Behavior of Driven Ultrahigh-Performance Concrete H ... - Sri Sritharan

Behavior of Driven Ultrahigh-Performance Concrete H-Piles Subjected to Vertical and Lateral Loadings Muhannad T. Suleiman, A.M.ASCE1; Thomas Vande Voort2; and Sri Sritharan, M.ASCE3 Abstract: In the United States, an estimated $1 billion is spent annually on repair and replacement of deep foundations. In a recent study, the possibility of using ultrahigh-performance concrete 共UHPC兲 for deep foundation applications was explored with the objectives of increasing the service life of deep foundations supporting bridges to 75 years and reducing maintenance costs. This paper focuses on field evaluation of two UHPC piles and references a steel H-pile. An UHPC pile with an H shape was designed to simplify the process of casting the pile and reduce the volume 共i.e., cost兲 of the material needed to cast the pile. Two instrumented UHPC piles were driven in loess on top of a glacial till clay soil and load tested under vertical and lateral loads. This paper provides a complete set of results for the field investigation conducted on UHPC H-shaped piles. The results presented in this paper prove that the designed UHPC piles can be driven using the same equipment used to drive steel H-piles through hard soil layers without a pile cushion. The vertical load capacity of the UHPC pile was over 80% higher than that of the steel H-piles. DOI: 10.1061/共ASCE兲GT.1943-5606.0000350 CE Database subject headings: Deep foundations; Driven piles; Load tests; Concrete. Author keywords: Deep foundation; Driven piles; UHPC; Load tests.

Introduction In 2005, the American Association of State Highway and Transportation Officials issued a strategic plan for bridge engineering, which included seven grand challenges. The first two challenges were extending service life and optimizing structural systems. One of the greatest challenges to achieving the 75-year service life for foundations with minimal maintenance is material deterioration, which is targeted for future typical bridges according to American Association of State Transportation and Highway Officials 共2005兲. The most common pile foundation systems used to support bridges and buildings are made of steel or concrete. Both of these pile materials have limitations, especially related to durability and drivability. Concrete piles 共precast or prestressed-precast兲 are susceptible to cracking during easy driving due to development of large tensile stresses—a main reason why pile cushion is used during driving. Concrete piles may also fail in compression during driving in hard soil due to development of excessive compres1

Assistant Professor, Dept. of Civil and Environmental Engineering, Lehigh Univ., 326 STEPS Building, Bethlehem, PA 18015; formerly, Assistant Professor, Dept. of Civil and Environmental Engineering, Lafayette College, 321 Acopian Engineering Center, Easton, PA 18042 共corresponding author兲. E-mail: [email protected] 2 Formerly, Research Assistant, Dept. of Civil, Construction and Environmental Engineering, Iowa State Univ., 176 Town Engineering, Ames, IA 50011-3232. E-mail: [email protected] 3 Wilson Engineering Associate Professor, Dept. of Civil, Construction and Environmental Engineering, Iowa State Univ., 406 Town Engineering, Ames, IA 50011-3232. E-mail: [email protected] Note. This manuscript was submitted on August 2, 2009; approved on March 2, 2010; published online on March 4, 2010. Discussion period open until March 1, 2011; separate discussions must be submitted for individual papers. This paper is part of the Journal of Geotechnical and Geoenvironmental Engineering, Vol. 136, No. 10, October 1, 2010. ©ASCE, ISSN 1090-0241/2010/10-1403–1413/$25.00.

sive stresses or due to the use of large driving hammers 共Salgado 2008兲. Cracked concrete intensifies corrosion of steel reinforcement, leading to eventual spalling and deterioration of the concrete pile, which reduces its axial and flexural capacities. Steel piles are vulnerable to local buckling during hard driving conditions and experience corrosion, which often occurs near the region that has the largest stresses due to combined axial load and flexural moment 共Huck and Hull 1971; Mekkawy 2004; Vande Voort et al. 2008兲. The maintenance challenges associated with conventional piles are emphasized by Pando et al. 共2006兲, who reported that the United States spends an estimated $1 billion annually on maintenance and replacement of pile foundations. To achieve the target service life, minimize drivability challenges, and ensure durability, the use of pile foundation elements made of ultrahigh-performance concrete 共UHPC兲, with its high compressive and tensile strengths and excellent durability, was investigated. An H-shaped UHPC section, with dimensions and weight similar to that of a commonly used HP 10⫻ 57 共depth = 254 mm and weight= 86 kg/ m兲 steel pile, was designed. Instrumented UHPC full-scale piles were driven near a bridge construction site with a soil profile of loess on top of a glacial till clay 共a common soil profile in the Midwestern United States兲 and tested under vertical and lateral loads. Furthermore, laboratory 3 tests were conducted on 2.4-m-long specimens scaled to 4 of the cross section of the piles used in the field, and drivability studies were also performed. In addition to summarizing the properties and drivability of precast-prestressed normal concrete 共NC兲 and steel H-piles, Vande Voort et al. 共2008, 2009兲 summarized the properties and durability of the UHPC, the section design of the UHPC piles, and the theoretical drivability analysis results of the pile in several soil conditions using different driving equipment. The drivability analysis showed that stresses induced in the UHPC pile during driving are not significantly higher than those in NC piles, and the ratio of the calculated stresses to the permissible stress limits was significantly lower for the UHPC than for

JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGINEERING © ASCE / OCTOBER 2010 / 1403

Downloaded 01 Oct 2010 to 129.186.167.15. Redistribution subject to ASCE license or copyright. Visithttp://www.ascelibrary.org

Fig. 1. Cross-sectional details of the UHPC pile compared with the steel HP 10⫻ 57 pile

Properties, Design, and Production of UHPC Piles

steel HP 10⫻ 57 pile, which is commonly used by many state department of transportation to support bridges; both piles have similar outer dimensions and weight per unit length. The design process of the UHPC pile cross section included developing the interaction diagram, which describes the axial load capacity of a beam-column element as a function of bending moment for the UHPC pile compared with the steel HP 10⫻ 57, and the momentcurvature relationship of several possible sections as well as a drivability study considering different soil conditions. Vande Voort et al. 共2008兲 presented more details on the design of the UHPC pile cross section.

