KSCE Journal of Civil Engineering (2013) 17(5):1008-1022 DOI 10.1007/s12205-013-0262-8
Geotechnical Engineering
www.springer.com/12205
Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity Maria Cecilia M. Marcos*, Yit-Jin Chen**, and Fred H. Kulhawy*** Received May 27, 2012/Accepted October 8, 2012
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Abstract An extensive evaluation of capacity interpretation criteria is presented for driven Precast Concrete (PC) piles under axial compression loading. A database of static load tests for round and square cross-section piles under both drained and undrained loading was developed and utilized in the analysis. The compiled data are generalized into four groups (round - drained, square drained, round - undrained, and square - undrained) to examine the differences in their behavior. Generally, similar trends for each criterion are noted for the four groups. Statistical analyses show a decreasing COV with increasing pile displacement, and the drained load tests show less variability than the undrained load tests. For drained loading, somewhat higher interpreted capacities are exhibited by round piles than by square piles. In addition, a slightly stiffer normalized load-displacement curve is shown for the drained loading compared to the undrained loading. The drained and undrained databases are further sub-divided by hammer type (drop, air/steam, diesel, and hydraulic) to explore the effects of driving energy and driving resistances on pile capacity. Based on these analyses, the relative merits and interrelationships of these criteria are established, and design recommendations for the use of these methods are suggested in terms of normalized capacity and displacement. Keywords: driven, PC piles, interpretation criteria, static load test, database ··································································································································································································································
1. Introduction Driven Precast Concrete (PC) piles typically are designed and constructed based on analytical models, empirical rules, or dynamic formulas. However, the stability and safety of structures supported by driven piles commonly are verified through pile load tests. While dynamic and statnamic pile tests are considered modern practices in piling, static load testing has been the traditional standard for evaluating the pile capacity under field loading conditions. More than one static load test often is needed to ensure the foundation reliability in larger piling projects. Although the pile behavior is substantially verified by a static load test, geotechnical issues, construction factors, and inevitable uncertainties are encountered, leading to significant variation of pile behavior to applied loads. As a result, relatively different shapes of the load-displacement curves will occur, and often the maximum resistance of the pile is not clearly defined. Therefore, the capacity needs to be defined by an “interpreted failure load”. Numerous interpretation criteria (van der Veen, 1953; Terzaghi and Peck, 1967; Chin, 1970; DeBeer, 1970; Fuller and Hoy, 1970; Davisson, 1972; O'Rourke and Kulhawy, 1985; Hirany
and Kulhawy, 1988, 1989, 2002) have been recommended for interpreting the capacity from load test results. However, as indicated in Table 1, different failure definitions are used by different methods, which lead to different recommendations for design. Research on this issue has been conducted (Fellenius, 1975; Duzceer and Saglamer, 2002; Fellenius, 2006) using limited numbers of field load tests on various driven pile types, while Hirany and Kulhawy (1988, 1989, 2002), Chen et al., (2008), and Chen and Fang (2009) focused on interpretation criteria for drilled shaft foundations. Drilled shafts and driven piles differ in construction procedure, and often in geometry, which can affect their relative behavior. Therefore, it is worthwhile to examine these criteria over a wider database. In this study, the representative interpretation methods listed in Table 1 are examined in detail to assess their relative merits and interrelationships. A broad database of axial compression load tests on driven PC piles is used for this purpose. Both drained and undrained loading conditions for round and square crosssection piles are examined. These piles were installed using drop, air/steam, diesel, and hydraulic hammers, so the influence of hammer type also was examined. The interpreted results are compared statistically and graphically, and recommendations are made for consistent use in geotechnical practice.
*Ph.D. Student, Dept. of Civil Engineering, Chung Yuan Christian University, Chung-Li 32023, Taiwan; Instructor, Dept. of Civil Engineering, Adamson University, Manila, Philippines (E-mail:
[email protected]) **Professor, Dept. of Civil Engineering, Chung Yuan Christian University, Chung-Li 32023, Taiwan (Corresponding Author, Email:
[email protected]) ***Professor Emeritus, School of Civil and Environmental Engineering, Cornell University, Ithaca, NY 14853-3501, USA (E-mail:
[email protected]) − 1008 −
Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
2. Database of Load Tests Static pile load test results on driven PC piles were collected primarily from the geotechnical literature and available load test reports and were compiled for this study. The resulting database includes 72 sites with 152 field compression load tests and covers a range of soil profiles, pile shapes, and construction methods. The soil profile is categorized herein as drained or undrained, based on the predominant soil condition along the pile depth. In general, when the soil along the pile depth is comprised of mainly cohesionless layers, the soil profile is categorized as “drained loading condition”. When the soil along the pile depth is comprised of mainly cohesive layers, the soil profile is categorized as “undrained loading condition.” The piles were round or square in cross-section. Using the two profile types and two cross-sections, the piles were grouped into four categories. For drained compression with round piles (DCR), there are 10 sites with 37 tests; for drained compression with square piles (DCS), there are 24 sites with 45 tests; for undrained compression with round piles (UCR), there are 18 sites with 36 tests; and for undrained compression with square piles (UCS), there are 20 sites with 34 tests. For cases with driving records, the load tests were sub-divided further based on hammer types. For the drained cases, there were 13, 11, 43, and 4 tests using drop, air/steam, diesel, and hydraulic hammers, respectively, while for the undrained cases, there were 7, 1, 20, and 28 tests using drop, air/steam, diesel, and hydraulic hammers, respectively. All of the load tests have almost complete geological data and load-displacement curves, and all were conducted on straightsided, driven PC piles. Based on the case history descriptions, the pile construction and test performance appear to be of high quality. Therefore, these load tests should be representative of common field situations. The basic information, the basic pile driving data records, and the interpreted capacities for the DCR, DCS, UCR, and UCS tests are summarized in Tables 2 to 5, respectively, while the reference sources are listed in Table 6. These load tests were
conducted throughout the world at different points in time and in various soil types. For convenience, the ranges of foundation geometry, compression capacity, and their Coefficients of Variation (COV), which is the Standard Deviation (SD) divided by the mean value, are summarized in Table 7. Obviously, the range of geometry is broad and the pile dimensions and capacities for drained load tests are relatively comparable. However, based on the available data for undrained load tests, the round section piles have somewhat larger diameter and capacity ranges.
3. Interpretation of Load Tests The eight interpretation criteria shown in Table 1 were used to evaluate the interpreted failure load or capacity Q from the loaddisplacement curve of each pile case. These criteria were selected because they represent a wide distribution of interpreted results from the lower, middle, and higher bounds as found in practice. The definitions of interpreted failure load or capacity by van
Fig. 1 Regions of Load-Displacement Curve
Table 1. Definition of Representative Compression Interpretation Criteria for Driven Piles Method
Category
Definition of interpreted capacity, Q
van der Veen (1953)
Mathematical model
QVDV is Pult that gives a straight line when log (1-P/Pult) is plotted versus total displacement.
Chin (1970)
Mathematical model
QCHIN is the inverse slope (1/m) of a line s/p = ms+c, where p = load and s = total displacement.
Fuller and Hoy (1970)
Settlement limit
QF&H is the minimum load that occurs at a rate of total displacement of 0.05 in. per ton (0.14 mm/kN).
Terzaghi and Peck (1967)
Settlement limit
QT&P is the load that occurs at 1.0 in. (25.4 mm) total displacement.
DeBeer (1970)
Settlement limit
QDB is the load at the change in slope on a log-log load-displacement curve.
Davisson (1972)
Graphical construction
Slope tangent Graphical construction (O’Rourke and Kulhawy 1985) L1 - L2 (Hirany and Kulhawy, Graphical construction 1988, 1989, 2002)
Vol. 17, No. 5 / July 2013
QDAV occurs at a displacement equal to the pile elastic compression line (PD/AE) offset by 0.15 in. (3.8 mm) + B (in. or mm)/120, where P = load, D = depth, A = area, E = Young’s modulus, B = pile dia. QST occurs at a displacement equal to the initial slope of the load-displacement curve offset by 0.15 in. (3.8 mm) + B (in. or mm)/120. QL1 and QL2 correspond to elastic limit and failure threshold loads, respectively, as shown in Fig. 1.
