0Proceedings of the Institution of Civil Engineers
Geotechnical Engineering 163
February 2010 Issue GEI Pages 13-22
doi: I0.I680/geng.20I0.I63.I.I3
Paper 900034
Received I6/04/2009 Accepted 2I/08/2009
Keywords: field testing & monitoring/foundations/piles & piling
Paul Doherty
PhD candidate, School of Architecture, Landscape & Civil Engineering, University College Dublin, Ireland
Kenneth Gavin
Lecturer, School of Architecture, Landscape & Civil Engineering, University College Dublin, Ireland
David Gallagher
Geotechnical Engineer, Golder Associates Pty Ltd, Hawthorn, Australia; formerly University College Dublin
Field investigation of base resistance of pipe piles in clay
P. Doherty be, K. Gavin PhD and D. Gallagher PhD
This paper presents the results of a series of field tests performed to study the effect of soil plugging during installation on the base resistance developed by an open-ended (pipe) pile in clay. Three instrumented pipe piles were installed in soft clay, and the installation resistance and soil plug development were recorded. The tests revealed that the annular base resistance was equal to the cone penetration test qc resistance and was independent of the soil plug development. In contrast, the soil plug resistance increased from a minimum when the pile was fully coring to a maximum when the pile was fully plugged. A simple expression is proposed that links the plug resistance to the cone penetration test qc value at the specified depth and the incremental filling ratio value developed during installation. This expression is shown to provide reasonable estimates of the plug resistance developed by a large-diameter pipe pile driven into stiff to hard clay in the North Sea. The proposed relationship will be particularly useful for modelling the installation resistance of monopile foundations.
NOTATION
q plug
qs Q t
annular area total base area
horizontal coefficient of consolidation
coefficient of consolidation
pile diameter
cone diameter
internal pile diameter
distance from pile tip
permeability constant
cone reduction factor
embedded pile length
slenderness ratio
plug length
bearing capacity factor
cone factor for undrained shear strength
atmospheric pressure
base load
plug load
shaft load
total pile resistance
annular end bearing pressure
pile end bearing pressure
average ultimate unit end bearing resistance
cone tip resistance
qcnet net cone tip resistance
qplug unit plug resistance; normal plug stress
qs closed-ended shaft stress
qsav average shaft resistance
qse external open-ended shaft stress
qsi internal open-ended shaft stress
s displacement
su undrained shear strength averaged over pile shaft sub undrained shear strength acting on pile base a empirical constant linking qs and undrained shear strength
av0 in situ vertical effective stress 1. INTRODUCTION
Open-ended pipe piles are commonly employed in offshore environments where large pile penetrations are necessary to resist uplift. The compressive axial load resistance developed by a pile is derived from a combination of external unit shaft resistance qse and unit base resistance qb (see Figure 1). For pipe piles the base resistance is composed of two parts: the annular resistance qann developed beneath the wall of the pile, and the internal shaft resistance qsi. The integral of the internal shaft resistance divided by the area beneath the soil plug is termed the unit plug resistance, qplug. The development of the soil plug during the installation of a pipe pile is described by the incremental filling ratio (IFR), given by
ALp A L
where ALp is the change in the length of the soil plug over a given increment of pile penetration AL. During the initial stages of driving the soil, plug length is coincident with the penetration depth, and the pile is said to be fully coring (IFR = 1). If, at some stage during installation, the internal shear stresses generate sufficient resistance to prevent further soil intrusion, the pile is said to be fully plugged (IFR = 0). Most pipe piles are driven in an intermediate condition: that is, the incremental rate of soil intrusion is somewhat smaller than the pile penetration, and the pile is said to be partially coring (0 < IFR < 1).
Most recent research on open-ended piles in clay (e.g. Jardine et al., 2005; Miller and Lutenegger, 1997; Xu et al., 2006) has concentrated on the development of shaft capacity, with the result that design methods for evaluating the base resistance of open-ended piles in clay have undergone limited development.
