Scholarly article on topic 'CPT-correlated design method of open-ended concrete pipe pile'

CPT-correlated design method of open-ended concrete pipe pile Academic research paper on "Civil engineering"

0
0
Share paper
OECD Field of science
Keywords
{""}

Academic research paper on topic "CPT-correlated design method of open-ended concrete pipe pile"

CPT-correlated design method of open-ended concrete pipe pile Liu, Zhang and Yu

ice | proceedings

Proceedings of the Institution of Civil Engineers

Geotechnical Engineering 169 June 2016 Issue GE3 Pages 240-249 http://dx.doi.org/10.1680/jgeen.15.00016 Paper 1500016

Received 21/01/2015 Accepted 31/07/2015

Published online 20/11/2015

Keywords: buildings, structures & design/field testing & monitoring/piles & piling

Published with permission by the ICE under the CC-BY license. http://creativecommons.org/licenses/by/4.0/

Institution of Civil Engineers

publishing

CPT-correlated design method of open-ended concrete pipe pile

J Junwei Liu Mingyi Zhang PhD

Lecturer, Qingdao University of Technology, Qingdao, Professor, Qingdao University of Technology, Qingdao, Shandong,

Shandong, People's Republic of China; Key Laboratory of People's Republic of China

Geotechnical and Underground Engineering (Tongji University), H Feng Yu PhD

Ministry of Education, Shanghai, People's Republic of China; Professor, School of Civil Engineering and Architecture,

Postdoctor, Zhejiang University, Hangzhou, Zhejiang, Zhejiang Sci-Tech University, Zhejiang, Hangzhou,

People's Republic of China People's Republic of China

An improved cone penetration test (CPT)-correlated design method is proposed for open-ended prestressed high-strength concrete pipe piles. The method takes the effects of compaction, plugging and friction fatigue into consideration. The evaluation proposal for shaft resistance introduces three parameters of stress reduction factor, compaction sensing coefficient and friction fatigue coefficient with their values addressed. The base resistance is divided into two parts of pile wall resistance and plug capacity, and is expressed in terms of the incremental filling ratio of the soil plug. The ratios of shaft and base resistance to CPT cone resistance, as well as the influence zone for end bearing, are determined. The rationality of this method is tested by comparing the calculated values with the results measured in field tests.

Notation qann pile annulus resistance

Ar area ratio of open-ended pipe pile ?b,b tip resistance of drilled pile

Arb,eff effective area ratio of open-ended pipe pile ?b,c tip resistance of close-ended pile

a stress reduction coefficient qc cone tip resistance

b compaction-sensitive coefficient qc,avg average value of CPT-qc

c friction fatigue coefficient qpk pile tip resistance

D external diameter of pipe pile qplg plug tip resistance

A internal diameter of pipe pile qplg,min plug tip resistance when IFR equals 1

Dr relative density degree of sand qsik pile shaft resistance

d pile diameter qt revised cone tip resistance

./s,avg average value of pile lateral friction qt,avg average value of cone penetration tip resistance

G shear modulus of soil Re equivalent pile diameter

H final plug length ß a dimensionless value (equal to qt,avg//s,avg)

h> vertical distance from pile tip to calculated S friction angle at pile-soil interface

point Ar radial displacement of pile-soil shear band

hb pile bearing stratum Äff' radial effective stress changes due to axial load

L pile embedment depth Xp plug effect coefficient

m ratio of qbc to qc ff' radial effective stress when pile-soil interface

n ratio of qplg,min to qc deteriorates

Pa reference pressure (=100 kPa) ff'c radial effective stress during static period after p

Q dimensionless parameter sinking

Qann pile total annulus resistance ffv vertical effective stress of soil

Qb,o pile total base resistance ffv0 initial vertical stress of soil

Qplg total plug capacity ffv0 initial effective vertical stress of soil

1. Introduction

Owing to high quality and low cost, prestressed high-strength concrete (PHC) pipe piles have been widely used as deep foundations in China and some other Asian countries. The open-ended form is frequently adopted owing to its favourable pile drivability. Unlike closed-ended piles, soil enters the pile and a soil plug is formed when open-ended piles are installed in the ground or on the seabed (riverbed). Thus, different bearing capacity characteristics can be observed between open-ended and closed-ended piles. The components of the bearing capacity of open-ended piles include the pile shaft resistance, base capacity provided by the pile annulus, and inner soil plug. Thus, the performance of open-ended piles becomes more complicated once the plugging action is considered.

