Vulnerability Assessment of Arizona's Critical Infrastructure
9th International Conference on Short and Medium Span Bridges Calgary, Alberta, Canada, July 15-18, 2014
EXPERIMENTAL INVESTIGATION ON TRANSVERSE STRENGTH OF PL-3 BARRIERS REINFORCED WITH SAND-COATED GFRP BARS Hamidreza Khederzadeh Ph.D Candidate, Ryerson University, Canada Khaled Sennah Professor, Ryerson University, Canada ABSTRACT One of the main factors concerning service life of steel-reinforced barriers is corrosion of steel bars especially when exposed to a de-icing salt in winter times. The use of glass giber reinforcing polymer (GFRP) bars as non-corrosive material has emerged as an innovative solution to corrosion-related problems, reduce the maintenance cost and increase the service life of bridge structures. A recent cost-effective design of PL-3 bridge barrier was developed at Ryerson University incorporating high-modulus GFRP bars with headed ends. This paper presents results of fullscale static tests to collapse performed on the developed PL-3 bridge barrier to investigate the ultimate load carrying capacity to be compared with the design factored loads specified in Canadian Highway Bridge Design Code, CHBDC. The experimental ultimate load carrying capacity of the barriers was observed to be far greater than CHBDC factored design transverse load. The failure pattern was initiated by a trapezoidal crack pattern at front face of the barrier, followed by punching shear failure at the transverse load location. Based on the punching shear failure developed in the barrier wall and comparison with available punching shear equations in the literature, an empirical punching shear equation is proposed to determine the transverse load carrying capacity of PL-3 bridge barrier reinforced with GFRP bars. 1. INTRODUCTION Until recently, the installation of GFRP bars was often hampered by the fact that bent bars have to be produced in the factory since GFRP bars cannot be bent at the site. In addition, bent GFRP bars are much weaker than straight bars, due to the redirection and associated rearrangement of the fibres in the bend. As a result, number of bent GFRP bars in bridge barrier design is increased and even doubled at such locations where bar bents are required. The use of headed-end GFRP bars, shown in Fig. 1, is intended to eliminate the expensive use of bar bends. Few authors dealt with the use of fiber reinforced polymers (FRP) in concrete barriers. Maheu and Bakht (1994) developed a new barrier wall using FRP- NEFMAC grids, with connection to the deck slab by means of double-headed steel bars spaced at 300 mm. This new barrier wall system was adopted in the Canadian Highway Bridge Design Code, CHBDC (CSA, 2006a and 2006b). An extensive research program investigating the use of GFRP bars in concrete bridge barrier was carried out by El-Salakawy et al. (2003). Their study focused on comparing the overall behavior and cracking pattern of barrier walls reinforced with GFRP with that of conventional steel reinforcement under static and impact loading. Also, El-Salakawy et al. (2004) conducted pendulum impact tests on GFRP-reinforced barriers to examine their crack pattern, stresses and deflection under impact loading. CHBDC specifies that bridge barriers should be crash tested to comply with certain criteria for structural adequacy, occupant risk, and vehicle trajectory after collision. For barrier-to-deck slab anchorage, CHBDC prescribes if crash test results for the anchorage are not available, the anchorage and the deck should be designed for the maximum bending moment, shear and pullout loads that can be transmitted to them by the barrier wall. This can be achieved by the manual calculation using a generally established theory or evaluation of a full-scale prototype by static load test. For PL-3 barriers, CHBDC specifies transverse, longitudinal and vertical loads of 210, 70 and 90 kN, respectively, that can be applied simultaneously over a certain barrier length. CHBDC prescribes that transverse load shall be applied over a barrier 280-1
length of 2400 mm for PL-3 barriers. Since transverse loading creates the critical load carrying capacity, both the longitudinal and vertical loads were not considered in the design of barrier wall reinforcement and anchorage between deck slab and the barrier wall. It should be noted that CHBDC specifies a live load factor of 1.7. Thus, the design impact load on PL-3 barrier wall over 2.4 m length is 357 kN. Given the fact that the material cost of the GFRP bars differs based on the type of the bar (i.e high modulus versus standard modulus), a new GFRP-reinforced barrier wall was developed as shown in Fig. 2, incorporating the use of High Modulus (HM) bars with headed ends. High-modulus (HM) GFRP bars with sand-coated surface were considered in this study. 12M (#4) high-modulus (HM) GFRP bars of specified tensile strength of 1312 MPa, modulus of elasticity of 65.6 2.5 GPa and strain at rupture of 2%, as listed in the manufacturers catalogue, were used in barrier reinforcement. Also, 15M (#5) highmodulus (HM) GFRP bars of specified tensile strength of 1184 MPa, modulus of elasticity of 62.5 2.5 GPa and strain at rupture of 1.89% were used in barrier reinforcement. These 12M and 15M bars have cross-sectional area of 126.7 and 197.9 mm2, respectively. This paper describes the experimental results of the proposed barrier under transverse static loads.
Figure 1. View of Sliced GFRP headed bar.
Figure 2. Proposed GFRP-reinforced barrier.
2. CONSTRUCTION OF THE DEVELOPED PL-3 BARRIER A prototype bridge barrier with 27.6 m long was built with four construction joints and then was crash tested by a remote control 36000V tractor trailer at location of the second construction joint (Sennah and Khederzadeh, 2012). The barrier was constructed over a one-meter length cantilever projecting from an existing concrete foundation. At front face of the barrier, M15 GFRP bars were placed as vertical reinforcement at 300 mm spacing. At the back face of the barrier, M13 GFRP bars were placed as vertical reinforcement at 300 mm spacing as shown in Fig. 2. All horizontal reinforcement shown in Fig.2 was made of M15 GFRP bars. At barrier ends and over a length of 2700 mm, vertical bars at front face of the barrier wall were doubled by reducing bar spacing to 150 mm. Fig. 3 shows view of the barrier before and after the placement of reinforcement. The barrier wall was cast subsequent to the casting of the deck portion. The characteristic concrete compressive strength based on concrete cylinders on the day of testing was 30.9 MPa. Figure 3(a) shows view of the barrier reinforcement and part of the formwork, while Fig. 3(b) shows view of the cast barrier after concrete hardening.
