Testing of full-scale concrete bridge deck slabs reinforced with fiber-reinforced polymer (FRP) bars

Testing of full-scale concrete bridge deck slabs reinforced with fiber-reinforced polymer (FRP) bars

Construction and Building Materials 25 (2011) 3956–3965 Contents lists available at ScienceDirect Construction and Building Materials journal homepa...

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Construction and Building Materials 25 (2011) 3956–3965

Contents lists available at ScienceDirect

Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat

Testing of full-scale concrete bridge deck slabs reinforced with fiber-reinforced polymer (FRP) bars K. Bouguerra, E.A. Ahmed, S. El-Gamal, B. Benmokrane ⇑ Dept. of Civil Engineering, University of Sherbrooke, Quebec, Canada

a r t i c l e

i n f o

Article history: Received 16 November 2010 Received in revised form 23 March 2011 Accepted 14 April 2011 Available online 10 May 2011 Keywords: Fiber-reinforced polymer (FRP) Concrete Bridge deck slab Design Punching strength

a b s t r a c t This paper presents an experimental study investigating the behavior of FRP-reinforced concrete bridge deck slabs under concentrated loads. A total of eight full-scale deck slabs measuring 3000-mm long by 2500-mm wide were constructed. The test parameters were: (i) slab thickness (200, 175 and 150 mm); (ii) concrete compressive strength (35–65 MPa); (iii) bottom transverse reinforcement ratio (1.2–0.35%); and (iv) type of reinforcement (GFRP, CFRP, and steel). The slabs were supported on two parallel steel girders and were tested up to failure under monotonic single concentrated load acting on the center of each slab over a contact area of 600  250 mm to simulate the footprint of sustained truck wheel load (87.5 kN CL-625 truck). All deck slabs failed in punching shear. The punching capacity of the tested deck slabs ranged from 1.74 to 3.52 times the factored load (Pf) specified by the Canadian Highway Bridge Design Code (CHBDC) CAN/CSA S6-06. Besides, the ACI 440.1R-06 punching strength equation greatly underestimated the capacity of the tested slabs with an average experimental-to-predicted punching capacity ratio (Vexp/Vpred) of 3.17. Ó 2011 Elsevier Ltd. All rights reserved.

1. Introduction Corrosion of steel reinforcement in concrete structures is a major durability problem that leads to structural degradation and consequent costly repairs. In a bridge system, concrete bridge decks deteriorate faster than any other bridge component due to direct exposure to environment, de-icing salts and ever-increasing traffic loads. A considerable amount of North America’s transportation infrastructure is in need of repair or replacement and constitutes a major problem when measured in terms of rehabilitation costs and traffic disruption [1]. The American Society of Civil Engineers (ASCE) estimates that over $1.6 trillion would be needed in the next five years in the U.S. to alleviate potential problems with civil infrastructure. Canada’s deficit for its municipal infrastructure, which represents 70 percent of the country’s total infrastructure was estimated to be $60 billion in 2004, and is expected to grow by $2 billion dollars per year (ISIS Canada [2]). Worldwide, governments and industrial firms are looking for infrastructure systems that are stronger, longer lasting, more resistant to corrosion and less costly to build and maintain. Engineers all over the world are searching for new and affordable construction materials as well as innovative approaches and systems to solve these problems. An effective solution to this problem is corrosion-resistant materials, such as fiber-reinforced polymer (FRP) composites. This

⇑ Corresponding author. E-mail address: [email protected] (B. Benmokrane). 0950-0618/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.conbuildmat.2011.04.028

cutting-edge technology has emerged as one of the most costeffective alternatives to traditional solutions. Recent advances in polymer technology have led to the development of the latest generation of FRP reinforcing bars (in particular glass FRP (GFRP) bars). These corrosion-resistant bars have shown promise as a way to further protect bridges and public infrastructure from the devastating effects of corrosion. With FRP being standardized through the recently published CSA specifications (CAN/CSA S807-10 [3]) and bars of the highest quality being produced, FRP bars are emerging as a realistic and cost-effective alternative to traditional steel reinforcement for concrete structures under severe environmental conditions. At the Department of Civil Engineering, University of Sherbrooke, a multi-phase experimental program investigating the structural performance of restrained FRP-reinforced concrete bridge deck slabs under different loading conditions has been completed. The first phase [4] investigated the restrained FRP-reinforced concrete bridge deck slabs under static loading condition acting at the center of the slab on an area equivalent to a wheel load from a CL-625 truck specified by the CHBDC CAN/CSA S6-06 [5]. The second phase [6] investigated the behavior of the restrained FRP-reinforced concrete bridge deck slabs under constant and variable amplitude fatigue loading. The load was also applied at the center of the slab on an area equivalent to a wheel load from a CL-625 truck specified by the CHBDC CAN/CSA S6-06 [5]. The aforementioned two phases [4,6] considered different reinforcement types and ratios maintaining a 200-mm slab thickness for all test specimens and using normal strength concrete.

