is obtained.

In practical design, an average mobilised shaft friction can be adopted, which

can be obtained by summing the area under the fmax/N graph and then averaging

the sum by the mobilised length of the pile. That is, instead of dividing the

shaft friction along the pile length into three zones, it is a common practice to

average the mobilised shaft friction and then apply a single FOS to obtain the

safe working load for the pile. This Information Paper therefore calculates the

average mobilised shaft friction, which is obtained by averaging the area under

the plot of fmax/N against the depth of the pile by the length where the shaft

friction has been mobilised. Table 4 shows the detailed calcualtion of the

average Fgeo (=fmax/N) with the mobilised length for these instrumented piles on

these sites. In summary, the following avearge fmax/N values are calculated as:

Site 1 – 3.57

Site 2 – 3.66

Site 3 – 2.70

Site 4 – 5.15

The average fmax/N for the eight instrumented piles of the four sites is 3.77.

Applying a FOS of 3 or 2, the design shaft friction can be taken as 1.25 to

1.88N respectively.

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(a) Yuen Long Site (b) Tung Chung Site

(c) Ma On Shan Site A (d) Ma On Shan Site B

Figure 21 Variation fmax/N values for frictional piles formed by CFA

(Source: Wong 2003)

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Table 4 fmax/N against pile depth and calculation of average value

Depth z

from pile

head (m)

fmax/N (kPa)

Yuen Long

Site

Tung Chung

Site

Ma On Shan

Shan Site A

Ma On Shan

Shan Site B pile no

1

pile no

2

pile no

1

pile no

2

pile no

1

pile no

2

pile no

1

pile no

2

-2 7.83 5.22 - - 1.74 1.74 18.48 13.05

-4 - - 11.67 13.39 - - - -

-5 - - - - 6.79 8.78 13.44 25.30

-6 6.17 12.57 - - - - - -

-7 - - 4.06 8.70 - - - -

-8 - - - - 6.96 6.40 17.39 5.80

-10 2.51 4.11 4.83 5.31 - - - -

-11 - - - - 5.57 3.71 4.89 6.52

-13 - - 3.99 3.26 - - - -

-14 2.69 0.79 - - 1.34 2.68 1.63 0.54

-16 - - 1.74 1.96 - - - -

-17 - - - - 0.79 0.79 3.16 1.58

-18 0.58 1.23 - - - - - -

-19 - - 1.45 1.86 - - - -

-20 - - - - 4.06 2.90 4.89 2.72

-22 4.25 1.02 4.08 3.53 - - - -

-23 - - - - 0.38 1.51 1.45 3.62

-25 0.14 0.88 - - - - - -

-26 - - - - 4.04 3.73 1.24 1.24

-28 - - 0.93 0.47 - - - -

-29 - - - - 1.09 1.52 1.30 1.96

-32 - - 1.19 0.40 0.18 0.72 0.61 1.01

-35 - - - - 1.09 1.09 1.11 0.67

-38 - - 0.39 0.10 - - - -

-45 - - - - 0.49 0.24 0.19 0.06

Average

Fgeo 3.57 3.66 2.70 5.15

Figure 22 plots the relationship of fmax/N and shear displacement at the pile-soil

interface. An approximate linear correlation is noted between the shear

displacement and the value of fmax/N (Figure 22) when the shear displacement

is less than 4mm, as the soil behaves elastically. For a shear displacement

exceeding 4mm, the soil starts to behave elasto-plastically, and for a shear

displacement exceeding 6mm, the soil behaves plastically. The observations

coincide with the idealised zoning in Figure 7 with the top portion of the piles

in plastic zone and the lower portion in elastic zone.

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Figure 22 Relationship between fmax/N values and shear displacement

4.4 Shaft friction for frictional mini-piles without shaft-grouted at City Hall Annex

Project

4.4.1 In order to investigate the shaft friction for the frictional mini-piles without

shaft-grouted, a pile in the City Hall Annex Project was instrumented to

measure the shaft friction along the pile by load-tested it to twice its working

capcaity of 560kN. Figure 23 plots the axial force along the depth of the

instrumented pile, and Figure 24 plots the shear displacement at the pile-soil

interface movement along the length of the pile. Figure 24 shows that up till a

depth of 22m from the pile head, only about 300kN out of the total load of

1100kN was taken up by the soil friction. Limited friction was developed at the

top 22m despite that there have been substantial movement at the top portion of

the pile. This was due to the fact that the temporary casing was not removed,

and hence the friction along this portion of the pile shaft is realtively smaller.

