Pile Design

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ICE Manual of Geotechnical Engineering © 2012 Institution of Civil Engineers www.icemanuals.com 803

doi: 10.1680/moge.57098.0803

CONTENTS

54.1 Introduction 803

54.2 Selection of pile type 803

54.3 Axial load capacity (ultimate limit state) 804

54.4 Factors of safety 814

54.5 Pile settlement 814

54.6 Pile behaviour under lateral load 816

54.7 Pile load testing strategy 818

54.8 Definition of pile failure 820

54.9 References 820

54.1 IntroductionPiles are generally required to transfer load from a superstruc-ture through weak or compressible strata, or through water, on to stiffer and less compressible soils and rock. Piles are also required to reduce both overall and differential settlements of the supported structure and may be required to enable con-struction processes such as top-down construction. They may also be required to resist uplift loads as well as compressive loads when used, for example, to support tall structures which are subject to overturning forces caused by wind. Where base-ments are to be constructed piles may also be subject to heave-induced tension. A more recent additional use of piles is within ground source heating and cooling systems as part of a devel-opment’s renewable energy strategy.

Piles used for marine structures are often subject to lateral loads from ship impact and wave forces. Many land-based piles are also subject to lateral forces; for example, in bridge works there are earth pressure forces, expansion and contrac-tion forces and braking and traction forces. Combinations of vertical and horizontal forces and bending moments are com-mon and pile sections have then to be designed according to the relevant structural codes in addition to geotechnical consider-ations. Foundations beneath machinery may also be subject to cyclic and vibration forces.

Before embarking on the design of a piled foundation, it is essential to have a thorough understanding of all the relevant factors that might affect the performance of the pile being considered.

The design of piled foundations has traditionally been based on a combination of simple empirical methods and local expe-rience, but over the past few decades there has been a grad-ual change towards more soundly based theoretical methods. Poulos (1989) provided a summary of the range of analysis and design methods available for piled foundations, ranging from simple empirical methods to the most sophisticated non-linear numerical analysis methods. When selecting the appropriate design approach, the following factors should be considered:

the quality of the site investigation and scope of geotechnical data ■■

available;

the design stage, e.g. conceptual design or detailed design;■■

the budget and timescale for foundation design;■■

the scale and sensitivity of the proposed structure, especially ■■

allowable total and differential settlements;

the complexity of the ground conditions and loading regime.■■

54.2 Selection of pile typeFoundation selection is discussed in Chapters 9 Foundation design decisions and 52 Foundation types and conceptual design principles.

The key factors that usually affect the choice of pile type and ultimately the design of the pile itself are as follows:

1. Ground conditions

Strength or weakness of the overlying and founding strata

Ground variability

Potential for downdrag or heave loading, or lateral loading

Presence of obstructions

Aggressive ground conditions

2. Construction constraints

Location

Access

Working area

Sensitivity of adjacent assets

Underground utilities and structures

3. Safety and environmental constraints

Risk to operatives and the public

Chapter 54

Single pilesAndrew Bell Cementation Skanska Ltd, Doncaster, UKChristopher Robinson Cementation Skanska Ltd, Doncaster, UK

This chapter considers the design of single piles, from selection of appropriate pile type(s) through to pile load testing strategies. Methods which are commonly used for estimating axial pile capacities in three principal ground types, as well as layered soils, are discussed, covering both empirical and theoretical methods. Other aspects of single pile design covered by this chapter relate to factors of safety, pile settlements and pile response to lateral loading.

Design of foundations

804 www.icemanuals.com ICE Manual of Geotechnical Engineering © 2012 Institution of Civil Engineers

objective of piles – if it is to reduce settlement/bending moments ■■

in a raft then consider piled raft (see Chapter 56 Rafts and piled rafts);

influence of any ‘global’ ground movements both during construc-■■

tion and long term (see Chapter 57 Global ground movements and their effects on piles);

testing strategy for checking integrity of pile;■■

testing strategy for checking vertical and/or lateral load ■■

behaviour;

potential for foundation reuse.■■

54.3 Axial load capacity (ultimate limit state)The approach described in this chapter to assess the resistance of piles to compressive loads is the static ‘soil mechanics’ method for piles loaded axially. Dynamic methods of design and analysis are available but have not been included in this section and attention will be limited to problems involving static or quasi-static axial loading of single piles only.

The general simplified formula for the design of bearing piles typically takes the form of

Qult = Qb + Qs – Wp, (54.1)

where Qult is ultimate pile capacity, Qb is base capacity, Qs is shaft capacity and Wp is pile self weight.

Furthermore, Qb = qb Ab and Qs = qs As, where qs is unit shaft resistance, qb is unit base resistance, Ab is area of pile base and As is area of pile shaft.

The diagram in Figure 54.1 shows this relationship, for circular piles.

The above approach, separating the evaluation of shaft and base capacity, forms the basis of all static pile capacity calcula-tions. These components can be broken down further to resis-tance per unit shaft or base area, and the relative magnitude is dependent on the combination of pile geometry, pile installa-tion method and ground conditions.

Typically, the shaft capacity of a pile is mobilised at much smaller displacements than the base capacity. Shaft capacity can be fully mobilised at movements of only 0.5% to 2% of pile diameter, whereas displacements of around 10% to 20% of pile diameter may be required to mobilise the base resis-tance, even more if the soil at the base of the pile is disturbed as part of the installation process. It is essential that these dif-ferences between the load deformation characteristics of the pile shaft and base are considered early during the pile design process; Figure 54.2 shows this as an idealised load displace-ment response curve. Generally, long slender piles are more efficient than short fat piles with regard to settlement control but limitations on the installation technique may often dictate the pile geometry.

There can be concerns about the axial pile capacity under tension loads. De Nicola and Randolph (1993) showed that shaft capacity under tension load is between about 70% and

Risks to adjacent structures and underground utilities assets

Risk of contamination to surrounding environment

Effects of noise

Effects of vibration

4. Nature and magnitude of loads

Direction of foundation loading (e.g. lateral, tensile, compression)

Source and characteristics of loadings (e.g. heave, down-drag, cyclic, static)

5. Economic considerations

What it will cost and whether there is a more economic solution

Prior to embarking on the design of a pile, it is essential that the designer first assesses the type of piles and construction meth-ods that are most appropriate for the particular site under con-sideration. There is clearly no sense in designing a driven pile in an area where noise and vibration will be an issue. Conversely, designing a continuous flight auger pile 35 m deep when avail-able plant will only reach 30 m is equally futile. Under cur-rent health and safety regulations (Health and Safety Executive, 2007), it is also essential for the designer to take into account the risks associated with the installation of the proposed pile solution during the design and specification process; to com-plete this satisfactorily the designer will need at least a general understanding of the various pile types available and the advan-tages, disadvantages and key safety considerations of each.

Before any design work is carried out it is essential that an assessment of pile buildability is carried out. Reference should be made to Chapters 52 Foundation types and conceptual design principles, 79 Sequencing of geotechnical works and 81 Piling problems in this respect.

It is important to recognise that each pile type and installa-tion method disturbs the ground in different ways, and as such the design method and considerations may vary between the different systems. This disturbance can improve or reduce both skin friction and end-bearing, and as such will affect the bear-ing capacity of the complete pile.

Piles of varying type and length may be used on a site pro-vided a suitable assessment of likely differential settlement is carried out. This should include:

pile size (diameter and length), i.e. geotechnical axial load capac-■■

ity in compression and tension of a single pile;

displacement at working load of a single pile;■■

differential settlement between adjacent foundations;■■

influence of lateral loading, assessment of geotechnical and ■■

structural capacity, and lateral displacement;

assessment of pile group behaviour under vertical and lateral ■■

loading (see Chapter 55 Pile-group design);

Single piles


The bearing capacity factor relevant for the depth of most piles is usually taken as 9 after Skempton (1951). Allowance should be made for the potential of a softened or disturbed base in some circumstances, particularly when considering under-reamed piles, with enlarged bases. In such cases a reduction in Su should be applied. A reduced value of Nc should also be considered when the pile tip only just penetrates the found-ing fi ne-grained strata. A value of Nc = 6 is recommended for founding in the top of such founding strata, increas-ing linearly to 9 after 3 pile diameters’ penetration into the stratum.

Shaft friction

Piles in fi ne-grained soils generally derive most of their capac-ity in shaft friction, with the base resistance often not mobil-ised under working loads, particularly in the short term.