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(a) View of barrier reinforcement (b) View of the constructed barrier Figure 3. Views of the constructed barrier before and after concrete casting. 3. TEST SETUPS AND SENSOR DESIGNATIONS Figure 4 shows schematic of the constructed barrier with test locations at interior and exterior regions. The barrier was loaded at the interior location first with the edge of the loading at the control joint. Then, it was loaded at the exterior location marked as (a), however, the test was not successful due to failure in the anchorage between the anchor bolts of the loading frame and the deck slab cantilever due to concrete pullout. Therefore, the second exterior location marked as (b) in Fig. 4 was selected to repeat the static test. The test setup shown in Fig. 5 consists of hydraulic jack that applied horizontal load over 2400-mm length of the wall through steel spread beams and 200-mm width trapezoidal timber plank attached to the barrier tapered surface. Before conducting the static test, the constructed barrier was instrumented with Linear Variable Displacement Transducers (LVDTs) and Potentiometers (POTs).at the loaded region as shown in Fig. 5.
Figure 4. Schematic diagram of the constructed barrier showing locations of static tests.
4. EXPERIMENTAL TEST RESULTS 4.1 Interior Location The barrier was tested under increasing monotonic load till failure. At 25-kN load increments, crack initiations and crack propagations were marked. The barrier was assumed failed when the deflection increased with no increase in the applied load. Figures 6(a) and 6(b) show view of crack patterns at front and back faces of the barrier wall, respectively. With increase in the applied load, horizontal cracks appeared at deck-wall junction as well as on the 280-3
tapered portion of the wall within 2400 mm length of the barrier. With further increase in the applied load, these flexural cracks extended diagonally outside the loading region and reached the top of the wall.
(a) At interior location (b) At exterior location Figure 5. Views of the test setup at interior and exterior locations. It should be noted that horizontal cracks represent the barrier wall behaving as a cantilever wall within the loaded region, while the diagonal cracks on each side of the load region represent a two-way slab action since the undeformed part of the barrier outside the loaded region resisted the deformation of the loaded length in addition to the cantilever action of the wall with deck slab. However, punching shear crack appeared at the top of the barrier at an applied load greater than 525 kN. With increase in the applied load, the punching shear cracks propagated through the wall thickness to the back face of the wall at an ultimate load of 654.9 kN. The sudden punching shear failure at the line load location may be attributed to the low stiffness of the GFRP bars and their linear elastic response till failure. Additionally, vertical cracks were observed at back face of the wall at load level of 500 kN. By observation, crack pattern in the wall was found in a trapezoidal shape rather than triangular shape failure pattern stipulated in AASHTO-LRFD Specifications (2012). Given the fact that the CHBDC factored design load is 357 kN, the factor of safety for such loading case is 654.9 / 357= 1.83. 4.2 Exterior Location The barrier wall was loaded at exterior location (b) shown in Figs.4 and 5. Similar trend of crack pattern with load increase at interior location was observed at exterior location. First, horizontal cracks appeared within the load region at the deck-wall junction as well as the tapered portion of the wall. With further increase in the applied load, diagonal cracks extended outside the loaded region and reached the top of the wall. Punching shear cracks were observed at top surface of the wall at a load of 525 kN. These cracks propagated through the wall thickness and continued to appear at back face of the wall. Figures 6(c) and 6(d) show views of the cracks pattern at front and back faces of the barrier wall after failure. The barrier wall failed at a load of 541 kN. When considering the CHBDC design factored load of 357 kN, a factor of safety of 1.52 in design was obtained. It can also be noticed that crack pattern contradicts with that specified in AASHTO-LRFD Specifications (2012), where the exterior portion of the barrier wall, approximately half the loaded length, acted as a cantilever while the rest of the loaded length acted as a two-way slab transferring the load to both the deck slab and the unloaded length of the barrier wall.
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Punching shear
(a) Front face- interior location
(b) Back face- interior location
Punching shear Punching shear
(c) Front face - exterior location (d) Back face- exterior location Figure 6. View of crack patterns in the tested barrier wall.
5. ANALYTICAL INVESTIGATION OF PUNCHING SHEAR OF PL-3 BARRIERS Experimental test results on the developed PL-3 at interior and exterior location showed the failure of the barrier by punching shear. Therefore, a punching shear strength equation is required to qualify PL-3 barrier wall design, in terms of barrier wall dimensions and amount and materials characteristics of GFRP bars. This section intends to correlate the experimental findings with the available punching shear strength equations in the literature. Due to differences in mechanical properties of steel and GFRP bars, punching shear equations derived for steelreinforced concrete structures cannot be employed directly to the GFRP-reinforced elements. Most of the current code provisions and empirical equations predicting punching shear strength of FRP-reinforced structures are modified forms of those available for steel-reinforced concrete structures to account for lower modulus of elasticity of FRP bars compared to steel bars. Experimental tests have shown that FRP-reinforced concrete member experienced reduced shear strength compared to steel-reinforced structures due to lower modulus of elasticity of FRP bars. The lower modulus of elasticity in turn results in larger deformation and developing wider and deeper cracks. In two-way reinforced concrete slabs, punching shear resistance is provided by the shear resistance of concrete in the compression zone, Vc. The shear resistance acts over an area equal to the critical perimeter, bₒ, of punching shear failure plane multiplied by effective depth, d, of the concrete section. The critical perimeter, bₒ, is specified in different design codes as either 0.5d or 1.5d. The FRP design codes, CSA-S806-12 (CSA, 2012), ACI 440-1R-06 (ACI, 2006) and JSCE Guidelines (1997) and other empirical punching shear equations developed by researchers (El-Ghandour et al. 1999, and 2000, Mattys and Taerwe, 2000, Ospina et al., 2003, El-Gamal et al., 2005 and Jacobson el al., 2005) considered the FRP flexural 280-5
reinforcement ratio in calculating punching shear strength of FRP-reinforced concrete slabs. A number of design standards and punching shear models provide design equations applicable to FRP-reinforced concrete slabs. However, the punching shear behavior of bridge barrier walls under the applied transverse loads has not yet been studied. As such, an attempt was made in this paper to determine the best punching shear prediction model for PL-3 GFRP-reinforced bridge barrier wall tested in this study. The following punching shear models have been selected to predict the punching shear capacities of barrier wall, which can then be compared to the test ultimate punching shear loads.
CSA-S806-12 The Canadian Standard “Design and construction of building structures with fibre reinforced polymers,” CSAS806-12, specifies the punching shear strength of FRP-reinforced concrete as the smallest of the following three equations. It can be noticed that these equations are the modified forms of those specified in the Canadian Standard “Design of Concrete Structures,” CSA-A23.3-04, to account for the FRP-reinforcing bar ratio. [1] Vc = (1 + 2 / βc).0.028λφc (Ef.ρf.fʹc) 1/3. bₒ,0.5d.d [2] Vc = [(αs.d / bₒ, 0.5d) + 0.19]. 0.147λφc (Ef.ρf.fʹc) 1/3. bₒ,0.5d.d [3] Vc = 0.056λφc (Ef.ρf.fʹc) 1/3. bₒ,0.5d.d Where, βc is the ratio of long side to short side of the concentrated load or loading patch, λ is a density factor (i.e for normal density concrete is equal to 1), φc is the concrete resistance factor, Ef is modulus of elasticity of FRP bars, ρf is the FRP tensile reinforcement ratio, fʹc is the concrete compressive strength in MPa, bₒ, 0.5d is the critical perimeter length measured at 0.5d from the loading patch, d is effective slab depth in (mm) and α s is a factor to adjust Vc for support dimensions that is equal to 4 for interior columns, 3 for edge columns and 2 for corner columns.