Focus will therefore be data for the lower portion of the pile to calibrate the

relationship among the measured shaft friction, SPT-N values and for these

instrumented piles.

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Figure 23 Axial load for the frictional pile at City Hall Annex site

Figure 24 Shear displacement along the length of pile

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4.4.2 method

Figure 25 shows the relationship of the value of at depth z for the

instrumented pile with the best-fit trend line shown alongside with the

equations of O’Neill and Reese (1999) for sand and gravelly sand. The best-

fit line is calibrated as:

= 1.2 – 0.175× z0.50

This equation differs substantially from that given by O’Neill and Reese (1999)

for CFA piles installed in gravelly sand, though it lies close to that given by

O’Neill and Reese (1999) for sand. This equation also differs from that

obtained from the eight instrumented PIP piles by Wong (2003). Moreover,

the values of (and hence the shaft friction) are less than those for PIP piles.

Figure 25 Variation of with depth z for instrumented frictional mini-

pile without shaft-grouted

Again, to investigate the value of with the shear displacement of the soil at

the pile-soil interface, a plot of against the measured shear displacement at

the interface is shown in Figure 26, and a linear correlation can be observed.

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Figure 26 Varation of with soil movement at pile-soil interface

4.4.3 N-value method

Figure 27 plot the variation of fmax/N and shear displacement at the pile-soil

interface with the depth. A linear correlation is noted between the shear

displacement and the value of fmax/N (Figure 28), as the lower portion of the

pile is the elastic zone, where soil at the pile-soil interface behaves elastically as

that predicted in the idealised pile-soil interface model. The average value of

fmax/N for the tested frictional mini-pile is 0.80. Table 5 shows the detailed

calcualtion of the average Fgeo(=fmax/N) starting from a depth of 22m from the

pile head for the instrumented pile. Applying a FOS of 2, the design shaft

friction can be taken as 0.40N.

Figure 27 Variation of fmax/N values and shear displacement with depth

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Figure 28 Relationship between fmax/N values and shear displacement

Table 5 Calculation of average fmax/N in mobilised length of

instrumented pile at City Hall Annex Project

Depth z from Pile Head (m) fmax/N (kPa)

-27.22 1.46

-29.22 0.97

-31.22 0.56

-35.22 1.01

-39.22 0.66

-43.22 0.16

Average Fgeo 0.80

4.5 Shaft friction for frictional mini-piles with shaft-grouted at Hollywood Road

Project

4.5.1 Two shaft-grouted instrumented piles at Hollywood Road Project were test

loaded to twice its working capacity of 1300kN and maintained for 72 hours.

Figure 29 plots the axial force along the depth of the piles at test load 650kN,

1300kN, 1950kN and 2600kN.

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(a) Pile no. 19 (a) Pile no. 22

Figure 29 Axial load for the instrumented piles with shaft grouted along their

lengths

4.5.2 method

Figure 30 shows the relationship of the value of at depth z from the ground

level for two instrumented piles at Hollywood Road Project, with the best-fit

trend line in black shown alongside with the equations of O’Neill and Reese

(1999) and Coleman and Arcement (2002). Again, it correlates with the above

equations with the greatest value at pile head decreasing with depth z. The data

of these two instrumented piles show similar trend and values, indicating that

in this particular site generally follows the same trend with depth z. The best-fit

line is thus:

= 2.0 – 0.3× z0.58

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Figure 30 Variation of with depth z for instrumented frictional mini-

piles with shaft-grouted

As the construction method is different from CFA piles, the equation therefore

differs substantially from those given by O’Neill and Reese (1999) and

Coleman and Arcement (2002), and also differs from that obtained from the

eight instrumented piles by Wong (2003). Thus, it confirms the observation of

Kulhaway and Chen (2007) that a single design equation could not be derived,

as is affected by shaft geometry, soil particle size, soil properties, the degree

of consolidation, construction method, etc, especially the fact that the

construction method in the present case is completely different from CFA.

Notwithstanding such limitation, as compared with that for PIP piles the values

of (and hence the shaft friction) in the present case shows greater value at the

top portion; but the trend line shows steeper slope. That is, it decreases at a

faster rate with depth as compared with that for PIP piles.

Figure 31 plots the variation of against the measured shear movement at the

interface and with the trend line added. The plot confirms that soil behaves

elastically with small shear displacement; but will then behave elasto-plastically

with the increase in the shear displacement at the pile-soil interface exceeding

about 4mm. Full plastic behaviour could not be achieved in the instrumented

piles, when the shear displacement is not adequate for full mobilisation of its

shear strength.