Figure 1. Modes of penetration for open-ended and closed-ended piles: (a) closed-ended; (b) fully coring; (c) partial plugging; (d) fully plugged
This is in stark contrast to open-ended pile design in sand, where concentrated research efforts (Chow, 1997; Lehane and Gavin, 2001; Paik and Salgado, 2003) have identified the critical role of soil plug development during installation on the ultimate base resistance, and have led to improvements in predictive design approaches, such as ICP-05 (Jardine et al., 2005) and UWA-05 (Lehane et al., 2005). The historical lack of research initiatives stems from the general opinion that, for typical pile geometries, the end bearing resistance contributes a relatively minor proportion of the ultimate axial resistance of a pile installed in clay (Jardine et al., 2005). Figure 2 uses traditional total stress design to show the impact of slenderness ratio (length/diameter) on the proportion of total resistance developed by the pile base. It is clear that for a typical closed-ended pile geometry (embedded length L — 10 m and diameter D — 0.3 m), with a slenderness ratio (L/D) of 30, less than 15% of the total pile resistance QT is developed by the base load Qb. However, for large-diameter monopiles, typical L/D ratios are likely to be significantly less than 30, and up to 45% of the resistance may be developed at the pile base. Furthermore, traditional design may also significantly underestimate the base component, as demonstrated by instrumented field tests in stiff boulder clay (Farrell et al., 1998), where a pile with a slenderness ratio of 24 developed a base load that was 61% of the total undrained resistance.
50 45 40 35 30 25 20 15 10 5 0
Qt = QS + Qb
Qs = 05su(nDL)
Qb = 9su(nD2/4)
Figure 2. Effect of slenderness ratio on the proportion of base to ultimate pile resistance
Using instrumented field model testing, this paper considers the mechanisms controlling the development of the base resistance of open-ended piles in clay. A simple empirical design approach is proposed for estimating the ultimate base resistance of pipe piles. This method is subsequently tested against a case history involving load tests on full-scale open-ended piles installed in the North Sea.
2. BACKGROUND
The American Petroleum Institute (API RP 2A) method (API, 1993), which is a derivation of the traditional total stress approach, is widely used for the design of offshore piles. The method considers two plug conditions, treating the pile as being either fully plugged (IFR — 0) or unplugged (IFR — 1). The plug condition is determined by assuming that qse and qsi are equal. The plug capacity qplug and normal plug stress qplug are given by
2a Qplug — q^ Di
q _ Qplug qplug nD2„
where Di is the internal diameter of the pile, qsi — qse — asu, c is an empirical reduction factor, and su is the average shear strength along the pile shaft. Recommended a values reduce from 1.0 for undrained strength ratios su/av0 < 0.25 to a minimum of 0.4 for realistically high strength ratios (Pelletier et al., 1993). The normal plug stress is compared with the average ultimate unit end bearing resistance qbu determined from the bearing capacity equation
qbu — Nc sub
exceeding the base resistance when the soil plug length exceeds 6Di. On this basis most field-scale piles will be assumed to be fully plugged.
Because of the uncertainties associated with determining a reliable su profile in some deposits, direct correlations with in situ tests, such as the cone penetration test (CPT) end resistance qc, have been suggested by several researchers. Bustamante and Gianeselli (1982) proposed the LCPC method: a direct relationship between qc, averaged 1.5D above and below the pile tip, and qb, given by
qb — kc qc
The cone reduction factor kc varies, depending on clay strength. No distinction is made between the end bearing resistance of open-ended or closed-ended piles.