− 1009 −
Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
der Veen (1953) and Chin (1970) are based on mathematical models that correspond to the asymptote of the load-displacement curve. The Fuller and Hoy (1970) and Terzaghi and Peck (1967) interpreted capacities are defined as the load at a rate of settlement and an absolute displacement, respectively, while
DeBeer (1970), which is also based on a settlement limit, determines capacity at a load occurring in the change of slope from the log-log plot of the load-displacement curve. The Davisson (1972) method is a graphical construction and defines capacity at the intersection of the load-displacement curve and
Table 2. Basic Information and Interpreted Results for Drained Compression Round Section (DCR) Tests Site & Pile No. DCR1 DCR2-1 DCR2-2 DCR2-3 DCR2-4 DCR2-5 DCR2-6 DCR2-7 DCR3 DCR4-1 DCR4-2 DCR4-3 DCR4-4 DCR4-5 DCR4-6 DCR4-7 DCR4-8 DCR5-1 DCR5-2 DCR5-3 DCR5-4 DCR5-5 DCR5-6 DCR5-7 DCR5-8 DCR5-9 DCR6-1 DCR6-2 DCR6-3 DCR7 DCR8 DCR9
Test Site/ Soil Description Arkansas; fine and silty sand
Dramen, Norway; medium to coarse sand
Spain; fine silty sand
GWTa Drb (m) (%) 1.0
1.7
0
67
Pile depth/ dia.(m) 13.7/0.41
Interpreted capacity, Qh (kN)
Hammer Final typec/ setg energy (bl/25 (kN-m) mm) B / 49
2 f
9 8 10 8 8 9 22 65 41 52 52 52 52 52 52 52 28 32 32 32 38 38 38 66 66 55 55 55
8.0/0.28 16.0/0.28 7.5/0.28 11.5/0.28 15.5/0.28 19.5/0.28 23.5/0.28 18.0/0.91 38.0/0.60 27.0/0.50 27.0/0.60 27.0/0.60 25.0/0.60 25.0/0.60 22.0/0.50 30.0/0.50 16.0/0.40 20.0/0.40 30.0/0.60 27.0/0.60 22.0/0.30 22.0/0.40 22.0/0.50 34.0/0.60 34.0/0.60 25.0/0.60 23.0/0.50 23.0/0.50
A / 7.5 A / 7.5 A / 7.5 A / 7.5 A / 7.5 A / 7.5 A / 7.5 -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d C / -- d
2 4f 2 2 4 4 8 -- d 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f 2f
QL1
QDB
QST
QDAV
QT&P
QL2
667
1334
1646
1731 >1780e 1732
106 226 208 213 249 252 203 391 365 387 445 445 59 153 160 162 >202e 189 166 216 245 245 296 212 178 330 360 366 418 412 267 445 488 505 565 566 445 578 632 682 745 773 2530 2900 3325 3180 4540 4260 2669 >4893e >4893e >4893e >4893e >4893e 1032 2322 2499 2917 2925 2920 2251 3125 3800 >5871e >5871e >5871e 3292 >4448e >4448e >4448e >4448e >4448e 2251 4270 4520 4645 4680 4804 2046 4272 4600 5015 4982 5159 1779 2002 2265 2310 2390 2265 2295 2295 4100 4367 4300 4220 213 427 445 425 445 445 320 600 638 638 641 641 641 1921 2030 2100 2150 2135 639 1281 1440 1500 1558 1494 641 1922 1770 2130 2110 2108 854 2135 2190 2255 2275 2242 1601 2882 3050 3130 3180 3160 1700 4448 4680 5200 5000 5275 2200 3114 3800 4050 3995 4040 2100 4450 4600 4950 4920 4893 1780 4448 4410 4580 4580 4448 1200 3558 3470 3600 3642 3642
QF&H
QVDV
QCHIN
1734
1802
2064
196 391 151 230 369 512 721 4950 >4893e 3096 >5871e >4448e 5100 5200 2446 4385 440 641 2135 1494 2108 2242 3202 5337 4080 5338 4620 4003
257 461 184 282 442 566 825 6800 6900 3096 6200 5500 5120 5650 2580 4900 459 670 2350 1691 2300 2349 3247 5400 4320 5338 4893 4011
270 476 216 313 464 642 910 10765 9607 3421 8807 9314 6107 6146 2656 5218 494 705 2553 1837 2788 2633 3660 6378 5059 6158 6129 4665
Miliao, Taiwan; silty sand
1.9
Kaohsiung, Taiwan; clayey silty sand
-- d
Changbin, Taiwan; silty sand
-- d
N. Carolina; sandy/clayey silts NRd; loose silt and sand College Station, Texas; sand
-- d
-- d
16.5/0.31
B / 44f
5
1033
1674
1847
1870
1875
1851
1875
1860
2318
-- d
-- d
40.0/0.30
-- d
-- d
712
1957
1223
2100
1623
2064
2135
2430
3087
7.5
-- d
10.4/0.91
-- d
-- d
1800
2500
2545
2530
2775
2875
2875
3020
3117
DCR10-1 46 27.0/0.60 C / 287f 3f 2000 3020 3805 4550 4610 4783 5300 5200 6254 f DCR10-2 30 10.0/0.50 C / 287 3f 490 800 805 580 820 900 840 903 1153 Chiayi, Taiwan; DCR10-3 1.0 30 30.0/0.50 C / 287f 3f 2390 3150 3390 4830 3750 5780 5800 7000 9189 silty sand DCR10-4 30 20.0/0.50 C / 287f 3f 1365 2500 2715 1610 2430 2780 2800 2950 3566 f DCR10-5 30 10.0/0.50 C / 287 3f 200 500 610 415 825 900 820 923 1212 a GWT = groundwater table; bDr = relative density; if not reported, it is inferred from standard penetration test N value (Terzaghi and Peck, 1967) c A = drop hammer; B = air/steam hammer; C = diesel hammer; D = hydraulic hammer d -- not reported; eload test was terminated before interpreted load; fvalue is deduced based on available information g final set = number of blows of pile hammer for final 25 mm of driving h QDB = DeBeer, QST = slope tangent, QDAV = Davisson, QT&P = Terzaghi and Peck, QF&H = Fuller and Hoy, QVDV = van der Veen, QCHIN = Chin − 1010 −
KSCE Journal of Civil Engineering
Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
Table 3. Basic Information and Interpreted Results for Drained Compression Square Section (DCS) Tests Test Site/ Soil Description
GWTa (m)
Drb (%)
DCS1
Calhoun, Florida; fine to coarse sand
2.3
60
DCS2
Fittja, Sweden; fine to medium sand
-- d
25
Hammer Final Interpreted capacity, Qh (kN) Pile c g / Type set depth/sec. energy (bl/25 QL1 QDB QST QDAV QT&P QL2 QF&H QVDV Q CHIN (m) (kN-m) mm) 18.8/0.762 D / 120 11 3500 3500 5920 6640 7130 7291 7000 7385 8813 12.8/0.235 A / 8 -- d 200 250 260 256 310 300 262 326 406
DCS3
Porto, Portugal; silty and clayey sand below tip
42
6.0/0.350
A / 40f
-- d
580
1300 1350 1345
1536 1440 1518
1606
65
12.1/0.300
A / 30
3
656
1156 1148 1166 >1227e 1223 1219 1400
1694
65
9.3/0.300
A / 30
3
534
623
897 >1023e 1012 1001 1150
1425
1.5
75
8.7/0.300 A / 37.5
10
1672 1957 1957 1875
2295
2242 2375 2315
3020
-- d
45
27.0/0.355 B / 33.15
6
2167 3000 3375 3760
3625
3606 3800 3830
4799
1.52
40
15.2/0.406
C / 48
32
2260 2372 2402 2375
2553
2669 2580 2555
3270
2.13
33
21.0/0.360
B / 44
1
623
1112 1459 1776
2171
2224 2224 1950
2819
0
25
21.0/0.360
B / 44
3
771
1836 1868 1937 >1959e 1898 1957 2500
2960
0
20
26.0/0.360
B / 85
1
1008 2668 3114 4130
4119
4131 4181 2800
4762
43
11.0/0.285 C / 33.64f
4
350
900
1000
1000
1000 1000 1025
1295
44
15.0/0.285 C / 33.64f
4
464
1000 1080 1516
1550
1500 1500 1625
2040
Site & Pile No.
DCS4-1
Salmiya, Kuwait; calc. silty sand
2.0
DCS4-2
DCS6
Shuwaikh, Kuwait; calc. gravelly sand Florida, USA; silty /clayey fine sand
DCS7
Georgia; fine to medium coarse sand
DCS5
DCS8-1 DCS8-2
Tidewater, Virginia; clayey silty sand
DCS8-3 DCS9-1
943
960
1437
Baghdad University Complex, Iraq; uniform sand with some silt
5.7
DCS9-2 DCS10
Shandong, China; sandy loam
-- d
22
25.0/0.450
A / -- d
6
1000 2500 2400 2741
2820
2885 2845 2915
3692
d
58
10.0/0.300
C / 65f
16
888
1650 2000 1988
2405
2660 2660 2750
3418
58
12.0/0.300
C / 65f
10
560
1500 1350 1305
1656
1867 1810 1905
2272
76
38.0/0.510 C / 104.5
38
2776 5012 5380 7100
5385
7295 7450 7620
9928
DCS13-1
80
11.4/0.305 C / 62.4
6
445
1477 1241 1415
1607
1603 1612 1660
1936
DCS13-2
80
11.2/0.305 C / 39.3
50
737
1334 1607 1539
1786
1786 1786 1780
2099
DCS13-3 Toronto, Southern Ontario-site A; DCS13-4 sand, silt and clay with some gravel
79
8.5/0.305
C / 39.3
12
377
1000
1051
1250
1236 1241 1271
1395
79
8.4/0.305
C / 62.4
4
377
1000 1071 1203
1473
1489 1482 1515
1681
DCS13-5
81
12.5/0.305 C / 39.3
48
1285 1334 1786 1780 >1786e 1786 1786 2000
2369
DCS13-6
81
15.1/0.305 C / 62.4
7
773
1334 1451 1528
1585
1604 1567 1604
1750
DCS14-1 Toronto, Southern Ontario-site B; DCS14-2 firm clayey silt and dense silty sand
0
68
34.8/0.305 C / 62.4
12
893
2224 1719 2486
2098
2815 2790 2700
3486
50
16.5/0.305 C / 62.4
12
829
2002 1897 2410
2478
2853 3013 3010
3426
DCS15-1
0
35
13.0/0.254 B / 20.3
5
247
445
505
640
515
610
865
DCS15-2
1.0
55
10.0/0.356 C / 35.6
49
890
1556 1646 1586
1653
1646 1664 1666
2547
DCS15-3 Lower Atlantic Coastal Plain; loose to dense sand, clayey fine sands and DCS15-4 some soft silt and clayey layers DCS15-5
2.3
46
16.5/0.356 B / 39.31
11
502
1112 1512 1566
1750
1656 1823 1900
2264
0
35
19.4/0.356 B / 26.4
13
614
1601 1824 2201 >2277e 2144 2277 2300
2947
0
32
19.2/0.457 B / 59.3
9
899
1557 2037 2045 >2140e 2011 2140 2200
2737
0
15
