IFCEE 2015 © ASCE 2015
Current Limitations and Challenges of Driven Piles in Rock as Demonstrated using Three Case Studies in Wyoming Kam W. Ng1, Seyed Yashar Yasrobi2 and Todd A. Sullivan3 1
Assistant Professor, P.E., Department of Civil and Architectural Engineering, University of Wyoming, Laramie, WY 82071, USA (E-mail address:
[email protected]; Tel: 307 766 4388; Fax: 307 766 2221) (Corresponding Author) 2 Research Assistant, Department of Civil and Architectural Engineering, University of Wyoming, Laramie, WY 82071, USA 3 Engineering Geologist, P.G., Geology Program, Wyoming Department of Transportation, Cheyenne, WY 82009-3340, USA ABSTRACT: There are currently no static analysis methods available for estimating the geotechnical resistance of a driven pile in rock. According to the American Association of State Highway and Transportation Officials (AASHTO), the limiting factor for a pile in rock is its structural capacity. The estimation of its structural axial capacity depends strongly on an effective length factor (K) and its unbraced length (L). These parameters significantly depend on its soil confinement along the pile and rock support at its toe, which lead to a large discrepancy between estimated and measured resistances. In this paper three case studies of completed pile projects in Wyoming are presented to highlight the limited knowledge and challenges pertaining to present design and construction practices of driven piles in rock. Steel H-piles were installed at Burns South, Casper and Torrington sites in Wyoming. Static analysis method were used to estimate the geotechnical resistances of these piles. Wave Equation Analysis Program (WEAP) and Case Pile Wave Analysis Program (CAPWAP) were used to verify their performances during construction. Structural capacities of these piles were also calculated. The results of the studies show that the static analysis methods and structural analyses yield inconsistent pile resistance estimations. Several recommendations are proposed: 1) develop an analytical method for piles in rock; 2) conduct full-scale load test of piles in rock; 3) characterize rock properties; and 4) develop a relationship of percent pile bracing, soil properties and embedded pile length.
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INTRODUCTION Deep foundations are typically used to support bridges in Wyoming with driven steel piles being the preferred choice due to its high driving durability in rock. Because of a shallow bedrock stratigraphy in Wyoming, piles are often driven in rock materials. Hence, the total axial resistance of these piles consists of a combination of shaft resistance and end bearing. To attain the required resistance, especially in a soft overburden soil, the pile would have to rely on its end bearing in a stiff rock or Intermediate Geo-Materials (IGM). Unfortunately, there are currently no pragmatic static analysis methods available for estimating the end bearing of a driven pile in rock. For example, the characteristic lines method proposed by Serrano and Olalla (2002) based on Hoek and Brown’s non-linear failure model requires rock parameters that are not readily available for an ultimate end bearing estimation. They concluded that the proposed method is acceptable for piles bearing in poor rock with ultimate end bearings less than 418 to 627 ksf (20 to 30 MPa). However, the method is believed to be unconservative for end bearings greater than 627 ksf (30 MPa). Hannigan et al. (2006) acknowledged that the pile-rock contact area, penetration depth, and rock quality are usually not available for pile resistance estimation during a design state. Thus, the resisting performance of these piles depends on driving observations, dynamic and static load tests, and local experience. They noted that a pile resistance will usually be governed by its structural capacity when it is driven to end bearing in rock of fair to excellent quality based on Rock Quality Designation (RQD) values. On the other hand, a pile supported on soft weathered rock should be designed based on pile load test results, because 1) the rock strength would govern the pile resistance, and 2) pile resistance could decrease due to relaxation in soft weathered rock near pile toe (Thompson and Thompson, 1985). AASHTO LRFD Bridge Design Specifications The American Association of State Highway and Transportation Officials (AASHTO) Load and Resistance Factor Design (LRFD) Bridge Design Specifications (2012) provide the following recommendations for piles driven in rock: (1) Piles driven in soft rock shall be designed in the same manner as soil; (2) The nominal geotechnical resistance of piles driven to end bearing in hard rock shall be governed by the structural limit state described in Article 6; (3) There are no well acceptable approaches to differentiate soft from hard rocks. However, local experience with driving piles in rock shall be applied to define the quality of rock; (4) Piles shall be driven based on locally developed criteria to prevent pile damage. Dynamic analysis methods should be used to evaluate pile drivability, control pile driving and detect pile damage; and (5) Article 6.5.4.2 provides the resistance factors (ϕ) for the strength limit state design. The resistance factor of an axial compression steel pile is 0.90 and the resistance factor during pile driving is 1.0. However, the application of these ϕ values are confused by another two ϕ values recommended for compression piles in the same section: a) ϕ of 0.50 for steel H-piles subject