Jardine et al. (2005) proposed an alternative CPT design method known as IC-05. Plugging during static loading is assumed to occur if
—— + 0-45-^-
DCPT patm
where DCPT is the diameter of the CPT cone (0.036 m) and patm is 100 kPa.
where Nc, the bearing capacity factor, is 9, and sub is the undrained strength at the pile base. If the plug stress calculated using Equation 2 exceeds the ultimate bearing capacity of the soil determined using Equation 3 the pile is deemed to be plugged, and the base load is calculated over the entire base area Ab, as
In contrast, if the base resistance exceeds the plug resistance, the pile is assumed to be fully coring and the base resistance is assumed to develop only over the annular area of the pile, Aann
q b — Nc sub Aa:
Recent assessment of the predictive capacity of the API approach suggests large scatter in the ratio of predicted to measured base capacity, with Nc factors for closed-ended piles measured as high as 17 for load tests by Farrell et al. (1998). Chow (1997) suggests that much of the scatter may result from uncertainty with respect to the design su value; however, overall conservatism is apparent throughout the database, with Nc — 9 forming an approximate lower bound to the closed ended data. There is also a dearth of information for open-ended piles, which may lead to further uncertainty when applying Nc values to pipe piles with various degrees of plugging. Consideration of the inequality between Equations 2 and 3 suggests that in soil deposits where the undrained strength profile is uniform with depth (s„b « su), and adopting a typical a value of 0.5, results in the plug resistance
The base load is given by Equation 8a for plugged piles and Equation 8b for unplugged piles, where kc — 0.45 or 0.65 for undrained or drained loading of plugged piles, and 1.0 or 1.6 for undrained or drained loading of unplugged piles respectively.
qb — kc qc Ab
qb — kc qc Aa
It is clear from the current literature that although the effects of plug development on the base resistance developed by pipe piles in clay are incorporated into some design methods, these tend to consider only the extreme case of fully plugged or unplugged piles. The techniques used to asses the likelihood of piles becoming fully plugged are highly empirical, and there is a dearth of measurements of plug response with which to calibrate these methods. With this in mind, a series of instrumented open-ended pile tests were performed to investigate the mechanisms of plug development and quantify the base stress changes that occurred during soil plugging.
3. SOIL CONDITIONS
The pile tests considered in this paper were conducted at the Kinnegar geotechnical research site on the shores of Belfast Lough. Soil deposits at the site comprise a layer of fill typically 1 m thick, consisting mainly of building rubble, over a natural soft sandy silt deposit of variable thickness (approximately 0.7-2 m thick), which in turn overlies 6 m of soft clayey silt. The soft silt deposits are locally termed sleech, and are the
— Nc s
strata of primary concern for the pile tests considered. The soft deposits contain sandy silt overlying clayey silt, and are referred to as the upper and lower sleech respectively. The sleech is underlain by a uniform medium dense sand. The site is tidal, with a water table fluctuating between 0.8 and 1.3 m below ground level.
The geotechnical properties of the sleech have been the subject of extensive laboratory and field investigations described by McCabe (2002), Lehane et al. (2003) and Gallagher (2006). The upper sandy sleech, located between 1 m and 3 m below ground level (bgl), has proportions of sand to clay of 20% and 10% respectively; below this level the proportions reverse, with a maximum clay fraction of 38%. The natural moisture content of the lower sleech is 60 ± 10%, with a liquid limit and plasticity index of 65 ± 10% and 35 ± 5% respectively. The material plots above the A-line in the Casagrande chart, identifying it as an intermediate- to high-plasticity clay. The low permeability (k — 1.5 X 10~10 to 5 X 10~10 m/s), vertical coefficients of consolidation Cv in the range 0.5-3 m2/year (at an effective stress of 100 kPa) and coefficient of consolidation in the horizontal direction, Ch, that ranged from 7 m2/year to 12 m2/year, are more typical of clay than of silt, and indicate the dominant nature of the clay size fraction in influencing the engineering behaviour. The sleech is lightly overconsolidated, with an overconsolidation ratio (OCR) decreasing with depth from 1.6 at 3 m to 1 at 8 m.
Cone penetration tests (CPT) and piezocone tests (CPTu) were performed at the site, with a typical trace shown in Figure 3(a). The CPT profiles were noticeably variable near ground level, but became much more uniform when the cone entered the lower sleech, where the qcnet value is seen to increase gradually from 175 kPa at -2.3 mbgl to 240 kPa at 6 mbgl. The undrained strength (su) profile measured in vane tests is compared in Figure 3(b) with values back-figured from the qcnet profile (assuming Nkt — 12.5), and those measured in triaxial compression tests.