Material Properties

Production of UHPC Piles

The compressive strength of the UHPC ranges from 180 to 220 MPa, while its direct tensile strength ranges from 8.4 to 9.3 MPa 共Tuchlinski et al. 2006; Graybeal 2006兲. When compared to NC, UHPC shows more than five times the compressive strength, four times the tensile strength, and two times the elastic modulus. Increases in the compressive and tensile strengths of concrete pile foundations allow contractors to use larger hammers during construction without exceeding the tensile and compressive strengths of concrete, which also results in higher penetration rates and reduced construction time. The very low porosity of the UHPC improves its durability compared to NC and high-performance concrete 共HPC兲. The results summarized by Vande Voort et al. 共2008兲 show that UHPC has a chloride ion penetration resistance that is 28 times smaller than that of the NC and a reinforcement corrosion rate that is 120 times slower than that of the NC. These favorable properties will reduce the required maintenance cost and extend the life of foundation systems.

Two identical 10.7-m-long UHPC piles were cast at Coreslab in Omaha, Neb. Wooden forms with Styrofoam inserts were used for casting the two piles. The test units were instrumented with vibrating wire strain gauges 共sister bars兲 and normal strain gauges. Then, the prestressing strands were stressed, and the UHPC was mixed and poured into the forms. The cast UHPC piles were covered with plastic to prevent moisture loss. The UHPC members were initially cured at 30° C until they reached the release strength of 97 MPa. The strands were then released, and the UHPC piles were steam cured for 48 h at 90° C. The measured compressive strength of the UHPC material used to construct the UHPC pile test units ranged from 182 to 199 MPa on the day of full-scale testing.

the NC, and was comparable to those obtained for steel H-piles. The analyses further showed that UHPC may be driven with no pile cushion in most situations with a wide range of hammers and strokes. This paper focuses on the behavior of instrumented UHPC piles tested in the field and presents the experimental results obtained during driving as well as vertical and lateral load test behaviors compared with theoretical responses and with the behavior of a steel H-pile subjected to vertical loads.

Section Design A tapered H-shaped prestressed-precast section was designed to avoid the risk of forming air pockets during production. The UHPC pile section was reinforced with ten 13-mm-diameter prestressing strands with no mild steel or shear reinforcement. The high strength and durability of the UHPC allowed a reduction in the concrete cover required on the prestressing strands from 32 to 19 mm. A 51-mm center-to-center spacing between the prestressing strands was maintained to allow the UHPC to flow freely through the section during casting. The final UHPC section is a 254⫻ 254-mm tapered H-shaped section shown in Fig. 1 next to a

Instrumentation of Test Units The first test unit 共UHPC Pile 1兲 was instrumented with 3 pairs of sister bars and 12 normal strain gauges, while the second unit 共UHPC Pile 2兲 was instrumented with 11 normal strain gauges. The locations of strain gauges and sister bars along the length of the pile and on the pile cross section are shown in Figs. 2共a–d兲. For UHPC Pile 1, the locations of the sister bars and strain gauges were chosen to obtain estimates of the skin friction along the length of the pile and to the end bearing at the tip of the pile during vertical axial load testing 关see Figs. 2共e and f兲兴. These gauges were also used to determine the moment distribution along the length of the pile when subjected to lateral loading. For UHPC Pile 2, most of the strain gauges were installed in the top 3 m of the pile, in which the location of the maximum moment was predicated by LPILE 共Reese et al. 2000兲 when subjected to

1404 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGINEERING © ASCE / OCTOBER 2010


203

280

890 7, 8 13-15 635 9-12

927 2298 610 445 470

890 1, 2 16-17

1, 2 3-6

Strand 2 Sister bars 1, 3, 5

Strand 2 Gauge 9

Strand 4 Gauge 7,10,13,16

7-9

10 11,12

Strand 1 Sister bars 2, 4, 6

Strand 5 Gauges 8,11,14


3683

(b)

(c) Strand 1 Gauge 3

3, 4

Strand 2 Gauge 1,4,7,10,11

3543

Strain Gauge Sister Bars (Pairs)

5, 6

UHPC 1

UHPC 2



(d)

(a)

Fig. 2. Instrumentation used for test units UHPC Pile 1 and UHPC Pile 2: 共a兲 locations of various gauges along test units measured from the top of the pile with all dimensions in mm; 共b兲 strands instrumented with sister bars for UHPC Pile 1; 共c兲 instrumented prestressing strands on UHPC Pile 1 and 共d兲 instrumented prestressing strands on UHPC Pile 2; 共e兲 sister bars used in UHPC Pile 1; and 共f兲 strain gauges used on strands of UHPC Pile 1

lateral loads. Threaded rods were also installed near the top of UHPC Pile 2 to mount the acceleration and strain gauges of the pile driving analyzer 共PDA兲 during pile driving.

Site Characterization The two 10.7-m-long UHPC piles and steel HP 10⫻ 57 pile were installed next to a bridge site near Oskaloosa, Iowa. The bridge at this site is a three-span bridge carrying future expansion of northbound U.S. 63 across Union Pacific railroad. The soil at the bridge was initially characterized using standard penetration tests 共SPTs兲 conducted by the Iowa Department of Transportation 共2007兲. In the field, the research team characterized the soil at the site using the SPT and piezocone penetration tests 共CPTs兲. Disturbed and relatively undisturbed soil samples were also collected using Shelby tubes to characterize the soil in the laboratory. The soil at the Oskaloosa bridge site consists of 4.8 m of Wisconsinan loess overlaying a pre-Illinoian hard glacial till, which is a well-graded mixture of clay, silt, and sand with occasional gravel and boulders. The soil profile of loess on top of a glacial till is very common in the Midwestern region of the United States. Using laboratory tests and CPT results, the Wisconsinan loess soil was classified into three soil layers consisting of 1.5 m of a low plastic silt 共ML兲 overlaying a 1.2 m of a low plastic clay 共CL兲 and 2.1 m of a low plastic clay with sand 共CL兲. The pre-Illinoian hard glacial till consists of layers of clayey sand, sandy clay, and well-graded sand with gravel 共see Fig. 3兲. The results of two CPTs, which were conducted at distances of 1.2 and 12 m from UHPC Pile 2, indicate a 0.9-m sand layer at a depth of 4.65 m, with an average CPT tip resistance of 15.6 MPa and 1.35 m deep of a very stiff well-graded silty sand layer at a depth of 7.6 m with an average CPT tip resistance of 26.7 MPa. The measured pore-water pressure during the CPTs indicated that the water table was located at approximately 3 m below the ground surface. The CPTs were terminated at approximately 11 m due to refusal indicating large inclusions at this depth. The SPT