15.5/0.457 B / 43.4
12
288
890
1281 1168
1735
1246 1743 1780
2215
DCS11-1 Information Center of Beijing; DCS11-2 sandy loam and silty sand Waldport, Oregon, USA; sand and silt DCS12
DCS15-6
--
3.0
7.93
982
534
658
DCS16
Netherland; coarse sand and gravel
1.0
20
9.3/0.250
C / -- d
2
305
450
460
440
600
600
600
510
753
DCS17
Toronto; clayey silt and silty sand
9.8
41
14.7/0.305 C / 62.4
6
606
780
826
838
973
947
951
1020
1259
-- d
35
46.8/0.400 Df / -- d
7
800
3000 2050 3200
3110
3071 3210 3500
3694
39
48.5/0.400 Df / -- d
31
1425 3200 2020 4568
3620
3900 4410 4700
5547
-- d
13.4/0.600
-- d
-- d
2400 3000 >3400e 3400 >3400e >3400e >3400e 7500
8835
-- d
10.5/0.600
-- d
-- d
2205 3200 3650 3420
3900
3990 4100 4160
4806
-- d
13.4/0.600
-- d
-- d
1600 2600 2770 2715
3110
3300 3700 3690
3993
-- d
-- d
8.6/0.300
-- d
-- d
460
863
850
8.1
44
9.4/0.508 D / 85.4
-- d
70
30.6/0.380
-- d
-- d
70
30.3/0.380
--
d
d
70
30.7/0.380
-- d
-- d
21.5/0.270
-- d
DCS18-1
Kelantan, Malaysia; clayey silty sand
DCS18-2 -- d
DCS19-1
Pensacola, Florida; DCS19-2 loose to dense sand DCS19-3 NRd; sand DCS20
850
950
1660 4005 4213 4213 >4213e 4118 4175 4360
6312
1600 2400 3600 4380
3582 >4380e >4380e 6700
8185
1690 3200 2900 4360
3973 >4600e >4600e 4820
7155
-- d
1896 2400 3970 4300
3810 >4300e >4300e 6800
9339
-- d
500
958
1030
1145
Hampton, Virginia; silty fine sand 1.2 41 18.0/0.610 C / 108.6 4 1200 3000 3060 3082 >3100e 3060 3082 4200 DCS24 a b GWT = groundwater table; Dr = relative density; if not reported, it is inferred from standard penetration test N value (Terzaghi and Peck, 1967) c A = drop hammer; B = air/steam hammer; C = diesel hammer; D = hydraulic hammer d -- not reported; eload test was terminated before interpreted load; fvalue is deduced based on available information g final set = number of blows of pile hammer for final 25 mm of driving h QDB = DeBeer, QST = slope tangent, QDAV = Davisson, QT&P = Terzaghi and Peck, QF&H = Fuller and Hoy, QVDV = van der Veen, QCHIN = Chin
4399
DCS21
Sussex County, Virginia; clayey sand
DCS22-1
Rotterdam Harbour, Netherlands; DCS22-2 clayey and silty sand DCS22-3 Fitja, Stockholm; loose silty sand DCS23
Vol. 17, No. 5 / July 2013
-- d
− 1011 −
4 --
595
750
720
830
719
885
1036
843
936
Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
the pile elastic compression line offset by 3.8 mm + B/120. The slope tangent method by O’Rourke and Kulhawy (1985) proposed use of the initial slope instead of the elastic line because of overconservatism at smaller depth/diameter (D/B) ratios. The L1-L2 method (Hirany and Kulhawy 1988, 1989, 2002) also is a graphical construction. This method employs the fact that the load-displacement curve generally can be simplified into three distinct regions: initial linear, curve transition, and final linear, as illustrated in Fig. 1. Point L1 (elastic limit) corresponds to the load (QL1) and butt displacement (ρL1) at the upper end of the initial linear region, while L2 (failure threshold) corresponds to the load (QL2) and butt displacement (ρL2) at the initiation of
the final linear region. QL2 is defined as the “interpreted failure load” or “interpreted capacity” because, beyond QL2, a small increase in load gives a significant increase in displacement. The slope tangent and L1 - L2 methods were developed using drilled foundations and are adopted herein to assess their suitability for driven PC piles. The interpretation results are given in Tables 2 to 5 for the DCR, DCS, UCR, and UCS tests, respectively. However, some load tests were terminated before reaching the interpreted values. The interpreted results for these cases are denoted as greater than (>) the terminated load, except for van der Veen and Chin because of their definitions. Extrapolation of the load-displacement
Table 4. Basic Information and Interpreted Results for Undrained Compression Round Section (UCR) Tests Hammer Final Interpreted capacity, Qh (kN) Pile b a c g GWT s / Type set u depth/dia. Test Site/ Soil Description (m) (kN/m2) energy (bl/25 QL1 QDB QST QDAV QT&P QL2 QF&H (m) (kN-m) mm) UCR1 East Boston; marine clay -- d 124 45.5/0.41 C / 98.5 12 1382 2250 >2777e >2777e >2777e >2777e >2777e UCR2 Downtown, Boston; marine clay -- d 53 41.8/0.31 C / 98.5 4 750 >1500e >1500e >1500e >1500e >1500e >1500e d d d UCR3 Philippines; silty clay & fine sand -41 57.0/0.41 C / --712 >1157e >1157e >1157e >1157e >1157e >1157e UCR4-1 184 21.7/0.30 C / 35.7 8 893 1786 2000 2092 2030 >2130e >2130e San Juan, Puerto Rico; silty clay UCR4-2 3.0 209 19.8/0.30 C / 35.7 12 446 1079 642 650 1080 1205 1339 with sand UCR4-3 261 22.9/0.30 C / 35.7 11 446 1700 1500 1545 1580 >2139e >2139e d d d UCR5 Brazil; silty clay --40.0/0.42 D / 40.5 -750 >2000e 1900 >2000e >2000e >2000e >2000e UCR6-1 -- d 90 13.0/0.46 -- d -- d 800 1400 1850 1830 1999 1830 2040 lake Charles, LA; sandy silty clay UCR6-2 155 17.7/0.46 -- d -- d 2100 2400 2425 2390 2320 2545 2545 UCR7 College Station, Texas; very stiff clay 6.0 -- d 31.0/0.32 -- d -- d 1750 2905 2880 3355 3250 3350 3690 d d d UCR8 Louisiana; tan, gray silty clay --15.9/0.76 --- d 3600 6000 8610 >8850e >8850e >8850e >8850e UCR9 Brazil; clayey silt 18.0 35 14.0/0.18 D / 8.1 2 140 250 248 257 262 259 258 UCR10-1 China; clayey silt 1.0 -- d 7.2/0.40 C / 130 23 2405 3000 >4060e >4060e >4060e >4060e >4060e UCR10-2 China; muck clay 1.0 -- d 38.2/0.40 C / 124 22 2400 3000 4990 >5100e 4882 >5100e >5100e UCR10-3 China; sandy clay 1.0 -- d 24.8/0.40 C / 100 22 3540 3980 >5225e >5225e >5225e >5225e >5225e UCR10-4 1.0 -- d 19.6/0.30 C / 100 17 1300 2995 3000 >3000e >3000e >3000e >3000e d d UCR11 Canada; silty clay/clayey silt 2.1 62 36.0/0.36 --995 1500 2500 >2500e >2500e >2500e >2500e d d UCR12 Chesapeake, U.S.; clay 0 67 8.2/1.37 --1410 3114 3895 3845 4030 3850 4092 UCR13-1 1.5 24 35.5/0.25 Df / -- d 1f 240 500 557 635 637 611 608 UCR13-2 1.5 13 14.5/0.25 Df / -- d 1f 161 200 >268e >268e >268e >268e 268 Malaysia; marine clay UCR13-3 1.5 34 23.5/0.25 Df / -- d 1f 162 230 >423e >423e >423e >423e >423e UCR13-4 1.5 16 11.5/0.25 Df / -- d 1f 115 213 217 216 221 218 210 UCR14-1 Malaysia; clayey silt and silty clay -- d 99 14.5/0.30 D / 41.2 13 480 1250 1148 1150 1395 1600 1600 UCR14-2 Malaysia; sandy clay -- d 73 14.2/0.35 D / 55 11 800 1190 1725 1450 1872 1940 >2200e d UCR14-3 Malaysia; silty and sandy clay -128 18.7/0.25 D / 41.2 9 382 510 641 620 953 1010 1000 UCR15-1 Perak, Malaysia; clayey silt -- d 203 4.5/0.30 D / 27.5 125 1025 1530 >1530e >1530e >1530e >1530e >1530e UCR15-2 115 28.5/0.40 D / 61.8 14 900 1600 2105 2120 2108 2000 >2200e UCR16-1 0 134 43.0/1.00 D / 122.8 10 5395 7400 8800 9000 8050 9000 9000 UCR16-2 Malaysia; soft marine clay 144 57.5/1.00 D / 145 10 2820 8000 5220 8190 6995 8560 9143 UCR16-3 90 33.8/1.00 D / 137.9 6 2500 6500 6615 6520 6850 7000 7000 UCR17 Jimah, Malaysia; clayey and silty clay 0 52 38.9/0.60 D / 129.6 21 3000 3000 6000 6230 5800 6350 6638 UCR18-1 Bangkok; soft to stiff clay with sand -- d -- d 28.0/0.60 D / 78.5 2 3000 4546 4610 4900 4793 5209 5500 UCR18-2 -- d 29.0/0.80 D / 92 2f 3300 7995 8636 9210 8640 10090 10600 UCR18-3 -- d 38.0/0.80 D / 85.8 2 2446 6727 7450 8800 8156 8509 >9150e UCR18-4 -- d 24.6/0.80 D / 85.8 2f 2400 5900 6878 7260 7874 6900 >8182e UCR18-5 -- d 29.4/0.80 D / 92 3 4210 10073 9600 10600 9580 11300 11900 a GWT = groundwater table; bsu = undrained shear strength; if not reported, it is inferred from standard penetration test N value (Terzaghi and Peck, 1967) c A = drop hammer; B = air/steam hammer; C = diesel hammer; D = hydraulic hammer d not reported; eload test was terminated before interpreted load; fvalue is deduced based on available information g final set = number of blows of pile hammer for final 25 mm of driving h QDB = DeBeer, QST = slope tangent, QDAV = Davisson, QT&P = Terzaghi and Peck, QF&H = Fuller and Hoy, QVDV = van der Veen, QCHIN = Chin Site & Pile No.