Several approaches have been developed to estimate open-ended pile capacity in recent decades. Most of the fundamental methods employ empirical correlations (Randolph, 2011). Among these methods, the cone penetration test (CPT) is preferred in establishing relationships between cone resistance and pile capacity because of the similarity between the cone penet-rometer and a pile foundation. Several methods have been proposed to investigate the relationship between cone resistance and pile capacity, such as the Marine Technology Directorate Ltd (MTD) approach and the recently improved Imperial College pile (ICP) method developed by Jardine et al. (2005); the The University of Western Australia (UWA) design method of Lehane et al. (2005); Norwegian Geotechnical Institute (NGI)-05 design method by the Norway Geotechnical Engineering Society (2005); and the The University of Hong Kong (HKU) method developed by Yu and Yang (2012a, 2012b). These methods have improved the accuracy of the assessment of the capacity of open-ended steel pipe piles.

Nevertheless, these methods cannot be directly applied to assess the bearing capacity of open-ended concrete pipe piles because of the differences in pile dimensions and skin frictions. Moreover, most methods have not considered the soil plug effect, and thus the research on predicting the bearing capacity of open-ended concrete pipe piles is still at the initial stage. In this paper, a new bearing capacity prediction method is developed based on CPT-correlated design methods, and the CPT cone tip resistance is selected as the fundamental parameter.

2. Current design methods

2.1 ICP design method

In the ICP method (Jardine et al., 2005) equation, when one of the conditions in Equation 1 is satisfied, the pipe pile is assumed to be in a plugged state. Otherwise, the pile is considered to be in an unplugged state

where Di is the internal diameter of the pipe pile, Dr is the relative density degree of the sandy soil around the pile tip, qcavg is the average value of CPT-qc adjacent to the pile tip (MPa), and DCPT equals 0-036 m.

In a plugged state, the pile tip resistance is given by

qc,avg

0-5 - 0-25 log ( ——), 0-15, Ar dcpt j

where Ar is the area ratio of the tip-open pipe pile, which is calculated by

3 a-\-Dl

3 Ar - 1 D2

In an unplugged state, the pile tip resistance is

qc,avg

■ = Ar

The unit shaft resistance is obtained by Coulomb's law, which is given by

qs — ci tan 5 — (cré + Aci) tan 5

where 5 is the friction angle at the pile-soil interface. c is the radial effective stress (kPa) when the pile-soil interface breaks, composed of the radial effective stress, c4, during the static period after pile sinking and the radial effective stress changes, A ci, due to axial load. The radial effective stress and the corresponding change are calculated by Equations 6 and 7, respectively

6. — 0-029 ( —

qc \P:

7. Aci- —

4GAr D

where av is the vertical effective stress of the soil at some depth before pile sinking, Pa is the reference pressure (=100 kPa), R is the equivalent pile diameter according to the cross-sectional area of the pile wall, h0 is the vertical distance from the tip of the pile to the calculated point and Ar is the radial displacement (m) of the pile-soil shear band due to the axial load, which is related to the shear band thickness and the dilatancy of sandy soils. The typical value of Ar for the steel-sand interface equals 0-02 mm (Randolph, 2003). G is the shear modulus (kPa) of the soil surrounding the pile at the calculated depth, which is similarly expressed by the value of CPT-qc at the same depth

1. Di < 0-02(Dr - 30) or Di < 0-083

qc,avg\

8 G= 185(qc/Pa

qc (cv/Pa)-

2.2 UWA design method

Lehane et al. (2005) presented improved formulations for CPT-based axial capacity of typical (large-diameter) offshore piles in sand. In their formulations, the effective area ratio Arb>eff was incorporated into the UWA method to measure the degree of soil compaction. The design value of the pile tip resistance is

qc,avg

■ = 0-15 + 0-45Arb e

Similarly to the ICP and UWA methods, the unit pile shaft resistance in the HKU method is determined by Coulomb's law. Calculation of the effective radial stress in the HKU method is similar to that of the UWA method. However, the HKU method employs PLR as the parameter, which can be measured more easily in experiments, as shown in Equation 15. Furthermore, the effects of PLR are considered when calculating the radial stress change AoT caused by dila-tancy, as expressed by Equation 16.