was conducted at approximately 75 m from the CPT located at 12 m from UHPC Pile 2 and showed hard layers at depths of 7.6 m 共N ⬎ 100兲 and 11 m 共N ⬃ 75兲, respectively. Observations during driving confirmed the locations of the hard layers at 4.65 and 7.6 m, as will be discussed in a later section. Fig. 3共d兲 also shows the estimated undrained shear strength and friction angle of soil layers using the CPT-based empirical relationships recommended by Lunne et al. 共1997兲. The average undrained shear strength of the loess soil in the top 2.7 m was 60 kPa underlain by a sandy clay having an undrained shear strength of 136 kPa. The friction angles of the sandy soils ranged from 34° to 43° with estimated relative densities varying between 77 and 87%. Fig. 3共d兲 also included the strain at 50% of the strength for cohesive soil and the modulus of subgrade reaction for cohesionless soil based on values recommended by Reese et al. 共2000兲.

Predicted Pile Capacity Using the soil properties presented in Fig. 3 and pile dimensions, static analysis methods were used to estimate the capacity of the UHPC and steel piles. Because the soil profile at the site consisted of mixed soils 共clay and sand兲, static analysis methods developed for both materials were used to provide estimates of pile capacity. Assuming no soil plugging, Table 1 summarizes the predicted pile capacity using SPT-Meyerhof 共Meyerhof 1976兲, ␣ method 共Tomlinson 1986兲, ␤ method 共Esrig and Kirby 1979兲, Nordlund 共1963兲 method, and Iowa Blue Book method 共Dirks and Kam 1989兲, which is an in-house method developed by the Iowa Department of Transportation 共2007兲 for the local soil conditions. The estimated capacity of the UHPC pile ranged from 725 to 1,299 kN while the estimated capacity of the steel pile ranged from 479 to 755 kN. The difference in calculated pile capacity depends mainly on estimated soil properties and assumptions made during the development of several static analysis methods. For example, the Nordlund 共1963兲 method accounts for the volume of the pile and the interface between soil and pile material 共i.e., steel or con-



Tip resistance (qT in MPa) -1

1 2 3

Depth (m)

4 5 6 7 8 9 10 11 12

Low plastic silt  = 18.8 kN/m3 w =21.2% (ML) LL = 42.1%, PI = 10.4%

0 1

3

Low plastic clay  = 18.5 kN/m  w =7.6% (CL) LL = 44.4%, PI = 17.9%

Clayey sand  = 20.5 kN/m3 w =4.33% (SC) LL = 32.5%, PI = 17.7% 3

Low plastic clay  = 20.4 kN/m  w =4.83% LL = 36.7%, PI = 19.2% with sand N (SPT) = 7 (CL) Well graded sand (SW) Low plastic clay (CL) Well graded sand with gravel (SW)

 = 20.6 kN/m

Ground surface UHPC H-Pile

Water table

2

3

Low plastic clay  = 18.5 kN/m  w =12.7% with sand LL = 27.9%, PI = 7.4% (CL) N (SPT) = 8

3

3

4

Loess

4

6

Undrained shear strength, Su (kPa) 8

10

12

0

Glacial Till

0

400

Ground surface

0

1

1

2

2

3

3

4

4

Water table

6

Su  Su = 60 kPa 50 = 0.007 Su = 136 kPa 50 = 0.005  = 41o k = 33.9 kN/m3

5 6

7

7

7

8

8

8

 = 35o k = 6.7 kN/m3  = 42o k = 33.9 kN/m3

N (SPT) >100  = 20.4 kN/m  = 20.4 kN/m

3

3

N (SPT) ~ 75

800 1200 1600 2000

-1

5

5 6

2

-1

-1

Properties

Depth (m)

0

Classification

0

10 20 30 40 50 60 70

Depth (m)

0

Fr (%)

9

9

9

10

10

10

11

11

11

12

12

12

Su = 800 kPa 50 = 0.004

0

10

20

30

40

50

Friction angle,  (Degrees)

(a)

(b)

(c)

(d)

Fig. 3. Properties of soil at the test site: 共a兲 soil profile; 共b兲 CPTu tip resistance; 共c兲 CPTu friction ratio; and 共d兲 estimated undrained shear strength for cohesive soils and friction angle for cohesionless soils including the values of the undrained shear strength 共Su兲 and the strain at 50% of the strength 共␧50兲 for cohesive soils and the friction angle 共␾兲 and the modulus of subgrade reaction 共k兲 used to generate the p-y curves for the LPILE analysis investigating the response of the UHPC pile subjected to lateral loads

crete兲, while the ␣ method assumes a total stress analysis 共i.e., short-term capacity兲, and the ␤ method uses effective stress analysis 共i.e., long-term capacity兲. The vertical load tests on UHPC and steel piles were planned to be conducted within 1 week after the end of pile driving and assembly of the loading

frame. However, due to very low cold temperatures and significant snow accumulation experienced in Iowa, the vertical load tests were conducted approximately 3 months after driving. Based on this, it was expected that the ␤ method will provide the best estimate of the load test results.