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Recently, there is an increase in the use of high-strength concrete (HSC) in bridges due to its superior strength and stiffness. Carpenter [7] reported that the use of high-strength concrete (HSC) in bridge applications can result in: (1) greater compressive strength per unit cost, per unit weight, and per unit volume; (2) higher modulus of elasticity resulting in reduced deflection; and (3) increased tensile strength. The enhanced properties which includes compressive and tensile strengths, stiffness, and modulus of the HSC would contribute to achieving the design capacity with smaller dimensions which reduces the load that has to be resisted by the other supporting elements. Thus, the third phase of this experimental program was planned to investigate the effect of slab thickness, reinforcement type and ratio, and the concrete strength (normal and high-strength concretes) on the structural performance and capacity of restrained FRP-reinforced bridge deck slabs. The test results of this phase are presented and discussed herein. 2. Experimental program 2.1. Test specimens A total of eight full-scale concrete deck slabs of 3000-mm long  2500-mm wide  (150, 175, and 200-mm) thick, as shown in Fig. 1, were constructed and tested. All slabs were reinforced with two mats of orthogonal reinforcing bars. Three different diameter of glass FRP (GFRP) bars were used and designated as No. 13 (12.7 mm), 15 (15.9 mm), and 20 (19.1 mm) according to the CAN/CSA S807-10 [3] and one diameter of carbon FRP (CFRP) bars designated as No. 10 (9.5 mm) [8]. The bars were fabricated using a pultrusion process with a fiber content of 73% and 78% by weight in a vinyl ester resin for the GFRP and CFRP bars, respectively. The FRP bars had a sand-coated surface to enhance bond performance between the bars and surrounding concrete. In addition, 10 M steel bars were used

in the reference slab for comparison. The tensile properties of the reinforcing FRP bars were determined by testing of representative specimens in accordance with ACI 440 3R-04 B.2 [9] Test Method ‘‘Test Method for Longitudinal Tensile Properties of FRP Bars.’’ Table 1 summarizes the tensile properties of the reinforcing bars used in this study. The test slabs were divided into four groups. The first and second groups were designed to investigate the effect of slab thickness and concrete compressive strength, respectively. The third and fourth group, however, were designed to investigate the effect of reinforcement ratio and type. Slab designation is as follows: The first letter indicates reinforcement type (S for steel, G for Glass FRP, and C for Carbon FRP), the middle number indicates the thickness of the slab followed by a letter representing the strength of the concrete (N for normal strength and H for high strength). Numbers at the end (only in Group 3) represent the reinforcement ratio when not equal to 1.2%. The first group, Group 1, comprised three normal strength concrete deck slabs (G-200-N, G-175-N and G-150-N) with three different thicknesses (200, 175 and 150 mm), respectively. Group 2 comprised a deck slab; G-175-H with identical reinforcement to the G-175-N in Group 1; however 64.8 MPa HSC was used for the G175-H slab. For the slabs of these two groups (Group 1 and Group 2), the bottom transverse GFRP reinforcement ratio was 1.2% while the GFRP reinforcement ratio in all other directions was set to 50% of the bottom transverse reinforcement which yield a reinforcement ratio of 0.6%. Group 3 comprised two slabs reinforced with GFRP bars: G-175-N-0.7 and G-175-N-0.35 with bottom transverse reinforcement ratios equal to 0.7 and 0.35%, respectively. The fourth group comprised two slabs, (S-175-N and C-175-N), reinforced with different types of reinforcing bars; steel and CFRP, respectively. The reinforcement ratio in the other directions for the FRP-reinforced slabs of Group 3 and 4 was maintained to 0.35%. The dimensions and reinforcement details of the deck slabs are shown in Fig. 1 while the complete characteristics of the deck slab are listed in Table 2. 2.2. Test setup and procedure Similar to the first two phases of this research program [4,6] and to simulate bridge deck slabs with restrained edges, the slabs were supported on two steel girders spaced at 2000 mm center-to-center. The girders were connected together by

2500 mm

GFRP No. 13 @ 255 mm

GFRP No. 15 @ 200 mm

25 150 25

200

25 125 25

GFRP No. 20 @ 150 mm

175 GFRP No. 15 @ 200 mm

G-200-N

G-175-N-0.7 GFRP No. 13 @ 255 mm

GFRP No. 15 @ 230 mm

600

A

175

25 125 25

GFRP No.15 @ 115 mm

175 GFRP No. 13 @ 255 mm

G-175-N

250

A

25 125 25

G-175-N-0.35 Steel 10M @ 230 mm

GFRP No. 15 @ 270 mm

25 100 25

150

25 125 25

GFRP No. 15 @ 140 mm

175 Steel 10M @ 230 mm

G-150-N

S-175-N GFRP No. 15 @ 230 mm

GFRP No. 15 @ 230 mm

25 125 25

175

25 125 25

GFRP No. 15 @ 115 mm

175 CFRP No. 10 @ 125 mm

G-175-H

(a) Typical dimensions of the tested slabs

C-175-N

(b) Section A-A

Fig. 1. Dimensions and reinforcement details of the tested slabs.