4. INSTRUMENTATION AND EXPERIMENTAL PROCEDURE
Several instrumented pile tests have been performed by University College Dublin (UCD) researchers at the Kinnegar test site between 2003 and 2008. The piles were developed specifically to investigate the influence of penetration mode on pile capacity in clay, and therefore adopted two different end conditions: a closed-ended (CE) and an open-ended (OE) geometry. In total, seven pile installations were conducted, the locations of which are illustrated in Figure 4. The primary focus of this paper is on the base resistance developed by open-ended piles, but the results of some closed-ended pile tests are used for comparative reference.
The UCD open-ended model pile (UCD-OE) comprises a heavily instrumented section 2 m long, capable of penetrations up to a maximum depth of 6 m through the addition of extension pieces of equal external diameter. A twin-wall construction was adopted for the heavily instrumented lower section to allow the resistance developed by the pile annulus, shaft and plug to be determined independently of each other. This involved joining two steel tubes: a larger outer tube with an external diameter of 168 mm and wall thickness of 3.34 mm,
Figure 3. Belfast sleech: (a) a typical qc profile; (b) undrained shear strength
Kinnegar sewage treatment plant
111 f-'_.----,------,---:--T-------r--T---, -
Previous tests conducted by McCabe (2002)
aX—P|? O i ■!
■ÎT .UCD-CE-4
.^""DTIE y.
ippOrrj UCD-OE-1 UCD-OE-3
P__I _
SuppOrt platforms |
_ .. L. . i I I
UCD-CE-1
1 ■'■' 1 UCD "OF 2 ' UCD-CE-3 UCD OE 2
.1. I I I I,
UCD-CE-2
I . J -__J___I___1__1____I___I
the instrumentation layout, construction details and design requirements are provided by Gallagher (2006).
The UCD instrumented closed-ended stainless steel tubular model pile consists of an instrumented stainless steel tube of 73 mm external diameter and 1.7 m long, capable of penetrations up to a maximum depth of 7 m through the addition of extension pieces of equal external diameter. The 1.7 m lower section adopted an instrumentation layout that included total stress and pore pressure sensors at three separate locations along the pile shaft, at h/D of 1.5, 5.5 and 10, mimicking the OE pile. Base stress measurements were performed using a pressure sensor mounted in the pile tip plate.
Figure 4. UCD pile layout at the Kinnegar test site
and an inner tube with an external diameter of 154 mm and a wall thickness of 2 mm. The void space between the tubes that accommodated the instrumentation was 3.66 mm: thus the pile has a total wall thickness of 9 mm and a D/t ratio of 18.7. The instrumentation layout is shown schematically in Figure 5. Kyowa PS-5KA sub-miniature pressure transducers were used to measure the radial stress and pore pressure at multiple locations. They were positioned diametrically opposite each other at three levels along the pile shaft, at a distance h from the pile tip normalised by the pile diameter D of 1.5, 5.5 and 10.5. A pair of Kyowa PS-5KB total stress and pore pressure sensors were used in the annulus. TML brand PFL-10-11, 120 Q uniaxial foil strain gauges were used to measure the axial load distribution in each tube. Additional information on
168 mm
Three separate installations of the UCD OE pile (termed UCD-OE-1 to UCD-OE-3) were performed, and four installations of the UCD CE instrumented model pile were conducted (termed UCD-CE-1 to UCD-CE-4). A schematic of all the piles, indicating their final penetration depth and instrument locations, is shown in Figure 6.
The piles were installed using a hydraulic jack, from the base of pre-formed starter holes to their final depths in strokes of 100 mm length for UCD-OE-1 and UCD-OE-2; the jacking stroke length was increased to 250 mm for UCD-OE-3. The vertical reaction was provided by a specially fabricated steel frame anchored to a concrete kentledge founded on deep piles bearing in the underlying dense sand. The pile head accommodated a hollow load cell between the hydraulic jack and the driving cap, providing a continuous load profile during installation. A schematic illustration of the test set-up and load cell detail is presented in Figure 7. The load cell, strain gauges and pressure sensor data were logged at 0.1 s intervals throughout installation, using a System 5000 datalogger.