15. — = 0-03

1 - \ PLR

ArVff = 1 - FFR

where FFR is the average value of the incremental filling ratio (IFR) of the soil plug at the ultimate 3D pile penetration during installation. The pile shaft resistance is still calculated according to Coulomb's law, given by

qs = «c + Aff ) tan ¿c

where the effective radial stress oTc caused by pile installation is expressed as Equation 12, and the effective stress change is calculated by Equation 7, the same as that used in the ICP method

12. — = 0-03

1 - ( d I ifr

max I ■

2.3 HKU design method

In this method (Yu and Yang, 2012a, 2012b), the effect of the pile length/diameter ratio on the base resistance provided by the open-ended pipe pile wall is considered. Equation 13 was proposed to calculate the pile annulus resistance based on a comparison of the current research results. The soil plug capacity is calculated by Equation 14.

qc;avg

■= 1-063

0-0451 — I > 0-46

qc;avg

= 1 -063 exp(-1-933 PLR)

where L is the embedment depth of the piles and PLR is the plug length ratio defined by H/L, where H is the final plug length.

16. Лот' =

^/(D2 - PLRd2)

2.4 Design method in Chinese technical code (JGJ94-2008)

In the Chinese technical code for building pile foundations (JGJ94-2008 (MCC, 2008)), the pile bearing capacity is divided in to two parts of shaft resistance and end capacity, as Equation 17

Quk = qsikli + qpk(Aj + XpAp1 )

18. Ap =

0,6 d @ < 5 0-8 ('±>5

where qsik and qpk are the unit shaft resistance and end resistance for different soils, respectively, the values of which can be obtained directly from the tables in code JGJ94-2008 based on the soil properties. Xp is a plug effect coefficient that can be calculated by Equation 18 according to the ratio of the embedment depth in the bearing stratum, hb, to the pile diameter, d.

3. New CPT-correlated design method of open-ended concrete pipe pile

3.1 Pile shaft resistance

White and Deeks (2007) and Karlsrud (2012) used earth pressure theory (with the p method as the main representative) and CPT (or standard penetration test) as the prevalent design methods for examining shaft resistance; they eventually considered CPT to be more suitable. Therefore, the cone penetration tip resistance is incorporated into the Coulomb

calculation in the proposed design method. A common formula suitable for all kinds of soil is given by

?s = (""re + A^r) tan S

gc.avg tan S

(Arf )b (ii)

+Aoi tan S

where S is the pile-soil interfacial friction angle (degrees); or is the radial effective stress (kPa) when the pile-soil interface deteriorates, consisting of the radial effective stress oTc during the static period after pile sinking and the radial effective stress changes, Aoi due to axial load. The pile shaft resistance is composed of two parts. The first part pertains to the frictional resistance caused by radial stress during pile installation, and the second part is the friction caused by the increase of radial effective stress due to axial loads. The second part is only available for sandy soil with dilatancy and is neglected for clay and silt.

The friction caused by radial stress, o^, is controlled by three parts, which correspond to the marks (i), (ii) and (iii), and represent undetermined parameters in Equation 19, as discussed below.

Part (i) is the maximum friction of close-ended or solid piles at a certain depth caused by pile installation, and is also the pile shaft resistance at a certain depth, in which piles pass without fatigue degradation. The radial stress surrounding the pile tip is smaller than the probe tip resistance, qc. The radial stress is determined by the soil distortion feature, particularly unloading rigidity (White and Deeks, 2007), which is implied by the parameter a (namely, stress reduction coefficient) in Equation 19. The rigidity of sandy soils is generally higher than that of both silt and clay. Thus, the stress reduction coefficient of sandy soils is larger than that of silt and clay. The value of a for vertical axially loaded piles is always 33, and is determined by Equation 20 proposed by Schneider and White (2007) for other piles.