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QVDV QCHIN 5750 2350 2400 2700 1500 2700 3100 2010 2565 3690 9350 260 4350 9500 8500 6500 3300 4060 647 280 450 226 1600 2310 1290 2500 2270 9001 9144 7120 7450 5386 10500 9150 7972 12150
6461 3186 3596 3683 1940 4008 4530 2469 3305 4144 12903 275 5853 13080 10184 8167 5444 4439 752 397 652 256 2141 3323 1426 3906 3399 11328 10481 8012 11364 6279 12505 11404 10571 14531
KSCE Journal of Civil Engineering
Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
Table 5. Basic Information and Interpreted Results for Undrained Compression Square Section (UCS) Tests Hammer Final Interpreted capacity, Qh (kN) Pile Typec/ setg Test Site/ Soil Description depth/sec. energy (bl/25 QL1 QDB QST QDAV QT&P QL2 QF&H (m) (kN-m) mm) UCS1-1 19 26.0/0.280 C / 75 15 1643 >1643e >1643e >1643e >1643e >1643e >1643e Singapore; silty clay, marine clay -- d UCS1-2 18 30.0/0.260 C / 75 17 1000 1600 1750 >2110e >2110e >2110e >2110e UCS2-1 203 18.2/0.300 D / 33.6 10 890 1427 1700 >1779e >1779e >1779e >1779e UCS2-2 203 17.4/0.350 D / 52.5 10 1557 2400 >2491e >2491e >2491e >2491e >2491e Malaysia; silt and clay (stiff, very UCS2-3 stiff and hard cohesive soils) -- d 203 18.6/0.350 D / 52.5 10 1334 1779 >1779e >1779e >1779e >1779e >1779e UCS2-4 203 14.7/0.300 D / 33.6 10 1489 2002 >2495e >2495e >2495e >2495e >2495e UCS2-5 203 16.8/0.250 D / 21.7 10 617 >1245e >1245e >1245e >1245e >1245e >1245e d d UCS3 Mexico; clayey soil 2.0 34 15.0/0.300 --291 450 525 335 590 538 540 UCS4-1 58 9.0/0.400 C / 45f -- d 419 500 520 502 620 676 600 Tianjin, China; soft clayey soil 2.0 UCS4-2 80 21.0/0.400 C / 45f -- d 1240 3200 2630 3215 3218 3190 3544 UCS5-1 143 25.1/0.280 C / 69.85f 25 1241 2224 2340 2402 2375 2387 2375 Singapore; sandy clay and silt -- d UCS5-2 169 25.6/0.325 A / 70f 13 1449 2669 2927 3300 3158 3228 3300 UCS6 Northern Ireland; clayey silt 1.4 24 6.0/0.250 D / 22.1 -- d 19 55 58 56 >64e 63 45 f e e e UCS7-1 Richmond, California; very 119 21.0/0.355 C / 100 6 1250 >1489 >1489 >1489 >1489e >1489e >1489e 2.0 UCS7-2 stiff lean to fat clay 103 15.0/0.355 C / 100 1 f 800 1373 >1600e >1600e >1600e >1600e >1600e UCS8 Mexico City; soft clay 2.0 34 15.0/0.300 -- d -- d 200 344 490 329 550 492 518 UCS9 Argentina; clay and silty sand -- d 44 19.0/0.400 C / -- d -- d 1510 >4719e >4719e >4719e >4719e >4719e >4719e UCS10 Elba Is., Georgia; soft/sandy clay 1.5 89.5 22.9/0.457 C / 113 -- d 2000 3750 3680 3480 3750 3750 3750 UCS11-1 534 712 828 787 857 857 857 below 154 8.5/0.305 A / 13.5 20 Barnet, London; London clay tip UCS11-2 154 4.4/0.305 A / 13.5 4 249 311 385 370 450 390 390 UCS12-1 London; London clay -- d 13.1/0.400 -- d -- d 817 1139 1190 1185 >1201e 1201 1201 UCS12-2 London; soft alluvium -- d -- d 12.8/0.360 -- d -- d 667 1334 1040 1040 1334 1557 1556 UCS12-3 -- d 12.2/0.360 -- d -- d 600 667 812 667 951 1245 963 d e UCS13-1 -39.0/0.350 D / 44 1 1500 3750 3456 >4000 3405 3806 >4000e Melbourne, Australia; silty clay -- d d e e UCS13-2 -42.0/0.275 D / 35.3 1 1500 1564 2670 >3083 >3083 >3083e >3083e d d UCS14 Louisiana; deltaic clays and silts -154 25.0/0.356 B / -8 670 1785 1750 1718 >1786e 1786 1786 d f d UCS15-1 Depok, Indonesia; silt-clay -39 5.5/0.250 A / 15 -120 200 308 300 402 300 300 UCS15-2 residual soil 37 11.5/0.250 A / 15f -- d 200 500 510 513 563 500 500 UCS16 India; soft to stiff clay -- d 163 20.0/0.400 -- d -- d 560 1423 1580 1600 1657 1655 1601 UCS17 Illinois, Carbondale; silty clay 1.52 73 6.1/0.305 -- d -- d 335 >893e >893e 893 >893e >893e >893e UCS18 Kakinada, India; soft to stiff clay -- d 79 19.5/0.400 -- d -- d 1038 1334 1420 840 1400 1446 >1557d UCS19 Penang, Malaysia; med to stiff clay -- d 79 60.3/0.406 C / 81.5f 4 2180 3558 >4448e >4448e >4448e >4448e >4448e d UCS20-1 Sint-Katelijne-Waver, Belgium; 82 7.4/0.350 A / 40 -250 750 960 955 >972e 972 972 -- d d boom clay UCS20-2 113 11.6/0.350 A / 40 -750 1150 1630 1622 1648 1648 1648 a GWT = groundwater table; bsu = undrained shear strength; if not reported, it is inferred from standard penetration test N value (Terzaghi and Peck, 1967) c A = drop hammer; B = air/steam hammer; C = diesel hammer; D = hydraulic hammer d not reported; eload test was terminated before interpreted load; fvalue is deduced based on available information g final set = number of blows of pile hammer for final 25 mm of driving h QDB = DeBeer, QST = slope tangent, QDAV = Davisson, QT&P = Terzaghi and Peck, QF&H = Fuller and Hoy, QVDV = van der Veen, QCHIN = Chin Site & Pile No.
GWTa (m)
sub (kN/ m2)
curve is not done herein. All interpreted results greater than the terminated load were eliminated from further evaluation to prevent bias.
4. Evaluation Results for Drained Load Tests The statistics for the interpreted capacities and the corresponding displacements for drained compression tests are summarized in Tables 8 and 9, respectively. The tables include summaries for the drained-round section (DCR), drained-square section (DCS), and all drained data combined. All of the interpreted capacities are normalized by the Chin interpreted capacity. The Chin method always projects the highest asymptote value, well above more conventional interpreted capacities, and could interpret all Vol. 17, No. 5 / July 2013
QVDV QCHIN 3985 3000 2150 4800 3397 2495 1970 600 690 3401 2410 3400 65 3750 1700 550 4000 4000 845 410 1210 1670 1200 4000 4400 2100 360 578 1663 1100 1685 6000 1010 1750
6009 4191 2731 5976 3988 3643 2269 945 724 3966 2836 5084 74 4529 2560 835 6028 5483 1207 519 1597 1953 1507 5846 6009 2334 447 618 1867 2593 2246 9072 1217 2428
the load test cases. Therefore, it was adopted as a base for comparing the interpretation criteria. Results are compared below to evaluate the interrelationships and characteristics of the methods. The statistical data are also presented to illustrate the quality of the data set results and the comparisons. The QVDV for some cases are beyond the maximum applied load. The displacements for these cases were not extrapolated resulting to fewer “n” values for δVDV (Tables 9 and 11). Also, note that QL1 is included for reference only. It is not an interpreted failure load or capacity; it is the elastic limit. The results in Table 8 show mean load ratios ranging from 0.67 to 0.85 for round piles and 0.58 to 0.81 for square piles when compared to QCHIN, with COV values of 11 to 22% (round) and 9 to 22% (square). The general trends of mean ratios for both
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Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
Table 6. Reference Sources of Pile Load Tests in Tables 2 to 5 Pile no.