Table 1. Estimated Capacities 共in kN兲 of UHPC and Steel Piles Using Static Analysis Methods Compared to Static Load Test Results UHPCa Capacity

Method

Skin friction

Steelb End bearing

Capacity

Skin friction

End bearing

c

SPT-Meyerhof 725 370 355 479 374 105 794 715 79 746 714 32 ␣ methodd 1,299 519 780 755 525 230 ␤ methode 897 302 595 642 466 176 Nordlundf 932 380 552 618 320 298 Iowa Blue Bookg 925 259 666 641 545 96 PDAh 894 384 510 578 434 144 CAPWAPh Static load test 1,640 1,053 587 881 NA NA a Cross-sectional area equals to 36, 600 mm2 and cross-sectional surface length equals to 1,016 mm. b Cross-sectional area equals to 10, 800 mm2 and cross-sectional surface length equals to 1,028 mm. c After Meyerhof 共1976兲 and based on SPT corrected N values. d Adhesion factor 共␣兲 after the American Petroleum Institute 共API兲共1989兲; the Su calculated based on the SPT corrected N values using the empirical correlation 关0.29N0.72兴 after Hara et al. 共1974兲. e ␤ coefficient after Burland 共1973兲; the ␾ calculated based on SPT corrected N values using the empirical correlation 关54− 27.6 exp共−0.014N160兲兴 after Peck et al. 共1974兲. f After Nordlund 共1963兲; using ␾ 共same ␾ as in ␤ method兲. g Iowa Blue Book 共after Dirks and Kam 1989兲; mainly use a combination of the SPT-Meyerhof method 共Meyerhof 1976兲 and ␣-Tomlinson method 共Tomlinson 1986兲. h Using the procedures summarized in Hannigan et al. 共1997兲. 1406 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGINEERING © ASCE / OCTOBER 2010


Cumulative Blows 0

50

100

150

200

250

300

0 1

Blows/0.3m of Steel HP 10x57 Blows/0.3 m of UHPC pile 2 Cum blows for steel HP 10x57 Cum blows for UHPC pile 2

Fig. 4. Disintegration of pile cushion as the UHPC pile penetrates the hard soil layer at a depth of 7.9 m

Penetration Depth (m)

2 3 4 5 6 7 8 9 10

Driving of Test Piles

11

The anchor piles were driven first followed by the steel HP 10 ⫻ 57 pile, UHPC Pile 1, and UHPC Pile 2 in December 2007 using a Delmag D19-42 hammer with a 51-mm-thick hammer cushion made of aluminum and micarta. The PDA was used to measure strains and accelerometers in UHPC Pile 2 and steel pile during driving. Although theoretical calculations using the one-dimensional wave equation analysis program 共GRLWEAP兲 indicated that the UHPC piles could be driven with no pile cushion 共Vande Voort et al. 2009兲, the contractor requested the use of plywood cushion during driving mainly to follow the general practice used for driving precast-prestressed concrete piles. UHPC Pile 2 was driven first with a 57-mm-thick plywood pile cushion. The cushion disintegrated with approximately 1.5 m of driving remaining 关see Figs. 4共a and b兲兴. UHPC Pile 1 was then driven with 95 mm of cushion, which also disintegrated with about 0.6 m of driving remaining. After pile cushions were damaged, both UHPC test units were driven with no pile cushion. The depth of the test units when driven with no pile cushion corresponded with the depth of the hard soil layer starting at a depth of about 7.6 m. After the end of driving, the top of both UHPC piles were examined carefully and found no visible signs of damage. The UHPC piles and the steel pile were driven to a penetration depth of 9.9 m. The driving log that included the number of blows required to penetrate each 0.3-m depth during driving and the cumulative number of blows are shown in Fig. 5 for UHPC Pile 2 and steel HP 10⫻ 57 piles. The figure shows that the number of blows required per 0.3-m depth penetration increases at a depth of approximately 7.6 m and, to a lesser extent, at 4.65 m, both of which correspond to the hard layers indicated by the CPT test results. A total of 275 blows were required to drive UHPC Pile 2 to its final depth, while the steel pile required 175 total blows. Although the UHPC pile required more blows to reach the target depth, because of their higher compressive and tensile strengths, UHPC piles can be driven with larger hammers and strokes reducing the required total number of blows and improving the efficiency of the driving process especially when compared to precast-prestressed concrete piles 共see Vande Voort et al. 2009兲. PDA strain gauges and accelerometers were used to monitor UHPC Pile 2 and steel pile test units during driving. The PDA results near the end of driving show typical response of force and velocity with the two curves diverging from each other because of skin friction provided by the soil 关see Figs. 6共a and b兲兴. The significant increase in velocity and decrease in force at time of 2L / C for steel pile, where L is embedded pile length and C is the wave velocity in the pile, indicate a small end bearing resistance. Furthermore, the PDA results, with velocity smaller than the force and no intersection between the force and velocity curves along

12 0

5

10

15

20

25

Number of Blows/ 0.3m penetration

Fig. 5. Driving logs for test units UHPC Pile 2 and HP 10⫻ 57 showing blows/0.3 m and cumulative blow count

the length of the pile, indicate that neither pile was damaged during driving. PDA results at the end of driving indicated that the UHPC pile capacity was 925 kN with 259 kN 共i.e., 28%兲 of the load resisted by skin friction and 666 kN 共72%兲 by end bearing, while the steel pile capacity was predicted to be 641 kN with 545 kN 共85%兲 of the load resisted by skin friction and 96 kN 共15%兲 by end bearing. The results of CAse Pile Wave Analysis Program 共CAPWAP兲 analysis at the end of driving show that the UHPC pile capacity of 894 kN with 384 kN 共43%兲 resisted by skin friction and 510 kN 共57%兲 by end bearing. For the steel pile, CAPWAP analysis indicated a capacity of 578 kN with 434 kN 共75%兲 carried by skin friction and 144 kN 共25%兲 by end bearing. The results of fully instrumented steel H-piles driven in a similar soil profile in Iowa with PDA data collected at the end of driving and during several restrikes show that the CAPWAP results of the latest restrike correlate well with the static vertical load test. Ng et al. 共2010兲 concluded that a pile capacity difference between static load test results and CAPWAP results was less than 5% when conducted within 3 days of the latest restrike. Static analysis using the Nordlund 共1963兲 method, which accounts for the effects of soil-pile material interface and volume of driven pile, resulted in 33.7 and 72.6% of the load resisted by skin friction for UHPC and steel pile, respectively, which are consistent with estimates provided by PDA 共28 and 85%兲 measurements and CAPWAP analysis 共43 and 75%兲.