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Figure 31 Varation of with soil movement at pile-soil interface

4.5.3 N-value method

Figure 32 plots the relationship of fmax/N and shear displacement at the pile-soil

interface with the depth along the piles. Again, a strong linear correlation is

noted between the shear displacement and the value of fmax/N, as the lower

portion of the pile is the elastic zone, where soil at the pile-soil interface

behaves elastically as that predicted in the idealised pile-soil interface model. A

linear correlation is noted between the shear displacement and the value of

fmax/N (Figure 33) when the shear displacement is less than 2mm, and with a

shear displacement exceeding 2mm, the soil starts to behave elasto-plastically,

and for a shear displacement exceeding 6mm, the soil behaves fully plastic. The

maximum fmax/N is 9.78, and the average fmax/N is 4.22 for the first

instrumented pile and 3.89 for the second instrumented pile. Table 6 shows the

detailed calcualtion of the average Fgeo(=fmax/N) with the mobilised length for

these two instrumented piles. Applying a FOS of 2 to 3, the design shaft

friction can be taken as 1.9N or 1.2N respectively.

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(a) Pile no. 19 (b) Pile no. 22

Figure 32 Variation fmax/N values for frictional piles with shaft grouted

Figure 33 Relationship between fmax/N values and shear displacement

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Table 6 Calculation of average fmax/N in mobilised length of instrumented

piles at Hollywood Road Project

Depth z from

Pile Head (m)

fmax/N (kPa)

Pile no. 19 Pile no. 22

-0.95 4.99 4.29

-5.95 5.35 6.25

-10.95 8.25 9.78

-15.95 7.62 5.02

-20.95 2.52 1.22

-25.95 0.56 0.49

-30.95 0.27 0.17

-35.95* 0.15 0.04

-40.95* 0.05 0.00

-45.95* 0.01 -0.01

Average Fgeo 4.22 3.89

* The values of fmax/N are neglected in calculating the

average value, as this portion is assumed not have been

mobilised under the loading test.

5. Summary of Findings

The above paragraphs are a literature review together with the in-situ

measurements of the instrumented piles in ArchSD projects of the shaft friction

for cast in-situ piles. Table 7 summarises the average mobilised shaft friction

over the length where shear resistance has been mobilised for different types of

cast in-situ replacement piles and the suggested maximum design shaft friction

for different types of frictional piles with or without post-grouting along pile

shaft. It is suggested that N-value method using SPT-N value is used for the

design for the frictional capacity. It should however be noted that shaft friction

along a pile is hard to be estimated accurately as different construction method,

concreting/grouting operation etc will have different effects of modifying or

remoulding the soil along the pile shaft, and hence the suggested values have to

be further verified by trial piles before construction.

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Table 7 Suggested average shaft friction for different types of

replacement piles for design in sandy soil #

Pile type

Average mobilized shaft friction fmax

(kPa)

Suggested average

design shaft friction

f (kPa)

values from the

studied sites*

N-value

method

Bored

piles - 0.8 to 1.4N

[1] -

Frictional

mini-piles

without

post-

grouting

fs = σ’v 150kPa [2]

where

σ’v = average effective

overburden pressure,

and

= 1.2 – 0.175× z0.50

0.8 to 2.0N

150kPa [1]

0.4 to 0.7N

50kPa (FOS 3.0)

or

60kPa (FOS 2.5) [3]

Frictional

mini-piles

with post-

grouting

fs = σ’v 270kPa[4]

where

σ’v = average effective

overburden pressure,

and

= 2.0 – 0.3× z0.58

1.4 to 5.5N

270kPa [3]

1.0 to 1.5N

90kPa [5]

PIP piles

installed

by CFA

fs = σ’v 200kPa [6]

where

σ’v = average effective

overburden pressure,

and

= 7.5× z-1.2

3.0 to 4.8N 1.0 to 1.6N, where

N40 [7]

# The values quoted in Table 7 are only applicable to sandy soil, and are not applicable to

clayey soil. * PSE should particularly note that the equations derived for method are -specific for

the studied sites and the values are different from the values obtained from

literature. It is not recommended for use as the method is only related to depth and

does not correlate with construction methods and soil properties, e.g. soil strength and

degree of consolidation. It can be used only when it has been calibrated for the

specific site and construction method. [1]

The range of values is suggested by GEO (2006). [2]

fs is limited to 150kPa with reference to Micropile Design and Construction Guidelines

(2000). [3]

A factor of safety of 3 has been included to get the average design shaft friction; but a

smaller FOS can be adopted. [4]

fs is limited to 270kPa with reference to Micropile Design and Construction Guidelines

(2000) and Lui et al (1993). [5]

A factor of safety of 3 has been included to get the average design shaft friction. [6]

fs is limited to 200kPa with reference to O’Neill and Reese (1999) and Brown et al (2007). [7]

Limit of N-value is quoted in SE Instruction No. 04/2010, which is based on the PIP piles

patented by Intrusion Prepakt (Far East) Limited.