Shaft friction, in terms of total stress, is generally calculated using an empirical factor α, where

qs = αSu. (54.3)

The value of α adopted for bored piles typically ranges from 0.3 to 0.6 for straight-shafted piles; higher values have been proven for ribbed or irregular pile shafts. Values of 0.45 to 0.6 are generally adopted for bored piles in stiff overconsolidated clays but can be as high as 0.7 for driven displacement piles. The value of α depends on a combination of the pile type and installation method and the nature of the fi ne-grained soil. In the UK, the LDSA London District Surveyors Association’s Guidance Note 1 (2009) gives clear recommendations on appropriate values for α for bored and CFA piles founding in London Clay. It is also important for more complex and vari-able fi ne-grained soils, such as glacial till, that reference is made to specifi c guidance such as Weltman and Healy (1978) and Trenter (1999). Figure 54.3 details the variation of α with

90% of the shaft capacity under compression loads (depend-ing on the relative soil/pile compressibility). However, most codes require higher factors of safety for piles subject to tension loads than those for piles subject to compression. If these higher factors of safety are used, then there is no need to reduce the shaft capacity calculated by routine methods, since the potential adverse effects are taken into account by the use of a larger factor of safety. For piles subject to tension, the base resistance is ignored.

54.3.1 Fine-grained soils54.3.1.1 Total stress

End-bearing

Whilst the long-term ‘effective stress’ drained end-bearing capacity of a pile in clay will be signifi cantly larger than the undrained capacity, unfortunately the movements required to mobilise such capacity would be unacceptable for most struc-tures. Furthermore, the pile will require suffi cient capacity in the short term to provide the initial load-carrying capability required, and the rate of build-up of load over time will depend on the speed of construction, something which continues to increase as construction technology advances. For these rea-sons, the end-bearing of piles in fi ne-grained soils is generally considered using total stress methods, in terms of the undrained shear strength of the clay, Su, and a bearing capacity factor, Nc, where

qb = NcSu. (54.2)

Figure 54.1 Pile under vertical loadCourtesy of Cementation Skanska

Figure 54.2 Idealised load displacement responseCourtesy of Cementation Skanska



An important application of effective stress methods is for piles below or adjacent to deep excavations (say in excess of 4 m deep), when total stress methods may be unsafe.

End-bearing

The use of effective stress methods to calculate the end-bearing of piles in clay soils is, in general, not recommended. It would only be appropriate to adopt such an approach based on the results of suitable pile load testing or appropriate experience in similar ground conditions.

Shaft friction

Assuming loading takes place under fully drained conditions, where remoulding and drainage at the pile–soil interface

undrained shear strength in glacial till, after Weltman and Healy (1978).

Poulos (1989) summarised some of the most commonly used methods in calculating α for driven piles, an extract of which is included in Table 54.1.

It is generally considered good practice to limit the average skin friction in fi ne-grained soils to 100 kN/m2 unless proven higher by pile load testing or by suitable experience of pile performance in similar ground condition.

The Su value in equation (54.3) is really an empirical index value, closely linked to historical sampling/lab test methods to-gether with back analysis of pile test data, rather than a ‘fun-damental’ soil parameter. In the UK, the method is mainly based on experience in London Clay, with the results of Quick Undrained Triaxial (QUT) tests on U100 samples being com-pared with the results of maintained load tests (refer to Patel, 1992 for further background). Shaft failure requires shearing to take place through a large volume of ground; hence, the Su value selected for design is a ‘mean’ value (where the mean is based on representative data, excluding spurious high or low values).

54.3.1.2 Effective stress (fi ne-grained soils)

Historically, the use of effective stress calculations for the design of piles in fi ne-grained soils has been limited, but recent research has provided more accurate methods of using this approach. The Imperial College design methods (Jardine et al., 2005) provide recommendations on appropriate meth-ods of calculation for both end-bearing and shaft resistance for driven steel piles in terms of effective stress. Chapter 22 Behaviour of single piles under vertical loads provides guid-ance on effective stress design to assess the capacity of single piles subjected to vertical loading. It is important with such methods to limit the calculated values of end-bearing and shaft friction to those achievable in the fi eld for the pile type and pile construction methods being employed.

1.2

1.0

0.8

0.6

0.4

0.2

Adh

esio

n fa

ctor

α

0

60 80 100 120

Undrained shear strength cu, kN/m3

140 160 180 200 220

Reduced values for drivenpiles where L < 10B and tillB and tillBis overlain by soft clay

Bared pileBared pileBared piles

Driven and cast in-place pilesen and cast in-place piles

Figure 54.3 Adhesion factors, , for piles in glacial tillReproduced from CIRIA PG5, Weltman and Healy (1978), www.ciria.org

Table 54.1 Summary of commonly used factors for driven piles in fi ne-grained soils

Alpha factors Reference

α = 1.0 (Su ≤ 25 kN/m2)

α = 0.5 (Su ≥ 70 kN/m2)

API (1984)

Linear variation in betweenα = 1.0 (Su ≤ 35 kN/m2)

α = 0.5 (Su ≥ 80 kN/m2)

Linear variation in between

Length factor applies for L/d > 5

Semple and Bigden (1984)

Fleming et al. (1985)

c c c

nc

u

v

u

v

0.5

u

v

for 10 5.

c c

nc

u

v

u

v

0.250 5.

for 1u

v

c

Single piles


history and also depends on stress changes during pile instal-lation. Further discussion can be found in Bown and O’Brien (2008).

The β approach is particularly useful for normally con-solidated or lightly overconsolidated soils where K0 can be assumed to be close to 1 − sin φ′. Then, for example, in the case of a driven pile, if the angle of effective friction is taken as, say, 21°, the ultimate friction becomes 0.25σ′v. For a very wide range of normally consolidated soils, the value of β is likely to be between 0.2 and 0.3. This value is often used as the basis for calculation of downdrag or ‘negative friction’.

54.3.2 Coarse-grained soils

Approximate design methods for piles in coarse-grained soils are well documented and reference should be made to Fleming et al. (2008) and Tomlinson (2004) for a wider understanding of the range of design methods available. Recent research has focused on steel-driven piles, aimed predominantly at the off-shore, industry, and reference should be made to the work by Randolph and White (2008) and Jardine et al. (2005).

destroys all cohesion so that c′ = 0, then the unit shaft friction can be calculated using

qs = σ′h tan δ, (54.4)

where σ′h is horizontal effective stress following pile installa-tion and δ is effective interface friction angle.

For bored piles δ would usually be assumed to equal the critical state angle of friction, φ′cv, whereas for driven piles it would be assumed to equal the residual angle of friction, φ′res.

A simplifi cation can be made if σ′h is assumed equal to σ′vK, in which case qs = σ′vK tan δ, further simplifi ed, where K tan δ = β, to qs = βσ′v.v Values of β can vary signifi cantly depend-ing on consolidation history (i.e. normally consolidated or overconsolidated) and method of pile installation (driven or bored). Examples of typical values of β and K are given in Tables 54.2 and 54.3.

The main challenge in applying equation (54.4) above is to assess the appropriate values of σ′h or K. As shown in Figure 54.4 the value of K is very sensitive to the site’s stress

0 100 200 300

Stress (kPa) K0K0K

400 500 600 00

10

2020

10

0D

epth

bel

ow c

lay

surf

ace

(m)

30

Clay at surface; hydrostatic pressure

100 kPa surcharge; hydrostatic pressure

Clay at surface; underdrainage-water pressure half hydrostatic100 kPa surcharge; underdrainage

30

3 41 2

σ ′v

σ ′v

Figure 54.4 Infl uence of stress history on K0 and ’h in overconsolidated clayReproduced from Burland et al. (1979)



54.3.2.1 End-bearing

The end-bearing capacity of piles in coarse-grained soils can be calculated by the formula

qb = σ′vBNq, (54.5)

where σ′vB is effective vertical stress at the pile toe and Nq is bearing capacity factor.

Nq is often derived based on a correlation with φ. A com-monly used method is that by Berezantzev et al. (1961) (Figure 54.5). The main problem with this approach is that it is unduly sensitive to small changes in φ′, and does not allow for the reduction in mobilised φ′ as effective stress levels increase. It also implies that end-bearing pressure will continue to increase proportionally with depth. An alternative approach is that given in Fleming et al. (2008), which relates Nq to relative density, effective stress and the critical state angle of friction φ′cv, shown graphically in Figure 54.6 for φ′cv = 30°. φ′cv values for a range of sands are summarised by Stroud (1989) and Bolton (1986). Relative density can be assessed from standard penetration test (SPT) ‘N’ or cone penetration test (CPT) ‘qc’ values (section 54.3.4).