Table 1 Mechanical properties of the reinforcing bars. Bar typea Steel GFRP

CFRP a b

Bar designation 10 M No. 13 No. 15 No. 20 No. 10

Diameter (mm) 11.3 12.7 15.9 19.1 9.5

Areaa (mm2) 100 129 199 284 71

FRP bars are designated according to the CSA S807-10 [3]. fy and ey are the yield strength and yield strain of the steel bars, respectively.

Modulus of elasticity (GPa) 200 41.0 ± 2.1 41.6 ± 1.5 44.5 ± 1.3 122 ± 2.4

Tensile strength (MPa) b

fy = 453 769 ± 33 778 ± 16 637 ± 15 1444 ± 18

Ultimate strain (%)

eyb = 0.23 1.87 ± 0.02 1.87 ± 0.01 1.37 ± 0.03 1.23 ± 0.07

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Table 2 Concrete strength and reinforcement details of the tested deck slabs. Group

Slab

Slab thickness (mm)

fc0 ðMPaÞ

Reinforcement ratio (%)

Reinforcement configuration Bottom transverse direction (axial stiffness, MN/m)

Other directions

1

G-200-N G-175-Na G-150-N G-175-H G-175-N-0.7 G-175-N-0.35 S-175-N C-175-N

200 175 150 175 175 175 175 175

49.1 35.2 35.2 64.8 53.1 53.1 42.3 40.3

1.20 1.20 1.20 1.20 0.70 0.35 0.30 0.40

No. 20 @ 150 mm (84.2) No. 15 @ 115 mm (72.0) No. 15 @ 140 mm (59.1) No. 15 @ 115 mm (72.0) No. 15 @ 200 mm (41.4) No. 13 @ 255 mm (20.7) 10 M @ 230 mm (86.9) No. 10 @ 125 mm (69.3)

No. 15 @ 200 mm No. 15 @ 230 mm No. 15 @ 270 mm No. 15 @ 230 mm No. 13 @ 255 mm No. 13 @ 255 mm [email protected] mm No. 15 @ 230 mm

2 3 4 a

Used for comparison with Groups 2, 3, and 4.

two cross frames spaced at 3000 mm. Each cross frame consisted of an X-shaped bracing with 55  55  6 mm angles, as shown in Fig. 2a. Each slab was bolted to the top flange of the steel girders through two rows of holes at each edge using 25-mm diameter steel bolts and two steel channels. The two rows of the holes were

spaced 180 mm apart with a pitch of 250 mm. To ensure equal and uniform edge clamping force, all bolts were hand tightened to an equal torque of 160 N m using a torque wrench. In addition, a 3-mm thick neoprene pad was used between the steel sections (channel and top flange of girder) and the concrete slab. The steel

Fig. 2. Test setup: (a) Supporting system. (b) Restrained edge details.

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Fig. 3. Testing of a slab specimens.

bolts replace shear studs, which provide edge restraint to the slab in both longitudinal and transverse directions [4]. Fig. 2b shows the details of the edge-restraint configuration that were utilized herein. The deck slabs were tested up to failure using one concentrated load over a contact area of 600  250 mm to simulate the foot print of the truck wheel load; CL-625 Truck (CAN/CSA S6-06) acting on the center of each slab. In all tests, the load was applied through a 60-mm thick steel plate over a 10-mm thick neoprene pad. The load was monotonically applied at a load controlled rate of 5 kN/min. During loading, the formation of cracks on the bottom surface of the deck slabs were marked and recorded. Fig. 3 shows a photo of the test setup during the testing of a slab specimen.

3. Test results and discussion The test results are summarized in Table 3 and the final crack patterns of the slabs at failure are shown in Fig. 4. However, the following section will discuss the results and the effect of the different parameters on the behavior of the tested slabs. The design service load, Pser, of the deck slabs was taken as 1.4  0.9  87.5 = 110.25 kN, where 87.5 is the maximum wheel load of the design truck (CL-625 Truck – CAN/CSA S6-06 [5]), 1.4 is the impact coefficient, and 0.9 is the live load combination factor. The design factored load was taken as Pf = 1.4  1.7  87.5 = 208.25 kN, where 1.7 is the live load combination factor (CAN/CSA S6-06 [5]). 3.1. Ultimate capacity and mode of failure For all tested deck slabs, the failure mode was punching shear around the loaded area. The top surface of the failure zone had an elliptical shape passing through the corners of the loaded area. The bottom surface had an approximately circular shape with a