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to damage due to severe driving conditions where use of a pile tip is required; and b) ϕ of 0.60 for steel H-piles under good driving conditions where use of a pile tip is not required. Following these recommendations, the structural compressive resistance of a steel H-pile shall be considered as the smallest value based on any applicable buckling failure modes. The structural resistance of each failure mode depends on the effective length of a pile (i.e., a product of effective length factor (K) and the unbraced pile length (L)). However, the effective pile length (KL) is influenced by the rotational and translational bracing along its length, contributing from rarely known overburden soil confinement as well as the rock fixity at its toe. Recommendations to address this issue are not provided in the AASHTO (2012). Neglecting the pile-geomaterial interaction and assuming an unrealistic KL value, the pile resistance will not be accurately estimated in terms of its structural compressive strength, leading to a potentially large discrepancy between estimated and measured resistances. Wyoming DOT Specifications and Practices The Wyoming Department of Transportation (WYDOT) adapts the AASHTO LRFD Bridge Design Specifications (2012) and applies local experiences to design and construct pile foundations for highway bridges. While the bridge design manual for substructure is currently under preparation, the WYDOT Standard Specifications for Road and Bridge Construction (2010) as well as the WYDOT Construction Manual (2014) Section 504 describe the following important construction control requirements for furnishing and driving steel bearing piles: (1) The adequacy of a pile hammer will be evaluated using a wave equation analysis method; (2) Pile driving stresses shall not exceed 90% of the minimum yield strength for steel piling; and (3) A dynamic formula, wave equation, or both will be used to establish pile driving criteria. If necessary, Pile Driving Analyzer (PDA) with subsequent signal matching analyses using the CAse Pile Wave Analysis Program (CAPWAP) will be used to determine and verify the required pile resistance during construction. Pile restrikes at 24 hours after the end of driving (EOD) will be required to further ensure that the desired pile resistance is achieved. A site investigation is performed at every bridge project to determine its subsurface profile and geomaterial properties. Standard Penetration Test (SPT) is the most commonly used in-situ field test in Wyoming. At the same borehole for SPT test, a drivepoint penetration test is performed by driving a 2-in (50-mm) diameter drivepoint into the ground using a 140-lb (63.5 kg) hammer at a drop height of 30 in (760 mm). Hammer blow counts to penetrate the drivepoint 1 ft (305 mm) into the ground are recorded. The main purpose of the drivepoint penetration test is to determine the depth of an adequate bearing layer, such as unweathered bedrock, for end bearing piles. Soil samples collected from the site investigation will be classified in accordance with the Unified Soil Classification System (ASTM D2487, 2012) and the AASHTO Classification System (M 145, 2012). Occasionally, unconfined
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compressive tests will be performed on cohesive soil samples to determine their respective undrained shear strengths. If a bedrock layer is encountered, a rock coring will be performed to determine the RQD value, and rock samples will be tested for the uniaxial compressive strength (qu). Based on past experiences, the WYDOT has developed a table of typical properties of compacted soils, which facilitates the quantification of geotechnical resistance of piles. However, locally calibrated unit shaft resistance and unit end bearing, especially unit end bearing of piles driven into a rock, are not available to directly and accurately estimate pile resistances. Nordlund (1963) method is frequently used to estimate the geotechnical resistance of piles in cohesionless soils, while Cone Penetration Test (CPT) method by Nottingham and Schmertmann (1975) is used for cohesive soils. Steel H-pile is the most commonly used steel pile in Wyoming, because of its suitability for hard driving in rock. Pile driving is mostly performed using locally available diesel hammers. A refusal blow count of 160 blows per ft is used by the WYDOT to prevent overstressing and damage to the pile. Pile driving will be terminated when a target nominal pile resistance is achieved at the planned depth, verifying using Wave Equation Analysis Program (WEAP) and/or PDA with CAPWAP analysis. Limitations and Challenges The aforementioned sections clearly highlight the current limitations with the design and construction control of piles driven in rock. These limitations are summarized as follows: (1) Static analysis methods are not available to estimate end bearing in rock; (2) Advanced rock parameters are not readily available for complex pile analyses; (3) Pile-rock interaction is usually not known; (4) Limited knowledge of rock quality as rock is normally characterized based only on the RQD and qu values; (5) Overburden soil is usually not well characterized. It is often estimated based on correlations developed from historical data; (6) No clear definition of soft and hard rocks. They are normally differentiated based on local experiences; (7) Limited knowledge about the bracing of a pile from its surrounding soil and rock materials prohibits an accurate structural analysis; (8) The performance of piles in rock is typically verified using dynamic analysis methods, which are not a proof load test. Furthermore, static load test, which is expensive and time consuming, is not performed to verify the pile resistance and calibrate the dynamic analysis methods; and (9) Inadequate recommendations are provided by current guidelines and specifications to address this problem. These limitations create the following challenges in the design and construction of pile foundations in rock, especially in the State of Wyoming: (1) The performance of piles in rock has a relatively high uncertainty, which incurs many challenges to construction management;