Routinely used methods that can be used to calculate end-bearing are summarised in Table 54.4.

With any design approach, it is very important to limit the calculated end-bearing to achievable values to ensure service-ability limits are achieved; this is discussed further in section 54.3.5 below. It is possible with theoretical bearing capacity based methods to signifi cantly overestimate the end-bearing resistance, particularly for piling methods that may disturb the soil at the pile toe or for short piles that are highly dependent on end-bearing.

The end-bearing capacity of a pile in a uniform layer of coarse-grained soils will not continue to increase proportion-ately with depth. Research has shown that the end-bearing pressure soon reaches a limiting value (Vesic, 1977), typi-cally for pile lengths in excess of about 20D (where D is the pile diameter). For this reason end-bearing is often limited

Table 54.2 Typical K values for bored piles in fi ne-grained soils

K value Reference

K is the smaller of K0 and 0.5(1 + K0) Fleming et al. (1985)

K/K0 = 2/3 to 1; K0 is a function of OCR Stas and Kulhawy (1984)

Table 54.3 Typical values for driven piles in fi ne-grained soils

value Reference

β = (1 − sin φ′) tan φ′(OCR)0.5 Burland (1973)

Meyerhof's value Meyerhof (1975)

1500

500

200

100

60

30

1025 30 35

ϕ (°)

40 45

NqNqN

Figure 54.5 Variation of Nq with wReproduced from Berezantsev et al. (1961)

10

1 2 3 5 7 10

qb (MN/m)2

20 30

0.5

0.25

20

30

50

70

100

200

300

500

0.750.750.75

IDD = 1

Figure 54.6 End-bearing capacity, coarse-grained soils, with φ’cv = 30°Reproduced from Fleming et al. (2008), Taylor & Francis Group

Single piles


For full displacement driven piles Ks can be estimated as Nq/50; this relationship gives a typical value of Ks = 1.2 for such piles, and limiting values between 0.7 and 2.0 are suggested. For partial displacement piles Ks values will be lower than for full displacement, and the Ks value is usually reduced by 20%. For driven cast in situ piles some reduction in horizontal stress will occur when the casing is extracted, and normally a Ks value of 1.0 is assumed (if wet concrete is placed).

For auger or rotary displacement piles, there are examples of Ks values derived from pile load testing well in excess of 1.2 (Bell, 2010). Such installation methods can produce signifi-cant improvements in shaft friction over conventional bored or driven piles but the use of such high values must be used with caution and in conjunction with suitable pile load testing. The shaft friction of such piles is also highly dependent on the shape of the drilling tool, of which there are numerous variants in the marketplace.

The basic formula for calculating qs can be further simpli-fied to qs = βσ′v and a range of typical β values are included in Tables 54.5 and 54.6.

It is important, particularly with long piles in coarse-grained soils, to limit the calculated skin friction to achievable values: a sensible limit to average shaft friction of 110 kN/m2 should be applied. Values well in excess of this can be achieved for various installation techniques, but should ideally be proven by suitable pile load testing or relevant experience in similar ground conditions.

The ultimate shaft friction of driven piles in coarse-grained soils will approach a limiting value for pile lengths in excess of 20 pile diameters, a similar phenomenon to that for end-bearing. The limiting value is dependent on several fac-tors (including installation methods, due to friction fatigue effects). White and Lehane (2004) provides a useful refer-ence for this in addition to the work carried out by Jardine et al. (2005).

It is also important, particularly for bored and CFA piles in sands, especially silty sands, not to place too much reliance on the base capacity of the pile at working loads. It is common practice to design such piles with a factor of safety greater than 1.0 on the shaft capacity alone to limit settlements at working

conservatively to between 10 and 15 MN/m2 for dense sands or gravels, depending on the pile installation method, and to about half these values for medium-dense sands or gravels. Higher values can be adopted based on the results of suitable pile load testing or appropriate experience in similar ground conditions. For loose coarse-grained soils very little end-bear-ing resistance may be mobilised at displacements which are small enough for normal foundation requirements.

54.3.2.2 Shaft friction

The ultimate shaft friction of piles founding in sands or gravels is generally calculated using

qs = Ks σ′v tan δ, (54.6)

where Ks is earth pressure coefficient after pile installation, σ′v is effective vertical overburden pressure and δ is angle of skin friction of the pile.δ is generally derived from φ′cv, with the correlation depen-

dent on the roughness of the pile–soil interface, resulting from the combination of pile installation method and soil charac-teristics, generally ranging from 0.75 to 1.0. For most bored piles the angle of skin friction may be taken as approximately equal to the critical state angle of friction of the soil. φ′cv values are typically between 30° and 33° for uniformly graded sands; 34° and 37° for well-graded sands and gravels; lower values are relevant for coarse-grained soils with rounded particles, and higher values if particles are angular; the above assume clay contents are low, say less than 5–10%. Reference should be made to Fleming et al. (2008) for further details. For full displacement driven piles, assuming δ equals 0.75−1 φ′cv would usually be appropriate.

For most bored piles, Ks = 0.7 is appropriate, but for CFA piles the value of Ks is more sensitive to the soil characteristics and appears to be dependent on technique. The following val-ues are recommended for the design of CFA piles:

Clean medium/coarse sand Ks = 0.9

Fine sand Ks = 0.7–0.8

Silty sand Ks = 0.6–0.7

Interlayered silt and sand Ks = 0.5–0.6

Table 54.4 Methods of calculating end-bearing

Material Nq Value Reference

Silica sand Nq = 40

Nq plotted against φ′

Nq related to φ′, relative density and mean effective stress

Nq from cavity expansion theory, as a function of φ′ and volume compressibility

API (1984)

Berezantzev et al. (1961)

Fleming et al. (2008)

Vesic (1972)

Uncemented calcareous sand Nq = 20

Typical range of Nq = 8–20

Nq determined for reduced value of φ’

Datta et al. (1980)

Poulos (1988)

Dutt and Ingram (1984)



When the pile toe is within the zone of infl uence of the adja-cent weaker layer, the end-bearing capacity can be assessed from Figure 54.7 by interpolating between the capacity for the coarse-grained founding layer and the adjacent weaker layer.

A key practical issue for pile design in layered deposits is the reliable identifi cation of the level, thickness and continuity of the proposed coarse-grained founding layer. Hence, a large number of boreholes/CPTs may be required to verify that the founding layer does not thin out, and is not in the form of iso-lated lenses.

If there is signifi cant uncertainty, which can be a particular issue in glacial deposits, then it may be appropriate to base pile design on the worst credible stratifi cation unless proven otherwise by full-scale testing. A common approach is to seek a factor of safety on shaft capacity alone in excess of 1.0, based upon the worst-case stratigraphy in terms of shaft capacity, and to place a limit on end-bearing, e.g. assuming a clay soil is present at the pile toe.

54.3.4 Correlations with SPT/CPT

In general, it is possible to use either ‘indirect’ or ‘direct’ cor-relations. In the ‘indirect’ approach the SPT or CPT is used to correlate the measured in situ test value with laboratory test parameters or other indices of soil behaviour (such as rela-tive density for sands/gravels). The methods outlined in sec-tion 54.3.1 or 54.3.2 are then used. In the ‘direct’ approach, the measured in situ test value is directly correlated with shaft fric-tion or end-bearing capacity. Both approaches can be of value, especially when there is a high level of uncertainty associated with predicting the ultimate pile capacity, e.g. driven piles in coarse-grained soils.

loads to acceptable levels as determined through specifi ed ser-viceability limits. If base resistance needs to be mobilised at working loads (even for driven piles), then a separate service-ability check must be carried out to assess if the working load can be mobilised at tolerable settlements.

54.3.3 Layered soils

From sections 54.3.1 and 54.3.2 above it should be clear that piles in coarse-grained soils may mobilise relatively high end-bearing capacity compared with fi ne-grained soils. Hence, when designing piles in interbedded sand and clays the loca-tion of the pile toe is of fundamental importance. A common practical situation when designing driven piles would be to attempt to take advantage of the potentially high end-bearing resistance of coarse-grained layers. However, if the pile toe is close to the interface with adjacent fi ne-grained layers then the pile will not develop its full end-bearing capacity (i.e. the end-bearing resistance which would be calculated by the methods outlined in 54.3.2 above).

Meyerhof (1976) fi rst considered this important issue, and suggested that end-bearing capacity would be reduced if the pile toe was within 10B (where B is pile diameter) of the adja-cent weaker layer (Figure 54.7). Subsequent work by Meyerhof and Sastry (1978) and Matsui (1993) indicates that Meyerhof’s original guidance is, in general, overconservative; nevertheless it remains a useful guide for preliminary design.