diameter equal to the spacing between the two girders as shown in Fig. 4. Generally, the carrying capacities of the tested deck slabs ranged from 362 to 732 kN representing 1.74–3.52 times the factored design load (Pf = 208.25 kN) specified by the CHBDC CSA S6-06 [5]. The high level of conservativeness is due to the general behavior of the restrained bridge deck slabs and mode of failure. Hewitt [10], Hewitt and Batchelor [11]; and Perdikaris and Beim [12] reported that reinforced concrete restrained bridge deck slabs subjected to concentrated wheel loads failed in punching shear rather than flexure as assumed by conventional design. Furthermore, Hewitt and Batchelor [11] reported that when a deck slab is restrained, there is no need for reinforcement to resist the wheel loads due to the compressive membrane action which is similar to the arching action in the reinforced concrete beams. A minimum amount of reinforcement was proposed for serviceability reasons. This concept was adopted by the Canadian Highway Bridge Design Code (CAN/ CSA S6-06 [5]) through its empirical method. However, this method does not account for any design parameters other than the effective depth of the slab and the modulus of elasticity of the reinforcing bars such as the concrete strength or the effective slab span. Ahmed and Benmokrane [13] concluded that the empirical method in most cases overestimates the required bottom transverse reinforcement. To evaluate the design methods of the Canadian Highway Bridge Design Code CAN/CSA S6-06 [5], the bridge deck slabs are designed using the flexural design method and the empirical design method. As the deck slab specimens were edge restrained, they were designed as a continuous slab over girders. The required amount of reinforcing bars determined by those methods is listed in Table 4. From this table it can be noticed that, the empirical design method in most cases yields higher reinforcement amount than the flexural design method which was also reported by Ahmed and Benmokrane [13]. This is referred to the fact that the empirical design method is affected only by the thickness of the concrete section and the tensile modulus of the reinforcing bars. Increasing the depth to enhance the performance or decreasing the effective spacing between the girders will not be reflected on the required amount of reinforcement. On the other hand, changing the any of those parameters will affect the reinforcement amount resulted from the flexural design method. However, the design of the bridge deck slab is often controlled by the crack width criterion when the flexure design method is employed [13]. From Table 4 it can be noticed that, even though the failure load of the tested slabs was higher than the factored load of the CAN/ CSA S6-06 [5], the design of the bridge deck slabs using flexural design method gave more reinforcement than what was provided in some specimens. Of interest, the flexural design of slabs G-175-N and G-175-H yielded the same amount of reinforcement as the high-strength concrete did not affect the calculated amount of reinforcement. This is due to the fact that the design of the section

Table 3 Summary of the test results. Group

1

2 3 4

*

Slab

G-200-N G-175-N* G-150-N G-175-H G-175-N-0.7 G-175-N-0.35 S-175-N C-175-N

Cracking load (kN)

115 113 107 130 118 98 121 103

Used for comparison with Groups 2, 3, and 4.

Ultimate load (kN)

732 484 362 704 549 506 550 530

Ultimate load/Pf

3.52 2.32 1.74 3.38 2.64 2.43 2.64 2.55

Maximum deflection (mm)

Maximum strain at service load (le)

Service

Failure

Bars

Concrete

0.45 0.72 1.49 0.45 0.48 1.04 0.52 0.78

22.9 17.9 17.1 21.7 20.4 26.4 23.8 16.7

624 214 552 214 223 1293 78 905

180 200 200 140 120 298 146 276

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(a) G-150-N

(b) G-175-N

(c) G-150-N

(d) G-175-N

(e) G-175-H

(f) G-175-N-0.35

(g) C-175-N

(h) S-175-N

Fig. 4. Crack pattern of the slab specimens at failure.

was based on failure in flexure when in fact the failure mode was punching shear. Considering the actual mode of failure in design may lead to reducing the FRP reinforcement amount which will result in reducing the initial costs of the FRP-reinforced concrete bridge deck slabs. Furthermore, the slab G-15-N with a 150 mm thickness, which is less than the CHBDC (CAN/CSA S6-06 [5]) minimum allowable thickness of 175 mm, gave very high reinforcement ratio (2%) when designed using the flexural method. Also, the empirical method of the CAN/CSA S6-06 [5] does not account for the concrete strength. Of interest, increasing the slab thickness resulted in increasing the reinforcement amount even

though the sections are designed to resist the same straining actions. Besides, for slabs with thickness of 150 mm there was a big difference between the reinforcement amount calculated from the flexural and empirical methods. The test results revealed that the ultimate carrying capacity of the slab specimens was affected by the slab thickness, reinforcement ratio, and concrete strength. Table 3 shows that the punching capacities of the tested slabs were proportional to the slab thickness. In Group 1, when the slab thickness was decreased by 12.5% (from 200 to 175 mm) for slabs G-200-N and G-175-N and the ultimate capacity decreased by about 34%. Similarly when

Table 4 Design of the bridge deck slabs according to the CHBDC (CSA S6-06) [5]. Slab

Thickness (mm)

fc0 ðMPaÞ

Bottom transverse direction Flexural design method

Bottom transverse direction Empirical design method

G-200-N G-175-N G-175-H G-150-N

200 175 175 150

49.1 35.2 64.8 35.2

GFRP GFRP GFRP GFRP

GFRP GFRP GFRP GFRP

No. No. No. No.