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(2) The estimated pile resistance based on the structural limit state is often not attainable during construction. Pile restrikes, pile extensions, increase in pile size and/or enlarging pile caps are required to achieve the required pile resistance. This will incur additional construction duration and cost; (3) It is difficult to prepare an accurate tender document for bridge foundations driven in rock to eliminate any potential construction cost overruns; (4) It is likely to create conflicts between owners and piling contractors. These conflicts usually result in change-orders to the contract for additional money and time to achieve the required pile performance; and (5) The operational cost of owners will increase due to the repeated trips back to the field during pile construction in order to evaluate the pile performance. The paper presents three case studies of completed pile projects in Wyoming to highlight the limitations and challenges pertaining to present design and construction practices of driven piles in rock. Steel H-piles were installed at Burns South Road, Casper Streets and Torrington Streets project sites. Static and dynamic analysis methods were performed to estimate the geotechnical resistances of these piles based on the available soil, rock, and pile driving data. Furthermore, structural capacities of these piles were determined for comparative purposes. CASE STUDY NO. 1 (BURNS SOUTH ROAD) A 430-ft (131-m) four span reinforced concrete bridge was constructed over an existing Union Pacific Rail Road (UPRR) near the intersection of Interstate-80 and Burns South Road in Laramie County, Wyoming. The construction of bridge foundation began late fall 2011 and continued in spring 2012. The bridge consisted of two abutments and three piers as shown in Figure 1. Five, Grade 50, HP 14×73 (HP 360×109) piles were installed in a pile group at each abutment while a pile group of twenty-one, same type of piles was constructed at each pier. Total LRFD factored loads per pile based on a combination of dead and live loads were 322 kips (1,432 kN) and 258 kips (1,148 kN) at the abutment and pier, respectively. Delmag D16-32 single acting diesel hammer was used to install all piles. Pile driving criterion established based on bearing graphs generated by WEAP was used to verify the performance of all piles. Pile No. 1 at Abutment No.1 and Pile No. 1 at Pier No. 3 were selected for dynamic tests using a PDA with subsequent CAPWAP analyses. The subsurface profile was generally silty sand (SM) overlying a dense to very dense silty, fine grained, non-cemented sandstone. Figure 2 shows the subsurface profiles along with the measured SPT N-values near the two test piles. Test Pile at Pier No. 3 Geotechnical and structural analyses were performed on the test pile at Pier No. 3 driven in the 37.4 ft (11.4 m) silty sand layer and 1.6 ft (0.50) into the sandstone bedrock with a total embedded length of 39 ft (11.9 m). Soil samples were collected for classification tests but not shear strength test. Using the SPT N-values and soil properties estimated from the WYDOT’s soil table, SPT-Meyerhof (1976), Nordlund (1963), DRIVEN (Mathias and Cribbs 1998) and β-method (Burland 1973) were used to estimate the nominal pile resistances as summarized in Figure 3. DRIVEN is a
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Figure 1. Plan view of Burns South Road project.
(a) Pile No. 1 at Pier No. 3 (b) Pile No. 1 at Abutment No. 1 Figure 2. Subsurface profile of test piles at Pier No. 3 and Abutment No.1 (Burns South Road project). program that estimates pile resistance based on a combination of Nordlund (1963) method for cohesionless soil and α-method (Tomlinson 1980) for cohesive soil. The RQD and qu values of the sandstone bedrock were not available while its SPT blow count was determined to be 100 per 10 in (254 mm). Since static analysis methods and locally calibrated unit end bearing are not available to estimate end bearing in rock, the rock material was treated as a soil with the maximum possible SPT N-value of 60. Dynamic tests using a PDA were performed on the test pile at the EOD on November 17, 2011 and at restrike the next day. Based on driving logs, hammer blows increased by more than 40 blows per ft when sandstone was encountered. The measured hammer blow counts were 100 and 108 blows per ft at the EOD and the
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restrike, respectively. The nominal pile resistances estimated using CAPWAP are presented in Figure 3. The percent shaft resistances estimated at the EOD and the restrike by CAPWAP were 40% and 63%, respectively. Using the driving, pile, soil and hammer information, driveability analyses were performed and bearing graphs were generated using the WEAP. Two soil input procedures, soil type based method (ST) and SPT N-value based method (SA), were used in the WEAP analysis, and their results are summarized in Figure 3.