Table 54.5 values for driven piles (where qs = v’)

Material value Reference

Silica sand β = 0.15–0.35 (compression)

β = 0.10–0.24 (tension)

β = 0.44 for φ′ = 28°

β = 0.75 for φ′ = 35°

β = 1.2 for φ′ = 37°

McClelland (1974)

Meyerhof (1976)

Uncemented calcareous sand

β = 0.05–0.1 Poulos (1988)

Table 54.6 values for bored piles (where qs = v’)

Material value Reference

Silica sand β = 0.1 for φ′ = 33°

β = 0.2 for φ′ = 35°

β = 0.35 for φ′ = 37°

β = F tan (φ′ – 5°)

where F = 0.7 (compression) & F = 0.5 (tension)

Meyerhof (1976)

Kraft and Lyons (1974)

Uncemented calcareous sand

β = 0.5–0.8

limiting shaft resistance = 60 to 100 kN/m2

Poulos (1988)D

H

qoqoq

qp

q l

~10

B

Weak soil

Dense sand

Weak soil

Figure 54.7 Relation between ultimate point resistance of pile and depth in thin sand overlying weak soilReproduced from Meyerhof (1976), with kind permission of ASCE

Single piles


It is usually appropriate to also use SPT ‘N’ values to assess relative density, with:

overburden correction, ■■ N1 = CNN;

overburden correction factor, ■■ CN = 2/[1 + (σ ′v/100)];

relative density, ■■ Dr = [(N1)60/60]0.5.

Here, N is measured SPT ‘N’, σ′v is vertical effective stress, and (N1)60 is the SPT blow count normalised to a vertical effective stress of 100 kN/m2 and corrected to 60% of free-fall energy. The above relationships are for uncemented, normally consoli-dated sands (with a median particle size of less than 0.5 mm). For soils of large particle size (i.e. gravels/cobbles) the above correlations are unconservative. CIRIA report 143 provides further guidance.

Meyerhof (1976) proposed a set of direct correlations between SPT ‘N’ values and end-bearing resistance/shaft fric-tion as follows:

qb = 0.4 ‘N’ (MN/m2) for sands, when P > 6D, (54.9)

qb = 0.3 ‘N’ (MN/m2) for silts, when P > 6D. (54.10)

Note: P is penetration of pile toe beneath surface of coarse grained soil. The ‘N’ value is averaged over about 3 pile diam-eters above/below the pile toe. For shaft friction, Meyerhof recommends the following:

qs = 2 ‘N’ (MN/m2) for driven piles, (54.11)

qs = 1 ‘N’ (MN/m2) for bored piles.

54.3.6 CPT54.3.6.1 Indirect methods – fine-grained soils

As with the SPT a number of correlations between in situ undrained shear strength, cu and CPT cone resistance, qc, have been proposed.

Many of the most recently developed driven pile design methods directly use CPT data to derive the pile capacity, and the available evidence indicates that these methods are more reliable than the older empirical methods. When driven piles are proposed it is strongly recommended that CPTs are used to the maximum practical extent.

54.3.5 SPT54.3.5.1 Fine-grained soils

A number of correlations between SPT results (‘N’ value) and in situ undrained strength, Su, of silts and clays have been pro-posed. The most commonly adopted correlation within the UK for heavily overconsolidated clays is that proposed by Stroud and Butler (1975): Su = 4 to 6 × ‘N’ kN/m².

Other published correlations are available, but these can be unreliable. Some of the differences in correlations may result from variations in the methodology adopted in the field whilst undertaking SPTs, or the types of, or interpretive methods used for, pile tests rather than any significant differences in the soils being tested. It is to be noted that any correlations should be compared to all other available data, e.g. results from triaxial tests, field shear vane test results, etc.

There are direct methods of calculating single pile capaci-ties from the results of SPTs undertaken in fine-grained soils, although there is limited experience with their use for UK conditions, and they should therefore be used cautiously.

Poulos (1989) has proposed the relationship

qs = α + βN, (54.7)

where qs is the unit ultimate shaft friction, N is the SPT pen-etration resistance, and α and β are constants depending upon soil and pile type.

The relationship between SPT N value and unit ultimate end bearing proposed by Poulos is

qb = KN, (54.8)

where K is a constant depending upon soil and pile type and N is the SPT penetration resistance.

The values of α, β, and K are highly dependent on the type of pile and method of pile construction. CIRIA Report 143 and Poulos (1989) give suggested values of α, β and K.

54.3.5.2 Coarse-grained soils

Correlations between SPT ‘N’ and peak friction angle are avail-able, and the correlation by Peck et al. (1967) is often used and is usually conservative. This correlation is given in Table 54.7, and although fairly crude (it ignores many fundamental fac-tors which govern mobilised friction in coarse-grained soils), it is useful for the preliminary assessment of the end-bearing resistance of small diameter (say less than 0.5 m) driven piles of moderate length (say 10–15 m long). The SPT ‘N’ values are uncorrected. Table 54.7 SPT ‘N’ correlation to w′

SPT ‘N’ w′ (degrees)

< 5 28

5 29

10 30

15 31.5

20 33

25 34.5

30 36

35 37

40 38

> 40 38

N.B. use φ′ > 38° with caution



ultimate base capacities measured in pile tests against mea-sured values of CPT qc. They found that

qbu = 0.9 qcm, (54.13)

where qbu is ultimate, or ‘plunging failure’, capacity (refer to section 54.8 below) and qcm is the weighted average of qc val-ues within the zone of influence of the pile toe (up to 8D above and 4D below the toe, where D is pile diameter), based upon the ‘Dutch’ method.

Lehane et al. (2005) carried out further analysis on a rela-tively large pile test database and found that

qb0.1 = 0.6 qcm, (54.14)

where qb0.1 is bearing capacity mobilised at a pile settlement equal to 10% of the pile diameter and qcm is as given above in equation (54.13).

In both the above studies, the important influence of adjacent weaker layers was emphasised, and this is readily taken into account by the selection of qcm based upon the Dutch method.

Shaft resistance is correlated to qc in both the ICP design and UWA design methods (Lehane et al., 2005).

54.3.7 Rock

The consideration of pile design in weak rocks differs widely for replacement piles (bored and CFA) and displacement piles (driven or bored). This is predominantly because it is possible to install significant rock sockets with replacement techniques, particularly when using modern high-torque piling rigs, whereas displacement systems typically meet refusal at the top of the weak rock strata or at a very short penetration into the rock. Generally strong rocks underlying a development site would not require bearing piles to penetrate them by any significant dis-tance. There are of course occasional exceptions to this general rule, for example where adjacent railway running tunnels are required to be isolated from bearing pile loads with slip liners.

54.3.7.1 Shaft capacity, i.e. rock sockets (generally applicable to replacement piles only)

Transfer of pile loads by shear transfer along the pile-shaft-to- rock interface is relatively complex and depends on a number of factors including:

rock strength;■■

rock socket roughness;■■

rock–concrete bond;■■

degree of ‘contamination’ of the rock socket with overlying super-■■

ficial deposits;

potential degradation (smearing/polishing) of rock socket through ■■

the boring/drilling process.

Comprehensive details of load transfer mechanisms can be found in CIRIA report R181 (Gannon et al., 1999).

The general relationship is given by

cu = qc/Nk + p0, (54.12)

where Nk is a cone factor, qc is the unit cone end resistance (kN/m²) and p0 is the total overburden pressure (kN/m²).

It is important to note that some correlations include the effect of overburden pressure, p0, within the cone factor, Nk. Some correlations are specifically for use with the piezone.

The cone factor, Nk, is a variable which is affected by many factors, including the rate of penetration, strength anisotropy, soil fabric (e.g. fissuring, laminations), the degree of overcon-solidation of the clay deposit and the reference tests which the CPT is compared against. Overconsolidated clays generally tend to exhibit significantly higher values of Nk (especially if fissured) than normally consolidated clays.

Comprehensive guidance for the selection of appropriate values of Nk can be found in Lunne et al. (1997).

54.3.6.2 Direct methods – fine-grained soils

Fully mobilised unit end-bearing, qbu, of a pile can be directly correlated to the unit cone end resistance, qc, determined by CPT testing. Effects of scale and rate of testing mean that the ratio qbu/qc is usually less than one (Poulos et al., 2000).