15 15 15 15

@ @ @ @

135 mm 120 mm 120 mm 85 mm

No. No. No. No.

15 15 15 15

@ @ @ @

100 mm 115 mm 115 mm 140 mm

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Expiremental punching capacity, Vexp (kN)

700

100

600 80

500

60

400 300

40

200

Vexp

100

Vexp /√f'c

0 125

150

175

20

Normalized Punching √ f'c Capacity, Vexp /√

120

800

0 225

200

Slab thickness (mm)

(a)

Expiremental punching capacity (kN)

800 700 600 500 400 300 200 100 0

0

20

40

60

80

100

Concrete compressive strength (MPa)

560

84

550

82

540

80

530

78

520

76

510

74

500 490 480

72

Vexp

70

Vexp /√f'c 0

0.25

0.5

0.75

1

1.25

1.5

Normalized Punching Capacity, Vexp /√f'c

Expiremental punching capacity, Vexp (kN)

(b)

68

Reinforcement ratio %

(c) Fig. 5. Experimental punching capacity versus: (a) Slab thickness (Group 1). (b) Concrete compressive strength (Group 2). (c) Bottom transverse reinforcement ratio (Group 3).

the slab thickness was decreased by 25% (from 200 to 150 mm) for slabs G-200-N and G-150-N the ultimate capacity decreased by about 50%. The punching capacity versus the slab thickness relationship in Fig. 5a also shows that the punching capacity is directly proportional to the slab thickness. The higher the slab thickness the higher the punching capacity. This is referred to that increasing the slabs thickness results in increasing the surface area that resists the punching shear stresses which yields higher punching shear capacity. For slabs of Group 2, the punching capacity increased with the increase in concrete compressive strength as shown in Fig. 5b. For slabs G-175-N and G-175-H, increasing the concrete compres-

sive strength by 83% (from 35.4 to 64.8 MPa) increased the punching capacity by 45% (from 484 to 704 kN) as shown in Table 3. This is due to the dependency of the punching shear strength of reinforced concrete slabs on the concrete strength. However, this is not reflected in the flexural design method incorporated in the Canadian Highway Bridge Design Code (CSA S6-06 [5]). For the slabs pffiffiffiffi of Group 3, the normalized punching capacity curve (V exp = fc0 ) in Fig. 5c shows that punching capacity is proportional to reinforcement ratio. It can be noticed that increasing the reinforcement ratio increased the punching capacities of the tested slabs. For Group 4, it can be noticed that the type of reinforcement

K. Bouguerra et al. / Construction and Building Materials 25 (2011) 3956–3965

800

800

700

700

600

600

500

500

400

400

Load (kN)

Load (kN)

3962

300

G-200-N

300

200

G-175-N

200

G-175-N

G-150-N

100

G-150-N

100

G-200-N

0

0 0

0.25

0.5

0.75

1

1.25

1.5

1.75

2

2.25

2.5

-3000 -1500

0

1500

3000

Crack Width (mm)

7500

9000 10500 12000

800

800

700

700

600

600

Load (kN)

Load (kN)

6000

(a)

(a)

500

500

400

400

300

300 G-175-N

200

G-175-H

100

200

G-175-N

100

G-175-H

0

0

-3000 -1500 0

0.25

0.5

0.75

1

1.25

1.5

1.75

2

2.25

0

1500

2.5

3000

4500

6000

7500

9000 10500 12000

Strains (micro-strain)

Crack Width (mm)

(b)

(b) 800

800

700

700

Load (kN)

600

600

Load (kN)

4500

Strains (micro-strain)

500

500

400

400

300

300 200 100

200

G-175-N-0.7

G-175-N-0.7

100

G-175-N-0.35

G-175-N-0.35

0 -3000 -1500

0 0

0.25

0.5

0.75

1

1.25

1.5

1.75

G-175-N

G-175-N

2

2.25

0

1500

2.5

3000

4500

6000

7500

9000 10500 12000

Strains (micro-strain)

Crack Width (mm)

(c)

(c) 800 700

700

600

600

500

500

400

400

300

300

Load (kN)

Load (kN)

800

S-175-N

S-175-N

200

G-175-N

200

G-175-N

100

C-175-N

100

C-175-N

0 -3000 -1500

0 0

0.25

0.5

0.75

1

1.25

1.5

1.75

2

2.25

0

1500

Fig. 6. Load–crack width relationships for test deck slabs: (a) Group 1. (b) Group 2. (c) Group 3. (d) Group (4).