ACI 440-1R-06 The American Standard “Guide for the design and construction of structural concrete reinforced with FRP bars,” ACI 440-1R-06, specified the equation below for calculating punching shear capacity of FRP- reinforced concrete slab; [4] Vc = (5k/2)0.33. √fʹc . bₒ,0.5d.d Where, k =[
- ρf.nf] and nf is modular ratio equal to (Ef /Ec).
JSCE-1997 The Japanese Standard “Recommendation for Design and Construction of Concrete Structures Using Continuous Fiber Reinforcing Materials,” (JSCE, 1997) specifies that the punching shear strength can be determined from the following equation; [5] Vc = βd. βp. βr. (fpcd/ γb). bₒ,0.5d.d Where, βd =
1.5 (d in mm), βp = (100 ρf Ef / Es)1/3
1.5, βr = 1 +
; uₒ is the perimeter of
concentrated load area, fpcd = 0.2√ fʹc 1.2 in MPa and γb is a partial safety factor to account for concrete compressive strengths below 50 MPa (1.3) and above 50 MPa (1.5). However, γ b was set equal to 1 to determine an un-factored prediction capacity to be compared with experimental ultimate strength.
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El- Ghandour-1999 From the experimental tests performed on FRP-reinforced flat slabs, El-Ghandour et al. (1999) proposed a modification to the punching shear strength specified in the ACI code “Building Code Requirements for Structural Concrete,” ACI- 318-05, (ACI, 2005) by applying the term (E f/Es)1/3 to the predicted punching shear strength as follow; [6] Vc = 0.33√fʹc . (Ef/Es)1/3 bₒ,0.5d.d
El- Ghandour-2000 El- Ghandour et al. (2000) proposed a modification to the British Code “Structural use of concrete,” BS 8110-1, (British, 2002) by applying strain correction factor (0.0045/ε y) to the equivalent reinforcement ratio (ρs = ρfEf/Es) so that a strain limit of 0.0045 is assumed for FRP reinforcements. ε y is the yield strain of steel reinforcements typically equal 0.002. Therefore, El- Ghandour et al. proposed the following equation to determine the punching shear strength of FRP-reinforced concrete slabs. [7] Vc = 0.79[100 ρf(Ef/Es). (0.0045/εy)]1/3. (fcu/25)1/3(400/d)1/4 bₒ,1.5d.d Where, fcu is the concrete cube strength equal to (fcu = fʹc/0.8 MPa) and bₒ,1.5d is the critical perimeter length measured at a distance 1.5d away from the loading patch.
Mattys and Taerwe (2000) Mattys and Taerwe (2000) conducted experimental study on punching shear strength of concrete slabs reinforced with FRP grids. From the experimental investigations, they proposed the following modification to the provisions of BS- 8110-1 Standard to account for the use of FRP bars. [8] Vc = 1.36[100 ρf(Ef/Es)]1/3(fʹc)1/3(1/d)1/4. bₒ,1.5d.d
Ospina et al. (2003) Ospina et al. performed experimental tests on flat slabs reinforced with FRP bars and grids and proposed a modification to the punching shear strength suggested by Mattys and Taerwer (2000) as follow. [9] Vc = 2.77 (ρf fʹc)1/3(Ef/Es)]1/2 bₒ,1.5d.d
El-Gamal el al. (2005) El- Gamal et al. (2005) proposed modification to the ACI 318-05 punching shear equation by applying a new parameter (α) as follows. [10] Vc = 0.33√fʹc .bₒ,0.5d.d.α Where, α = 0.5(ρf Ef)1/3.(1 + 8d/ bₒ,0.5d) and Ef is in GPa.
Jacobson et al. (2005) Jaconson el al. (2005) conducted experimental investigation on punching shear capacity of double layer FRP gridreinforced slabs and proposed a new model which is a modification of the empirical approach suggested by Mattys andTaerwe (2000) as follows.
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[11] Vc = 4.5(ρf fʹc)1/3. (1/d)1/4. bₒ,1.5d.d Where, ρf is to be calculated as the average of the two reinforcement ratios in both longitudinal and transverse directions. All above-mentioned prediction models consider the reinforcement ratio of tension reinforcement in the direction of applied tension force for edge loading, except JSCE-1997 and Jacobson el al. (2005) that considers the average reinforcement ratios in both directions at tension face of slab. The punching shear strength of the developed GFRPreinforced PL-3 barriers in this study was calculated using various equations mentioned earlier at the interior and exterior locations. Table 1 provides the predicted capacities of the PL-3 barrier in accordance with above-mentioned punching shear equations, Vc, along with the ratio between the experimental and predicted punching shear strength (Vc,test/Vc). Traditionally, the Vc,test/Vc ratio of 1 presents perfectly predicted test capacity, while ratios greater than 1 provide conservatism in designing punching shear capacity of the barrier walls. Ratios less than 1 show that the predicted punching shear models overestimate the experimental shear capacity of the barrier walls making the design unsafe. From the punching shear prediction models reported in Table 1, the predication by Mattys and Taerwe’s equation as well as the equation by Jacobson et al. yielded reasonable punching shear strength for the tested barrier. On the other hand, CSA-S806-12 equation yielded the lowest predicted punching shear strength, followed by the ACI 440.1R-06 equation. Other equations presented in Table 1 provided unsafe prediction of the barrier punching strength at least at either the interior or exterior location.