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6. Loading Tests

6.1 Trial pile

In order to verify the design assumptions and parameters, loading test on a trial

pile is required prior to the installation works. The intent of the loading test on

a trial pile is to:

a) establish and/or verify installation means and methods of the contractor,

and

b) verify the design parameters and hence the load carrying capacity of the

pile.

6.2 Installed piles

After the completion of the installation of the piles, loading tests on a number of

the completed piles are required to verify that the contractor is producing

acceptable piles. In the General Specification of Building 2012 of ArchSD, 1%

of the piles are required to be to be load tested to twice the theoretical safe

loading capacity. However, it should be noted that the variation of the load

carrying testing of frictional piles is expected to be higher than that for end-

bearing piles, as the design assumptions and parameters are only provided by

the ground investigation and initial loading test on the trial pile. Moreover,

even for sites with uniform soil properties, the integrity of the piles is affected

by the workmanship of the contractor. Thus, it is prudent to adopt a higher

testing frequency for installed piles. Micropile Design and Construction

Guidelines published by the US Federal Highway Administration specifies 5%

of the installed micropiles to be subjected to loading tests; whilst EN 14199:

Execution of Special Geotechnical Works - Micropiles (BSI 2005) specifies 2%

of the first 100 installed micropiles to be tested and 1% for each next 100

installed piles thereafter. This Information Paper therefore suggests specifying

3% of the first 100 installed piles and 1% for each next 100 installed piles

thereafter for loading test.

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7. Pile Group Settlement

7.1 Settlement of pile is always a concern for floating piles with their toes stop in

the overburden soil. The increase of stress in the underlying soil below the pile

toe may cause undue settlement; the thicker the soil layer, the larger the

settlement will be. It is therefore specified in our GS that a settlement analysis is

required for such piling systems. Frictional mini piles, PIP piles are considered

as floating piles.

7.2 Piles installed in a group to form a foundation can give rise to interaction

between individual piles. The overlapping of stress and strain fields can result in

the pile-soil-pile interaction and this will not only affect the capacity of the piles

but also the settlement behaviour of the pile group. Interference between zones

of influence causes a pile within a group to settle more than a single isolated

pile, as a result of pile-soil-pile interaction. Figure 34 shows the zone of

influence for a single pile and a pile group. Further details on pile group

settlement are discussed in the following paragraphs.

Figure 34 Zone of Influence for a Single Pile and a Pile Group

(Source: Brown et al 2007)

7.3 If the building or column load is not high, the supporting piles are not closely

spaced. However, if piles are closely spaced, the pile groups behave differently

from single isolated piles because of pile-soil-pile interactions that take place in

the group. For a group of closely spaced frictional piles which is required to

support loading from heavily loaded columns, the pile group will form a

network of reticulated piles that create a system of confined soil composite with

the piles acting as reinforcing elements. Unlike a single pile system where the

soil at the pile-soil interface is modified and the lateral stresses may also be

increased during the pile installation stage, such effect cannot happen along the

perimeter of the pile group (Figure 35). In this case, the block settlement of a

group with closely spaced piles should be considered which may often exceed

that predicted from a single pile analysis of an individual pile at the same load.

This pile group settlement is sometimes a few times more than a single pile; that

means the settlement obtained from load testing of a single pile cannot give the

settlement of the pile group/building. The settlement of pile group can be

assessed by various methods such as equivalent raft methods or computer

modelling by PLAXIS 3D FOUNDATION. SEB promulgated SEB Guidelines

SEBGL-PL14: Guidelines on Assessment of Pile Group Settlement (available:

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http://asdiis/sebiis/2k/resource_centre/) on the methods to assess the pile group

settlement. If necessary, it is recommended to carry out settlement monitoring

after a building is constructed.

Figure 35 Pile Group with Closely Spaced Piles

7.4 For drilled placement (DD) piles or shafted grouted frictional mini-piles, the

required pile length will be shorter than that those ordinary CFA piles or

frictional mini-piles without post-grouting under the same required loading.