The Imperial College Pile (ICP) design method (Jardine et al., 2005) provides methods for calculating base capacity for both closed and open-ended driven piles directly from CPT tests. The ICP design method does not, however, provide a direct correlation to pile shaft friction.

Estimates of pile shaft friction for driven piles can be made from CPT sleeve friction, e.g. qs = fc, but a high degree of cau-tion is required, since there may be large differences in the degree of consolidation of the soil around the pile when com-pared to the CPT test. Results of pile tests have indicated that long-term shaft friction values can be as much as 500% greater than those measured during driven pile installation (Fleming et al., 2008).

Correlations between unit cone end resistance, qc, and unit shaft friction, qs, generally yield more reliable results than those derived from CPT sleeve friction. These correlations themselves have a reasonably wide degree of variation, e.g. qs = qc/10 (Fleming and Thorburn, 1983) and qs = qc/40 (Thorburn and McVicar, 1971).

A more reliable method for assessing pile shaft friction is usually attained from the methods outlined in section 54.3.1.

54.3.6.3 Coarse-grained soils

As noted in Chapter 22 Behaviour of single piles under vertical loads, where a pile toe is located within a granular stratum, the CPT cone tip can be regarded as a model of the pile toe, and so the base capacity, qb, will correlate with the cone resistance, qc. Care needs to be taken to suitably derive an appropriate value of qc given the typical variability of the magnitude of qc. White and Bolton (2005) have carried out a careful comparison of

Single piles


area or rock formation for similarly constructed piles to those under consideration.

It must be noted that particular care needs to be taken when calculating design values for rock socket friction for mud-stones. Experience has shown that the calculated values for unit shaft friction based on mudstone rock strengths are rarely achieved in practice, largely due to softening/polishing effects during rock socket construction.

The most recent published guidance for pile shaft friction in chalk is based on an effective stress approach. Reference should be made to CIRIA report PR86 (Lord et al., 2003) for specific guidance.

54.3.7.2 Base capacity (applicable to both displacement and replacement piling techniques)

Where piles are driven to practical refusal on strong rock the limiting factor on pile capacity is usually the maximum allow-able direct stress on the pile cross-section. In the UK this is typically taken as 0.25 × fcu. Care should be taken when driv-ing pre-cast concrete piles to ensure that the piles are not overdriven, as this could result in damage to the pile section. When driving steel piles, care should be taken not to overdrive the pile, as this could cause shattering of the bearing stratum, resulting in a reduced pile capacity. Care should also be taken to ensure that the design assumptions relating to rock material (e.g. strength) are compatible with the rock mass (e.g. joint infilling, joint orientation, joint aperture, solution features, weaker underlying strata).

Bored piles can also be constructed to practical refusal on strong rock. The definition of practical refusal should be agreed in advance with the engineer where it is anticipated this will occur. In certain circumstances the definition of practi-cal refusal may be qualified subject to undertaking satisfactory maintained load tests.

For both driven and bored piles constructed within rock, the ultimate unit end-bearing capacity is often taken as the design UCS value for the rock stratum, i.e. allowable unit end-bearing qb all = UCS/FoS. The factor of safety is usually a reasonably high value, typically 2.0–2.5, which will ensure that the rock present at the toe of the pile will not become overstressed, even for predominantly end-bearing piles, again subject to maintaining compatibility between rock material and rock mass.

Where piles derive a significant proportion of their capacity from shaft friction, there may be less concern over consider-ation of the end-bearing component of capacity in terms of the pile load–settlement behaviour, particularly if the pile has a factor of safety on shaft capacity alone of 1.5 or greater.

Where a significant proportion of the pile capacity is being derived from the end-bearing component, care must be taken to ensure that the pile toe remains in intimate contact with the rock. Factors such as soil heave (from overburden removal or driving further displacement piles) and disturbance due to breaking down the piles can lead to lifting of the pile toe,

Driven piles will generally not penetrate significant dis-tances into rock of moderate strength. For driven piles in weak to moderately strong rock, skin friction can be calculated from the equation

qs = 0.5 Ks σ′vo tan δ, (54.15)

where Ks is an earth pressure coefficient, σ′vo is the effective overburden pressure and δ is the angle of friction between the rock and pile shaft.

In general, pile driving into mudstone will create a ‘skin’ of remoulded material around the shaft, and mobilised shaft friction will only be a small fraction of that mobilised around bored (replacement) piles.

Skin friction in weathered argillaceous rocks, particularly weathered mudstones/siltstones, can be calculated using meth-ods described in section 54.3.1 – pile capacity in fine-grained deposits – although care is needed to suitably account for remoulding of material at the pile–ground interface.

For bored cast-in-place piles (replacement piles) a number of correlations between the unconfined compressive strength of the rock and rock socket skin friction have been proposed. The most common correlations are those proposed by Horvath (1978), Rosenberg and Journeaux (1976), Williams and Pells (1981) and Rowe and Armitage (1987).

Whitworth and Turner (1989) summarised the various pro-posed formulae for calculating ultimate unit shaft friction (see Table 54.8):

Table 54.8 Typical correlations between qwc and ultimate shaft friction

Pile design method Ultimate unit shaft friction, qs (kN/m²)

Horvath (1978) 0.33(quc)0.5

Horvath and Kenney (1979) 0.20–0.25(quc)0.5

Meigh and Wolski (1979) 0.22(quc)0.6

Rowe and Armitage (1987) 0.45(quc)0.5

Rosenberg and Journeaux (1976) 0.375(quc)0.515

Williams and Pells (1981) αβ(quc)

here, quc is unconfined compression strength, α is a reduction (adhesion) factor relating to unconfined compression strength and β is a correction (rock socket side friction) factor related to discontinuity spacing within the rock mass.

Clearly, the above correlations can yield a wide range of val-ues for ultimate shaft friction. Some of these differences may be due to the construction methods adopted for test piles, dif-ferent back analysis methods to derive the correlations, or dif-ference of the rock materials and/or rock masses in which back analysed test piles were constructed. A common approach is to take an average value from each of the correlations to gain a starting position in terms of design values and use additional information from historical pile testing in a given geographical



the potential for soft spots and dissolution features beyond the pile toe. Similarly the design of bored or CFA piles in weak rocks may often be dictated by ensuring a factor of safety as high as 1.5 on the shaft capacity alone to allow for the poten-tial for machining/polishing or softening/remoulding of the pile shaft, or disturbance of the pile base, and to limit settle-ments at working load.

If a pile relies predominantly on end-bearing, then great care will be required both for design and construction, and more specialist techniques and expertise will normally be required. Site-specific preliminary pile tests may be needed unless there is significant previous experience in similar ground conditions and using similar piling techniques.

54.4.2 Partial safety factors

Partial factors are commonly used in structural design and these are applied to the particular parameters and loading con-ditions to reflect more accurately where the uncertainties lie. Most modern design codes, in particular the Eurocodes such as EC7 (BS EN1997 Part 7), recommend pile design based on the use of partial safety factors. This code does allow reduced fac-tors of safety where there are several pile tests on a site. Such an approach encourages the use of preliminary pile testing.

The factor of safety, whether partial or a lump factor, is a factor on strength. However, as reported by Atkinson (2000), the key parameter that governs pile behaviour is stiffness of both the ground and the pile. The factor of safety is only of relevance in determining the likelihood of satisfactory per-formance beyond working load. For example, an end-bearing pile can fail to meet the required performance at working load because of low stiffness and still have the prescribed factor of safety. Conversely it can perform at working load and still not meet the required factor of safety. In either case it is necessary to determine whether the behaviour is of sufficient concern to be detrimental to the structure.

The main difficulty is in selecting criteria for settlement and factors of safety that are compatible and allow for the likely variability of the ground. This can only be done by making a prediction of pile behaviour based on the best estimate of the soil parameters and allowing a reasonable margin in the permit-ted test criteria to allow for natural variability in the ground.

54.5 Pile settlementStructures do not generally suffer due to total settlement per se. It is differential settlement of, and the vulnerability of services connected to, structures that give rise to most concern. Refer to Chapter 52 Foundation types and conceptual design principles.

One of the reasons for introducing a lumped factor of safety when considering allowable pile capacity (such lumped safety factors applied to calculated ultimate pile capacities are most often in the range 2.0–3.0) is to ensure that settlement of single piles under working load (i.e. at serviceability limit state) is within tolerable limits for most structures.

which will result in greater settlements upon loading than would otherwise have been experienced. For driven piles it is often considered to be good practice to re-drive piles (subject to programme and access constraints) to ensure the toe of the pile is in good contact with the bearing stratum.