Testing Plan and Setup The testing plan at the site consisted of conducting two vertical static load tests on UHPC Pile 1 and the steel HP 10⫻ 57 piles using the same loading frame system, as shown in Fig. 7共a and b兲. Following the vertical load tests, UHPC Pile 1 and UHPC Pile 2 were simultaneously subjected to lateral load testing using a hydraulic jack placed between the two units.

Vertical Load Tests Test Procedure Due to very low cold temperatures and significant snow accumulation on the ground experienced in Iowa during the winter of



500 kips

17.0 f/s

500 kips

16.7 f/s

F

V

F

V

51.2ms

51.2ms 4.00 ms

4.20 ms

(a)

(b)

Fig. 6. PDA force and velocity curves at the end of driving for 共a兲 UHPC Pile 2; 共b兲 steel pile

2007–2008, the static load test was conducted in March 2008, approximately 93 days for UHPC and 102 days for steel pile after driving. Vertical load tests were conducted on UHPC Pile 1 and steel HP 10⫻ 57 using the setup shown in Fig. 7. Given that the predicted vertical load capacity of the UHPC pile was larger than that of the steel pile, and it ranged from 725 to 1,299 kN with an average estimated capacity of 930 kN, the load frame was designed to sustain a maximum vertical load of 1,668 kN 共1.8 times the predicted capacity兲. A hydraulic jack with 1,780-kN capacity and a load cell of 1,335-kN capacity were used. In addition to monitoring applied axial load and the strains along the length, the vertical settlement of UHPC Pile 1 was measured using four displacement transducers. For the steel HP 10⫻ 57, only the applied axial load and the vertical settlement were monitored. The tests were conducted in accordance with the quick test procedure out-

Reaction Pile

Reaction Pile

Reaction Beam

1.98 m

1.98 m

1.52 m Unit UHPC 1

Steel Pile

1.22 m Reaction Pile

Reaction Pile

Unit UHPC 2

(a)

Steel H Pile

UHPC 1 UHPC 2

(b)

Fig. 7. Setup used field testing of UHPC Pile 1, UHPC Pile 2, and HP 10⫻ 57 steel pile: 共a兲 plan view; 共b兲 elevation view

lined in ASTM D1143/D1143M-07 共ASTM 2007a兲. For UHPC Pile 1, load increments of 31 kN 关i.e., 5% of the predicted pile capacity using initial soil testing conducted by the Iowa Department of Transportation 共2007兲兴 were used and sustained between 4 and 8 min to allow pile settlement to stabilize. After exceeding the predicted pile capacity, the load steps were increased to 111 kN. For loads greater than 890 kN, the measurements of strain gauges and sister bars were not recorded. For the steel H-pile, load increments of 27 kN, which is also 5% of the predicted pile capacity, were used.

Pile Load Capacity The load-displacement responses of UHPC Pile 1 and steel HP 10⫻ 57 pile obtained from the vertical load tests are shown in Fig. 8. The Davisson 共1972兲 failure criterion, which is shown by the dashed line in Figs. 8共a and b兲, was used to define the capacity of both piles. Although several methods could be used to determine the pile capacity using the measured pile response during a load test, the Davisson 共1972兲 criterion was used because it is the major pile capacity determination method used in the load and resistance factor design of deep foundations. Since the study focuses on comparison of pile capacities, the choice of using Davisson 共1972兲 criterion has less impact on the outcomes of the study. The load test on the UHPC pile had to be stopped at a load of 1,335 kN, which was the capacity of the load cell used for this test. At this load, the Davisson 共1972兲 criterion was not reached for UHPC Pile 1. To estimate the UHPC Pile 1 capacity, the procedure recommended by Paikowsky and Tolosko 共1999兲 was used to extrapolate the load-displacement curve, resulting in a Davisson 共1972兲 pile capacity of 1,640 kN. The steel pile experienced failure during the vertical load test as the pile vertical displacement increased significantly without any significant increase in the applied load, as shown in Fig. 8共b兲. The steel pile capacity estimated using the Davisson 共1972兲 criterion was 881 kN. Given that the static load tests were conducted about 3 months after driving the piles, the ␤ method, which is based on effective stress analysis, provided the best prediction of the UHPC and steel pile capacities. Furthermore, the ␤ method also closely predicted the ratio of the UHPC pile capacity to the steel pile capacity, which was 1.72 compared to the ratio based on static load test results, which was 1.86. The ratio of the UHPC pile capacity to the steel pile capacity is not only a function of the ratio of crosssectional area 共which was 3.4兲 but also a function of the interface properties between the soil and the material of the piles. In comparison, the Nordlund 共1963兲 method, which accounts for the pile



Load (kN) 0

Load (kN)

200 400 600 800 1000 1200 1400 1600 1800

0

0

4 Extrapolation

12

Dav isso nF

ailu re C rite rion

16

Displacement (mm)

Displacement (mm)

4

8

200 400 600 800 1000 1200 1400 1600 1800

0

8

Da vis s Dav isso n Fa ilure Crit er

12

on Pil eC ap ac

ity =

88 1k

N

ion

16 Davisson Pile Capacity = 1640 kN

20

20

(a)

(b)