Figure 36 illustrates the equivalent raft model for the estimation of pile group

settlement. It can be observed that the settlement of the pile group of friction

mini-piles with shaft grouted will be greater than those piles without post-

grouting due to the shorter pile length and hence a thicker depth of compressible

soil beneath the base of the pile group.

Figure 36 Equivalent Raft Model for Estimation of Pile Group Settlement

7.5 Thus, it should be noted that the advantages of friction mini-piles in achieving

high loading capacity at shallow depth may be offset by settlement

considerations. PSEs should therefore make an assessment of the effect of pile

group settlement on the superstructure if frictional mini-pile is considered to be

a feasible piling option. This is especially the problem for high-rise buildings or

under heavily-loaded columns, not only because of inadequate total pile

capacity, but also the resulting settlement may be too excessive leading to the

infeasibility of using frictional mini-piles as the foundation.

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8. Method of Procurement of Frictional Piles

8.1 Either an engineer-design or the traditional deign-and-build arrangement can be

adopted for procuring the frictional mini-piles. For an engineer-design, the

design in the form of the location, sizes, founding depth and method of

installation of the piles will be specified in the contract, and the contractor is

only required to carry out the works according to the specified design.

However, this Information Paper has shown that one of the key parameters

affecting the load carrying capacity of frictional mini-piles is the construction

method (e.g. during the post grouting works and operating the CFA for PIP

piles). The parameters used in the design of the length of piles also show that

the length of the piles is very site-specific, and the only means to check the

assumed parameters is through the initial trial pile. Thus, adopting an engineer-

design arrangement cannot demarcate the defaults in design and in

workmanship should a dispute arise. It is therefore recommended that the

traditional design-and-build arrangement should still be adopted.

8.2 Adopting the traditional design-and-build arrangement, the contractor should

then be held responsible to design and construct each frictional pile to a capacity

to meet the contract requirements (including the loading specified in the loading

schedule), according to his own construction method. PSE should specify the

loading points, and minimum length and size of the piles in the contract. The

contractor can then design the number of piles, the size of the piles and the

length of the piles based on their chosen design shaft friction, which can vary

depending on his proposed plant and experience of the crew; but such chosen

values shall be validated on site by trial piles and loading test. To specify the

minimum length for the piles, the PSE should normally adopt the N-value

method for assessment the length. In case the method is used, the limitations

of method as stated in Table 7 should be noted. Annex A provides an

example to estimate the minimum length of piles in a typical site. Annex B

provides a sample particular specification for the design and installation works

of frictional mini-piles with steel sections in pre-bored holes formed by the

Concentric or Symmetrix system (i.e. a system with the pilot bit set back from

the ring bit during drilling or other drilling systems) and PSE may vary it to suit

individual site and project, and when steel rebars are used in lieu of the steel

section.

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Foundations in Hong Kong”, in Dennis, N D et al (eds), New Technological and Design

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O'Neill, M W and Reese, L C (1999), Drilled Shafts: Construction Procedures and Design Methods,

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Wong, H Y (2003), “Design and Construction of Friction Bored Piles in Hong Kong with Particular

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Annex A

Estimation of the Length of Frictional Mini-Piles

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Table A1 calculates of the length of a frictional mini-pile of 305mm using N-value method with load carrying capacity of 1000kN to be

installed in a hypothetical site by using the SPT-N values of the soil shown. The water table is at 34mPD. Two cases will be considered:

frictional mini-pile with and without shaft grouted. For frictional mini-pile without shaft grouted, the design average shaft friction of 0.4N is

adopted, whilst an average value of 1.0N is adopted for frictional mini-pile with shaft grouted. The calculation shows that the required lengths

of the pile without and with shaft grouted are 42m and 28m respectively.

Table A1 Calculation of length of pile using N-value method

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Annex B

Sample Particular Specification for

Design and Construction of Frictional Mini-Piles

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1. General

A Frictional Minipile for the Contract is defined as a pile which is built-up H-

section from steel plates, or rolled steel H-section [delete where appropriate] to

be installed in a prebored hole formed into soil with a temporary steel casing

with minimum internal diameter of [insert dimension] and then filled with

cement grout. The requirements on the construction of the pile are as shown in

drawing no. [insert drawing number].

The Contractor is required to engage a specialist contractor to design and

construct the piling works, who shall be in the List of Approved Suppliers of

Materials and Specialist Contractors for Public Works in the Category of Land

Piling Group II and eligible to carry out the registered piling system of Rock-

socketed Steel H-pile in Pre-bored Hole or Minipile.