6. PROPOSED PUNCHING SHEAT EQUATION FOR PL-3 BARRIERS The punching shear capacity models presented in the codes were empirically derived for two-way FRP-reinforced concrete slabs with simply-supported or some level of edge restraint. Table 1 provides a correlation between the experimental findings and the available punching shear equations for FRP-reinforced members. However, none of these equations were specifically derived for punching shear failure of GFRP-reinforced bridge barrier walls under transverse applied load. So, none of these equations appeared to be almost equal to those obtained experimentally as depicted in Table 1. Therefore, a new punching shear prediction model was proposed as follows (more details of experimental and analytical results can be found elsewhere (Khederzadeh and Sennah, 2014), considering the punching shear plane located at a distance 1.5d from the patch load area, where d is the effective depth of the reinforcement at the punching shear plane. [12] Vc, pro. = (1 + 2 / βc). 0.136. λ.φc. {(ρf.Ef.fʹc)1/3)/(d1/4)} . bₒ,1.5d.d Where, βc is the ratio of long side to short side of loading patch equal to (β c = Lt / W). ρf is average reinforcement ratio in x and y direction of the wall. Ef is the average modulus of elasticity of transverse and longitudinal GFRP reinforcement at tension face of the barrier wall. b ₒ,1.5d is perimeter of the punching shear plane. λ is 1 for normal density concrete, φc is concrete resistance factor and d is effective depth of wall thickness. Table 2 presents the predicted punching shear strength of the barrier wall, denoted as Vc, proposed, based on the above-mentioned proposed equation. It can be observed that the ratio between the experimental and proposed punching shear strength of the barrier wall at interior and exterior locations, Vc,test / Vc,proposed, are 1.19 and 1.11, respectively, representing a good agreement between the experimental findings and the theoretical prediction.
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Table 1 : Punching shear capacities of the tested PL-3 GFRP-reinforced bridge barriers using available design codes and previous research. Interior location* Exterior location* (Vc,test = 654.90 kN) (Vc,test = 541.20 kN) Code Designation Vc Vc,test / Vc ratio Vc Vc,test / Vc ratio CSA-S806-12 (Vc, S806) 319.30 2.05 346.60 1.56 ACI 440.1R-06 (Vc, ACI 440) 364.90 1.79 432.80 1.25 JSCE (Vc, JSCE) El-Ghandour et al. (1999), (Vc, EGA 1999) El-Ghandour el al. (2000), (Vc, EGA 2000) Mattys and Taerwe, (2000), (Vc, MT) Ospina et al. (2003), (Vc, OSP) El-Gamal et al. (Vc, EGM) Jacobson et al. (Vc, JCOB) * See Fig. 4 for test locations
637.95 854.40 561.30 447.53 613.80 551.60 506.05
1.03 0.77 1.17 1.46 1.07 1.19 1.29
702.40 736.60 558.15 445.05 610.40 630.60 446.96
0.77 0.73 0.97 1.22 0.88 0.86 1.21
Table 2 : Experimental and proposed punching shear capacities of the developed PL-3 barrier. Load location Vc,test (kN) Vc, proposed (kN) Vc,test / Vc,proposed ratio Vc,proposed / VCHBDC** Ratio Interior location* Exterior location (b)*
654.90 541.20 * See Fig. 4 for transverse load locations,
550.40 486.10
1.19 1.11
1.54 1.36
**V CHBDC = 357 kN
Table 3 : CHBDC check requirements for FRP reinforced concrete structures. Developed barrier CHBDC limit 4 Deformability, J 4.43 4 83 kN.m/m (interior location) Resisting moment, Mr (KN.m/m) 215.5 0.7 Crack Width at wall base (mm) 0.72
The ratios between the proposed punching shear strength and the 357kN factored design load, Vc,proposed / VCHBDC, is calculated as 1.54 and 1.36 for interior and exterior locations, respectively. Although the ratio between the experimental and design values for punching shear strength should be at least 1 for safe design, it may be advisable to consider a durability factor for environmental effects that would lead to degradation of GFRP bars. Per CHBDC, this factor would be 0.75 in case of GFRP bars. As such, the ratio between the experimental and design values for punching shear strength of PL-3 barrier wall should be greater than 1.33 in lieu of 1. In addition to the abovementioned punching shear requirement, CHBDC clause 16.8.2.1 requires that for concrete component reinforced with FRP, the deformability factor, J, should be greater than 4. CHBDC clause 16.8.2.3 also requires crack controls for FRP reinforced concrete when the maximum tensile strain in FRP under full service loads exceeds 0.0015. Table 3 compares the above two requirements to CHBDC limits, in addition to the correlation between the experimental moment and CHBDC factored design moment at the barrier-deck junction. Good agreement regarding safe design of the developed barrier was observed.
7. CONCLUSIONS The experimental findings showed that the failure pattern of the developed GFRP-reinforced barrier is punching shear and the major flexural crack pattern is trapezoidal in shape rather than the triangular shape stipulated in AAHSTO-LRFD Specifications. The ultimate load carrying capacities of the barrier wall at interior and exterior locations were greater than CHBDC limits with a factor of safety of 1.83 and 1.52, respectively. A punching shear prediction model was proposed for barrier wall design. The proposed equation yielded a good correlation with the 280-9
experimental test results. In addition, the developed barrier reinforcement detailing provided adequate deformability factor in accordance with the CHBDC limit. 8. ACKNOWLEDGEMENTS The authors acknowledge funding from Pultrall Inc. and Government of Quebec to support this research.
9. REFERENCES AASHTO-LRFD. “AASHTO-LRFD Bridge Design Specifications”. Third Edition, American Association of State Highway and Transportation Officials, Washington, DC, 2012. ACI 440.1R-06. 2006. “Guide for the Design and Construction of Structural Concrete Reinforced with FRP Bars”. American Concrete Institute, Farmington Hills MI, USA. CSA. 2006a. “Canadian Highway Bridge Design Code. CAN/CSA-S6-06”. Canadian Standard Association, Toronto, Ontario, Canada, 2006. CSA. 2006b. Commentary on CAN/CSA-S6-06, “Canadian Highway Bridge Design Code”. Canadian Standard Association, Toronto, Ontario, Canada, 2006. CSA-S806-12. “Design and construction of building structures with fibre reinforced polymers”. Canadian Standards Association, Rexdale, Ontario, 2012. El-Gamal, S., El-Salakawy, E., and Benmokrane, B. 2005. “A New Punching Shear Equation for Two-Way Concrete Slabs Reinforced with FRP Bars”. ACI Special Publication, SP-230-50, pp. 877-894. El-Ghandour, A., Pilakoutas, K., and Waldron, P. 1999. “New Approach for Punching Shear Capacity Prediction of Fiber Reinforced Polymer Reinforced Concrete Flat Slabs”. ACI journal, 135-144. El-Ghandour, A., Pilakoutas, K., and Waldron, P. 2000. “Punching Shear Behavior and Design of FRP RC Flat Slabs”. Proceedings of the international workshop on punching shear capacity of RC slab, Stockholm: TRITABKN Bulletin 57, p. 359– 66. El-Salakawy, E., Benmokrane, B., Masmoudi, R., Brière, F., and Breaumier, E. 2003. “Concrete Bridge Barriers Reinforced with Glass Fiber-Reinforced Polymer Composite Bars”. ACI Structural Journal, 100(6): 815–824. El-Salakawy, E., Masmoudi, R., Benmokrane, B., Briére, F., and Desgagné, G. 2004. “Pendulum Impacts into Concrete Bridge Barriers Reinforced with GFRP Composite Bars.” Canadian Journal of Civil Engineering, 31(4): 539-552. Jacobson, D., Bank, L., Oliva, M., Russell, J. 2005. “Punching Shear Capacity of Double Layer FRP Grid Reinforced Slabs”. ACI Special Publication, SP- 230-49, pp. 857-875. Japan Society of Civil Engineers (JSCE). 1997. “Recommendation for Design and Construction of Concrete Structures Using Continuous Fiber Reinforcing Materials”. Concrete Engineering Series, No. 23, 325 pp. Maheu, J., and B. Bakht. 1994. "A New Connection Between Concrete Barrier Walls and Bridge Decks", Proceedings of the 22nd CSCE Annual Conference, Winnipeg, Manitoba, Vol. II, pp 224-229. Matthys, S., and Taerwe, L. 2000. “Concrete Slabs Reinforced with FRP Grids, Part II: Punching Resistance”. ASCE Journal of Composites for Constructions, 4(3): 154-161 Ospina, C., Alexander, S., and Cheng, R. 2003. “Punching of Two-Way Concrete Slabs with Fiber-Reinforced Polymer Reinforcing Bars or Grids”. ACI structural Journal, 100(5): 589-598. Sennah, K. and Khederzadeh, H. 2012. “Crashworthiness of PL-3 Concrete Bridge Barrier Reinforced with SandCoated GFRP Bars.” Proceedings of 6th International Conference on Advanced Composite Materials in Bridges and Structures, Kingston, Ontario, pp. 1-10. Khederzadeh, H.R., Sennah, K. 2014. Development of Cost-Effective Concrete Bridge Barrier Reinforced with GFRP Bars. Static Load Test, Canadian Journal of Civil Engineering,41(4) pp. 368-379.