77 46
321
309 185
535
677 406
310 155 302 151 317 159 307 154 214 128
461 231
398 259 505 252
370 241
387
636 286
97
200
82
400
272
600
216
479
800
CAPWAP
EOD Driveability
Restrike
K = 1.2
K = 2.0
Full Strength
50% Strength
50% Braced
Fully Unbraced
50% Braced
Fully Unbraced
SA
ST
SA
ST
SA
ST
Restrike
EOD
β Method
DRIVEN
Nordlund
0
SPT - Meyerhof
Nominal and Factored Resistance (kips)
1000
1070
Nominal Resistance Factored Resistance Factored Load (258 kips)
1200
642
1400
Fully Braced
Bearing Graph WEAP
Structural Capacity
Figure 3. Summary of pile resistances at Pier No. 3 (Burns South Road project). Considering the test pile as a compression member that experienced only an axial compressive load, the nominal structural capacity of the test pile (Pn) was taken as the smallest value based on the applicable modes of flexural buckling, torsional buckling and flexural-torsional buckling (AASHTO 2012). For a steel H-pile section without slender elements, the flexural buckling was considered while the torsional buckling was neglected due to a greater torsional resistance contributed from its surrounding soil. Hence, the structural capacity (Pn) was estimated as follows: Pn = [0.658
P ( o) Pe ] P o
Pn = 0.877Pe if Pe =
if
Pe ≥ 0.44 Po
Pe < 0.44 Po
π2 E KL 2 (r ) s
Ag
(1) (2) (3)
where, Ag is the gross-sectional area of a pile, Fy is the specified minimum yield strength of a steel pile, Po is the equivalent nominal yield resistance = ψFyAg, ψ is the slender element reduction factor (taken as 1.0 for a pile without slender elements), K is the effective length factor in the plane of buckling, L is the unbraced pile length in the plane of buckling, and rs is the radius of gyration about the axis normal to the plane of buckling. Among these variables, the effective pile length (KL) depends on
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the soil confinement along its length and rock fixity at its toe, which were not easily known in this study. Based on the experience gained, Tscheotarioff (1973) believed that buckling of centrally loaded vertical end-bearing piles should not be a concern as the surrounding soil or even soft clay provides adequate lateral support. Since the top of test pile was embedded 12 in (305 mm) into a concrete pilecap, the pile top-end condition was assumed to be rotation fixed and translation free. Two extreme pile toeend conditions, fixed and pinned supports, were assumed with K values of 1.20 and 2.0, respectively, as illustrated in Figure 3. The pile length was assumed to be fully unbraced (L is the total pile length) and 50% braced (L is half of the total pile length). Based on these assumptions, the structural resistances of the test pile were calculated as summarized in Figure 3. The full yield strength (FyAg) of the steel test pile and its 50% value are also included for comparison. Applying the resistance factors (ϕ) recommended in the AASHTO (2012) summarized in Table 1 to the respective nominal pile resistances (R), the factored pile resistances (ϕR) were determined. To evaluate the LRFD strength limit state (γQ ≤ ϕR), computed factored resistances were compared with the factored load (γQ) of 258 kips (1,148 kN) in Figure 3. Observations are described as follows: (1) Static analysis methods provide high variation in computed pile resistances. The DRIVEN provided the highest pile resistance while the SPT method gave the lowest pile resistance; (2) The pile performance based on the LRFD strength limit state was not satisfied when WEAP was used as the construction control method. However, it was satisfied using CAPWAP at the 1-day restrike test. Thus, piles at Pier No. 3 were accepted; (3) Apparent pile setup was observed one day after the EOD; (4) Since static load test was not performed, the following observations resulted from structural analyses were illustrated based on the best estimated nominal pile resistance of 370 kips at the EOD using CAPWAP. The test pile was certainly not fully unbraced (i.e., braced length is not equal to embedded pile length). The 370 kips (1,646 kN) determined by CAPWAP was about 20% higher than 309 kips (1,375 kN) for a 50% braced pile with a pinned support at toe and about 45% lower than that based on a fixed support. This observation implies that the pile toe condition is neither pinned nor fixed. The yield strength of the steel pile was not fully mobilized, which implied that the geotechnical strength not the structural pile strength governed the axial pile resistance. Table 1. Summary of resistance factors recommended in the AASHTO (2012) Analysis Method SPT-Meyerhof Nordlund DRIVEN β-Method CAPWAP WEAP Structural Capacity
Resistance Factor (𝛟) for β = 2.33 0.30 0.45 0.45 for cohesionless soil 0.25 0.65 0.50 0.60
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Test Pile at Abutment No. 1 Similar geotechnical and structural analyses were performed on the test pile at Abutment No. 1 driven in the 31.55 ft (9.7 m) backfill, 38.4 ft (11.7 m) silty sand layer and 2.05 ft (0.6 m) into the sandstone bedrock with a total embedded length of 72 ft (22 m). Since backfill properties were not known, the shaft resistance in the backfill layer was neglected. The nominal and factored resistances along with the factored load of 322 kips (1,432 kN) are shown in Figure 4. The equivalent measured hammer blow counts were 452 and 600 blows per ft at the EOD and the restrike, respectively. The percent shaft resistances at the EOD and the restrike estimated by CAPWAP were 20% and 36%, respectively. Similar observations were obtained in this study, except this test pile satisfied the LRFD strength limit state at the EOD based on the CAPWAP analysis. 1070 535