Reference source Mansur, C.I., and Hunter, A.H., “Pile test-Arkansas River project,” Journal of Soil Mechanics and Foundation Division, ASCE, DCR1 Vol. 96, No. SM5, pp. 1545-1582 (1970). Gregersen, O.S., Aas, G., and DiBiagio, E., “Load tests on friction piles in loose sand,” Proceeding of the 8th International ConferDCR2 ence on Soil Mechanics and Foundation Engineering., Vol. 2, pp. 109-117 (1973). Mey, R., Oteo, C. S., Sanches del Rio, J., and Soriano, A., “Field testing on large driven piles,” Proceeding of the 11th International DCR3 Conference on Soil Mechanics and Foundation Engineering., San Francisco, Vol. 3 pp. 1559-1564 (1985). DCR4-1 Diagnostic Engineering Consultant, Limited. “Load test report of Formosa Petrochemical Corporation Plant,” Taiwan, (1998) DCR4-2 Diagnostic Engineering Consultant, Limited. “Load test report of HI building,” Mailiao, Taiwan, (1998) Chen, C. H., Perng, T. D., Hwang, J. H., and Chang, L. T. “Analyses for load test data of PC piles on West-Coast reclaimed areas of DCR5, DCR6 Taiwan,” Journal of the Chinese Institute of Civil and Hydraulic Engineering, Vol. 12, No. 1, pp. 51-62 (2000). DCR7 Abe, S., Likins, G., and Morgano, C. M. “Three case studies on static and dynamic testing of piles,” Geotechnical News, pp. 26-28 (1990). Fellenius, B., “Test loading of piles and new proof testing procedure, “Journal of the Geotechnical Engineering Division, ASCE, DCR8 Vol. 101, No. GT9, pp. 855-869 (1975) DCR9, DCS19, McVay, M., Kuo, C. L., and Guisinger, A. L. “Calibrating resistance factors for load and resistance factor design for statnamic load UCR6-UCR8 testing,” Project No. 4910450482312, Florida Department of Transportation, Tallahassee, Florida (2003). Hsu, S. T., Lin, S. Y., and Hung, C. J. “Analyses for axially loaded behavior of pre-stressed concrete piles,” Journal of the Chinese DCR10 Inst. of Civil and Hydraulic Engineering, Vol. 19, No. 1, pp. 13-23 (2007). Hussein, M.H., Sharp, M.R., and Knight, W.F., “The use of superposition for evaluating pile capacity,” Deep Foundations 2002, DCS1 Geotechnical Special Publications No. 116, Vol. 1, pp. 6-21 (2002). Fellenius, B.H., “Determining the true distribution of load in instrumented piles,” ASCE International Deep Foundation Congress, DCS2 Geotechnical Special Publications No. 116, Vol. 2, pp. 1455-1470 (2002). Fellenius, B. H., Santos, J. A., and Fonseca, A. V., “Analysis of piles in a residual soil-The ISC'2 prediction,” Canadian GeotechniDCS3 cal Journal, Vol. 44, pp 201-220 (2007). Ismael, N.F., “Analysis of load test on piles driven through calcareous desert sands,” Journal of Geotechnical and GeoenvironmenDCS4, DCS5 tal Engineering., ASCE, Vol. 125, No. 10, pp. 905-908 (1999). Preim, M.J., March, R., and Hussein, M., “Bearing capacity of piles in soils with time dependent characteristics,” Proceeding of the DCS6 3rd International Conference on Piling and Deep Foundations, London, England, pp. 363-370 (1989). Vesic, A.S., “Tests on instrumented piles, Ogeechee river site,” Journal of the Soil Mechanics and Foundations Division., ASCE, DCS7 Vol. 96, No. SM2, pp. 561-582 (1970). Martin, R.E., Seli, J.J., Powell, G.W., and Bertoulin, M., “Concrete pile design in Tidewater Virginia,” Journal of Geotechnical DCS8 Engineering, ASCE, Vol. 113, No.6, pp. 568-585 (1987). Altaee, A., Fellenius, B. H., and Evgin, E., “Axial load transfer for piles in sand: I. Tests on an instrumented precast pile,” CanaDCS9 dian Geotechnical Journal, Vol. 29, pp. 11-20 (1992). Li, D. Z., Liu, X. L., and Chen, F., “Pile bearing capacity, dynamic test, dumping factor,” Proceeding of the 12th International ConDCS10, DCS11 ference on Soil Mechanics and Foundation Engineering., Rio de Janeiro, Brazil, pp. 9-12 (1989). Riker, R. E., and Fellenius, B. H., “A comparison of static and dynamic pile test results,” Proceeding of the 4th International ConDCS12 ference on the Application of Stress-Wave Theory to Piles, Balkema, pp. 143-152 (1992). Thompson, C. D., and Thompson, D. E., “Effects of pile driving systems on drivability and capacity of concrete piles,” Proceeding DCS13, DCS14 of the Symposium on Deep Foundations, ASCE, Atlanta, Georgia, pp. 420-443 (1979). Roos, C. J., and Lingo, E. W., “Behavior of driven piles on soft limestone,” Proceeding of the Symposium on Deep Foundations, DCS15 ASCE, Atlanta, Georgia, pp. 365-395 (1979). Viergever, M. A., “Relation between cone penetration and static loading of piles in locally strongly varying sand layers,” ProceedDCS16 ing of the 2nd European Symposium on Penetration Testing, Amsterdam, pp. 927-932 (1982). Thompson, C. D., and Devata, M., “Evaluation of ultimate bearing capacity of different piles by wave equation analysis,” ProceedDCS17 ing of the International Seminar on the Application of Stress-Wave Theory on Piles, Stockholm, pp. 163-195 (1980). DCS18, UCR 14, Liew, S. S., Ng, H. B., and Lee, K. K., “Comparison of HSDPT and SLT results of driven piles in Malaysian residual soils,” UCR15 www.gnpgeo.com (2004). DCS20 Sahajda, K., “Calculation of piles based on CPT results in Poland,” www.aarsleff.com. Federal Highway Administration, “A laboratory and field study of composite piles for bridge substructures,” Report No. FHWADCS21 HRT-04-043 (2006). de Gijt, J. G., van Dalen, M., and Middendorp, P., “Comparison of statnamic load test and static load tests at the Rotterdam harDCS22 bour,” Proceeding of the 1st International Statnamic Seminar, Vancouver, (1995). Holm, G., Jansson, M., and Moller, B., “Dynamic and static load testing of friction piles in a loose sand,” Proceeding of the 2nd DCS23 International Conference on the Application of Stress-Wave Theory on Piles, Stockholm, pp. 240-243 (1985). Pando, M., Filz, G., Ealy, C., and Hoppe, E., “Axial and lateral load performance of two composite piles and one prestressed conDCS24 crete pile,” Transportation Research Record, No. 1849, pp. 61-70 (2003). Federal Highway Administration, “Design and construction of driven pile foundations-Lessons learned on the Central Artery/TunUCR1, UCR2 nel project,” Report No. FHWA-HRT-05-159 (2006). Advanced Geotechnical Engineering Services, “Load test report of SBMA administration building,” Subic Bay Freeport Zone, UCR3 Philippines, (2005) − 1014 −
KSCE Journal of Civil Engineering
Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
Table 6. (Continued) Pile no. UCR4 UCR5 UCR9 UCR10 UCR11 UCR12 UCR13 UCR16 UCR17 UCR 18 UCS1 UCS2 UCS3 UCS4 UCS5 UCS6 UCS7 UCS8 UCS9 UCS10 UCS11 UCS12 UCS13 UCS14 UCS15 UCS16 UCS17 UCS18 UCS19 UCS20
Reference source Salem, H. S., Fellenius, B. H., and Lavergne, H. R., “Using dynamic pile testing to overcome surprising soil variations,” Proceeding of the 33rd Annual Conference on Deep Foundations and 11th International Conference on Piling and Deep Foundations, New York, 7 p., (2008). Aoki, N., “Discussion”, Proceeding of the 12th International Conference on Soil Mechanics and Foundation Engineering, Re de Janeiro, Vol. 2, pp. 2977-2979 (1989). de Albuquerque, P. J. R., and de Carvalho, D., “Dynamic load test and elastic rebound analysis for estimation of the bearing capacity of piles in residual soil,” Proceeding of the 6th International Conference on the Application of Stress-Wave Theory to Piles, Brazil, pp.677-681, (2000). Zheng, Y. M., Zheng, J. M and Chen, B., “Correlation analysis of dynamic and static loading tests for nine piles,” Proceeding of the 6th International Conference on the Application of Stress-Wave Theory to Piles, Brazil, pp.651-656, (2000). Amini, A., Fellenius, B. H., Sabbagh, M., Naesgaard, E., and Buehler, M., “Pile loading tests at Golden Ears bridge,” Proceeding of the 61st Canadian geotechnical Conference, Edmonton, 8 p., (2008). McVay, M. C., Badri, D., and Hu, Z. “Determination of axial pile capacity of prestressed concrete cylinder piles,” Project No. 4910450487712, Florida Department of Transportation, Tallahassee, Florida (2004). Liew, S. S., and Kowng, Y. M., “Design, installation and verification of driven piles in soft ground,” Proceeding of the 11th International Conference of the International Association for Computer Methods and Advances in Geomechanics, Torino, Italy, (2005). Karunanidee, V. “Load resistance behavior and installation assessment of driven spun pile” Master Thesis, Department of Civil Engineering, University Teknologi Malaysia, (2010) Chen, W. C., “Prediction of ultimate load bearing capacity of driven piles” Master Thesis, Department of Civil Engineering, University Teknologi Malaysia, (2006) Kamal Uddin, M., and Tungsanga, K., “Dynamic pile testing and its correlation with static load test” Journal of Civil Engineering, Bangladesh, Vol. CE29, No.1 (2001) Leung, C. F., Radhakrishnan, R., and Tan, S. A., “Performance of precast driven piles in marine clay,” Journal of Geotechnical Engineering, ASCE, USA, Vol. 117, pp. 637-657 (1991). Rajasvaran, K., “Comparison of ultimate bearing capacity obtained by pile driving analyzer and maintained load test,” Master Thesis, Department of Civil Engineering, University Technology, Malaysia, (2007) Jaime, A. P., Romo, M. P., Ponce, J. A., and Mitre, A. M., “Static tests on friction piles in Mexico clay,” Proceeding of the 12th International Conference on Soil Mechanics and Foundation Engineering, Re de Janeiro, Vol. 2, pp. 1141-1146 (1989). You-zai, X. and Yabuuchi, S., “Bearing capacity of precast nodular piles in Tianjin soft clayey soil,” Proceeding of the 4th International Conference on Piling and Deep Foundations, Italy, pp. 319-323 (1991). Tan, S.B., Tan, S. L., Chin, Y. K. and Lor, B. L., “Dynamic pile testing in Singapore,” Proceeding of the 9th Southeast Asian Geotechnical Conference, Bangkok, Thailand, pp. 6-241-6-252 (1987). McCabe, B. A. and Lehane, B. M., “Behavior of axially loaded pile groups driven in clayey silt,” Journal of Geotechnical and Geoenvironmental Engineering., ASCE, Vol. 132, No. 3, pp. 401-410 (2006). Adib, M. E., “Load tests on prestressed precast concrete and timber piles,” Journal of Geotechnical and Geoenvironmental Engineering, ASCE, Vol. 127, No. 12, pp. 1043-1050 (2001). Trochanis, A. M., Bielak, J. and Christiano, P., “Three-dimensional nonlinear study of piles,” Journal of Geotechnical Engineering, ASCE, Vol. 117, No. 3, pp. 429-447 (1991). Goldemberg, H. G., and Goldemberg, J. J., “Is DLT the final word? Correlation between DLT and SLT,” Proceeding of the 6th International Conference on the Application of Stress-Wave Theory to Piles, Brazil, pp.719-723, (2000). Hajduk, E. L., Lin, G. and Yang, W., “Comparison of dynamic, static, and statnamic axial load testing on concrete piles in Savannah, GA.,” www.wpceng.com. Meyerhof, G. G., and Murdock, L. J., “An investigation of the bearing capacity of some bored and driven piles in London clay,” Geotechnique, Vol. 3, No. 7, pp. 267-282 (1953). Whitaker, T., and Cooke, R. W., “A new approach to pile testing,” Proceeding of the 5th International Conference on Soil Mechanics and Foundation Engineering, Paris, France, Vol. 2, pp. 171-176 (1961). Chapman, G. A. and Wagstaff, J. P., “The effect of bitumen slip coating on the drivability of precast concrete piles,” Proceeding of the 4th International Conference on Piling and Deep Foundations, Italy, pp. 193-199 (1991). Hegazy, Y. A., Cushing, A. G., and Lewis, C. J., “Driven pile capacity in clay and drilled shaft capacity in rock from field load tests,” Proceeding of the 5th International Conference on Case Histories in Geotechnical Engineering, New York, N. Y., pp. 1-8 (2004). Prakoso, W. A., and Hadiwardoyo, S. P., “CPT-based ultimate capacity of driven piles in residual soil,” Proceeding of the 2nd International Symposium on Cone Penetration Testing, California, Paper No. 3-20 (2010). Khare, M. G., and Gandhi, S. R., “Performance of bituminous coats in reducing negative skin friction,” Proceeding of the 10th International Conference on Piling and Deep Foundations, Amsterdam, The Netherlands, pp. 475-483 (2006). Kumar, S., Alarcon, C., Schmitt, B. R., and Kort, D., “Construction and full-scale testing of precast concrete piles made with coal combustion products,” Electronic Journal of Geotechnical Engineering, 10(A1), 2005. Raju, V. S., and Gandhi, S. R., “Ultimate capacity of precast driven piles in stiff clay,” Indian Geotechnical Journal, Vol. 19, No. 4, pp. 273-289 (1989). Kee, C. F., “Behaviour of piles in the soft organic clays in Southeast Asia,” Proceeding of the Geotechnical Engineering in Southeast Asia, pp. 111-126 (1985). Holeyman, A., Maertens, J., Huybrechts, N., and Legrand, C., “Result of an international pile dynamic testing prediction event,” Proceeding of the 6th International Conference on the Application of Stress-Wave Theory to Piles, Brazil, pp.725-739, (2000).