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EXPERIMENTAL INVESTIGATION ON TRANSVERSE STRENGTH OF PL-3 BARRIERS REINFORCED WITH SAND-COATED GFRP BARS Hamidreza Khederzadeh Ph.D Candidate, Ryerson University, Canada Khaled Sennah Professor, Ryerson University, Canada ABSTRACT One of the main factors concerning service life of steel-reinforced barriers is corrosion of steel bars especially when exposed to a de-icing salt in winter times. The use of glass giber reinforcing polymer (GFRP) bars as non-corrosive material has emerged as an innovative solution to corrosion-related problems, reduce the maintenance cost and increase the service life of bridge structures. A recent cost-effective design of PL-3 bridge barrier was developed at Ryerson University incorporating high-modulus GFRP bars with headed ends. This paper presents results of fullscale static tests to collapse performed on the developed PL-3 bridge barrier to investigate the ultimate load carrying capacity to be compared with the design factored loads specified in Canadian Highway Bridge Design Code, CHBDC. The experimental ultimate load carrying capacity of the barriers was observed to be far greater than CHBDC factored design transverse load. The failure pattern was initiated by a trapezoidal crack pattern at front face of the barrier, followed by punching shear failure at the transverse load location. Based on the punching shear failure developed in the barrier wall and comparison with available punching shear equations in the literature, an empirical punching shear equation is proposed to determine the transverse load carrying capacity of PL-3 bridge barrier reinforced with GFRP bars. 1. INTRODUCTION Until recently, the installation of GFRP bars was often hampered by the fact that bent bars have to be produced in the factory since GFRP bars cannot be bent at the site. In addition, bent GFRP bars are much weaker than straight bars, due to the redirection and associated rearrangement of the fibres in the bend. As a result, number of bent GFRP bars in bridge barrier design is increased and even doubled at such locations where bar bents are required. The use of headed-end GFRP bars, shown in Fig. 1, is intended to eliminate the expensive use of bar bends. Few authors dealt with the use of fiber reinforced polymers (FRP) in concrete barriers. Maheu and Bakht (1994) developed a new barrier wall using FRP- NEFMAC grids, with connection to the deck slab by means of double-headed steel bars spaced at 300 mm. This new barrier wall system was adopted in the Canadian Highway Bridge Design Code, CHBDC (CSA, 2006a and 2006b). An extensive research program investigating the use of GFRP bars in concrete bridge barrier was carried out by El-Salakawy et al. (2003). Their study focused on comparing the overall behavior and cracking pattern of barrier walls reinforced with GFRP with that of conventional steel reinforcement under static and impact loading. Also, El-Salakawy et al. (2004) conducted pendulum impact tests on GFRP-reinforced barriers to examine their crack pattern, stresses and deflection under impact loading. CHBDC specifies that bridge barriers should be crash tested to comply with certain criteria for structural adequacy, occupant risk, and vehicle trajectory after collision. For barrier-to-deck slab anchorage, CHBDC prescribes if crash test results for the anchorage are not available, the anchorage and the deck should be designed for the maximum bending moment, shear and pullout loads that can be transmitted to them by the barrier wall. This can be achieved by the manual calculation using a generally established theory or evaluation of a full-scale prototype by static load test. For PL-3 barriers, CHBDC specifies transverse, longitudinal and vertical loads of 210, 70 and 90 kN, respectively, that can be applied simultaneously over a certain barrier length. CHBDC prescribes that transverse load shall be applied over a barrier 280-1
length of 2400 mm for PL-3 barriers. Since transverse loading creates the critical load carrying capacity, both the longitudinal and vertical loads were not considered in the design of barrier wall reinforcement and anchorage between deck slab and the barrier wall. It should be noted that CHBDC specifies a live load factor of 1.7. Thus, the design impact load on PL-3 barrier wall over 2.4 m length is 357 kN. Given the fact that the material cost of the GFRP bars differs based on the type of the bar (i.e high modulus versus standard modulus), a new GFRP-reinforced barrier wall was developed as shown in Fig. 2, incorporating the use of High Modulus (HM) bars with headed ends. High-modulus (HM) GFRP bars with sand-coated surface were considered in this study. 12M (#4) high-modulus (HM) GFRP bars of specified tensile strength of 1312 MPa, modulus of elasticity of 65.6 2.5 GPa and strain at rupture of 2%, as listed in the manufacturers catalogue, were used in barrier reinforcement. Also, 15M (#5) highmodulus (HM) GFRP bars of specified tensile strength of 1184 MPa, modulus of elasticity of 62.5 2.5 GPa and strain at rupture of 1.89% were used in barrier reinforcement. These 12M and 15M bars have cross-sectional area of 126.7 and 197.9 mm2, respectively. This paper describes the experimental results of the proposed barrier under transverse static loads.