5. EXPERIMENTAL RESULTS
5.1. Plug development
The soil plug length (as reflected by the IFR value) measured during installation of the three pipe piles is shown in Figure 8. The plug development of all three piles was similar, reflecting the consistent installation procedure used. Paikowsky and Whitman (1990) conducted a comprehensive study of plug development, and noted that the plugging behaviour of driven piles was difficult to predict. However, it was observed that during quasi-static penetration, which occurs during free run (at the start of installation, when the pile penetrates under self-weight), piles became fully plugged within a penetration of 10-20 pile diameters in a range of soil types from soft to stiff clay. The UCD piles that were jacked into place remained fully coring (IFR = 1.0) until L/Di ratios in the range 7-10. Thereafter, the piles penetrated in a partially plugged mode until the normalised penetration exceeded 16-18, and they became fully plugged (IFR = 0). Pile UCD-OE-1, which had a final penetration depth z of 4.04 m, remained only partially plugged throughout installation, with a final IFR value of 0.44.
5.2. Shaft resistance
Although this paper primarily addresses the effects of plugging on the base resistance developed by pipe piles in clay, it is instructive to compare the external and internal shaft
Figure 5. Open-ended pile instrumentation layout
□ □
LU LU LU
□ □ □
Total stress and pore pressure sensors
150 m i
4 25 m I 1
|203 m
I 4 04 m
Total stress and pore pressure sensors
Total annular stress and pore pressure sensors
■83 m
Fill *
Sleech
Pre-formed starter holes
Figure 6. Relative pile lengths for UCD tests
resistance developed by pipe piles with that mobilised by closed-ended piles installed in the same deposit. The average shaft resistance qsav (— total shaft load/area of shaft) developed on the external shaft area of the open (qse) and closed-ended piles (qs), and the internal shaft (qsi) area of the open-ended pile, are compared in Figure 9. The initial variability over the first metre of penetration may in part be related to the site variability and the transition from the upper to the lower sleech. Despite the scatter, the data suggest that the closed-ended piles form an approximate upper bound to the shaft stresses, and qs reduces slightly with depth or with increasing pile length. The external shear stresses on the open-ended pile are slightly lower than those on the closed-ended piles for depths to 4 m bgl, although other factors, such as rate of installation and the occurrence of pause periods during installation, may also affect these results, as discussed by Gavin et al. (2010). The average internal shear stresses provide a lower bound to the data over the initial 2 m of penetration
(when the pile is unplugged). These observations (qsi < qse < qs) contrast with the general assumption of the API method (Equations 2 and 3), which suggests that the internal and external shear stresses mobilised by pipe piles are equal. For deeper pile penetrations (> 4 mbgl), when the pipe piles are fully plugged, the shear stresses developed on all piles are approximately equal (qsi « qse « qs).
5.3. Base resistance
The twin-walled construction of the pile meant that the total base load was easily discernible, as it was exclusively transferred to the inner pile wall. The distribution of load on
Figure 9. Average shaft stress comparison for all pile installations
Figure 11. Internal shear stress distribution for UCD-OE-3
Figure 10. Internal load distribution during installation of UCD-OE-3
the inner wall of the pile UCD-OE-3 during installation is shown in Figure 10. It is clear that the annular load developed at the pile toe (level one, L1) remained practically constant throughout installation. In contrast, the load developed by the pile plug (difference between L4 and L1) increased significantly as the soil core length increased and the IFR value reduced.