4500

6000

7500

9000 10500 12000

Strains (micro-strain)

2.5

(d)

Crack Width (mm)

(d)

3000

Fig. 7. Load–maximum strains relationships for tested slabs: (a) Group 1. (b) Group 2. (c) Group 3. (d) Group 4.

3.2. Cracking behavior did not show a significant effect on the punching capacities of the tested slabs.

All tested slabs had almost similar cracking patterns. The first cracks appeared directly under the loaded area and were oriented

K. Bouguerra et al. / Construction and Building Materials 25 (2011) 3956–3965

800.0 700.0

Load (kN)

600.0 500.0 400.0 300.0 200.0

G-200-N

100.0

G-175-H

0.0 0.00

0.25

0.50

0.75

1.00

1.25

1.50

1.75

2.00

Crack Width (mm)

(a) 800.0 700.0

Load (kN)

600.0 500.0 400.0 300.0 200.0

G-200-N

100.0

G-175-H

0.0 0

2.5

5

7.5

10

12.5

15

17.5

20

22.5

25

Deflection (mm)

(b) 800

Load (kN)

700 600 500 400 300

with a lower cracking load was observed in slab G-175-N-0.35, the slab with the lowest reinforcement ratio. The slabs of Group 4, with different types of reinforcement, also showed similar crack patterns. In this group, slab C-175-N had a lower cracking load than the two other slabs. The lower cracking loads observed in slabs G-175-N-0.35 (Group 3) and C-175-N (Group 4) could be explained by the presence of hairline shrinkage cracks that were observed in the slabs before testing. Fig. 6 shows the variation of maximum measured crack widths against applied loads for all tested slabs. The load-crack width relationships were almost linear for all FRP-reinforced concrete slabs. Slab S-175-N reinforced with steel showed a bilinear curve due to the yielding of the steel bars. Fig. 6a shows the load-crack width curves of the slabs of Group 1. It can be noticed that slabs G-200-N and G-175-N had almost similar crack width values. Higher crack width values were recorded in slab G-150-N. This indicates that decreasing the thickness from 200 to 175 mm did not affect crack widths. However, decreasing the thickness to 150 mm, which is less than the CHBDC (CAN/CSA S6-06 [5]) minimum allowable thickness of 175 mm, resulted in wider cracks. For the two slabs of Group 2, Fig. 6b shows that the load-crack width relationships of slabs G-175-N and G-175-H were similar up to failure. This indicates that the concrete strength did not have significant influence on the crack widths. Fig. 5c shows the load-crack width curves of Group 3 slabs. It can be clearly seen that decreasing the bottom transverse reinforcement ratio by increasing the spacing between bars significantly increased crack width values. This is in good agreement with several investigations which directly specify the maximum bar spacing to control cracking [14,15]. The load-crack width curves of the slabs of Group 4 are shown in Fig. 6d. The three slabs of this group had similar axial stiffness of the bottom transverse reinforcement. Slabs of the same axial stiffness may demonstrate the same behavior, however, slab S-175-N reinforced with steel, showed higher crack widths after steel yielding.

3.3. Strains in reinforcement and concrete

200

G-200-N

100

G-175-H

0 -3000 -2000 -1000

3963

0

1000 2000 3000 4000 5000 6000 7000 8000

Strains (micro-strain)

(c) Fig. 8. Comparison between G-200-N and G-175-H: (a) Crack width. (b) Deflection. (c) Strains.

in the longitudinal direction parallel to the supporting beams. As the load increased, subsequent cracks propagated in the radial direction away from the loaded area. For all test slabs, the cracking loads ranged between 98 to 130 kN as listed in Table 3. For the three slabs of Group 1, the cracking loads were greater than the service load (Pser), except for slab G-150-N with a 150-mm thickness. It should be mentioned that 150 mm is below the CAN/CSA S6-06 [5] minimum slab thickness of 175 mm. For the two slabs of Group 2, it was noticed that the slab G-175-H had more cracks on its tension face. This could be related to its HSC which increased the bond strength between the GFRP bars and the concrete components, consequently, increasing the number of cracks. It was also observed that increasing the concrete strength from 35.2 MPa (G175-N) to 64.8 MPa (G-175-H) increased the cracking load from 114 to 130 kN. For the slabs of Group 3, with different bottom transverse GFRP reinforcement ratios, less cracks accompanied