20. -i- « minf2 Q + 5,78 tan S \ 3

where dimensionless parameter Q = (qt - ov0)/ovo, in which qt is the revised cone tip resistance; and ov0 (0y0) is the initial (effective) vertical stress for different types of soils. The values of Q for clay, silt/sand and gravel are 5-15, 50 and 100-200, respectively. Correspondingly, the dimensionless values of a/tan S are 8-15, 43 and 71-78, respectively.

ratio, a/tan S, is about 1/2-1/3 of p, and the typical value of p is determined by the LCPC method proposed by Bustamante and Gianeeselli (1982). Note that the Laboratoire Central des Ponts et Chaussees (LCPC) method is regarded as the most common CPT method. Therefore, according to the proposed values of p in the LCPC method, the ranges of a/tan S for steel piles in cohesive soil, silt and sandy soil are 10-40, 40-60 and 40-100, respectively. In the UWA and HKU methods, the value of a/tan S is approximately 60 with a = 33 and S = 29°, in the range of the value of p from 10 to 100. However, Bustamante and Gianseslli (1982) and Niazi and Mayne (2013) suggested that the value of p for concrete piles should be smaller than that of steel piles. Thus, the values of a/tan S for concrete piles in cohesive soil, silt and sandy soil are 10-30, 20-40 and 30-75, respectively.

Part (ii) indicates the influence of the plugging effects on the lateral friction of open-ended pipe piles. If more soil is squeezed into a pile - that is, a long soil plug - the resulting soil compaction density and the lateral normal stress and friction will be smaller. The effects of the soil plug on the soil compaction are represented by the effective area Areff

21. ArVff = 1 - IFR d2

where Di and D are the internal and external diameters of a concrete pipe pile, respectively. If the incremental fill ratio (IFR) of the soil plug is difficult to determine, the PLR can be employed as a substitute for IFR. The PLR can either be measured on site or evaluated by the expression below by Carter et al. (1979).

22. PLR

1-18 - 0-18 ln - > 0 (drivenpile)

1-83 - 0-16 ln - > 0 (jackedpile)

In Equation 19, parameter b reflects the sensitivity of pile shaft resistance to soil compaction, which is called the compaction-sensitive coefficient in this paper. A large b indicates a greater influence of the plugging effect on shaft resistance. When b is zero, regardless of the variation in IFR, this item has no further effect. Using the core extension theory, White et al. (2005) conducted a theoretical study on the compaction-sensitive coefficient b of clay and sandy soils, and suggested b = 0-1 and 0-3 in sandy soil and clay, respectively, which was exactly the case with the UWA and HKU methods. As the properties of silt fall between clay and sandy soil, b = 0-2 was suggested in silt.

In many CPT methods, the empirical parameter p equals the ratio of the average value qt avg of cone penetration tip resistance to the average value /s>avg of pile lateral friction, that is, qt,avg//s,avg=p. According to Schneider and White (2007), the

Part (iii) reflects the influence of the shaft fatigue effect, which has been considered in the ICP, UWA and HKU methods. However, the difference in the three methods is primarily reflected by the friction fatigue coefficient, c. A small value of

0-38 is suggested in the ICP method according to the measured data of jacked piles in sandy soils. A value of 0-50, which is suitable for driven piles in sandy soils, is proposed in the UWA and HKU methods. Lehane et al. (2000) reported that 0-2 is appropriate for the friction degradation coefficient of driven piles in clay. The friction degradation of jacked pre-stressed piles is weaker than that of hammered piles, because jacked piles experience fewer pile installation stroke cycles. Accordingly, the fatigue coefficient of jacked piles in clay is smaller than 0-2, and is proposed to be 0-15 in this study. The value c for silt is suggested to be 0-35 (driven piles) and 0-25 (jacked piles), respectively, which are between the range of the values for clay and sandy soils.

Finally, the proposed values of the parameters a/tan 5, b and c are summarised in Table 1.