2. Piling Design

2.1 Design Requirements

The theoretical safe loading capacity of the individual Frictional Minipile shall

be the allowable axial force of the built-up H-section from steel plates, or rolled

steel H-section [delete where appropriate]). The maximum allowable axial

working stress of the built-up H-section from steel plates, or rolled steel H-

section [delete where appropriate]) shall be 45% of the yield stress, and the

combined stresses due to axial load and bending moments shall be limited to

50% of the yield stress. When the calculations of stresses are based on all loads

including wind loads, the permissible stress shall be increased by 25% of the

above stresses.

The theoretical safe loading capacity of each individual pile shall not exceed

[insert number]kN.

The Contractor shall be responsible for the design and construction of the

Frictional Minipiles including the length and the number of Frictional Minipiles

to support the loading in the loading schedule of drawing no. [insert drawing

number]. The founding level shall be at [insert number]m below the cut-off

level or ground level [delete as appropriate].

[Guidance notes: the minimum length of the pile is to be calculated according

to the GI, and refer to the example in Annex A of this information paper for a

worked example to estimate the minimum length.]

The frictional resistance between the pile shaft and the soil above [insert

number]mPD shall not be considered. End-bearing capacity of the pile shall be

ignored. The Contractor shall satisfy himself that his method of calculating the

design frictional resistance, which shall provide a sufficient factor of safety in

his design.

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Pile design submission shall be in accordance with Clause 5.02 of the GS.

2.2 Pile Head Details

The Contractor shall provide capping plate and dowel bars in accordance with

the detail as given in the GS for steel H piles..

2.3 Cover

The minimum clearance (cover) between casing and the steel section shall be

40mm. The Contractor shall submit his proposed spacer details with his pile

design submission.

2.4 Minimum Segment Length of Steel Sections

The minimum length of each segment of steel sections forming the whole length

of Frictional Minipile shall be 10 m except the uppermost section.

2.5 Provision of Shear Key

The Contractor shall provide shear bars to steel sections in accordance with the

details shown in Appendix.

3. Submissions for Piling Works

In addition to the submissions stated in GS Clause 5.02, the Contractor shall

submit 2 copies of each of the following information with the design

submissions:

a. details of built-up H-section from steel plates, or rolled steel H-section

[delete where appropriate];

b. details of grout mix;

c. method of installation including equipment to be used, sequence of

operations, drilling methods, temporary casing installation and extraction,

and time of grouting;

d. details of grouting operation, taking into account of the subsoil condition

and ground water fluctuation during the day;

e. method of piling operation to overcome underground obstruction, if

encountered;

f. spacers details;

g. a report on the existing conditions of the adjacent structures (including

existing underground pipe, existing retaining walls and any other

structures nearby). This initial condition survey report shall include

information and record photos of these existing structures with particular

attention to those aspects that may be adversely affected by the proposed

works. Foundation types of the adjacent structures shall also be included;

h. proposal on precautionary measures and actions to be taken so as to

prevent the above adjacent structures from being adversely affected by the

works;

i. any other requirements specified in this particular specification.

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No piling works shall commence on site unless the submissions are approved by

the SO in writing.

4. Drilling Method

Unless otherwise agreed by the SO, Frictional Minipile shall be formed with a

temporary casing used to stabilise the surrounding soil, and shall not be installed

with the use of bentonite slurry or other drilling muds. Temporary casing of

approved quality shall be lowered at the same time when the hole is made.

Unless otherwise approved by the SO, the Concentric or Symmetrix system or

other drilling systems shall be used to form the pile hole of the Frictional

Minipiles.

Temporary casing shall be free from distortion, internal projections and

hardened grout.

5. Flushing Medium

Air or water shall be used as the flushing medium during the drilling operation.

The Contractor’s attention is drawn to the formation process of the pile shaft

using air flushing where special care shall be taken to avoid disturbance to

adjacent ground of soil during forming of the pile shaft.

6. Tolerances

The maximum deviation of the centre of the head of each finished Frictional

Minipile from the designed centre point shall not be more than 50 mm in any

direction. The maximum deviation from the vertical axis of the pile through the

centroid of the cross section at the cut off level at any level of the finished pile

shall not be more than 1 in 75.

[Guidance note: If a group of reinforcement bars, instead of steel section, is

used, the maximum deviation of the centre of the head shall be limited to 15mm

and the maximum deviation from the vertical axis shall be limited to 1 in 100.]

7. Founding Level

Founding level of each Frictional Minipile shall not be higher than those

specified in Clause 2.1.

8. Cutting Off of Pile Heads

All Frictional Minipiles shall be grouted to a minimum level of 300 mm above

the specified pile cut-off level.