Sloping bedrock can pose particular risks to pile construc-tion. Driven and bored piles may be deflected by steeply bed-ded rock surfaces. Driven piles may be installed to a refusal criterion yet upon loading cause instability of the inclined rock mass due to downward sliding of blocks of rock. Infilled quarries can pose a particular risk to piling: the position of the highwall is critical in assessing the required pile lengths and suitability of particular construction techniques. It may not be possible or practicable to construct piles in the zone where highwalls are located.

When constructing bored piles within rock, care should be taken to ensure that the base of the pile bore is thoroughly cleaned, so that the design end-bearing capacity can be achieved without adversely affecting the piles’ load–settlement behaviour. This can only be realistically achieved when large diameter bored piling techniques are adopted, facilitating the use of cleaning buckets. Typically, large diameter bored piling ranges from 600 mm to 3 m diameter.

Reference should be made to CIRIA report PR86 (Lord et al., 2003) for specific guidance on constructing piles within chalk.

54.4 Factors of safety54.4.1 Global safety factors

Historically, the design of piles in the UK has adopted global factors of safety where the safe working load, SWL, is given as

SWL = Qb/Fb + Qs/Fs. (54.16)

Although overall factors of safety are not necessarily the best way to govern pile performance, the numbers traditionally adopted are:

2.0 overall minimum – usually with load test verification on ■■

expendable piles;

2.5 overall or 1.5 (on shaft) with 3.0 (on base) – usually with ■■

working pile load tests;

3.0 overall – may not require any load testing.■■

According to BS8004, for spread foundations a factor of safety of 3 is usually adopted, and for piles a value of between 2 and 3 is normally used, the lower value representing condi-tions where there is reliable information on the ground and pile behaviour. The above factors of safety should only be used as a guide and different values and combinations can be used in conjunction with suitable experience.

There are a number of ground conditions where different factors are used, placing more reliance on either end-bearing or shaft friction. Examples include piles in chalk, where large factors on end-bearing are adopted, as high as 10, to allow for

Single piles


are predominantly frictional piles (i.e. piles which carry the majority of applied load by shaft friction). Patel (1992) has summarised pile test data for friction piles in London Clay (Figure 54.8), which indicates that at typical working loads (factor of safety of 2 or more) the single pile settlement would be expected to be less than about 1.0% of the pile diameter. Similar behaviour would be expected for comparable stiff over-consolidated fi ssured clays (Gault Clay, Weald Clay, etc.).

An additional factor which must be considered in any assess-ment of pile settlement is the magnitude of pile head displace-ment caused by elastic shortening of the pile under load. This can be a signifi cant proportion of the total pile head settlement under load, but becomes increasingly signifi cant when consid-ering long slender piles constructed through deep weak depos-its (i.e. transmitting a signifi cant proportion of applied load to competent strata).

A specifi ed factor of safety does not guarantee a specifi c settle-ment and it is important to review the risk of excessive settlement as part of the pile design process. For most piling techniques in a range of ground conditions, except those that rely predominantly on end-bearing, it is good practice to ensure a suitable factor on the shaft capacity alone to limit settlements under working loads, i.e. to ensure suitable serviceability performance. It is recommended that the settlement of end-bearing piles is always assessed. Group settlements will need to be subject to separate consideration to ensure that pile group settlement is tolerable for the structure (refer to Chapter 55 Pile-group design).

It is widely acknowledged that more settlement is required to mobilise end-bearing than shaft friction. It is also gener-ally to be expected that piles which rely signifi cantly on end-bearing (carrying a low proportion of the total applied load in shaft friction), will settle more under load than piles which

1.0Straight shafted piles

I = Total pile length in groundd = Shaft diameter

Envelope of piletest results

Loadfactor

1.25

1.5

1.671.77

2.0

2.25

2.5

10

5.0

l /l /l d = 10 to 31d = 10 to 31d

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

00 0.2 0.4 0.5 0.0

Settlement ratio

Envelope of pilelope of piletest resulttest results

I /d = 31 to 43d = 31 to 43d

d%

δrm1.0 1.2 1.4 1.6

Figure 54.8 Results of loading tests on straight shafted bored piles in London Clay – normalised load settlement chartReproduced from Patel (1992)



54.6 Pile behaviour under lateral load54.6.1 Introduction

Piles can be subjected to lateral loads arising from different sources. These sources typically fall into two main categories: active loading (where external forces are applied to the pile) and passive loading (where movement of the soil immediately adjacent to the pile imposes bending stresses on to the pile). A more refined classification of the sources of lateral load may involve the following (after Elson, 1984):

Static: e.g. structural reactions, lateral earth pressure.■■

Dynamic – cyclic: e.g. from operating reciprocating equipment, ■■

earthquake, cyclic wind loading (wind turbines), wave action.

Dynamic – transient: e.g. general wind loading, berthing, vehicu-■■

lar braking (impact) loads.

Other: e.g. undrained soil movements, soil consolidation adjacent ■■

to piles, thermal movements, soil creep, soil shrinkage/expansion (soil moisture changes).

Geotechnical capacity is rarely an issue even in weak/soft soils. The issues are usually either ones of serviceability (i.e. pile deflection/deformation) or structural capacity of the pile. Passive loading is a particularly high risk in soft soils, where large global ground movements may readily develop (refer to Chapter 57 Global ground movements and effects on piles).

Piles subjected to cyclic loading require particularly close attention since cyclic loading will tend to modify soil behaviour and can lead to progressive failure. Consideration of dynamic cyclic and dynamic transient loads falls outside the scope of this chapter of the manual (refer to Chapter 60 Foundations subject to cyclic and dynamic loads).

54.6.2 Deformation/failure modes due to ‘active’ loading

Piles subjected to lateral loads will experience an increase in normal stress in front of the pile (in the direction of the lat-eral load), and a decrease behind the pile. Pile deformation is usually limited to approximately 10 pile diameters below ground level; the concept of a critical pile length (which is dependent upon the ratio of soil/pile stiffness) is particularly useful in this context (reference should be made to Fleming et al., 2008).

There are two principal failure modes for unrestrained (or free headed) single piles subjected to ‘active’ lateral loading (Figure 54.9). Unrestrained short rigid piles will fail by rota-tion when the passive resistance of the soil (in front of the pile above the point of rotation, and behind the pile below the point of rotation) are exceeded. Long piles will tend to fracture at some depth down the pile shaft. For the purposes of analysis a plastic hinge is assumed to develop which is capable of trans-mitting shear. The upper portion of the pile shaft above the plastic hinge will deform whilst the lower pile shaft, below the hinge, will not move significantly.

A number of different methods have been proposed over recent years: elastic approaches (e.g. Poulos and Davis, 1968; Tomlinson, 2004), simplified empirical methods (e.g. Burland et al., 1966), finite element/boundary element analyses, and hyperbolic analytic methods (e.g. Fleming, 1992).

Burland’s proposed simplified empirical approach is spe-cifically for the assessment of settlement of end-bearing under-reamed piles in London Clay (Burland et al., 1966). Burland showed that where plate load tests are taken to fail-ure, the normalised load settlement curve can be plotted, and provided the base factor of safety exceeds 3, the settlement of the pile base can be obtained for any diameter of pile base. Full details of this empirical method are provided in Burland et al.’s paper. Alternatively reference may be made to Tomlinson (2004).

In his paper, Fleming (1992) noted that sophisticated ana-lytical techniques require sophisticated input data (e.g. certain geotechnical parameters) which are not usually available from routine site investigations. The method described by Fleming combines hyperbolic functions describing both the shaft and base components of pile capacity, together with the compo-nent of pile settlement derived from elastic shortening of the pile. The method takes advantage of the fact that the hyper-bolic function adopted requires definition only of its origin and either its initial slope or a single point. The shaft flexibil-ity parameter, Ms, does not vary significantly, with Ms values typically between 0.0010 and 0.0025 for many pile types and ground conditions. However, as expected the base stiffness parameter, Eb, varies across a wide range for different pile and soil/rock types, and is particularly sensitive to pile construc-tion methods and workmanship. Eb, should not be confused with other ‘elastic’ soil parameters which are quoted in the lit-erature: it is essentially an empirical parameter derived from back analysis of pile tests and is as sensitive to pile construc-tion method as to the fundamental soil properties. Guidance on appropriate parameters is given in Fleming (1992). The definition of axial capacity adopted within Fleming’s method is the plunging-failure (or asymptotic) definition, i.e. that load at which the soil resistance is fully mobilised and where set-tlement continues indefinitely for the load magnitude reached. Corke et al. (2001) provide interesting practical examples of Fleming’s method, in the context of developing pile testing performance criteria.