321 91 54
252 151 50% Braced
23 14
214 128 Fully Unbraced
424 212
428 214
207
414
422 211
457 228
533 267
550 358
505 328
440
293
110
200
92
400
306
600
253
562
800
CAPWAP
EOD Driveability
Restrike
K = 1.2
K = 2.0
Full Strength
50% Strength
50% Braced
Fully Unbraced
SA
ST
SA
ST
SA
ST
Restrike
EOD
β Method
DRIVEN
Nordlund
0
SPT - Meyerhof
Nominal and Factored Resistance (kips)
1000
652
Nominal Resistance Factored Resistance Factored Load (322 kips)
1200
642
1400
Fully Braced
Bearing Graph WEAP
Structural Capacity
Figure 4. Summary of piles resistances at Abutment No. 1 (Burns South Road project). CASE STUDY NO. 2 (CASPER STREETS) A 505-ft (154-m) three span reinforced concrete bridge was constructed over North Platte River for the Casper Streets in Natrona County, Wyoming. The construction of bridge foundation began late Fall 2012 and continued in Spring 2013. The bridge consisted of two abutments and two piers as shown in Figure 5. Fourteen, Grade 50, HP 14×73 (HP 360×109) piles were installed in a pile group at each abutment. A pile group of twenty-one, Grade 50, HP 14×89 (HP 360×132) piles was installed at each pier location. Total LRFD factored loads per pile based on a combination of dead and live loads were 169.4 kips (754 kN) and 372 kips (1,655 kN) at the abutment and pier, respectively. Mississippi Valley Equipment company (MVE) M-19 single acting diesel hammer was used to install all piles. Pile driving criterion established based on bearing graphs generated by WEAP was used to verify the performance of all piles. Pile No. 1 at Abutment No.2 was selected for dynamic tests at the EOD and 1-day restrike using a PDA with subsequent CAPWAP analyses. Based on driving logs,
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hammer blows increased by more than 40 blows per ft when sandstone was encountered. The equivalent hammer blow count was 84 blows per ft at both the EOD and restrike.
Figure 5. Plan view of Casper Streets project. The subsurface profile for the Abutment No. 1 and Pier No. 1 was poorly graded sand (SP) overlying unweathered shale bedrock. However, the subsurface profile for Pier No. 2 and Abutment No. 2 was colluvium, loose to dense, silty, pea-gravelly sand overlying unweathered sandstone bedrock. The intersection of these two bedrocks was not known based only on four available boreholes. Figure 6 shows the subsurface profile along with the SPT N-values near the test pile.
Figure 6. Subsurface profile of the test pile at Abutment No. 2 of Casper Streets project.
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Geotechnical and structural analyses were performed on the test pile driven in the 19.9 ft (6.1 m) backfill, 11.5 ft (3.5 m) silty to gravelly sand and 1.6 ft (0.5 m) into the unweathered sandstone with a total embedded length of 33 ft (10.1 m). No soil samples were collected from the borehole closest to the Abutment No. 2 for laboratory testing. However, five sandstone samples were collected for uniaxial compressive tests, and the measured qu values ranged from 459 ksf (22 MPa) to 1,888 ksf (90 MPa) as illustrated in Figure 6. The RQD values are also included in the same figure. Since backfill properties were not known, the shaft resistance of the test pile was neglected. The estimated nominal and factored resistances along with the factored load of 169.4 kips (754 kN) are summarized in Figure 7. The percent shaft resistances at the EOD and the restrike estimated by CAPWAP were 56% and 64%, respectively. Important observations are described as follows: (1) All four static analysis methods cannot satisfy the LRFD strength limit state. This is primarily due to a relatively small shaft resistance for only 11.5 ft (3.5 m) of embedment and underestimated end bearing by treating the sandstone as soil with a maximum SPT N-value of 60; (2) The test pile performance was satisfactory based on both CAPWAP and WEAP analyses; (3) CAPWAP based on the restrike test showed about 11% increase in pile resistance while no increase in hammer blow counts was recorded. (4) The WEAP bearing graph analysis provided relatively higher nominal resistances than that from CAPWAP; (5) Based on the nominal pile resistance of 340 kips estimated at the EOD using CAPWAP, the test pile was certainly not fully unbraced. Structural analysis based on the assumed 50% braced at both pile toe conditions overestimated the pile resistance. The yield strength of the steel pile was not fully mobilized, which implied that the geotechnical strength not the structural pile strength governed the axial pile resistance.
200
89 27 45 20 88 40 50 13
642
321
431
535
771
258
400
108 65
600
340 221 380 247 156 78 221 110 449 225 452 226 449 225 452 226 300 180
800
CAPWAP
EOD
Driveability
Restrike
K = 1.2
K = 2.0
Full Strength
50% Strength
50% Braced
Fully Unbraced
50% Braced
Fully Unbraced
SA
ST
SA
ST
SA
ST
Restrike
EOD
β Method
DRIVEN
Nordlund
0
SPT - Meyerhof
Nominal and Factored Resistace (kips)
1000
1070
Nominal Resistance Factored Resistance Factored Load (169.4 kips)
1200
463
1400
Fully Braced
Bearing Graph WEAP
Structural Capacity
Figure 7. Summary of pile resistances at Abutment No. 2 (Casper Streets project).