3.2 Pile base resistance

The base resistance (2b,o) of open-ended pipe piles is composed of the pile wall resistance (Qann) and the plug capacity (Qplg), which can be expressed as

Aannqan

The base resistance (qb,c) of close-ended piles is almost proportional to the cone tip resistance (qb,c = mqc), which has been previously determined in many experiments, such as Lehane and Gavin (2001). The unit pile annulus resistance qann is close to the unit base resistance of closed-ended or solid piles qb,c because of the similarity in the load-bearing mechanism. Thus, qann is estimated as qb,c, which is the same assumption as in the UWA-05 design method.

The plug tip resistance qplg is closely related to the value of IFR. When an open-ended pile is in a fully plugged state (IFR = 0), its bearing capacity properties are similar to those

Types of soils a/tan ä b c Jacked Driven

Cohesive soil Very soft; soft 15 0-1 0-15 0-2

Firm; stiff 20

Hard 30

Silt Loose 20 0-2 0-25 0-35

Medium dense 30

Dense 40

Sandy soil Loose 40 0-3 0-38 0-5

and gravel Medium dense 60

Dense 70

Table 1. Proposed values of parameters a/tan ä, b and c

of closed-ended piles and the plug capacity reaches the maximum. When IFR increases, qplg approximately linearly decreases to qplg,min (Ghionna et al., 1993; Xu et al., 2008, Doherty et al., 2010). Lehane et al. (2005) assumed that the plug tip resistance qplg,min is identical to the tip resistance of drilled piles qb,b when IFR equals 1. Thus, the proportional relationship (i.e. the coefficient is n) between qplgimin and qc could also be established. The relationship among the tip resistance of pile walls qann, the plug tip resistance, qplg, and the average value of cone penetration tip resistance can be expressed as

g, max (qplg,max 1

g,max — mqc,avg

b,b — nqc,avg

Therefore, a universal expression of the unit base resistance of open-ended piles in various types of soils can then be obtained

>,c — «ic,avg + (m -

- «K,

Ghionna et al. (1993) suggested that the base resistance of drilled piles was found within the range of (0-15-0-23) CPT-qc. In the UWA design method, the suggested value of n is 0-15 for sandy soil. Alsamman (1995) proposed the following expression for calculating the bearing capacity of drilled piles.

Non-cohesive soil

0-15qc 0-05qc + 10 < 30

(qc < 100tsf) (qc > 100tsf)

Cohesive soil

29. qb,c — 0-25(qc - 0V0) < 25

According to the calculation expression for drilled piles proposed by Fleming (1992), the unit tip resistance is 15 to 20% of the qc value when the tip deflection is 0-1D (D is pile diameter). The value of n ranges from 0-15 to 0-25, and the value for cohesive soil is slightly larger than that for sandy soil and silt. Thus, 0-2, 0-15 and 0-15 are the proposed values of parameter n for cohesive soil, sandy soil and silt, respectively.

Lehane and Gavin (2001) and Chow (1997) found that the value of m ranges from 0-6 to 1-0. A relatively conservative value of 0-6 is adopted in the UWA design method. Based on the field test (Liu et al., 2012), the value of m is proposed to be 0-6 and 0-8 for sandy soil and cohesive soil, respectively. (See Table 2 for proposed values of m and «.)

3.3 Influence zone of pile end bearing

In most design methods, the influence zone above the pile tip is not less than that below the pile tip, which is consistent with the results of the theoretical analysis of rigid plasticity shear failure. In this method, two situations of local embedment (embedment depth less than 8D) and total embedment (embedment depth not less than 8D) of the pile toe are considered, and different values are used for the influence zone. This design idea is consistent with the distribution law of the practical stress field of pile tip soils. The proposed values of the influence zone of end bearing in this method are listed in Table 3.

4. Case study

4.1 Case 1

The field test was conducted in Fuyang, Zhejiang Province, China. A series of geological explorations and laboratory tests was performed. The geological profile of the test site and CPT-qc curves are shown in Figure 1. PHC pipe piles that were 13 m long were used in that test programme. The piles were open-ended with an outer diameter of 400 mm and a wall thickness of 75 mm. The piles were jacked using a jacking rig with a capacity of 9000 kN. The test pile was instrumented with strain gauges along the shafts that facilitated the measurements of pile-load transfer subjected to axial loading. The static load test followed a sequence of multistage loading in which the applied load at a stage was maintained until the pile-head settlement was not larger than the target value.