Fig. 8. Vertical load-vertical displacement and Davisson 共1972兲 failure criterion for UHPC and steel test piles from static load test and CAPWAP analyses

material in estimating skin friction, resulted in a skin friction of the steel pile that is 1.54 times the skin friction of the UHPC pile, although both piles have similar surface areas. Neglecting the effects of pile driving and those of larger soil movement resulting from driving of the UHPC pile on its load bearing capacity, the UPHC pile provided larger end bearing capacity than the steel pile due to 3.4 times larger cross-sectional area. When this increase in capacity combined with a reduction in the skin friction component expected for concrete piles 共e.g., Nordlund 共1963兲 method estimates 1.54 times smaller skin friction for concrete piles than for steel H-piles兲, using the ␤ method, the UHPC pile capacity is expected to be about 1.72 times that of the steel H-pile, which is close to the ratio of 1.86 from the measured UHPC pile capacity to the steel pile capacity. Although ongoing research at Iowa State University on steel H-piles shows a difference between the CAPWAP predicted capacity and static load test results less than 5% when conducted within 3 days 共Ng et al. 2010兲, significantly larger differences were observed between CAPWAP estimated pile capacities and those determined from the static load tests in this study. These differences were 83% for the UHPC pile and 50% for the steel pile, which was attributed to pile setup. Using PDA measurements and CAPWAP analyses conducted at the end of driving and at several restrikes, and the results of static vertical load tests, Ng et al. 共2010兲 developed an equation to estimate steel pile capacity as a function of time, which shows that for a similar soil profile pile setup after about 3 months 共the time between the end of driving and testing for the presented research兲 approached 50% for steel H-piles. Load Transfer

static load test results and CAPWAP analysis may be attributed to the soil displacement at which these results were obtained where the UHPC pile was driven at a rate of 15 blows/0.3 m at the end of driving 共average of 20-mm displacement per blow兲 and the pile displacement at the Davisson 共1972兲 pile capacity was 15 mm. Fig. 9共b兲 exhibits the increase in unit skin friction for the two main soil types at the test site as a function of pile head displacement 共i.e., loess for the top 4.8 m and glacial till for the bottom 5.1 m along the length of the pile兲. This figure shows that the unit skin friction of the glacial till is higher than that of the loess soil with a ratio ranging from 1.54 to 1.85. Although the unit skin friction is expected to reach a maximum limit value as the pile head displacement increases, such behavior was not observed during the test, indicating that the maximum skin friction limit was not mobilized in either soil. Extrapolation of the unit skin friction curves indicates that the maximum unit frictions between the UHPC pile and loess and glacial till were 47 kN/m 共3.22 kips/ft兲 and 94 kN/m 共6.44 kips/ft兲. Fig. 9共c兲 shows the increase in total, skin friction and end bearing as a function of pile head displacement, which indicates that the percent of the load resisted by skin friction decreases as the load increases 共i.e., the difference between applied load and the skin friction curves increases兲. The figure also indicates that the curve of skin friction may become constant at a top pile displacement of about 7 mm 共2.7% of the pile cross-sectional dimension兲.

Lateral Load Test Test Procedure and Observations

Using section and material properties, the measured strains from sister bars and strain gauges were used to calculate the load in UHPC Pile 1 at different locations along the length of the pile, as shown in Fig. 9. Fig. 9共a兲 shows that when the applied load was 876.7 kN 共i.e., corresponding displacement at the top of the pile was 5.3 mm兲, 23.6% of the applied load was resisted by end bearing. At smaller pile displacements 共i.e., beginning of the test兲, 6% of the applied load was resisted by end bearing. CAPWAP analysis results indicated that at the end of driving 57% of the load was resisted by end bearing and 43% of the load resisted by skin friction. In addition to pile setup, the difference between the

The lateral load test on the UHPC piles was conducted using a 445-kN capacity hydraulic actuator with a 46-cm stroke, which was used to simultaneously push test units UHPC Pile 1 and UHPC Pile 2 laterally against each other. The actuator was placed at 0.61 m above the ground surface 共i.e., 0.25 m below the top of the pile兲. The lateral load test followed the “standard loading” procedure outlined in ASTM D3966-07 共ASTM 2007b兲. Accordingly, for all loading steps, the load was kept relatively constant for duration of at least 10 min. At the end of the test, the piles were unloaded in four equal load steps. For both the loading and



Vertical Load (kN) 200

400

600

800

1000

183.7 207.8 234.4 261.7 292.0 313.7 345.2 369.5 396.6 425.7 451.3 477.3 502.7 529.2 554.8 616.6 700.3 790.0 876.7

Ground Surface

0 2 4 5 6 7 8

Glacial Till

Depth (m)

3

Loess

1

UHPC pile

9 10 11

100

Load (kN)

-1

Points at 9.9 m are based on extrapolation

90

Loess soil (top 4.8 m) Glacial till (bottom 5.1 m)

80

Unit friction (kN/m)

0

70 60 50 40 30 20 10 0 0

12

1

2

3

4

5

6

Displacement at the top of the pile (mm)

(a)

(b) 1000 Total applied load Skin friction End bearing

900 800

Load (kN)

700 600 500 400 300 200 100 0 0

1

2

3

4

5

6

Displacement at the top of the pile (mm)

(c) Fig. 9. Measured vertical load transfer as a function of depth for UHPC Pile 1 obtained from strain measurements: 共a兲 load as a function of depth with points at 9.9 m determined using a constant slope of the measurements within the glacial till soil; 共b兲 skin friction per unit length for the two main soil layers; and 共c兲 skin friction and end bearing as a function of pile top displacement

unloading stages, deflections, strains, and loads were recorded at 1, 5, and 10 min after the load was applied and at 5-min intervals for any remaining duration. During the lateral load test, unexpected shear cracks and shear failure developed in UHPC Pile 1 near the point of lateral load application. Although section analysis showed that the shear capacity of the UHPC pile section was 191 kN, UHPC Pile 1 experienced extensive shear cracking at a lateral load of 101.3 kN. In a posttest investigation, it was concluded that the shear cracking and premature shear failure in this pile were caused by a reduction in shear capacity due to the presence of a 20-mm-thick bundle of instrumentation wires from 6 sister bars and 11 strain gauges through the web of the pile, as shown in Figs. 10共a and b兲. Force-Displacement Response The force-displacement responses of UHPC Pile 1 and UHPC Pile 2 from the lateral load testing are shown in Fig. 10共e兲, where both units show similar responses as the applied lateral load in-