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CASE STUDY NO. 3 (TORRINGTON STREETS)
Abutment No. 2
Pier No. 3
Pier No. 2
Pier No. 1
Abutment No. 1
The Torrington Streets bridge project focuses on building a new 300-ft (91.4-m) four span reinforced concrete bridge over the Burlington Northern & Sante Fe Railroad. This overpass connected US 85 to US 26 near F Street in downtown Torrington, Wyoming. The bridge construction work began in mid-2012 and completed in the summer 2013. The bridge consisted of two abutments and three piers as shown in Figure 8. Each abutment was supported by nine, Grade 50, HP 14×73 (HP 360×109) piles while each pier was supported by a shallow foundation. Total LRFD factored load per pile at the abutment was 216 kips (961 kN). MVE M19 single acting diesel hammer was used to install all piles. Pile driving criterion established based on bearing graphs generated by WEAP was used to verify the performance of all piles. Pile No. 1 at Abutment No.2 was selected for dynamic tests using a PDA with subsequent CAPWAP analyses. The subsurface profile from the ground consisted of 21.67 ft (6.6 m) backfill, 23 ft (7 m) well graded sand (SW), 45 ft (13.7 m) poorly graded sand (SP), 35.5 ft (10.8 m) well graded gravel (GW) overlying a weathered claystone. Figure 9 shows the subsurface profile along with the measured SPT N-values near the test pile. Soil samples were collected for classification tests. No claystone samples were collected for compressive test.
Figure 8. Plan view of Torrington Streets project. Geotechnical and structural analyses were performed on the test pile initially driven to an elevation of 4017.67 ft (1224.6 m) with a total embedded length of 100 ft (30.5 m). The hammer blow counts were 68 blows per ft and 108 blows per ft at the EOD and the restrike, respectively. The percent shaft resistances at the EOD and the restrike estimated by CAPWAP were 62% and 75%, respectively. Figure 10 summarizes the nominal and factored resistances along with the factored load of 216 kips (961 kN) based on the 100 ft (30.5 m) embedded length. Important observations are described as follows: (1) Static analysis methods provided high variation in estimated pile resistances; (2) The test pile performance based on the LRFD strength limit state was not satisfied when CAPWAP was used as the construction control method. However, its performance was satisfied using WEAP at the restrike test; (3) CAPWAP showed about 13% increase in pile resistance one day after the
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EOD. The increase in hammer blow counts from 68 to 108 blow per ft validated the 23% increase in pile resistance estimated by the WEAP; (4) Based on the nominal pile resistance of 239 estimated at the EOD using CAPWAP, the test pile was neither fully braced nor the yield strength of the steel pile was fully mobilized, which implied that the geotechnical strength not the structural pile strength governed the axial pile resistance. It was further validated due to the fact that the pile was bearing on well graded gravel (GW) instead of the claystone layer at which the structural strength can be effectively mobilized. For this reason, the structural capacities estimated based on both pile toe conditions were significantly low.
Figure 9. Subsurface profile of the test pile at Abutment No. 2 of Torrington Streets project
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642
535 47 28 50% Braced
321 12 7 Fully Unbraced
33 20
226
131 78
452
465
369 184 SA
233
382 191 ST
825 413
912 270 176 Restrike
456
239 155 EOD
551 138
260
218 65
400
579
800
419
931
1200
1070
Nominal Resistance Factored Resistance Factored Load (216 kips)
1600
CAPWAP
EOD
Driveability
Restrike
K = 1.2
K = 2.0
Full Strength
50% Strength
50% Braced
Fully Unbraced
SA
ST
SA
ST
β Method
DRIVEN
Nordlund
0 SPT - Meyerhof
Nominal and Factored Resistance (kips)
513
Fully Braced
Bearing Graph WEAP
Structural Capacity
Figure 10. Summary of pile resistances at Abutment No. 2 (Torrington Streets project). DISCUSSIONS AND RECOMMENDATIONS The three case studies illustrate the current limitations and challenges on steel Hpiles in rock. A combination of limited geomaterial properties, absence of a pragmatic static analysis method and unknown pile-geomaterial interaction leads to inconsistent estimation of pile resistances. These limitations underline the importance of pile construction control methods in the pile acceptance evaluation. Table 2 summarizes the pile performance acceptances for three case studies using either WEAP or CAPWAP as the construction control method. A pile performance is accepted when the factored pile resistance is greater than the required factored load. The performance of the test pile at Pier No.3 of the Burns South Road based on WEAP, with factored resistances about 40% less than the factored load, was not accepted. When CAPWAP was used as the construction control method, its performance was barely accepted at the restrike. Likewise, the results of the test pile at Abutment No. 1 of same project gave the similar observation. It is important to recognize that the dynamic test using a PDA with CAPWAP analysis is not a proof load test. Furthermore, CAPWAP based on a routine, indeterminate signal matching technique produces non-unique pile resistance estimations that are influenced significantly by the magnitude of dynamic soil parameters and resistance distribution (Ng and Sritharan, 2013). Therefore, a static load test should be incorporated into a pile construction control program to verify its performance and calibrate the dynamic test. The static load test can be eventually eliminated when experiences and confidences accumulate.