The distribution of internal shear stress during installation of UCD-OE-3 is shown in Figure 11. The tendency for high plug resistance to be concentrated near the toe of the pile is illustrated by the shear stress value at h/Di = 0.5, which developed to a maximum of ~35 kPa soon after the start of installation, and thereafter remained relatively constant at 20-30 kPa. The increase in internal plug resistance occurs as load was transferred to the upper plug (at h/Di > 2), with the qsi value at h/D — 0.5 remaining relatively independent of IFR. Lehane and Gavin (2001) compiled a database of similar measurements for pipe piles in sand, and noted that the
internal shear stresses near the pile tip were directly related to the degree of plugging, and the majority of load was shed over the bottom 3D. This suggests that the mechanism of plug development during undrained installation of pipe piles in clay is slightly different from that for the drained installation of piles in sand, allowing more load transfer in the upper regions of the plug. This is perhaps unsurprising, as the base resistance developed by pipe piles in sand is controlled by the very large effective stresses developed near the pile toe, and by arching of radial effective stresses.
Although quantification of the exact value and distribution of internal shear stress is an extremely complex problem, the normal plug stress qplug is given by the integral of the internal shear stress, and simple empirical correlations have been developed between qc, IFR and q,lug for piles in sand. Similar relationships are now considered for pipe piles in clay. The variation of qplug and qann during the installation of a typical pile (UCD-OE-2) is compared with the qb value for a closed-ended pile (UCD-CE-1) and CPT qc values in Figure 12. The plug stress is low during the initial stages of installation, and increases significantly as IFR reduces and the pile becomes fully plugged. In contrast, the qann profile is similar to those of qc and qb throughout installation, suggesting that it is independent of the degree of plugging experienced by the pile.
The data from all open-ended pile installations are considered in Figure 13, where the variation of the normalised maximum plug resistance (qplug/qc) and annular resistance (qann/qc) mobilised during each jacking stroke are plotted against the corresponding IFR value. As suggested previously by Figure 12, the annular stress is independent of the IFR, and can be approximated by the CPT qc resistance. In contrast, the data suggest a direct relationship between qplug, qc and the IFR of the form
9plug = 9c (0'8 - 0-6IFR) > Qpiugmin
Figure 12. A comparison of the stresses developed under the pile tip
6. CASE HISTORY
Clarke et al. (1985) describe the installation and static load testing of two pipe piles 762 mm in diameter and 18-3 m long, installed in the North Sea at the British Petroleum 'WC' block platform- One pile (pile A) had a driving shoe 457 mm long and 19 mm thick, which increased the annular thickness of the pile to 51 mm; a second pile (pile B) was driven without a shoe. The ground conditions at the site consisted of approximately 13 m of very stiff to hard glacial till over a layer of hard Lias clay, which extended to the base of the borehole (24-4 mbgl). Both the till and the Lias clay had occasional partings of fine sand and silt- The glacial till had a plastic limit of 15 and a plasticity index in the range 20-25. The natural moisture content reduced from 16% near the top of the deposit to 12% at depth. The su and qc profiles are shown in Figure 14. The Su and qc values were quite variable in the till, where Su increased with depth from 200 kPa to 700 kPa and qc ranged from 3 MPa to 12 MPa, with inferred Nt (= qc/su) factors in the range 15-25.
14 12 10 I 08 1 06 04 02 0
0ô«8 0o
- - ù4- - -
Range of qb/qc for closed ended piles
■ _ o_
A UCD-OE-1 O UCD-OE-2 □ UCD-OE-3
60 IFR: % (a)
□ □
10 O o 0
A UCD-OE-1 O UCD-OE-2 □ UCD-OE-3
□ 100
60 IFR: % (b)
Figure 13. (a) Annular stress and (b) plug stress dependence on IFR
qplugmin 0 '2 qc
9b = qc
Equation 9 implies that the plug resistance of a pipe pile increases by 400% as the pile moves from the fully coring to the fully plugged condition. The base resistance measurements in Figure 13 correspond to maximum values mobilised at relatively large pile head displacement (typically at s/D « 67%), but the base stress mobilisation curves during individual jacking strokes indicated full base mobilisation before s/D — 10% movement.