Figure 1. View of Sliced GFRP headed bar.
Figure 2. Proposed GFRP-reinforced barrier.
2. CONSTRUCTION OF THE DEVELOPED PL-3 BARRIER A prototype bridge barrier with 27.6 m long was built with four construction joints and then was crash tested by a remote control 36000V tractor trailer at location of the second construction joint (Sennah and Khederzadeh, 2012). The barrier was constructed over a one-meter length cantilever projecting from an existing concrete foundation. At front face of the barrier, M15 GFRP bars were placed as vertical reinforcement at 300 mm spacing. At the back face of the barrier, M13 GFRP bars were placed as vertical reinforcement at 300 mm spacing as shown in Fig. 2. All horizontal reinforcement shown in Fig.2 was made of M15 GFRP bars. At barrier ends and over a length of 2700 mm, vertical bars at front face of the barrier wall were doubled by reducing bar spacing to 150 mm. Fig. 3 shows view of the barrier before and after the placement of reinforcement. The barrier wall was cast subsequent to the casting of the deck portion. The characteristic concrete compressive strength based on concrete cylinders on the day of testing was 30.9 MPa. Figure 3(a) shows view of the barrier reinforcement and part of the formwork, while Fig. 3(b) shows view of the cast barrier after concrete hardening.
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(a) View of barrier reinforcement (b) View of the constructed barrier Figure 3. Views of the constructed barrier before and after concrete casting. 3. TEST SETUPS AND SENSOR DESIGNATIONS Figure 4 shows schematic of the constructed barrier with test locations at interior and exterior regions. The barrier was loaded at the interior location first with the edge of the loading at the control joint. Then, it was loaded at the exterior location marked as (a), however, the test was not successful due to failure in the anchorage between the anchor bolts of the loading frame and the deck slab cantilever due to concrete pullout. Therefore, the second exterior location marked as (b) in Fig. 4 was selected to repeat the static test. The test setup shown in Fig. 5 consists of hydraulic jack that applied horizontal load over 2400-mm length of the wall through steel spread beams and 200-mm width trapezoidal timber plank attached to the barrier tapered surface. Before conducting the static test, the constructed barrier was instrumented with Linear Variable Displacement Transducers (LVDTs) and Potentiometers (POTs).at the loaded region as shown in Fig. 5.
Figure 4. Schematic diagram of the constructed barrier showing locations of static tests.
4. EXPERIMENTAL TEST RESULTS 4.1 Interior Location The barrier was tested under increasing monotonic load till failure. At 25-kN load increments, crack initiations and crack propagations were marked. The barrier was assumed failed when the deflection increased with no increase in the applied load. Figures 6(a) and 6(b) show view of crack patterns at front and back faces of the barrier wall, respectively. With increase in the applied load, horizontal cracks appeared at deck-wall junction as well as on the 280-3
tapered portion of the wall within 2400 mm length of the barrier. With further increase in the applied load, these flexural cracks extended diagonally outside the loading region and reached the top of the wall.
(a) At interior location (b) At exterior location Figure 5. Views of the test setup at interior and exterior locations. It should be noted that horizontal cracks represent the barrier wall behaving as a cantilever wall within the loaded region, while the diagonal cracks on each side of the load region represent a two-way slab action since the undeformed part of the barrier outside the loaded region resisted the deformation of the loaded length in addition to the cantilever action of the wall with deck slab. However, punching shear crack appeared at the top of the barrier at an applied load greater than 525 kN. With increase in the applied load, the punching shear cracks propagated through the wall thickness to the back face of the wall at an ultimate load of 654.9 kN. The sudden punching shear failure at the line load location may be attributed to the low stiffness of the GFRP bars and their linear elastic response till failure. Additionally, vertical cracks were observed at back face of the wall at load level of 500 kN. By observation, crack pattern in the wall was found in a trapezoidal shape rather than triangular shape failure pattern stipulated in AASHTO-LRFD Specifications (2012). Given the fact that the CHBDC factored design load is 357 kN, the factor of safety for such loading case is 654.9 / 357= 1.83. 4.2 Exterior Location The barrier wall was loaded at exterior location (b) shown in Figs.4 and 5. Similar trend of crack pattern with load increase at interior location was observed at exterior location. First, horizontal cracks appeared within the load region at the deck-wall junction as well as the tapered portion of the wall. With further increase in the applied load, diagonal cracks extended outside the loaded region and reached the top of the wall. Punching shear cracks were observed at top surface of the wall at a load of 525 kN. These cracks propagated through the wall thickness and continued to appear at back face of the wall. Figures 6(c) and 6(d) show views of the cracks pattern at front and back faces of the barrier wall after failure. The barrier wall failed at a load of 541 kN. When considering the CHBDC design factored load of 357 kN, a factor of safety of 1.52 in design was obtained. It can also be noticed that crack pattern contradicts with that specified in AASHTO-LRFD Specifications (2012), where the exterior portion of the barrier wall, approximately half the loaded length, acted as a cantilever while the rest of the loaded length acted as a two-way slab transferring the load to both the deck slab and the unloaded length of the barrier wall.
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Punching shear
(a) Front face- interior location
(b) Back face- interior location
Punching shear Punching shear
(c) Front face - exterior location (d) Back face- exterior location Figure 6. View of crack patterns in the tested barrier wall.
5. ANALYTICAL INVESTIGATION OF PUNCHING SHEAR OF PL-3 BARRIERS Experimental test results on the developed PL-3 at interior and exterior location showed the failure of the barrier by punching shear. Therefore, a punching shear strength equation is required to qualify PL-3 barrier wall design, in terms of barrier wall dimensions and amount and materials characteristics of GFRP bars. This section intends to correlate the experimental findings with the available punching shear strength equations in the literature. Due to differences in mechanical properties of steel and GFRP bars, punching shear equations derived for steelreinforced concrete structures cannot be employed directly to the GFRP-reinforced elements. Most of the current code provisions and empirical equations predicting punching shear strength of FRP-reinforced structures are modified forms of those available for steel-reinforced concrete structures to account for lower modulus of elasticity of FRP bars compared to steel bars. Experimental tests have shown that FRP-reinforced concrete member experienced reduced shear strength compared to steel-reinforced structures due to lower modulus of elasticity of FRP bars. The lower modulus of elasticity in turn results in larger deformation and developing wider and deeper cracks. In two-way reinforced concrete slabs, punching shear resistance is provided by the shear resistance of concrete in the compression zone, Vc. The shear resistance acts over an area equal to the critical perimeter, bₒ, of punching shear failure plane multiplied by effective depth, d, of the concrete section. The critical perimeter, bₒ, is specified in different design codes as either 0.5d or 1.5d. The FRP design codes, CSA-S806-12 (CSA, 2012), ACI 440-1R-06 (ACI, 2006) and JSCE Guidelines (1997) and other empirical punching shear equations developed by researchers (El-Ghandour et al. 1999, and 2000, Mattys and Taerwe, 2000, Ospina et al., 2003, El-Gamal et al., 2005 and Jacobson el al., 2005) considered the FRP flexural 280-5
reinforcement ratio in calculating punching shear strength of FRP-reinforced concrete slabs. A number of design standards and punching shear models provide design equations applicable to FRP-reinforced concrete slabs. However, the punching shear behavior of bridge barrier walls under the applied transverse loads has not yet been studied. As such, an attempt was made in this paper to determine the best punching shear prediction model for PL-3 GFRP-reinforced bridge barrier wall tested in this study. The following punching shear models have been selected to predict the punching shear capacities of barrier wall, which can then be compared to the test ultimate punching shear loads.