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Table 2. Summary of pile performance acceptances Project Burns South (Pier 3) Burns South (Abutment 1) Casper (Abutment 2) Torrington (Abutment 2)
WEAP Restrike
CAPWAP Restrike
No. of Piles
EOD
21
No (-41%)
No (-40%)
No (-7%)
Yes (0.4%)
8%
5
No (-36%)
No (-34%)
Yes (2%)
Yes (11%)
9%
14
Yes (33%)
Yes (33%)
Yes (30%)
Yes (46%)
16%
9
No (-15%)
Yes (6%)
No (-28%)
No (-19%)
9%
EOD
% Gain
Yes−Satisfied the LRFD strength limit state; No−Did not satisfy the LRFD strength limit state, (%)−Percent larger (positive) or smaller (negative) than a factored load.
The test pile performance at the Casper Streets project was satisfied based on both methods while an opposite outcome was observed at the Torrington Streets project. To satisfy the LRFD limit state, all piles at Torrington were driven 2.83 ft (0.86 m) into the weathered claystone with a final embedded length of 128 ft (39 m) as indicated in Figure 9. No PDA/CAPWAP tests were performed as it was believed the LRFD strength limit state was satisfied by monitoring the hammer blow counts and using the WEAP bearing graph. Table 2 indicates that more piles were accepted when CAPWAP at one-day restrike was used as the construction control. This observation reveals the appropriateness of current WYDOT’s practice. Based on the author’s experience on pile setup (Ng et al., 2013), it is believed that the average 10% percent gain in pile resistances was primarily attributed to the increase in pile shaft resistance caused by the healing and quicker consolidation of the surrounding silty sand materials. However, this 10% percent gain was not as high as 50% to 70% observed by Ng et al. (2013) on steel H-piles driven in low plasticity clay (CL). Little or no increase in end bearing was observed. To facilitate the resistance estimation of piles driven into rock materials, it is indispensable to evaluate the pile-geomaterial interaction in terms of the amount of pile bracing from its surrounding soil. Matching the nominal resistance estimated from CAPWAP at the EOD to its structural compressive capacity calculated using Equations (1) through (3), the required percent pile bracing, a ratio of braced length to total embedded pile length in percentage, was determined for each test pile for both pile toe conditions, pinned and fixed supports. A relationship of percent bracing and embedded pile length was established in Figure 11 for steel H-piles driven in primary silty sand and into rock materials. The rationale of matching the resistances assumes that the pile resistance will be governed by its structural strength, although it is not always the case, while the geotechnical strength will be indirectly accounted for in terms of percent pile bracing. The authors believe that this approach will improve the current pile resistance estimation and alleviate the discrepancy between estimated and measured resistances. Recognizing limited data were available for this study, the following observations provide a basis for future investigations and should be further validated when more pile data become available: (1) The percent bracing increases with increasing embedded pile length up to
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about 72 ft (22 m) and decreases thereafter. This implies that the overall contribution of its soil confinement to the structural pile analysis decreases for an embedded pile length greater than 72 ft (22 m); (2) The required percent bracing based on a fixed support is about 14% larger than that of a pinned support at the embedded pile length of 72 ft (22 m). However, the percent difference increases to about 38% with a shorter embedded pile length; and (3) The maximum possible percent bracing increases from 65% for a pinned support to 80% for a fixed support.