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9. Grout for Piling Work

9.1 Grout Material

Grout shall consist of ordinary Portland cement and water with an approved

non-shrinkage additive. Where PFA is used, the maximum PFA content shall

not exceed 35% of the total cementitious content in the grout. Other admixtures

can be used when approved by the SO. The manufacturer’s guidance shall be

strictly followed. Cement sand mix is not allowed.

Grout shall have minimum cube strength of 30 MPa at 28 days.

Measurements for bleeding shall be taken at 15-minute intervals. The amount of

bleeding shall not exceed 2% at the end of the first 3 hours and no interim

readings shall exceed 4%. In addition the water must be reabsorbed by the

grout within 24 hours after mixing.

Free expansion of grout when measured at the end of 24 hours after mixing

shall have a figure between 0% and 5%. A negative percentage figure shall not

be accepted.

Any approved admixtures shall be chloride-free and comply with BS EN 934.

The maximum total chloride content, expressed as a percentage relationship

between the chloride ion and the cementitious content by mass in the grout shall

be 0.1%.

Water for grout shall be clean fresh water having a temperature not exceeding

30C nor less than 5C.

9.2 Grout Mixing

Grout material shall be mixed by weight batching. The amount of water used

shall be measured by a calibrated flowmeter or a measuring tank.

The mixing time in high-speed mixers shall be appropriate for the type of mixer

used.

After mixing, the grout shall be continuously agitated in a holding tank and

screened before injection. The grout shall be placed within the time limits

specified by the manufacturers of the additives.

9.3 Pressure Grouting and Extraction of Temporary Casing

9.3.1 Grout Pipes

Grouting shall be carried out with two non-flexible grout pipes, one at each side

of the web of the steel section.

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9.3.2 Grouting Methods

Before grouting, the bottom of the hole shall be cleaned by airlifting or an

alternative method approved by the SO.

[Delete either of the following options]

Either:

The hole shall be grouted by shaft grouted method. That is, upon extraction of

temporary casing after the completion of initial grouting to the body of the

whole Frictional Minipile, the surface of the shaft of each Frictional Minipile

shall be cracked by water within 24 hours of the initial grouting via Tube-a-

Manchettes or similar method approved by the SO. Subsequent pressurised

post grouting shall then be followed such that the pressurised grout expels itself

into the surrounding ground. The Contractor shall submit to the SO for

approval the detailed proposal of his shaft grouted method.

Or:

Grouting shall be carried out in an upstage sequence from the bottom of the

hole. After the initial pressure grouting of the bottom of the pile, the temporary

casing shall first be partially extracted upward to a predetermined level as

approved by the SO and then followed by pressure grouting. The bottom level of

the grout pipes shall in no case be higher than the bottom level of the temporary

casing. The above grouting sequences shall be repeated until the completion of

the grouting of the pile.

Unless otherwise permitted by the SO, grouting shall be carried out by injecting

the grout under pressure into each grouting stage of the hole until the grouting

stage refuses to take further grout.

The initial grouting shall be carried out in such a way that the lowest part of the

grout pipes shall be as close to the pile toe as possible.

9.3.3 Pressure Grouting

Grouting pressure shall be determined by the Contractor and shall in no case be

less than the overburden pressure unless otherwise approved by the SO.

Holes shall be grouted in a continuous operation at each grouting stage and

pressures as approved, and except during subsequent post grouting, shall not be

left partially grouted.

If, in the opinion of the SO, grouting of any hole or grouting stage has not been

completed due to low pressures, excessive leakage when compared to the

performance of the trial pile as required in Clause 11 or other causes, the hole

shall be redrilled or flushed out with water and re-injected with grout.

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The Contractor shall agree with the SO the method to measure the grout intake

volume.

Newly grouted piles shall be properly covered and fenced off.

9.4 Testing of Grout

The Contractor shall employ an approved laboratory to carry out the tests for

Bleeding, Free Expansion and Flow Cone Efflux and Crushing Strength of

grout.

9.4.1 Definition of Batch

A ‘batch’ of grout is any quantity of grout used for grouting in one continuous

operation in one day.

9.4.2 Test for Bleeding and Free Expansion

The Contractor shall provide one sample of the grout from each Frictional

Minipile after mixing and shall protect from changes in moisture content before

tests are carried out.

Each sample shall be divided into 3 specimens. Each specimen is to be placed

in a covered cylinder with a diameter of 100 10mm to a depth of 100 5 mm

and the amount of bleeding and free expansion is measured by a scale fixed to

the outside of the cylinder.