Fleming’s method is a powerful and simple method for single pile settlement analysis; it is especially useful for large diameter piles (say in excess of 1.0 m diameter) and for end-bearing piles.

Fleming’s method has a significant additional benefit, in that it can be used in reverse to back analyse the results of maintained load tests to determine shaft friction, end-bear-ing and elastic shortening of the pile, subject to the pile test having moved the pile sufficiently to mobilise end-bear-ing sufficiently, typically more than about 5% of the pile diameter.

Single piles


Free headed piles are free to rotate and are not subject to orthogonal restraint.

In free headed piles the bending moment in the pile at ground level is positive and acts in the same direction as the applied lateral load.

In fi xed headed piles the bending moment in the pile at ground surface is negative and acts in the opposite direction to the applied lateral load.

54.6.4 Lateral resistance of single piles

The ultimate lateral resistance of single piles can readily be esti-mated from the solution proposed by Broms (1964a, 1964b). Whilst Broms’s solution may be considered to be on the con-servative side, it is often recommended for routine design given the relative simplicity of the solution and the fact that the governing criterion is rarely the ultimate lateral soil resistance.

Other authors have provided more detailed/refi ned solu-tions. Full details of these solutions can be found in CIRIA Report 103 (Elson, 1984), Appendix A.

Once the bending moment and shear force envelopes have been derived adopting appropriate analysis methods, the pile can be designed in accordance with relevant structural design codes of practice (e.g. BS EN1997 Part 2).

54.6.5 Lateral deformation

When considering the lateral defl ection of laterally loaded piles, the near-surface soil parameters merit particular atten-tion. This can often cause some problems since the very near surface deposits (generally the upper 2–3 m) are often those about which very limited information is obtained during the ground investigation.

Lateral pile loading tests are occasionally specifi ed to be undertaken on piles. It is the authors’ experience that these are often specifi ed without full cognizance being taken of the head

For restrained (or fi xed headed) piles (Figure 54.10) there are three possible failure modes:

(a) translation;

(b) 1-hinge failure (at underside of pile cap/restraint);

(c) 2-hinge failure (at underside of pile cap/restraint and at some depth).

54.6.3 Fixed and free headed piles

Fixed headed piles are restrained at head level, typically by a pile cap (pile groups of three or more) or ground beams (pairs of piles or single piles) or potentially by a structural slab (if so designed). Piles need to be restrained in two orthogonal direc-tions to be considered truly fi xed headed.

Figure 54.9 Possible failure modes for free headed piles: (a) rotation; (b) hinge failure (Broms, 1964a,b)Courtesy of Cementation Skanska

Figure 54.10 Possible failure modes for fi xed headed piles: (a) translation; (b) 1-hinge failure; (c) 2-hinge failure (Broms, 1964a,b)Courtesy of Cementation Skanska



behaviour, proposed by Brettmann and Duncan (1996), is the ‘characteristic load’ method. Parameters are provided for sands and clays, and for free and fixed headed piles.

54.7 Pile load testing strategyStrategies for pile load testing often lack clear objectives. The requirements for testing are frequently driven by a desire to comply with regulations and to follow common practice. Testing is rarely seen as a part of a value engineering process, and therefore opportunities to optimise the foundation solution are often missed.

It is important that load testing of piles is considered as early as possible and included in the initial cost plan and programme stage. The programme must allow sufficient time for suitable tests to be carried out and for an objective appraisal of the test results and subsequent design revisions or value engineering to be carried out.

Without such an early review to set clear load testing objec-tives, combined with poorly specified requirements, avoidable problems will arise. Common problems include:

inappropriate test method specified and lack of flexibility in the ■■

testing regime;

final working conditions not modelled by the load tests;■■

piles not tested to failure or sufficient load to allow suitable ■■

evaluation;

insufficient time to carry out tests and to evaluate the results;■■

unrealistic/unachievable performance criteria specified;■■

no provision for value engineering;■■

unrealistic expectations of the possible benefits of load testing.■■

54.7.1 Objectives

The strategy for pile testing needs to be established at the time the piles are designed. For most projects the main purpose of pile testing is to validate the design before construction and to check compliance with the specification during construc-tion. However, in many cases there are considerable benefits in using testing for design development in order to identify the optimum solution. Testing strategies can therefore be divided into four main categories:

design development;■■

design validation;■■

quality control;■■

research.■■

The scope of testing will depend on the complexity of the founda-tion solution, the nature of the site and the risk of piles not meet-ing the specified requirements. The designer therefore needs to assess the risks and develop the testing regime accordingly. So what are the main considerations when assessing the risks?

Site investigation data – are they sufficient?■■

fixity of the piles in the final construction. The lateral load– deformation behaviour of a fixed head pile is very different from that of a free headed pile, and this must be taken into account when considering how the results of a lateral pile test (usually free headed) will be analysed and used for the design of permanent piles (usually fixed headed). The analysis method used for lateral pile test interpretation and permanent works design must be sophisticated enough to allow for:

(i) potential changes in pile head fixity;

(ii) potential differences in load eccentricity between pile test and permanent works;

(iii) potential changes in near-surface ground conditions between pile test and permanent works.

If there is a coherent testing and analysis strategy then lateral load testing of piles, particularly those whose load deflec-tion behaviour is deemed to be of significant importance, or where particularly high lateral loads will be applied to the pile (relative to the ground conditions and the pile geometry), can provide a very useful check on performance of piles and the analysis techniques adopted for pile design.

The two analysis approaches commonly adopted for calcu-lating the lateral displacement of single piles subjected to lat-eral loads are the sub-grade reaction approach and the elastic continuum approach. Lateral load–deformation behaviour is strongly nonlinear even under moderate lateral loads, and lin-ear elastic methods are usually inappropriate for piles subject to significant lateral loads. Ignoring nonlinear behaviour can lead to grossly inaccurate and unconservative predictions of lateral deformation.

The sub-grade reaction approach models the ground as a series of discrete springs along the length of the pile shaft, i.e. as a Winkler idealisation of the soil mass. A refinement of the early models allows the spring stiffness to vary along the length of the pile (Reese and Matlock, 1956) and a fur-ther refinement is the p–y form of the Winkler soil model (Matlock, 1970; Reese et al., 1974, 1975). The disadvantage of the p–y form of the model is that selection of reliable p–y curves requires careful judgement, although typical forms of the curves have been presented by Reese et al. (1974) in sand and by Matlock (1970) in clays, and are summarised in the API code (American Petroleum Institute, 2000).

Both Poulos (1971) and Randolph (1981) have published solutions relating to the behaviour of piles under lateral loads based on elastic continuum models. One principal advantage of the elastic continuum models over sub-grade reaction models is that the method can be extended to pile groups, since interac-tion between adjacent piles can be modelled. Where large lat-eral loads are required to be resisted by piled foundations, for example marine/offshore structures or complex/unusual super-structures, it will be more appropriate to assess lateral displace-ments by adopting nonlinear elastic plastic methods. A useful practical procedure for estimating lateral load–deformation

Single piles


54.7.3 Working pile tests

Working pile tests are carried out during the main piling works to verify that the workmanship and materials meet the speci-fied requirements. Testing on the working piles comprises a range of quality control checks. These checks will normally include load tests, integrity tests and materials testing.

The strategy for testing working piles should address the following issues:

load/settlement behaviour – if settlement criteria are specified ■■

then load tests should be carried out;

risk of integrity problems – the number of tests should reflect that ■■

certain ground conditions and pile systems may have increased risk;

installation criteria – use pile records, sets and dynamic testing ■■

where appropriate to confirm the installation criteria are being met.

The testing strategy for working piles should relate to the level of risk and the characteristics of the piling project. Test results and pile records should be continually reviewed during the works to reassess the risk level.

54.7.4 Number of tests

The need for preliminary testing in advance of the working piles, or in advance of new areas of piling on large projects, should be weighed against their cost and the time taken for installation, testing and analysis of results. It may be more economical and more logical to adopt a more conservative approach to design, for example by using larger or longer piles or reduced pile loads. Also, it may be more beneficial to undertake further site investigation to obtain information which is key to the success of the installation and hence achieve the necessary confidence in the piling. The value of preliminary testing undertaken to prove the design approach or pile performance is wasted unless the ground conditions are sufficiently well defined to allow the design to be applied to piles elsewhere on site.