Types of soils m n

Cohesive soil 08 0-2

Silt 0-6 0-15

Sandy soil 0-6 0-15

Table 2. Proposed values of m and n

Influence zone Partial embedment Total embedment

Above pile toe 4D 2D

Below pile toe 1D 4D

Table 3. Proposed values of influence zone

The calculated distribution of pile shaft resistance using the proposed method is shown in Figure 2. The calculated results reflect the influence of the soil property on the shaft resistance. In addition, the corrected shaft resistance when considering the degradation effect fits well with the distribution rule that the

Cone tip resistance: MPa

Figure 1. Change of CPT-qc with depth

Unit shaft resistance: kPa

Figure 2. Calculated unit shaft resistance using the proposed method

upper part is small and the lower part is large, even for one soil layer with similar properties.

A comparison between the calculated and measured average unit shaft resistance in different soil layers is shown in Figure 3. The calculated values are close to the measured ones. The difference between the measured and suggested values in the Chinese technical code JGJ94-2008 (MCC, 2008) is possibly due to the neglect of both the friction fatigue and plugging effects in JGJ94-2008. The depth effect may have caused the divergence for lower depth. When values are determined according to the empirical parameter method, the depth effect of the soil layers is disregarded. Hence, the shaft resistance at different depths in soils with identical physical standards is assumed to be the same, which contradicts reality and the commonly accepted p design method. However, these factors are included in the proposed comprehensive design method.

The measured values of total bearing capacity, pile shaft resistance, and pile base capacity are 800 kN, 661 kN, and 139 kN, respectively, as shown in Figure 4. The calculated values obtained using the proposed method are 831 kN, 705 kN and 126 kN respectively, and the corresponding deviations are +3-88%, +6-66%, and -9-35%, indicating high accuracy.

4.2 Case 2

Another field test was carried out at a construction site in Hangzhou, China to examine the applicability of this method. Figure 5 shows the soil profiles and CPT traces; the soil consisted of clay with silt. The groundwater table was 4-3 m. An open-ended PHC pile (26 m long, 500 mm outer diameter and 100 mm wall thickness) was employed in this test. Eighteen

Unit shaft resistance: kPa

0' 2 4

groups of strain gauges were instrumented on the level between two neighbouring layers to measure the accurate value of average unit shaft resistance for each soil layer. The final length of the soil plug at the end of installation is 3-23 m, thus the value of PLR is 0-124.

The calculated distribution of pile shaft resistance and average unit shaft resistance for each soil layer using the proposed method are shown in Figure 6 and Figure 7, respectively. With respect to the pile capacity (Figure 8), the predictions by this method provide the best agreement with the field measurements.

7 661 661 05

Y////X Measured value

gigig Calculated value by proposed method ÏÏ1 Suggested value by JGJ94-2008

139 126 v,—

Total capacity: kN Shaft resistance: kN Base capacity: kN

Figure 4. Comparison between calculated and measured pile capacity

Cone tip resistance: MPa

Silty clay Clayey silt

Figure 3. Comparison between calculated and measured average unit shaft resistance of each soil layer

0 3 6 9

E 12 h:

S 15 18 21 24 27

Figure 5. Change of CPT-qc with depth

4.3 Case 3

The third field test was also carried out at Hangzhou, China. Figure 9 shows the soil profiles and CPT traces. From the ground surface down to the depth of 14 m, the profile consisted of a stratigraphic sequence of fill, silty clay, sandy silt, clayey silt and sandy silt. The groundwater table was 3-2 m. The open-ended PHC pipe piles used in this test was 13 m long, 400 mm in outer diameter and had 75 mm wall thickness.

The calculated distribution of pile shaft resistance using the proposed method is shown in Figure 10, and their averaged

values of different soil layers and pile capacity are shown in Figure 11 and Figure 12. It is clear that the calculated values are close to the measured ones, indicating the rationality and the high accuracy of this proposed design method.