Fig. 7 shows the load-maximum concrete and reinforcement strains for the tested slabs. Before cracking, all the tested slabs showed almost similar concrete and reinforcement strains. After cracking, different strains were recorded. At failure, the maximum strains in the bottom transverse reinforcement were measured in slab G-175-N-0.35 (Group 3) which had the lowest reinforcement ratio and slab S-175-N (Group 4) due to the yielding of the steel reinforcement as listed in Table 3. For all FRP-reinforced slabs, the maximum measured strains in the bars at service and at factored load levels were about 9% and 20% of their ultimate strains, respectively. At failure, the maximum measured strains in those bars were about 66 % of their ultimate strains. These maximum strains were measured in slab G-175-N-0.35. Fig. 7a shows that increasing slab thickness decreased the measured strains in the GFRP bars after cracking. At the factored design load level (Pf = 208.25 kN), the maximum measured strains in the reinforcement were 1990, 2390 and 2495 micro-strain for slabs G-200-N, G-175-N, and G-150-N, respectively. These values were 30%, 38%, and 51% of the measured strains at failure (6690, 6224, and 4872 micro-strain). This provides an ample warning and safety factors of about 3.33, 2.63, and 1.96 before failure, respectively. Fig. 7b shows that lower strains were recorded in slab G-175-H compared to slab G-175-N. This indicates that as the concrete strength was increased, the measured reinforcement and concrete strains decreased. At the factored design load level, the maximum measured strains in the reinforcement were 2390 and 1575 microstrains for slabs G-175-N and G-175-H, respectively. These values

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K. Bouguerra et al. / Construction and Building Materials 25 (2011) 3956–3965 Table 5 Punching strength capacity models of FRP-RC elements. Reference

Equation

JSCE [16]

Vc = bdbpbrfpcdbo;0.5dd

El-Ghandour et al. [17,18] Matthys and Taerwe [19] Ospina et al. [20] El-Gamal et al. [21] ACI 440.1R-06 [22]

pffiffiffiffi bd = (1000/d)1/4 6 1.5, bp = (100qfEf/Es)1/3 6 1.5, br = 1 + 1/(1 + 0.25u/d), fpcd = 0.2 fc0 6 1.2 MPa pffiffiffiffiEf 1=3 bo;0:5d d V c ¼ 0:33 fc0 Es V c ¼ 1:36

ð100qEf =Es fc0 Þ1=3

bo;1:5d d 1=4 d qffiffiffiffi E V c ¼ 2:77ðqf fc0 Þ1=3 Efs bo;1:5d d pffiffiffiffi V c ¼ 0:33 fc0 bo0:5d da a = 0.5(qfEf)1/3(1 + 8d/bo;0.5d) pffiffiffiffi V c ¼ 45 fc0 bo;0:5d kd qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k ¼ 2qf nf þ ðqf nf Þ2  qf nf

Note: Vc is the punching shear capacity (N); Ef is the modulus of elasticity of the FRP reinforcement (MPa); Es is the modulus of elasticity of steel (MPa); d is the effective slab depth; qf is the FRP reinforcement ratio; u is the perimeter of the loaded area (mm); fc0 is the compressive strength of the concrete (MPa); bo;0.5d is the critical perimeter at a distance of 0.5d from the column face (mm); bo;1.5d is the critical perimeter at a distance of 1.5d from the column face (mm); nf is the ratio of modulus of elasticity of FRP bars to modulus of elasticity of concrete.

Table 6 Comparison between predicted and experimental punching shear strengths. Slab

Vexp (kN)

JSCE [16] Vexp/Vpred

El-Ghandour et al. [17,18] Vexp/Vpred

Matthys and Taerwe [19] Vexp/Vpred

Ospina et al. [20] Vexp/Vpred

El-Gamal et al. [21] Vexp/Vpred

ACI 440.1R-06 [22] Vexp/Vpred

G-200-N G-175-N G-150-N G-175-H G-175-N-0.75 G-175-N-0.35 C-175-N Mean S.D. COV (%)

732 484 362 704 549 506 530

1.51 1.32 1.29 1.92 1.79 1.99 1.49 1.62 0.28 18

1.44 1.28 1.21 1.37 1.18 1.04 0.93 1.21 0.18 15

1.59 1.53 1.45 1.82 1.81 2.02 1.65 1.70 0.20 12

1.23 1.06 1.06 1.26 1.26 1.39 0.97 1.18 0.15 13

1.13 1.03 1.02 1.10 1.13 1.26 1.08 1.11 0.08 7

2.57 2.68 2.55 3.28 3.43 4.18 3.47 3.17 0.60 19

Vexp: Experimental punching strength. Vpred: Predicted punching strength.