The level of testing of working piles should consider the vari-ability of the ground conditions from the point of view of the demands these place on the control of pile construction, area by area. Ensuring experienced supervision and monitoring, plus good pile records, can be preferable to increasing the level of control testing above the typical 1% of all piles installed.

54.7.5 Methods of load tests

The various available methods of testing piles are best charac-terised by the duration that the force is applied to the pile and the strain induced in the pile. Tests involving large forces applied for several minutes or hours, such as static load tests (also known as maintained load tests), are used to assess pile load capacity, and small-energy low-strain tests are used to assess pile integrity. In dynamic and kinetic tests, although the force is comparable in magnitude to a static test, it is applied over a much shorter period than in a static load test. Careful consid-eration is therefore needed in the interpretation of the dynamic

Previous knowledge – are there similar piles that have been tested ■■

on adjacent sites?

Time – is there enough time to verify the design? If not, a conser-■■

vative approach is needed.

Cost – can the testing provide cost savings? What are the cost and/■■

or construction programme implications of tests failing?

For simple structures on a site where the ground conditions are well understood and there are test data from adjacent sites which have used similar piling solutions then the risks are low and the testing can usually be restricted to routine checks for compliance with the specification. For situations where the ground conditions or structural requirements are complex and there is little experience of similar piling work, then careful evaluation of the piling proposals is essential prior to embark-ing on the main piling works. The testing regime therefore needs to be considered in two phases, comprising preliminary pile testing before the main piling works, and testing of work-ing piles.

54.7.2 Preliminary pile tests

Preliminary testing is carried out in advance of the main pil-ing works to verify that the design assumptions and construc-tion method will achieve the required performance. The testing strategy for preliminary pile testing should address a specific set of stated objectives, which should include the following:

minimise risk – uncertainties about ground conditions, contrac-■■

tor’s experience, novel techniques;

optimise the design – pile diameters, lengths and factors of ■■

safety;

confirm pile installation criteria – founding strata identification, ■■

pile set, refusal criteria;

assess buildability – site variability, pile integrity, pile uplift and ■■

displacement, soil remoulding;

check the pile performance meets the structural requirements – ■■

load/settlement behaviour, compression, tension and lateral loading;

confirm safety procedures – piling platform, working method, ■■

inspection, testing methods;

assess environmental impact – noise, vibration, pollution.■■

Where piles are required to carry very heavy loads, it may be impractical to carry out full-scale load tests. In such circum-stances, consideration should be given to carrying out tests on smaller diameter piles using the same method of construction, provided the results of the tests can be extrapolated with some degree of confidence to predict the load settlement characteris-tics of the larger piles. The test piles should be founded at the same level and in the same soil as the works piles.

For preliminary pile tests it is essential to have a borehole and/or CPT close to the test pile so that the test results can be reliably evaluated.



will be a function of the mobilised deformation modulus of the ground immediately beneath the pile toe.

54.9 ReferencesAmerican Petroleum Institute (2000). Recommended Practice for

Planning, Designing and Constructing Fixed Offshore Platforms: Working Stress Design. API RP 2A-WSD. Washington, DC: American Petroleum Institute.

Atkinson, J. H. (2000). Non-linear soil stiffness in routine design. Géotechnique, 50(5), 487–508.

Bell, A. G. (2010). Foundation solutions for the urban regeneration of Glasgow city centre. In Proceedings of the DFI/EFFC 11th International Conference on Urban Regeneration, 26–28 May, 2010, London.

Berezantsez, V. G. et al. (1961). Load bearing capacity and deforma-tion of piled foundations. In Proceedings of the 5th International Conference of the International Society for Soil Mechanics and Foundation Engineering. Paris: ISSMFE, vol. 2, pp. 11–12.

Bown, A. S. and O’Brien, A. S. (2008). Shaft friction in London Clay – modifi ed effective stress approach. In Foundations: Proceedings of the Second British Geotechnical Association International Conference on Foundations, ICOF 2008 (eds Brown, M. J., Bransby, M. F., Brennan, A. J. and Knappett, J. A.). Watford: IHS BRE Press, vol. 1, pp. 91–100.

Brettman, T. and Duncan, J. M. (1996). Computer application of CLM lateral analysis to piles and drilled shafts. ASCE, Journal of Geotechnical Engineering, 122(6), 496–498.

Broms, B. (1964a). Lateral resistance of piles in cohesive soils. Journal of the Soil Mechanics and Foundations Division; Proceedings of the American Society of Civil Engineers, 90(SM2).

Broms, B. (1964b). Lateral resistance of piles in cohesionless soils. Journal of the Soil Mechanics and Foundations Division; Proceedings of the American Society of Civil Engineers, 90(SM3).

Burland, J., Butler, F. G. and Dunican, P. (1966). The behaviour and design of large diameter bored piles in stiff clay. In Proceedings of the Symposium on Large Bored Piles, Institution of Civil Engineers and Reinforced Concrete Association. London: Institution of Civil Engineers, pp. 51–71.

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Corke, D. J., Fleming, W. K. and Troughton, V. M. (2001). A new approach to specifying performance criteria for pile load tests. In Symposium Proceedings of Underground Construction, pp. 401–410.

Elson, W. K. (1984). Design of Laterally Loaded Piles. CIRIA Report 103. London: Construction Industry Research and Information Association.

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effects in order to derive static load capacities. The various test methods and their applications are detailed in Chapter 98 Pile capacity testing of this manual. There are a number of pub-lished documents dedicated solely to the subject of pile testing, such as Weltman (1980) and Handley et al. (2006), and these should be referred to for more detailed information on the sub-ject of pile load testing. Chapter 98 Pile capacity testing of this manual focuses specifi cally on pile load testing.

54.8 Defi nition of pile failureThe defi nition of axial load capacity (or pile ‘failure’) needs to be clearly understood, since this can be the cause of con-fusion. Unambiguous defi nition becomes very important for piles which rely mainly on end-bearing resistance. Two defi ni-tions are commonly used:

(i) Ultimate capacity, ‘plunging failure’ – failure is defi ned as the load at which settlement continues indefi nitely, shown as Pu in Figure 54.11. This is often a theoretical concept rather than a practical reality as most pile tests are not taken to suffi cient load to reach this value. Calculation methods based on bearing capacity theory will usually derive a ‘plunging failure’ load.

(ii) Deformation controlled capacity, ‘10% failure load’ – because the ‘plunging failure’ load cannot often be practi-cally reached and measured during a pile test. Most codes/standards defi ne failure as the load which is measured at a settlement equal to 10% of the pile base diameter, shown as PF0.1 in Figure 54.11.

For friction piles, there may be little practical difference between the above defi nitions. However, for end-bearing piles there can be a considerable difference between the capacities derived by (i) and (ii) above. Referring to Figure 54.11 Pile A and B both have the same ‘ultimate’ capacity, but in terms of a deformation controlled capacity (consistent with most code defi nitions of failure) Pile A has a higher capacity than Pile B. For end-bearing piles, the deformation controlled capacity

Load

Pile A

Pile B Pile A, PuPuP = Pile B, PuPuPPile A, PF 0.1PF 0.1P > Pile B, PF 0.1

Both Pile A + B, PuPuP > PF 0.1PF 0.1P

Normalised settlement(settlement/pile diameter)

10%

NB.

PF 0.1PF 0.1P

Pu

‘plungingfailure’ loadailure’ loadailure’

PF 0.1PF 0.1P

Figure 54.11 Defi nitions of pile failure

Single piles


Reese, L. C., Cox, W. R. and Koop, F. D. (1974). Analysis of laterally loaded piles in sand. In Proceedings of the 6th Offshore Technical Conference, Houston, Texas, Paper 2080, pp. 473–483.

Reese, L. C., Cox, W. R. and Koop, F. D. (1975). Field testing and analysis of laterally loaded piles in stiff clay. In Proceedings of the 7th Offshore Technical Conference, Houston, Texas, vol. 12, pp. 671–690.

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Poulos, H. G. (1971). Behaviour of laterally loaded piles – I: single piles. Journal of the Soil Mechanics and Foundations Division; Proceedings of the American Society of Civil Engineers, 97(SM5).

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It is recommended this chapter is read in conjunction with

■ Chapter 19 Settlement and stress distributions

■ Chapter 22 Behaviour of single piles under vertical loads

■ Chapter 82 Piling problems

All chapters in this book rely on the guidance in Sections 1 Context and 2 Fundamental principles. A sound knowledge of ground investigation is required for all geotechnical works, as set out in Section 4 Site investigation.

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Pile Design