creased up to 92.7 kN. For UHPC Pile 2 at the end of testing, the maximum lateral displacement at the point of loading was 6.25 cm for a lateral load of 101.3 kN. Test units were expected to sustain larger lateral loads and displacements but could not be loaded further after the shear failure of UHPC Pile 1. At a load of 101.3 kN, UHPC Pile 1 started experiencing a significant increase in lateral displacement as the shear cracks developed to a shear failure. After the shear failure began, UHPC Pile 1 continued to displace to reach a total displacement of 20.1 cm while sustaining a lateral load of approximately 80 kN. The lateral load response of the UHPC pile was also predicted using the LPILE software 共Reese et al. 2000兲. The pile was subjected to an increasing lateral displacement. The soil resistance in the horizontal direction was represented with nonlinear springs 共p-y curves兲 located at the midheight of the elements modeling the pile foundation. The p-y curves for all soil layers were developed using the soil parameters 共i.e., undrained shear strength, Su, and strain at 50% of the strength, ␧50, for cohesive soil, and fric-



Fig. 10. Response of UHPC 1 and UHPC 2 when subjected to lateral loads: 关共a兲 and 共b兲兴 cracking and failure of UHPC 1 at the location of instrumentation wires through the web of the pile; 共c兲 the p-y curves representing the nonlinear soil response used in the LPILE analysis for the major soil layers along the UHPC pile; 共d兲 the moment-curvature relationship used to model the response of the UHPC pile subjected to flexural bending; and 共e兲 lateral load versus lateral displacement

tion angle, ␾, and modulus of subgrade reaction, k, for cohesionless soil兲 reported in Fig. 3共d兲 and using the default LPILE p-y curves shown in Fig. 10共d兲. The flexural responses of the beamcolumn elements representing the pile foundation were characterized by specifying the moment resistance of the pile section and the corresponding flexural stiffness 共EcIeff, where Ec is the elastic modulus of the UHPC material and Ieff is the effective moment of inertia of the UHPC pile cross section兲, which was generated by performing the moment-curvature analysis 关see Fig. 10共e兲兴. Fig. 10共c兲 shows the force-displacement response of the UHPC pile predicted by LPILE analysis compared with the measured force-displacement response of UHPC Pile 2, which shows that the analysis captured the overall lateral load response of the UHPC pile. LPILE analysis predicted the lateral capacity of the UHPC pile to be 141 kN occurring at a maximum displacement of 26.3 cm.

Strain Profiles The strain profiles established for longitudinal reinforcing strands in UHPC Pile 1 and UHPC Pile 2 are presented in Fig. 11. Fig. 11共a兲 presents the strain profiles combining measured strains for Strands 2 and 4 for UHPC Pile 1 and Fig. 11共b兲 shows the strain profile along Strand 3 for UHPC Pile 2; both of which show that, for a given lateral load, the largest strains were recorded at a depth of 1.43 m below the ground surface. The strain profiles also suggest that strains gradually reduced below 1.43 m, and insignificant strains were recorded at 5.12 m below the ground surface. Based on these observations and the LPILE analysis results, it was concluded that the maximum moment in test units UHPC Pile 1 and UHPC Pile 2 was at a depth of about 1.43 m below the ground surface 共i.e., depth approximately 5.6⫻ pile size兲. The differences in magnitude between recorded strains at similar



Strain (

Strain ( 0

200

400

600

800

0

1000

Ground Surface

0

400

600

800

1000

1

1

2

2

3 4

15.9 kN 27.9 kN 40.7 kN 55.9 kN 68.8 kN 77.8 kN 92.7 kN 101.2 kN

5 6 7 8 9 10

Depth =1.43 m

Ground Surface

0

Depth (m)

Depth (m)

200

-1

-1

3 15.9 kN 27.9 kN 40.7 kN 55.9 kN 68.8 kN 77.8 kN 92.7 kN 101.2 kN

4 5 6 7 8 9 10

(a)

(b)

Fig. 11. Measured strain profiles along the length of UHPC piles: 共a兲 along Strands 2 and 4 in UHPC 1; 共b兲 along Strand 3 in UHPC 2

depth in test units UHPC Pile 1 and UHPC Pile 2 were attributed to shear cracks and shear failure observed in UHPC Pile 1 during the lateral load test. The maximum measured strains along UHPC Pile 1 and UHPC Pile 2 result in bending moments of 125 and 101 kN m, respectively, which is smaller than the bending moment capacity of 198 kN m expected at zero axial load. These results confirm that UHPC Pile 1 failed due to shear as concluded previously.

Summary and Conclusions To achieve the target service life, avoid drivability problems, and ensure durability, the researchers investigated the use and behavior of deep foundation elements made of UHPC. A prestressedprecast H-shaped UHPC section, which was 254-mm deep with weight similar to that of a HP 10⫻ 57 steel pile, was designed with no mild steel reinforcement. The following are the major findings of this study. 1. H-shaped UHPC piles can be driven with the same driving equipment as steel piles of the same size and weight and can be successfully driven without a pile cushion through hard soil layers without causing any damage to the UHPC piles. Although the UHPC pile required more blows to reach the target depth, because of their higher compressive and tensile strengths, UHPC piles can be driven with larger hammers and strokes reducing the required total number of blows and improving the efficiency of the driving process especially when compared to precast-prestressed concrete piles. 2. CAPWAP analyses at the end of driving confirmed that UHPC piles resisted applied load by both skin friction 共43%兲 and end bearing 共57%兲 and steel H-piles resisted applied loads mainly by skin friction 共75%兲. 3. Although Nordlund 共1963兲 static analysis method indicates that the skin friction of steel piles is greater than that of UHPC piles, the axial load capacity of the UHPC pile was 86% greater than that of the steel pile. This suggests that the use of UHPC piles may reduce the total number of piles required for a bridge foundation. 4. When subjected to lateral loading, UHPC Pile 1 experienced unexpected failure due to the presence of wire bundle through the web of the pile. The maximum lateral load and displacement recorded for the unfailed test unit 共UHPC Pile

2兲 was 101.3 kN at 6.25 cm. The LPILE analysis of the UHPC pile closely matched the measured response of UHPC Pile 2 and predicted a lateral load capacity of 141 kN with a maximum moment location at a depth equals to approximately 5.6 times the cross section of the pile.