Vol. 17, No. 5 / July 2013
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Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
Table 7. Range of Geometry of Driven Piles for Analysis Data
Number of tests
DCR
37
DCS
45
UCR
36
UCS
34
Pile Geometry (m)
Statistics Range Mean COV Range Mean COV Range Mean COV Range Mean COV
Depth, D
Diametera, B
7.5-40.0 22.1 0.37 6.0-48.5 17.9 0.56 4.5-57.5 26.7 0.50 4.4-60.3 18.7 0.61
0.28-0.91 0.48 0.33 0.24-0.76 0.38 0.30 0.18-1.37 0.50 0.57 0.25-0.46 0.33 0.17
D/B
Interpreted capacity, QL2 (kN)
11.4-133.3 48.6 0.43 17.1-121.2 49.3 0.52 6.0-142.0 63.7 0.55 14.5-152.7 56.6 0.61
189-5780 2490 0.71 300-7295 2307 0.66 218-11300 4445 0.80 63-3806 1509 0.76
a
or width of square section Table 8. Summary of Interpreted Capacities for Drained Compression Tests Pile section
Data
QL1 QDB na 37 35 Mean 0.36 0.67 Round (DCR) SD 0.11 0.15 COV 0.30 0.22 na 45 45 Mean 0.32 0.58 Square (DCS) SD 0.12 0.13 COV 0.36 0.22 na 82 80 Mean 0.34 0.62 All Data SD 0.11 0.14 COV 0.33 0.23 a not including interpreted results with “>” symbol
QST 35 0.71 0.14 0.20 44 0.64 0.11 0.17 79 0.67 0.13 0.19
QDAV 34 0.75 0.15 0.20 45 0.69 0.10 0.15 79 0.71 0.13 0.18
Interpreted Q/QCHIN QT&P 32 0.79 0.13 0.17 36 0.75 0.12 0.16 68 0.77 0.13 0.16
QL2 34 0.80 0.10 0.13 41 0.78 0.09 0.11 75 0.79 0.10 0.12
QF&H 34 0.80 0.09 0.12 41 0.79 0.07 0.09 75 0.79 0.08 0.10
QVDV 37 0.85 0.09 0.11 45 0.81 0.08 0.10 82 0.83 0.09 0.11
QCHIN 37 1.00 45 1.00 82 1.00 -
Table 9. Summary of Interpreted Displacements for Drained Compression Tests Pile section
Data
δL1 δDB na 37 35 Mean 4.4 12.6 Round (DCR) SD 3.2 7.8 COV 0.73 0.62 na 45 45 Mean 4.5 11.0 Square (DCS) SD 3.3 5.7 COV 0.73 0.52 na 82 80 Mean 4.5 11.7 All Data SD 3.2 6.7 COV 0.73 0.57 a not including interpreted results with “>” symbol b by definition, δT&P = 25.4 mm, so there is no SD and COV c “n” values for δVDV < QVDV
δST 35 15.6 5.5 0.36 44 13.8 4.9 0.36 79 14.6 5.2 0.36
Displacement at Interpreted Criteria (mm) δDAV δT&Pb δL2 δF&H 34 32 34 34 20.0 25.4 25.3 29.0 10.7 − 11.7 18.2 0.54 − 00.46 0.63 45 36 41 41 19.7 25.4 26.5 28.6 12.8 − 13.9 15.1 0.65 − 0.53 0.53 79 68 75 75 19.8 25.4 25.9 28.8 11.9 − 12.9 16.5 0.60 − 0.50 0.57
sections are similar, and the COV values are comparable. However, for all the criteria, there are somewhat smaller mean ratios (Qχ / QCHIN) for the square piles at comparable mean
δVDV 23c 46.0 21.0 0.46 31c 36.1 14.7 0.41 54c 40.7 18.1 0.44
δCHIN 37 >43.8 20.6 0.47 45 >36.7 13.9 0.38 82 >39.9 17.5 0.44
displacements (Table 9). Among these methods, QDB is the lowest mean interpreted value, while QVDV is the highest value. By definition, the QCHIN ratio is 1.0 and is always above the
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Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
actual field measurements. The QL1 has the smallest mean ratio and displacement (Table 9), which implies that the initial linear region occurs within a very small displacement. The statistics of the mean ratios show that smaller displacements correspond to higher COVs, which may result from fluctuation during the initial loading or possible measurement sensitivity. The mean compression displacements for round and square piles shown in Table 9 follow the same order as the capacities. For round piles, the displacements at the interpreted failure load range from 12.6 mm at QDB to 25.3 mm at QL2 to >43.8 mm for QCHIN, while the square section piles range from 11.0 mm at QDB to 26.5 mm at QL2 to >36.7 mm for QCHIN. The COV values for these displacement data are high with a range of 36 to 63% (round) and 36 to 65% (square). For more direct comparison of the criteria and to observe the shape effects, the normalized load-displacement curves for round and square piles are presented in Fig. 2. The corresponding mean ratio of each interpretation method to QCHIN is plotted against the
mean displacement. Comparison of the normalized curves demonstrates a slightly stiffer behavior for round piles. This behavior might result from the non-uniform stress distribution across the face of the square pile, faster and more uniform set-up for round piles, or other reasons, including database differences. However, the differences are relatively small, especially beyond L2. For both sections, DeBeer, slope tangent, Davisson, and Terzaghi and Peck typically are located within the L1 to L2 transition, while Fuller and Hoy and van der Veen are close to or slightly beyond the failure threshold. Points L1 and L2 are convenient reference points within the curve because these points encompass the significant regions of the curve. However, it should be noted that measured loaddisplacement curves would not always possess a final linear region as indicated in the results (Tables 2 to 5). Nevertheless, these points are useful to demonstrate the general or first-order relationships among the criteria and provide capacity approximations. To illustrate these relationships, the round and square drained data are combined (Table 8) since, as previously noted, their differences are relatively small. Using L1 as a reference, the capacity for drained loading gives: QDB = 1.8 QL1, QST = 2.0 QL1, QDAV = 2.1 QL1, QTP = 2.3 QL1, QFH = 2.3 QL1, QL2 = 2.3 QL1, QVDV = 2.4 QL1, QCHIN = 2.9 QL1. Using L2, the capacity can be approximated for drained loading as: QDB = 0.78 QL2, QST = 0.85 QL2, QDAV = 0.90 QL2, QTP = 0.97 QL2, QFH = 1.0 QL2, QVDV = 1.05 QL2, QCHIN = 1.27 QL2. These ratios can be used to interrelate the methods where needed because of insufficient load-displacement data or for prematurely terminated load tests.
5. Evaluation Results for Undrained Load Tests
Fig. 2. Mean Load-Displacement for Drained Compression Loading
The statistics for the interpreted capacities and the corresponding displacements for undrained compression tests are summarized in Tables 10 and 11, respectively, which basically follows the same approach used for the drained tests. Both tables include the undrained round section (UCR), undrained square section (UCS), and all undrained loading data combined.