CSA-S806-12 The Canadian Standard “Design and construction of building structures with fibre reinforced polymers,” CSAS806-12, specifies the punching shear strength of FRP-reinforced concrete as the smallest of the following three equations. It can be noticed that these equations are the modified forms of those specified in the Canadian Standard “Design of Concrete Structures,” CSA-A23.3-04, to account for the FRP-reinforcing bar ratio. [1] Vc = (1 + 2 / βc).0.028λφc (Ef.ρf.fʹc) 1/3. bₒ,0.5d.d [2] Vc = [(αs.d / bₒ, 0.5d) + 0.19]. 0.147λφc (Ef.ρf.fʹc) 1/3. bₒ,0.5d.d [3] Vc = 0.056λφc (Ef.ρf.fʹc) 1/3. bₒ,0.5d.d Where, βc is the ratio of long side to short side of the concentrated load or loading patch, λ is a density factor (i.e for normal density concrete is equal to 1), φc is the concrete resistance factor, Ef is modulus of elasticity of FRP bars, ρf is the FRP tensile reinforcement ratio, fʹc is the concrete compressive strength in MPa, bₒ, 0.5d is the critical perimeter length measured at 0.5d from the loading patch, d is effective slab depth in (mm) and α s is a factor to adjust Vc for support dimensions that is equal to 4 for interior columns, 3 for edge columns and 2 for corner columns.
ACI 440-1R-06 The American Standard “Guide for the design and construction of structural concrete reinforced with FRP bars,” ACI 440-1R-06, specified the equation below for calculating punching shear capacity of FRP- reinforced concrete slab; [4] Vc = (5k/2)0.33. √fʹc . bₒ,0.5d.d Where, k =[
- ρf.nf] and nf is modular ratio equal to (Ef /Ec).
JSCE-1997 The Japanese Standard “Recommendation for Design and Construction of Concrete Structures Using Continuous Fiber Reinforcing Materials,” (JSCE, 1997) specifies that the punching shear strength can be determined from the following equation; [5] Vc = βd. βp. βr. (fpcd/ γb). bₒ,0.5d.d Where, βd =
1.5 (d in mm), βp = (100 ρf Ef / Es)1/3
1.5, βr = 1 +
; uₒ is the perimeter of
concentrated load area, fpcd = 0.2√ fʹc 1.2 in MPa and γb is a partial safety factor to account for concrete compressive strengths below 50 MPa (1.3) and above 50 MPa (1.5). However, γ b was set equal to 1 to determine an un-factored prediction capacity to be compared with experimental ultimate strength.
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El- Ghandour-1999 From the experimental tests performed on FRP-reinforced flat slabs, El-Ghandour et al. (1999) proposed a modification to the punching shear strength specified in the ACI code “Building Code Requirements for Structural Concrete,” ACI- 318-05, (ACI, 2005) by applying the term (E f/Es)1/3 to the predicted punching shear strength as follow; [6] Vc = 0.33√fʹc . (Ef/Es)1/3 bₒ,0.5d.d
El- Ghandour-2000 El- Ghandour et al. (2000) proposed a modification to the British Code “Structural use of concrete,” BS 8110-1, (British, 2002) by applying strain correction factor (0.0045/ε y) to the equivalent reinforcement ratio (ρs = ρfEf/Es) so that a strain limit of 0.0045 is assumed for FRP reinforcements. ε y is the yield strain of steel reinforcements typically equal 0.002. Therefore, El- Ghandour et al. proposed the following equation to determine the punching shear strength of FRP-reinforced concrete slabs. [7] Vc = 0.79[100 ρf(Ef/Es). (0.0045/εy)]1/3. (fcu/25)1/3(400/d)1/4 bₒ,1.5d.d Where, fcu is the concrete cube strength equal to (fcu = fʹc/0.8 MPa) and bₒ,1.5d is the critical perimeter length measured at a distance 1.5d away from the loading patch.
Mattys and Taerwe (2000) Mattys and Taerwe (2000) conducted experimental study on punching shear strength of concrete slabs reinforced with FRP grids. From the experimental investigations, they proposed the following modification to the provisions of BS- 8110-1 Standard to account for the use of FRP bars. [8] Vc = 1.36[100 ρf(Ef/Es)]1/3(fʹc)1/3(1/d)1/4. bₒ,1.5d.d
Ospina et al. (2003) Ospina et al. performed experimental tests on flat slabs reinforced with FRP bars and grids and proposed a modification to the punching shear strength suggested by Mattys and Taerwer (2000) as follow. [9] Vc = 2.77 (ρf fʹc)1/3(Ef/Es)]1/2 bₒ,1.5d.d
El-Gamal el al. (2005) El- Gamal et al. (2005) proposed modification to the ACI 318-05 punching shear equation by applying a new parameter (α) as follows. [10] Vc = 0.33√fʹc .bₒ,0.5d.d.α Where, α = 0.5(ρf Ef)1/3.(1 + 8d/ bₒ,0.5d) and Ef is in GPa.
Jacobson et al. (2005) Jaconson el al. (2005) conducted experimental investigation on punching shear capacity of double layer FRP gridreinforced slabs and proposed a new model which is a modification of the empirical approach suggested by Mattys andTaerwe (2000) as follows.