Precented Bracing
100%
Pinned Support (K=2.0) Fixed Support (K=1.2)
80%
60% 40% 20% 0% 0
20
40 60 80 100 Embedded Pile Length (ft)
120
Figure 11. A relationship of percent bracing and embedded pile length The limitations and challenges highlighted in this paper suggest the immediate need of research on piles driven in rock or very stiff geomaterials. Research should produce products addressing the limitations of current design and construction control of these piles. Some desirable products are suggested as follows: (1) A database of rock properties and load tests on piles driven in rock; (2) A static method to estimate the geotechnical resistance of a pile in rock; (3) A guideline for evaluating the structural pile capacity with considering the different surrounding soil confinements and pile toe-rock conditions; (4) A locally calibrated LRFD resistance factors; and (5) Adequate design and construction control recommendations in the AASHTO LRFD Bridge Design Specifications. CONCLUSIONS Motivated by the challenges faced by the WYDOT on steel H-piles driven in rock materials, detailed pile analyses of three case studies were performed to illustrate the limitations of current pile design and construction control procedures. It is envisioned that this paper will facilitate more research on piles driven in rock. From the outcomes of this study, the following conclusions were drawn: (1) Current static analysis methods provide inconsistent and potentially unconservative geotechnical resistance estimations of a driven pile in rock. DRIVEN and Nordlund method were likely to overestimate the pile resistance while the SPT method tended to underestimate the pile resistance;
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(2) Construction control using WEAP will produce a higher uncertainty than that based on CAPWAP. This finding agrees with the relatively lower resistance factors for WEAP than that for CAPWAP promulgated in the AASHTO LRFD Bridge Design Specifications; (3) Apparent pile setup was observed one day after the EOD, and most pile performances were satisfied during the restrike tests. Thus, it is recognized as a good practice by most agencies to include restrike tests in a pile construction control program; (4) Piles driven in overburden soil and into rock were neither fully unbraced nor braced. To facilitate the resistance estimation of piles driven into rock materials by assuming its structural strength governs the design while indirectly accounting for the geotechnical strength in terms of percent pile bracing, a relationship between percent bracing and embedded pile length was established. Percent pile bracing increases with increasing embedded pile lengths up to 72 ft (22 m) and decreases thereafter. The maximum percent bracing increases from 65% for a pinned support to 80% for a fixed support. Percent bracing for a fixed support is always higher than that for a pinned support. It is important to note that these observations were concluded based on limited pile data. However, it serves as a basis for future investigations and should be further validated when more pile data become available; and (5) Yield strength of the steel pile was not fully mobilized due to the constraint of its geotechnical resistance. ACKNOWLEDGEMENTS We express our gratitude to the Wyoming Department of Transportation, the Geology Program, for providing us valuable data for this study. REFERENCES American Association of State Highway and Transportation Officials. (2012). AASHTO LRFD Bridge Design Specifications, Sixth Edition, Washington, D.C. American Association of State Highway and Transportation Officials (AASHTO) M 145. (2012). Standard specification for classification of soils and soil-aggregate mixtures for highway construction purposes. AASHTO, Washington, D.C. American Society for Testing and Materials (ASTM) D2487. (2011). Standard Practice for Classification of Soils for Engineering Purposes. ASTM International, West Conshohocken, PA. Burland, J.B. (1973). Shaft friction of piles in clay. Ground Engineering, London, Vol. 6, No.3, 3042 p. Hannigan, P.J., Goble, G.G, Likins, G.E., and Rausche, F. (2006). “Design and construction of driven pile foundations-volume I.” FHWA-NHI-05-042. Final Report, Federal Highway Administration, Washington, D.C. Mathias D., and Cribbs, M. (1998). “DRIVEN 1.0: A Microsoft Windows based program for determining ultimate static pile capacity." Federal Highway Administration, Washington, D.C. Meyerhof, G. (1976). “Bearing capacity and settlement of pile foundations.” American Society of Civil Engineers, Journal of the Geotechnical Engineering
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Division, Vol. 102 (3): 195–228. Ng, K.W., Roling, M., AbdelSalam, S.S., Suleiman, M.T., and Sritharan, S. (2013). “Pile setup in cohesive soil. I: experimental investigation.” Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 139(2): 199-209. Ng, K.W., and Sritharan, S. (2013). “Improving dynamic soil parameters and advancing the pile signal matching technique.” Computers and Geotechnics, Elsevier, 54: 166-174. Nordlund, R.L. (1963). “Bearing capacity of piles in cohesionless soils.” Journal of Soil Mechanics and Foundation Engineering, Soil Mechanics and Foundation Division, ASCE, Vol. 89, SM 3: 1-36. Nottingham, L., and Schmertmann, J. (1975). An Investigation of Pile Capacity Design Procedures. Final Report D629 to Florida Department of Transportation, Department of Civil Engineering, University of Florida, 159 p. Serrano, A., and Olalla, C. (2002). “Ultimate bearing capacity at the tip of a pile in rock-part 1: theory.” International Journal of Rock Mechanics & Mining Sciences, 39: 833-846. Thompson, C.D. and Thompson, D.E. (1985). “Real and apparent relaxation of driven piles.” American Society of Civil Engineers, Journal of Geotechnical Engineering, Vol. 111 (2): 225-237. Tomlinson, M.J. (1980). Foundation design and construction. 4th Edition. Pitman Publishing Limited, 128 Long Acre, London WC2E 9AN, UK. Tschebotarioff, G.P. (1973). Foundations, Retaining, and Earth Structures. 2nd Edition, McGraw-Hill. Wyoming Department of Transportation. (2012). “Standard specifications for road and bridge construction.” 2010 Edition, Cheyenne, Wyoming. Wyoming Department of Transportation. (2014). “Construction manual.” 2014 Edition, Cheyenne, Wyoming.
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