Bleeding = 100% x H2 - Hg

H1

Free Expansion = 100% x H2 – H1

H1

where H1- initial height of grout sample

H2- height of sample measured at upper surface of water layer or

hardened grout surface if water is fully absorbed

Hg- height of grout portion of sample at upper surface of grout

The Contractor shall submit preliminary test results to the SO within 48 hours

after the mixing of grout.

If the result of the bleeding test of the grout for any pile does not comply with

the specified requirements or the free expansion of the grout for any pile is

greater than the specified upper limit, the Contractor shall propose changes to

improve the materials, grout mix or method of production, though the failure

does not constitute a failure of the pile.

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If the free expansion of the grout for any pile has a negative figure, the

Contractor shall carry out test(s) at their own expense to demonstrate that the

pile can fulfil the original design requirements.

9.4.3 Flow Cone Efflux Test

At least one sample from each pile shall be taken and tested in accordance with

ASTM C939 to determine the Flow Cone Efflux time. Agree with the SO the

frequency of the test.

Except with SO’s prior agreement for grout mixes containing additives, grout

having an efflux time of less than 15 seconds shall be rejected.

9.4.4 Test for Crushing Strength

The Contractor shall provide one sample of the grout for each Frictional

Minipile after mixing and shall protect it from changes in moisture content

before making test cubes.

Cubes shall be prepared using 100mm cube moulds.

The Contractor shall make two cubes from the sample. Strength compliance

requirements shall follow GS Clause 6.55.

10. Steel Sections

GS Clause 5.18 (iii), (v) and (vii) shall apply to Frictional Minipiles.

The Contractor shall employ an approved specialist firm to carry out and

interpret the inspection and testing of welds, and shall provide any necessary

labour and attendance. The Contractor shall submit evidence proving that

operators carrying out the inspection and testing have been trained and assessed

for competence in the inspection and testing of welds. In addition, the

Contractor shall submit certificates of competence from a recognised authority

for operators carrying out ultrasonic examination.

The welded joints of steel sections shall not be lowered into the pile shaft within

one hour after they are completed.

The maximum length of spliced steel sections in horizontal or inclined positions

shall be 24 m.

11. Trial Pile

After the approval of the design submission and before the commencement of

pile installation, one of the piles selected by the SO shall be installed as trial pile

to validate the design parameters, method of installation and grouting operation

proposed by the Contractor.

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The trial pile shall be subjected to a static loading test as included in Clause 12.

If the trial pile fails in the static loading test, the Contractor shall revise his

piling proposal and re-submit his revised design to the SO for approval. A new

trial pile shall then be installed in accordance with the revised proposal by the

Contractor and then subjected to the static loading test.

No installation works for the remaining piles shall be commenced until the trial

pile has passed the static loading test.

12. Static Loading Tests

Loading tests shall be required as instructed by SO in accordance with GS

Clause 5.28. Notwithstanding GS Clause 5.28, the SO may order 3% of the

installed piles or [insert drawing number] nos. of piles installed, whichever is

more, to be load tested to twice the loading capacity of the respective piles. In

determining the cross sectional area (A) of pile, the transformed section

(comprising the grout and steel section) shall be used.

The Young’s modulus of grout shall be taken as that of concrete of the same

strength as given in the GS.

[Guidance note: Allow for 3% (instead of 1%) of the first 100 installed piles

and 1% for each next 100 installed piles thereafter to be load tested to twice the

theoretical safe loading capacity]

13. Piling Records

The Contractor shall keep records of the installation of each pile and submit two

signed copies of these records to the SO not later than noon of the next working

day after the pile was installed.

The record shall give the following information in an approved format:-

a. Pile reference number;

b. Date and time of boring;

c. Soil samples taken and in situ test carried out, if any;

d. Date pile installed;

e. Pile type and size;

f. Date and time of drilling;

g Date of grouting;

h. Position of pile in the works and ground level at pile position;

i. Working level;

j Drilling rates and material encountered;

k. Depth from working level to pile toe;

l. Toe level;

m. Depth from working level to pile head level;

n. Length and toe level of temporary casing;

o. Length of steel section;

p. Grout mix;

q. Volume of grout in pile (actual and theoretical);

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r. Details of obstructions, delays and other interruptions to sequence of

work;

s. Flow rate and total time required for the grouting operation;

t. Grouting pressure used in each stage;

u. Any other data requested by SO.

On completion of all piling works, the Contractor shall submit to the SO two

copies of the record piling plans showing, as appropriate, the position, identity

number, size and top and bottom levels of each pile installed.

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Appendix

Typical Details of Shear Bars