5. Conclusions

In this paper, a new bearing capacity design method based on cone penetration tests that are suitable for open-ended concrete pipe piles is proposed. The method is derived from comparisons and analysis of common specifications and design methods. In the proposed method, CPT-qc was taken as the parameter; the influence of compaction, plugging and friction

o 16 -

Unit shaft resistance: kPa 40 60 80 100

2 2400

1 2000 ΠTO

™ 1600 c

ro 1200 CD _Q

27482774

Y///A Measured value

| ¡11111 Calculated value by proposed method ÏÏ1 Suggested value by JGJ94-2008

240 234 254

Total capacity: kN Shaft resistance: kN Base capacity: kN

Figure 8. Comparison between calculated and measured pile capacity

Figure 6. Calculated unit shaft resistance using the proposed method

0 5 10 15

Unit shaft resistance: kPa 60 80

Figure 7. Comparison between calculated and measured average unit shaft resistance of each soil layer

Cone tip resistance: MPa

2 3 4 5

Silty clay Sandy silt Clayey silt

Sandy silt

Figure 9. Change of CPT-qc with depth

Unit shaft resistance: kPa 10 20 30 40 50 60 70

aci 600

929 923razn892

Y///A Measured value

j1111111 Calculated value by ""lm proposed method 11111111 Suggested value by JGJ94-2008

230 220 220

Total capacity: kN Shaft resistance: kN Base capacity: kN

Figure 12. Comparison between calculated and measured pile capacity

Figure 10. Calculated unit shaft resistance using the proposed method

Unit shaft resistance: kPa 20 40 60 80 100

2 4 6 8 10 12 14

Calculated value by proposed method

Suggested value by JGJ94-2008

A<v&*6o

Figure 11. Comparison between calculated and measured average unit shaft resistance of each soil layer

fatigue effects on the bearing capacity were thoroughly considered. Design parameters for pile shaft resistance and pile base bearing capacity that are suitable for various soils were proposed. Finally, the comparisons with the field test results confirmed the rationality and the accuracy of the proposed design method.

Acknowledgements

The authors gratefully acknowledge the support of Specialized Research Fund for the Doctoral Program of Higher Education

(grant no. 20133721120004), Promotive Research Fund for Young and Middle-Aged Scientists of Shandong Province (grant no. BS2013SF004) and A Project of Shandong Province Higher Educational Science and Technology Program (grant no. J14LG04), China.

REFERENCES

Alsamman OM (1995) The Use of CPTfor Calculating Axial Capacity of Drilled Shafts. PhD thesis, University of Illinois, Urbana-Chapaign, IL, USA.

Bustamante M and Gianeeselli L (1982) Pile bearing capacity prediction by means of static penetrometer CPT. In Proceedings of the 2nd European Symposium on Penetration Testing (Verruijt A, Beringen FL and de Leeuw EH (eds)). Balkema, Amsterdam, the Netherlands, pp. 493-500.

Carter JP, Randolph MF and Wroth CP (1979) Some aspects of the performance of open- and closed-ended piles. In Proceedings of the International Conference on Numerical Methods in Offshore Piling. Thomas Telford Limited, London, UK, pp. 165-170.

Chow FC (1997) Investigations into the Behaviour of

Displacement Piles for Offshore Foundations. PhD thesis, Imperial College, London, UK.

Doherty P, Gavin K and Gallagher D (2010) Field investigation of the undrained base resistance of pipe piles in clay. Proceedings of the Institution of Civil Engineers -Geotechnical Engineering 163(1): 13-22, http://dx.doi.org/ 10.1680/geng.2010.163.1.13.

Fleming WGK (1992) A new method for single pile settlement prediction and analysis. Geotechnique 42(3): 411-425.

Ghionna VN, Jamiolkowski M, Lancellotta R et al. (1993) Base capacity of bored piles in sands from in situ tests. In Deep Foundations on Bored and Auger Piles - Bap II: Proceedings of the 2nd International Geotechnical Seminar (Impe WV (ed.)). AA Balkema, Rotterdam, the Netherlands, pp. 67-75.