Table 10. Summary of Interpreted Capacities for Undrained Compression Tests Pile section
Data
QL1 QDB na 36 33 Mean 0.30 0.54 Round (UCR) SD 0.11 0.18 COV 0.38 0.33 na 34 29 Mean 0.32 0.58 Square (UCS) SD 0.10 0.15 COV 0.30 0.26 na 70 62 Mean 0.31 0.56 All Data SD 0.10 0.16 COV 0.34 0.30 a not including interpreted results with “>” symbol Vol. 17, No. 5 / July 2013
QST 28 0.62 0.17 0.27 24 0.66 0.12 0.18 52 0.64 0.15 0.23
QDAV 23 0.68 0.17 0.25 21 0.64 0.17 0.26 44 0.66 0.17 0.25
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Interpreted Q/QCHIN QT&P 24 0.69 0.15 0.22 17 0.74 0.12 0.16 41 0.71 0.14 0.19
QL2 21 0.76 0.10 0.14 21 0.75 0.10 0.13 42 0.75 0.10 0.13
QF&H 18 0.80 0.09 0.12 19 0.73 0.09 0.13 37 0.77 0.10 0.13
QVDV 36 0.78 0.09 0.12 34 0.76 0.11 0.14 70 0.77 0.10 0.13
QCHIN 36 1.00 34 1.00 70 1.00 -
Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
Table 11. Summary of Interpreted Displacements for Undrained Compression Tests Pile section
Data
δL1 δDB na 36 33 Mean 5.1 14.3 Round (UCR) SD 3.1 8.2 COV 00.62 0.57 na 34 29 Mean 4.8 11.3 Square (UCS) SD 4.0 7.5 COV 0.83 0.67 na 70 62 Mean 4.9 12.9 All Data SD 3.5 8.0 COV 0.72 0.62 a not including interpreted results with “>” symbol b by definition, δT&P = 25.4 mm, so there is no SD and COV c “n” values for δVDV < QVDV
δST 28 19.6 7.5 0.38 24 15.0 6.5 0.43 52 17.5 7.3 0.42
Displacement at Interpreted Criteria (mm) δDAV δT&Pb δL2 23 24 21 24.2 25.4 33.1 10.0 − 14.5 0.41 − 0.44 21 17 21 13.9 25.4 23.4 9.2 − 19.0 0.66 − 0.81 44 41 42 19.3 25.4 28.2 10.8 − 17.4 0.56 − 0.62
The results for round and square section piles in Table 10 show mean load ratios ranging from 0.54 to 0.80 and 0.58 to 0.76, respectively, when compared to QCHIN, with COV values of 12 to 33% and 13 to 26%, respectively. These ranges are equally wide as those for drained loading with the same general trend observed from the methods. However, based on these data, the Fuller and Hoy method has fewer load test cases that could be interpreted, which is attributed to the definition of the method. In contrast to drained load tests, the mean load ratios of round and square piles generally are in good agreement. However, larger mean displacements occur for round piles (Table 11), which could be attributed to the larger diameter and capacity ranges of undrained-round section piles in the database. Among these methods, QDB is the lowest mean interpreted value, while QF&H is the highest value. The QCHIN ratio is 1.0 by definition and is always above the actual field measurements. The QL1 has the smallest mean ratio and displacement (Table 11). As in drained loading, the COV ranges of the mean load ratios are large for those criteria with relatively small displacements, which implies that a higher variability of load-displacement behavior can be expected during the initial loading for undrained conditions. The mean compression displacements shown in Table 11 follow a similar order as the capacities. The displacements at the interpreted failure load range from 14.3 mm at QDB to 33.1 mm at QL2 to >36.6 mm for QCHIN for round piles and QDB to 23.4 mm at QL2 to >26.5 mm for QCHIN for square piles . The COV values for these displacement data range from 38 to 66% (round) and 43 to 78% (square). The normalized load-displacement curves for round and square piles are presented in Fig. 3. As in the drained condition, similar general trends of interpretation criteria are observed from these curves. For both sections, DeBeer, slope tangent, Davisson, and Terzaghi and Peck typically are located within the L1 to L2 transition, while Fuller and Hoy and van der Veen are close to or slightly beyond the failure threshold.
δF&H 18 41.4 27.5 0.66 19 20.8 16.3 0.78 37 30.8 24.5 0.79
δVDV 18c 43.0 26.0 0.60 19c 30.9 22.0 0.71 37c 36.8 24.5 0.66
δCHIN 36 >36.6 24.1 0.66 34 >26.5 19.9 0.75 70 >31.7 22.6 0.71
Fig. 3. Mean Load-Displacement for Undrained Compression Loading
The combined round and square undrained data are likewise indicated in Table 10. Using L1 as a reference, the capacity for undrained loading gives: QDB = 1.8 QL1, QST = 2.1 QL1, QDAV = 2.1 QL1, QTP = 2.3 QL1, QFH = 2.5 QL1, QL2 = 2.4 QL1, QVDV = 2.5 QL1, QCHIN = 3.2 QL1. Using L2, the capacity can be approximated for undrained loading as: QDB = 0.75 QL2, QST = 0.85 QL2, QDAV = 0.88 QL2, QTP = 0.95 QL2, QFH = 1.03 QL2, QVDV = 1.03 QL2, QCHIN = 1.33 QL2.
6. Comparison of Drained and Undrained Load Tests Comparisons of drained and undrained load tests in Tables 8 to 11 and Fig. 4 show some interesting differences. First, the SD and COV values for the drained tests are slightly less than those
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Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
loading since it could not interpret many undrained cases. On average, QL1 = 0.32 QCHIN and QL2 = 0.77 QCHIN. The displacements at δL1 are consistent for both drained and undrained loading and occur at a small displacement of around 4 to 5 mm. Although δL2 gives a larger displacement for undrained loading, the difference between undrained and drained loading is quite small, having a range of 25 to 28 mm.
7. Effects of Installation Criteria on Driven Pile Performance
Fig. 4. Mean Load-Displacement Comparison for Drained and Undrained loading
for the undrained tests. Second, all of the capacity ratios (Qx/ QCHIN) in drained loading are larger than in undrained loading, although the differences are small. And third, the displacements for drained tests, especially for higher values, tend to be a bit higher. Whether these differences are fundamental or databaserelated is unclear. Overall, similar trends are shown by the interpreted results of all criteria for both drained and undrained loading as shown in Fig. 4. The DeBeer method represents the lower bound and is located within the nonlinear L1 to L2 transition, while Chin is the upper bound and is always above all measured results. Slope tangent, Davisson, and Terzaghi and Peck are also within the nonlinear L1 to L2 transition. In addition, Fuller and Hoy and van der Veen occur near or above L2. Among these criteria, L2, Fuller and Hoy, and van der Veen give the smallest COVs. However, Fuller and Hoy is more applicable for drained than undrained
The installation method is one of the important factors that affect the performance of driven piles. In the database, four general types of impact hammers were used to install the piles. The hammers exerted different driving energies and produced a wide range of pile resistances. The drained and undrained databases are further sub-divided by hammer type (drop, air/ steam, diesel, and hydraulic), as presented in Tables 2 to 5, to assess the hammer influence on load-displacement behavior. 7.1 Effects of Pile Hammer on General Load-Displacement Behavior The mean load-displacement curves for the different pile hammers are illustrated in Figs. 5(a) and 5(b) for drained and undrained loading, respectively. For drained loading, a stiffer load-displacement behavior is demonstrated by the drop hammer while the others are in good agreement. There was only one case using a steam hammer, so it was omitted to prevent undue bias. For undrained loading, a similar stiff behavior for the drop hammer was noted, while the most ductile behavior was exhibited by the diesel hammer. The higher stiffness of the drop hammer is likely due to the large pile sizes driven by the diesel hammer relative to the small piles driven by the drop hammer. This would indirectly affect the pore water pressure surrounding the pile, but the available data are insufficient to assess this point.
Fig. 5. Mean Load-Displacement Comparison of Different Hammers for: (a) Drained Loading and (b) Undrained Loading Vol. 17, No. 5 / July 2013
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Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
Table 12. Range of Driving Records for Different Hammers Drained Hammer type
Drop
Air/steam
Diesel
Hydraulic
Statistics n Range Mean COV n Range Mean COV n Range Mean COV n Range Mean COV
Rated energy (kN-m) 10 7.5 - 37.5 15.0 0.82 9 20.3 - 59 40 0.30 17 33.6 - 108.6 58.0 0.38 2 85.4 - 120 102.7 0.24
Undrained Final set (blows/25 mm) 10 2 - 10 4 0.66 9 2 - 13 4 0.55 17 4 - 50 18 0.94 2 4 -11 7 0.65
7.2 Effect of Hammer Type Driving on Pile Capacity The pile driving resistance, as expressed by the blow count, has been used to assess pile capacity during construction. Hammer rated energy contributes to this driving resistance and, therefore, capacity. To examine this issue, cases with complete driving records were analyzed. The ranges of these criteria for the different hammers are shown in Table 12. For general comparison, capacity estimates were made using the very simple Engineering News (EN) formula. The EN formula (Wellington, 1893) was derived from work-energy theory and is given as: QEN = Wrh/(s+C)
(1)
in which QEN = pile capacity, Wr = weight of ram, h = drop height of ram, s = penetration of pile per hammer blow, and C = constant (2.54 for air/steam hammers, and 25.4 for drop hammers). In this study, C = 2.54 is also adopted for diesel and hydraulic hammers. For single and double-acting hammers, the term Wrh can be replaced by EHE, in which E = hammer efficiency
Rated energy (kN-m) 3 13.5 - 70.0 32.3 1.01
Final set (blows/25 mm) 3 4 - 20 12 0.65
-
-
15 35.7 - 130 84.0 0.36 15 8.1 - 145.0 67.7 0.66
15 4 - 35 15 0.56 15 2 - 21 10 0.46
and HE = hammer rated energy. For simplicity, QL2 is used for comparison. The relationship between predicted capacity (QEN) and measured capacity (QL2) is shown in Fig. 6. The scatter is substantial. For drop hammers, QEN is underestimated for both drained and undrained loading, at about 50 and 60% of QL2, respectively. In contrast, for drained loading, QEN for air/steam, hydraulic, and diesel hammers is approximately 2, 2.5, and 4 times QL2, respectively. Similarly, for undrained loading, QEN for hydraulic and diesel hammers is roughly 2.5 and 3 times QL2, respectively. The tendency of the EN formula to overestimate pile capacity is a well-known fact, but it is underestimated for drop hammers in this study. One possible reason for the differences is the constant “C”. Wellington (1893) derived this constant based on the extra initial resistance to get the pile in motion again, so the values should vary depending on hammer type. Therefore, the “C” constants for drop, air/steam, diesel, and hydraulic hammers were backcalculated using the measured results (QL2). Table 13 shows the statistical results of back-calculated “C”
Fig. 6. Comparison of Predicted (QEN) and Measured (QL2) Capacity of Pile Hammers for: (a) Drained Loading, (b) Undrained Loading − 1020 −
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Evaluation of Compression Load Test Interpretation Criteria for Driven Precast Concrete Pile Capacity
Table 13. Statistical Results of Back-Calculated C Mode
Drained
Undrained
Hammer type
n
Mean
SD
COV
Drop
10
10.8
4.90
0.45
Air/steam
9
12.6
6.51
0.52
Diesel
17
19.0
5.61
0.30
Hydraulic
2
10.6
0.39
0.04
Drop
3
15.9
5.27
0.33
Diesel
15
14.4
6.83
0.47
Hydraulic
15
12.9
6.81
0.49
for the different hammer types and loading conditions. For drop hammers, the mean values are 10.8 and 15.9 for drained and undrained loading, respectively. These values obviously are smaller than the suggested value (25.4) by about 60% for drained and 40% for undrained loading. The back-calculated C values for air/steam, diesel, and hydraulic hammers in drained loading are 12.6, 19.0, and 10.6, respectively, while for undrained loading, diesel and hydraulic hammers have values of 14.4 and 12.9, respectively. These values are much larger than 2.54, but they can possibly lead to a better prediction of pile capacity. However, more data are needed to verify these results.