were about 38% and 21% of the measured strains at failure (6224 and 7566 micro-strain). Fig. 7c shows that decreasing the bottom transverse reinforcement ratio significantly increased the measured reinforcement strains after cracking. At service load level, much greater strains were recorded in slab G-175-N-0.35 because it was cracked at a load less than the service load. At factored design load level, the maximum measured reinforcement strains were 2390, 2969, and 3831 micro-strains for slabs G-175-N, G-175-N-0.7 and G-175-N0.35, respectively. For the slabs of Group 4, Fig. 7d shows that almost similar strains were recorded in the FRP-reinforced concrete slabs. For slab S-175-N, slightly lower strains were recorded before yielding of the steel bars. After yielding, greater strain values were recorded. 3.4. Comparison between G-200-N and G-175-H slabs The behaviors of slabs G-200-N and G-175-H were compared in deflection, cracking, and strain as shown in Fig. 8. From this comparison it can be seen that the two slabs with the same quantity of reinforcement but with different thicknesses and fabricated using normal and high-strength concrete respectively, showed the same behavior. It may be concluded that a reduction in the deck slab thickness may be allowed by increasing the concrete strength. However, this is not reflected in the design procedures of the CSA S6-06 as previously mentioned.

3.5. Comparison between experimental and predicted punching strengths The punching shear strengths of the tested bridge deck slabs were predicted using the available models in the literature that predict the punching strength of FRP-reinforced slabs. This includes the models of the Japan Society of Civil Engineers [16], ElGhandour et al. [17,18], Matthys and Taerwe [19], Ospina et al. [20], El-Gamal et al. [21], and ACI 440.IR-06 [22] Table 5 lists these models that were used to predict the punching capacity of FRPreinforced concrete slabs tested herein. The predicted punching shear strengths were compared to the experimental values as given in Table 6. It can be noticed that the model proposed by ElGamal et al. [21] yielded good yet conservative prediction with an average ratio of Vexp/Vpred os 1.11 and corresponding coefficient of variation of 7%. On the other hand, the ACI 440.1R-06 [22] underestimates the punching strength of bridge deck slabs reinforced with FRP bars. The average ratio of Vexp/Vpred is 3.17 with a coefficient of variation of 19%. 4. Conclusions This paper presented the experimental test results of eight fullscale bridge deck slabs reinforced with FRP bars. The slabs were divided into four groups to investigate the following parameters: slab thickness, concrete compressive strength, reinforcement ratio,

K. Bouguerra et al. / Construction and Building Materials 25 (2011) 3956–3965

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and reinforcement type. The test results were presented in terms of cracking, deflection, strains in concrete and reinforcement, mode of failure and ultimate capacity. The main findings of this investigation can be summarized as follows:

Also, many thanks for Pultrall Inc. (Thetford Mines, Quebec) for generously providing the FRP materials. Special thanks to François Ntacorigira, technician at the Department of Civil Engineering, University of Sherbrooke, for his help in the fabrication and testing.

1. The entire group of tested slabs failed in punching shear failure with a very similar cracking pattern except slab G-175-H which showed more cracks on its tension face. This could be related to the use of HSC which increased the bond strength between the GFRP bars and the concrete components. 2. The bottom transverse reinforcement ratio was the main parameter affecting crack widths. Decreasing the bottom transverse reinforcement ratio significantly increased crack widths. 3. The measured strains were affected by slab thickness, concrete strength, and reinforcement ratio. For all FRP-reinforced slabs, the maximum measured strains in the bars at service load level, factored load level, and at failure were about 9%, 20%, and 66% of their ultimate strains, respectively. These maximum strains were measured in slab G-175-N-0.35 with the smallest bottom transverse reinforcement ratio. 4. The punching capacities of the tested slabs were significantly affected by the slab thickness and the concrete compressive strength. Decreasing the deck slab thickness by 12.5% and 25% decreased the punching capacities by 34% and 51%, respectively. For the slabs of Group 2, increasing the concrete compressive strength by 83% increased the punching capacity by 45%. 5. The type of bottom transverse reinforcement, when similar axial stiffness was used, did not affect cracking, deflections, strains, or punching capacities of the tested deck slabs. 6. The behavior of the slab G-200-N was very similar to that of the slab G-175-H. This means that the reduction of the deck slab thickness can be recovered by increasing the concrete strength. 7. Considering the actual mode of failure which is the punching shear failure in the design of restrained bridge deck slabs may lead to a reduction in the amount of reinforcement which will result in a reduction in initial costs of the GFRP-reinforced concrete bridge deck slabs. 8. The current ACI 440.1R-06 [22] punching shear model underestimates the punching shear strength of the tested slabs. The average ratio of Vexp/Vpred was 3.1 and the corresponding COV was 19%. On the other hand, El-Gamal et al. [21] model yielded good yet conservative prediction with an average Vexp/Vpred of 1.11 and a corresponding COV of 7%.

References

Acknowledgements The authors acknowledge the financial support received from the Natural Science and Engineering Research Council of Canada (NSERC), the Fond Quebecois pour la Recherche en Nature et Technologie (FQRNT), and Ministry of Transportation of Quebec.

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