The Lias clay had a plastic limit of 27-28 and a plasticity index that increased from a value of 20 in the depth range
Figure 14. West Sole load tests: (a) cone resistance;
(b) undrained shear strength (adapted from Clarke et al., 1985)
13 - 18 m bgl to 25 below this depth. The natural moisture content was consistent with depth, at 17-18%. Although large scatter was evident in the measurements of su, values were typically > 400 kPa, with a maximum recorded value of 983 kPa. CPT qc values varied with depth from —8 MPa at the top of the deposit to —14 MPa at 24 mbgl.
Although the piles were uninstrumented, both static compression and tension load tests were performed on each pile at intervals of 3 m of pile penetration during installation (i.e. at pile penetration depths of 3 m, 6 m, 9 m, 12 m, 15 m and 18 m). The compression load tests were typically performed within 3 -12 h of the end of driving, and tension load tests typically followed within 3 h, resulting in the tests being essentially undrained. Delays in testing were noted for two tests: the first compression load test at 6 m bgl on pile A, which was performed 121 h after driving, and the first compression load test on pile B at 3 mbgl, which was performed 64 h after driving.
The IFR profiles, which were obtained by measuring the plug height prior to each static load test, are shown in Figure 15. These show that the presence of a driving shoe was effective in reducing plugging, with pile A exhibiting an IFR value in the range 0.8-1.0 throughout installation. In contrast, significant plugging occurred throughout the installation of pile B, with IFR values in the range 0.19-0.82. By measuring both the compression resistance Q C and tension resistance QT of the piles, Clarke et al. (1985) interpreted the unit base resistance qb as the difference between these measurements divided by the overall pile base area. By adopting the additional assumption that the annular resistance developed during the load tests was equal to the qc value averaged over the zone 1.5D above and below the pile base, the qplug values are inferred at each loading level. Although the accuracy of these assumptions is somewhat debatable, and specifically the inherent assumption that the shaft resistances in tension and compression are equal, and are unchanged by a previous load testing, the resulting qplug/qc values demonstrate a definite trend to increase in line with the IFR values, as illustrated in Figure 16. The West Sole data are seen to be a reasonable fit to Equation 9 developed from the UCD open-ended piles tests, despite the obvious difference of soil state at the two sites. The tendency for the
Figure 16. West Sole: plug stress dependence on IFR
normalised plug resistance measured on the full-scale field piles to increase with IFR gives confidence in the wider application of Equation 9.
7. CONCLUSIONS
Field tests that examined the effect of soil plugging on the base resistance developed during the installation of pipe piles in clay were presented. The following observations were made.
(a) The annular resistance of a pipe pile in clay was equal to both the closed-ended base stress and the CPT qc value, and was independent of the degree of plugging throughout installation.
(b) The plug capacity of a pipe pile in clay was found to be dependent on the plug formation, and when normalised by the CPT qc value acting at the pile tip could be related directly to the IFR value.
(c) The average internal shear stresses developed during installation formed a lower bound to the corresponding external shaft measurements at shallow penetrations, with both values tending toward a constant 5 kPa at depth. The distribution of shear stress within the pile plug suggests a mechanism for plug formation during undrained plugging in clay different from that for drained plugging in sand, as the internal shaft stress at the pile tip was found to be relatively independent of IFR.
(d ) A simple predictive equation based on the observations for Belfast soft clay was applied to a case history of offshore load tests with reasonable accuracy, implying a more widespread dependence of base stress on plug formation.
ACKNOWLEDGEMENTS
The authors wish to acknowledge the following people who contributed to the work: the technical staff at UCD (Frank Dillon and George Cosgrave), and Professor Barry Lehane, who developed the test site. In addition, David Igoe and Dave Cadogan are acknowledged for their contributions during the many weeks of field work. The first author is funded by a Sustainable Energy Ireland/IRCSET scholarship, and was a recipient of a Geotechnical Society of Ireland award. The third author's research at UCD was partly funded by an RPS-MCOS research scholarship and a postgraduate research award from the Geotechnical Society of Ireland.
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Figure 15. Incremental filling ratio back-figured for West Sole piles
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