Acknowledgments This study was funded through the high-risk research program by the Iowa Highway Research Board. The writers would like to thank the technical advisory committee of the research project. Special thanks are due to Iowa Department of Transportation, Coreslab Structures of Omaha, Nebraska, Lafarge North America, and Cramer and Associates for their help, in-kind support, and services. The writers would like to thank Sherif AbdelSalam, Kam Weng Ng, and Mathew Rolling for their help during field tests and interpretation of data.

References American Association of State Transportation and Highway Officials. 共2005兲. Grand challenges: A strategic plan for bridge engineering, AASHTO Subcommittee on Bridges and Structures, Washington, D.C. American Petroleum Institute 共API兲. 共1989兲. “Draft recommended practice for planning, designing and constructing fixed offshore platforms—Load and resistance factor design.” API RP2A-LRFD, American Petroleum Institute, Dallas. ASTM. 共2007a兲. “Standard test methods for deep foundations under static axial compressive load.” ASTM D1143-07, West Conshohocken, PA. ASTM. 共2007b兲. “Standard test methods for deep foundations under lateral load.” ASTM D3966-07, West Conshohocken, PA. Burland, J. B. 共1973兲. “Shaft friction of piles in clay.” Ground Eng., 6共3兲, 30–42. Davisson, M. T. 共1972兲. “High capacity piles.” Proc., Soil Mechanics Lecture Series on Innovations in Foundation Construction, ASCE, New York, 81–112. Dirks, K., and Kam, P. 共1989兲. “Foundation soils information chart-pile foundation.” Iowa Dept. of Transportation, Office of Roads Design, Ames, IA. Esrig, M. I., and Kirby, R. C. 共1979兲. “Soil capacity for supporting deep foundation members in clay.” Behavior of deep foundation, Vol. 670,



ASTM, West Conshohocken, PA, 27–63. Graybeal, B. A. 共2006兲. “Material property characterization of ultra-high performance concrete.” FHWA-HRT-06-103, Federal Highway Administration, Washington, D.C. Hannigan, P. J., Goble, G. G., Thendean, G., Linkin, G. E., and Rausche, F. 共1997兲. “Design and construction of driven foundations: Volume I and II.” Federal Highway Administration (FHWA) Rep. No. FHWAHI-97-013, Federal Highway Administration, Washington, D.C. Hara, A., Ohta, T., Niwa, M., Tanaka, S., and Banno, T. 共1974兲. “Shear modulus and shear strength of cohesive soils.” Soils Found., 14共3兲, 1–12. Huck, R. W., and Hull, J. R. 共1971兲. “Resonant driving in permafrost.” Foundation Facts, 7共1兲, 11–15. Iowa Department of Transportation. 共2007兲. “ASD/LFD bridge design manual.” 具http://www.dot.state.ia.us/bridge/manualasd.htm典共Jan. 10, 2008兲. Lunne, T., Powell, J. M., and Robertson, P. K. 共1997兲. Cone penetration testing in geotechnical practice, Taylor & Francis, London. Mekkawy, M. 共2004兲. “Management of water and backfill characteristics for improved bridge approach performance.” M.S. thesis, Iowa State Univ., Ames, IA. Meyerhof, G. 共1976兲. “Bearing capacity and settlement of pile foundations.” J. Geotech. Engrg. Div., 102共3兲, 195–228. Ng, K. W., Suleiman, M. T., and Sritharan, S. 共2010兲. “LRFD resistance factors including the influence of setup for design of steel piles using WEAP.” GeoFlorida 2010—Advances in analysis, modeling and design, Florida 2010, Geo-Institute of ASCE, West Palm Beach, FL. Nordlund, R. L. 共1963兲. “Bearing capacity of piles in cohesionless soils.”

J. Geotech. Found. Eng., 89共SM 3兲, 1–36. Paikowsky, S. G., and Tolosko, T. A. 共1999兲. “Extrapolation of pile capacity from non-failed load tests.” FHWA-RD-99-170, Federal Highway Administration 共FHWA兲, Washington, D.C. Pando, M. A., Ealy, C. D., Filz, G. M., Lesko, J. J., and Hoppe, E. J. 共2006兲. “A laboratory and field study of composite piles for bridge substructures.” Rep. No. FHWA-HRT-04-043, Federal Highway Administration 共FHWA兲, Washington, D.C. Peck, R. B., Hanson, W. E., and Thornburn, T. H. 共1974兲. Foundation engineering, Wiley, New York. Reese, L. C., Wang, S. T., Isenhower, W. M., and Arrellaga, J. A. 共2000兲. Computer program: LPILE version 4 technical manual, Ensoft, Austin, TX. Salgado, R. 共2008兲. The engineering of foundations, McGraw-Hill, New York. Tomlinson, M. J. 共1986兲. Foundation design and construction, Longman Scientific and Technical, Essex, England. Tuchlinski, D., Hegger, J., and Kommer, B. 共2006兲. “Studies on prestressed concrete beams made from UHPC.” Concrete Precasting Plant and Technology, 72共1兲, 14–20. Vande Voort, T., Suleiman, M. T., and Sritharan, S. 共2008兲. “Design and performance verification of ultra-high performance concrete piles for deep foundations.” Final Report, Iowa DOT, IHRB Project TR-558, CTRE Project 06-264, Iowa Dept. of Transportation, Ames, IA. Vande Voort, T., Suleiman, M. T., and Sritharan, S. 共2009兲. “Design, construction, and drivability of UHPC pile.” Int. Foundation Congress and Equipment Expo 09, ASCE and Geo-Institute, Orlando, FL.