Jardine R, Chow F, Overy R et al. (2005) /CP Design Methods /or Driven Piles in Sands and Clays. Thomas Telford Publishing, London, UK.

Karlsrud K (2012) Prediction o/Load-Displacement Behavior and Capacity o/Axially-Loaded Piles in Clay Based on Analyses and Interpretation o/Pile Load Test Results. PhD thesis, Norwegian University of Science and Technology, Trondheim, Norway.

Lehane BM and Gavin K (2001) Base resistance of jacked pipe piles in sand. Journal o/ Geotechnical and Geoenvironmental Engineering 127(6): 473-480.

Lehane BM, Chow FC, McCabe BA and Jardine RJ (2000) Relationships between shaft capacity of driven piles and CPT end resistance. Proceedings o/ the Institution o/ Civil Engineers - Geotechnical Engineering 143(2): 93-101, http://dx.doi.org/10.1680/geng.2000.143.2.93.

Lehane BM, Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand. In Frontiers in O//shore Geotechnics: Proceedings o/ the International Symposium on Frontiers in O//shore Geotechnics (/S-FOG2005) (Gourvenec S and Cassidy M (eds)). Taylor & Francis Group, London, UK, pp. 683-689.

Liu JW, Zhang ZZ and Yu F (2012) Case history of installing instrumented jacked open-ended piles. Journal o/ Geotechnical and Geoenvironmental Engineering 12(7): 810-820.

MCC (Ministry of Construction of the People's Republic of China)

(2008) Technical Code /or Building Pile Foundation Chinese technical code(JGJ94-2008). China Building Industry Press, Beijing, People's Republic of China.

Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single

piles. Journal o/ Geotechnical and Geoenvironmental Engineering 31(4): 979-1009.

Randolph MF (2003) Science and empiricism in pile foundation design. Geotechnique 53(10): 847-875.

Schneider JA and White DJ (2007) Back analysis of Tokyo port bay bridge pipe pile load tests using piezocone data. In Advances in Deep Foundations: International Workshop on Recent Advances o/Deep Foundations (/WDPF07) (Morikawa Y, Kikuchi Y, Otani J and Kimura M (eds)). Taylor & Francis Group, London, UK, pp. 183-194.

White DJ and Deeks AD (2007) Recent research into the behaviour of jacked foundation piles. In Advances in Deep Foundations: International Workshop on Recent Advances o/Deep Foundations (/WDPF07) (Morikawa Y, Kikuchi Y, Otani J and Kimura M (eds)). Taylor & Francis Group, London, UK, pp. 3-26.

White DJ, Schneider JA and Lehane BM (2005) The influence of effective area ratio on shaft friction of dis-placement piles in sand. In Frontiers in O//shore Geotechnics: Proceedings o/ the International Symposium on Frontiers in O//shore Geotechnics (/S-FOG2005) (S Gourvenec and M Cassidy (eds)). Taylor & Francis Group, London, UK, pp. 741-747.

Xu X, Schneider JA and Lehane BM (2008) Cone penetration test (CPT) methods for end-bearing assessment of open-and closed-ended driven piles in siliceous sand. Canadian Geotechnical Journal 45(8): 1130-1141.

Yu F and Yang J (2012a) Improved evaluation of interface friction on steel pipe pile in sand. Journal o/Per/ormance o/ Constructed Facilities 26(2): 170-179.

Yu F and Yang J (2012b) Base capacity of open-ended steel pipe piles in sand. Journal o/ Geotechnical and Geoenvironmental Engineering 138(9): 1116-1128.

WHAT DO YOU THINK?

To discuss this paper, please email up to 500 words to the editor at journals@ice.org.uk. Your contribution will be forwarded to the author(s) for a reply and, if considered appropriate by the editorial panel, will be published as discussion in a future issue of the journal.

Proceedings journals rely entirely on contributions sent in by civil engineering professionals, academics and students. Papers should be 2000-5000 words long (briefing papers should be 1000-2000 words long), with adequate illustrations and references. You can submit your paper online via www.icevirtuallibrary.com/content/journals, where you will also find detailed author guidelines.