8. Comparison of Driven Piles and Drilled Shafts To examine more general foundation behavior, a comparison was made of the load-displacement behavior of driven PC piles and drilled shafts. For driven piles, the round and square sections were combined for both drained and undrained loading since their behavior is comparable and the data are consistent. The drilled shaft results are from the recent study by Chen and Fang (2009). The normalized load-displacement curves are shown in Figs. 7(a) and 7(b) for drained and undrained loading, respectively. For comparison, the interpreted results are normalized by QCHIN. For ease in comparison, L1 and L2 are marked in Figs. 7(a) and 7(b).
For drained loading in Fig. 7(a), the initial linear region is comparable for both foundation types, but a more ductile behavior develops for drilled shafts after the initial linear region. The drilled shaft study indicated δL2 = 34.9 mm; this study has a mean δL2 = 25.9 mm. The drilled shaft database (Chen and Fang, 2009) indicated larger diameters, which requires larger displacements to mobilize the full capacity in drained soils. Another likely cause of larger displacement is the uncompacted tip for drilled shafts versus a compacted tip for driven piles. For undrained loading in Fig. 7(b), the load-displacement curve of drilled shafts shows a stiffer response until the end of the transition region of the curve. This may be attributed to more disturbance in undrained soils during pile driving where the effect of pore pressure is significant. The drilled shaft study indicated δL2 = 30.9 mm for undrained loading; this study has a mean δL2 = 28.2 mm.
9. Conclusions Eight representative interpretation criteria (van der Veen, Chin, Fuller and Hoy, Terzaghi and Peck, DeBeer, Davisson, Slope Tangent, and L1-L2) were evaluated based on a compiled database of driven PC piles. Axial compression static load tests on round and square cross-section piles in various soil profiles were used for this purpose. The interpretation criteria for both pile cross-sections under drained and undrained loading demonstrated approximately consistent trends. Based on these analyses for driven PC piles, the following conclusions are reached, and design recommendations are proposed for their use. 1. Of the eight criteria examined, the DeBeer method gives the lower bound, while the Chin method is the upper bound, almost by definition, and is always above the measured data. L1 is a consistent definition for the elastic limit, while L2 can be a useful definition for the interpreted failure load in driven piles. 2. Capacity approximation using L1 for drained and undrained loading, respectively, gives: QDB = [1.8, 1.8] QL1, QST = [2.0, 2.1] QL1, QDAV = [2.1, 2.1] QL1, QTP = [2.3, 2.3] QL1, QFH =
Fig. 7. Mean Load-Displacement Comparison of Driven Piles and Drilled Shafts for: (a) Drained Loading, (b) Undrained loading Vol. 17, No. 5 / July 2013
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Maria Cecilia M. Marcos, Yit-Jin Chen, and Fred H. Kulhawy
[2.3, 2.5] QL1, QL2 = [2.3, 2.4] QL1, QVDV = [2.4, 2.5] QL1, QCHIN = [2.9, 3.2] QL1. Using L2, the capacity can be approximated for drained and undrained loading, respectively, as: QDB = [0.78, 0.75] QL2, QST = [0.85, 0.85] QL2, QDAV = [0.90, 0.88] QL2, QTP = [0.97, 0.95] QL2, QFH = [1.0, 1.03] QL2, QVDV = [1.05, 1.03] QL2, QCHIN = [1.27, 1.33] QL2. The DeBeer method is a significant underpredictor, while the Chin method is a significant overpredictor. Both methods should be disregarded. Terzaghi and Peck, Fuller and Hoy, and van der Veen give results comparable to L2. Davisson and slope tangent are modest underpredictors of L2, but they are convenient and simple to use. The ratios given can be used to interrelate the methods where needed due to limited loaddisplacement data or prematurely terminated load tests. 3. The displacements at the interpreted failure load for the Chin method are >40 mm (drained) and >35 mm (undrained). The Terzaghi and Peck, L2, Fuller and Hoy, and van der Veen methods give “failure” displacements between 25 to 40 mm (drained) and 25 to 35 mm (undrained), while the DeBeer, slope tangent, and Davisson methods give even smaller “failure” displacements between 10 to 20 mm. At the elastic limit L1, the interpreted displacement is about 5 mm. 4. Statistical analyses show a decreasing COV with increasing pile displacement, and drained loading shows less variability than undrained loading. 5. The relatively small differences in behavior between round and square piles can be attributed to several factors. For drained loading, the somewhat smaller mean ratios (Qχ/QCHIN) for the square piles at comparable mean displacements might result from the non-uniform stress distribution across the face of the square pile, faster and more uniform set-up for round piles, or other reasons, including database differences. In contrast, the mean load ratios for both sections under undrained loading are generally in good agreement. However, larger mean displacements are produced by round piles which likely can be attributed to the larger diameter and capacity ranges of undrained-round section piles in the database. Meanwhile, all of the capacity ratios (Qx/QCHIN) in drained loading are larger than in undrained loading, so the drained load-displacement curve gives a stiffer response. 6. Piles driven by drop hammers demonstrate stiffer load-displacement behavior than diesel, air/steam, and hydraulic hammers for both drained and undrained loading. 7. The EN formula greatly overestimates the pile capacity. The back-calculated constants “C” for drained loading are 10.8, 12.6, 19.0, and 10.6 for drop, air/steam, diesel, and hydraulic hammers, respectively. For undrained loading, the values are 15.9, 14.4, and 12.9 for drop, diesel, and hydraulic hammers, respectively. 8. Somewhat smaller displacements are required to mobilize the full capacity of driven piles compared with drilled shafts in drained loading, while drilled shafts shows a stiffer loaddisplacement response until the end of the transition region of the curve in undrained loading.
9. In practice, if load tests are performed, a minimum safety factor of 2 is common for pile design. Using L2 as the definition for interpreted failure load (or the roughly comparable Terzaghi and Peck, Fuller and Hoy, or van der Veen methods herein, which are within + or - 5% of L2), with FS = 2, the mean displacements should be less than 10 mm, only a modest amount above L1. Davisson and slope tangent will give a lower interpreted failure load (by 10 to 15%) and slightly smaller displacement. These displacements should be tolerable for most structures and will be close to elastic in behavior.
Acknowledgements This study was supported by the National Science Council, Taiwan, under contract number: NSC 100-2221-E-033-073-MY3.
References Chen, Y.-J., Chang, H.-W., and Kulhawy, F. H. (2008). “Evaluation of uplift interpretation criteria for drilled shaft capacity.” J. Geotech. Geoenviron. Engrg., ASCE, Vol. 134, No. 10, pp. 1459-1468. Chen, Y.-J. and Fang, Y.-C. (2009). “Critical evaluation of compression interpretation criteria for drilled shafts.” J. Geotech. Geoenviron. Engrg., ASCE, Vol. 135, No. 8, pp. 1056-1069. Chin, F. K. (1970). “Estimation of the ultimate load of piles not carried to failure.” Proc., 2nd Southeast Asian Conf. Soil Engrg., Singapore, pp. 81-90. Davisson, M. T. (1972). “High capacity piles.” Proc., Lect. Series on Innov. in Found. Const., ASCE, Illinois Section, Chicago, p. 52. DeBeer, E. E. (1970). “Experimental determination of shape factors of sand.” Geotechnique, Vol. 20, No. 4, pp. 387-411. Duzceer, R. and Saglamer A. (2002). “Evaluation of pile load test results.” Proc., 9th Int. Conf. Piling and Deep Found., Deep Foundation Institute, Nice, France. Fellenius, B. H. (1975). “Test loading of piles-methods, interpretation and new proof testing procedure.” J. Geotech. Engrg. Div., Vol. 101, No. 9, pp. 855-869. Fellenius, B. H. (2006). Basics of foundation design, Electronic Ed. www.Fellenius.net. Fuller, F. M. and Hoy, H. E. (1970). “Pile load tests including quick load test method, conventional methods, and interpretations.” Research Record 333, Hwy. Res. Brd., Washington, pp. 74-86. Hirany, A. and Kulhawy, F. H. (1988). Conduct and interpretation of load tests on drilled shaft foundations: Detailed guidelines. Rpt EL5915(1), EPRI, Palo Alto, California. Hirany, A. and Kulhawy, F. H. (1989). “Interpretation of load tests on drilled shafts (2): Axial uplift.” Found. Engrg.,: Current Princ. & Pract. (GSP22), ASCE, New York, pp. 1150-59. Hirany, A. and Kulhawy, F. H. (2002). “On the interpretation of drilled foundation load test results.” Deep Found. 2002 (GSP 116), ASCE, Reston, pp. 1018-1028. O’Rourke, T. D. and Kulhawy, F. H. (1985). “Observations on load tests on drilled shafts.” Drilled Piers and Caissons II, ASCE, New York, pp. 113-128. Terzaghi, K. and Peck, R. B. (1967). Soil mechanics in engineering practice, 2nd Ed., Wiley, New York. van der Veen, C. (1953). “The bearing capacity of a pile.” Proc., 3rd Int. Conf. on Soil Mech. and Found. Engrg., Vol. 2, pp. 84-90.
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