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[11] Vc = 4.5(ρf fʹc)1/3. (1/d)1/4. bₒ,1.5d.d Where, ρf is to be calculated as the average of the two reinforcement ratios in both longitudinal and transverse directions. All above-mentioned prediction models consider the reinforcement ratio of tension reinforcement in the direction of applied tension force for edge loading, except JSCE-1997 and Jacobson el al. (2005) that considers the average reinforcement ratios in both directions at tension face of slab. The punching shear strength of the developed GFRPreinforced PL-3 barriers in this study was calculated using various equations mentioned earlier at the interior and exterior locations. Table 1 provides the predicted capacities of the PL-3 barrier in accordance with above-mentioned punching shear equations, Vc, along with the ratio between the experimental and predicted punching shear strength (Vc,test/Vc). Traditionally, the Vc,test/Vc ratio of 1 presents perfectly predicted test capacity, while ratios greater than 1 provide conservatism in designing punching shear capacity of the barrier walls. Ratios less than 1 show that the predicted punching shear models overestimate the experimental shear capacity of the barrier walls making the design unsafe. From the punching shear prediction models reported in Table 1, the predication by Mattys and Taerwe’s equation as well as the equation by Jacobson et al. yielded reasonable punching shear strength for the tested barrier. On the other hand, CSA-S806-12 equation yielded the lowest predicted punching shear strength, followed by the ACI 440.1R-06 equation. Other equations presented in Table 1 provided unsafe prediction of the barrier punching strength at least at either the interior or exterior location.
6. PROPOSED PUNCHING SHEAT EQUATION FOR PL-3 BARRIERS The punching shear capacity models presented in the codes were empirically derived for two-way FRP-reinforced concrete slabs with simply-supported or some level of edge restraint. Table 1 provides a correlation between the experimental findings and the available punching shear equations for FRP-reinforced members. However, none of these equations were specifically derived for punching shear failure of GFRP-reinforced bridge barrier walls under transverse applied load. So, none of these equations appeared to be almost equal to those obtained experimentally as depicted in Table 1. Therefore, a new punching shear prediction model was proposed as follows (more details of experimental and analytical results can be found elsewhere (Khederzadeh and Sennah, 2014), considering the punching shear plane located at a distance 1.5d from the patch load area, where d is the effective depth of the reinforcement at the punching shear plane. [12] Vc, pro. = (1 + 2 / βc). 0.136. λ.φc. {(ρf.Ef.fʹc)1/3)/(d1/4)} . bₒ,1.5d.d Where, βc is the ratio of long side to short side of loading patch equal to (β c = Lt / W). ρf is average reinforcement ratio in x and y direction of the wall. Ef is the average modulus of elasticity of transverse and longitudinal GFRP reinforcement at tension face of the barrier wall. b ₒ,1.5d is perimeter of the punching shear plane. λ is 1 for normal density concrete, φc is concrete resistance factor and d is effective depth of wall thickness. Table 2 presents the predicted punching shear strength of the barrier wall, denoted as Vc, proposed, based on the above-mentioned proposed equation. It can be observed that the ratio between the experimental and proposed punching shear strength of the barrier wall at interior and exterior locations, Vc,test / Vc,proposed, are 1.19 and 1.11, respectively, representing a good agreement between the experimental findings and the theoretical prediction.
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Table 1 : Punching shear capacities of the tested PL-3 GFRP-reinforced bridge barriers using available design codes and previous research. Interior location* Exterior location* (Vc,test = 654.90 kN) (Vc,test = 541.20 kN) Code Designation Vc Vc,test / Vc ratio Vc Vc,test / Vc ratio CSA-S806-12 (Vc, S806) 319.30 2.05 346.60 1.56 ACI 440.1R-06 (Vc, ACI 440) 364.90 1.79 432.80 1.25 JSCE (Vc, JSCE) El-Ghandour et al. (1999), (Vc, EGA 1999) El-Ghandour el al. (2000), (Vc, EGA 2000) Mattys and Taerwe, (2000), (Vc, MT) Ospina et al. (2003), (Vc, OSP) El-Gamal et al. (Vc, EGM) Jacobson et al. (Vc, JCOB) * See Fig. 4 for test locations
637.95 854.40 561.30 447.53 613.80 551.60 506.05
1.03 0.77 1.17 1.46 1.07 1.19 1.29
702.40 736.60 558.15 445.05 610.40 630.60 446.96
0.77 0.73 0.97 1.22 0.88 0.86 1.21
Table 2 : Experimental and proposed punching shear capacities of the developed PL-3 barrier. Load location Vc,test (kN) Vc, proposed (kN) Vc,test / Vc,proposed ratio Vc,proposed / VCHBDC** Ratio Interior location* Exterior location (b)*
654.90 541.20 * See Fig. 4 for transverse load locations,
550.40 486.10
1.19 1.11
1.54 1.36
**V CHBDC = 357 kN
Table 3 : CHBDC check requirements for FRP reinforced concrete structures. Developed barrier CHBDC limit 4 Deformability, J 4.43 4 83 kN.m/m (interior location) Resisting moment, Mr (KN.m/m) 215.5 0.7 Crack Width at wall base (mm) 0.72
The ratios between the proposed punching shear strength and the 357kN factored design load, Vc,proposed / VCHBDC, is calculated as 1.54 and 1.36 for interior and exterior locations, respectively. Although the ratio between the experimental and design values for punching shear strength should be at least 1 for safe design, it may be advisable to consider a durability factor for environmental effects that would lead to degradation of GFRP bars. Per CHBDC, this factor would be 0.75 in case of GFRP bars. As such, the ratio between the experimental and design values for punching shear strength of PL-3 barrier wall should be greater than 1.33 in lieu of 1. In addition to the abovementioned punching shear requirement, CHBDC clause 16.8.2.1 requires that for concrete component reinforced with FRP, the deformability factor, J, should be greater than 4. CHBDC clause 16.8.2.3 also requires crack controls for FRP reinforced concrete when the maximum tensile strain in FRP under full service loads exceeds 0.0015. Table 3 compares the above two requirements to CHBDC limits, in addition to the correlation between the experimental moment and CHBDC factored design moment at the barrier-deck junction. Good agreement regarding safe design of the developed barrier was observed.
7. CONCLUSIONS The experimental findings showed that the failure pattern of the developed GFRP-reinforced barrier is punching shear and the major flexural crack pattern is trapezoidal in shape rather than the triangular shape stipulated in AAHSTO-LRFD Specifications. The ultimate load carrying capacities of the barrier wall at interior and exterior locations were greater than CHBDC limits with a factor of safety of 1.83 and 1.52, respectively. A punching shear prediction model was proposed for barrier wall design. The proposed equation yielded a good correlation with the 280-9
experimental test results. In addition, the developed barrier reinforcement detailing provided adequate deformability factor in accordance with the CHBDC limit. 8. ACKNOWLEDGEMENTS The authors acknowledge funding from Pultrall Inc. and Government of Quebec to support this research.
9. REFERENCES AASHTO-LRFD. “AASHTO-LRFD Bridge Design Specifications”. Third Edition, American Association of State Highway and Transportation Officials, Washington, DC, 2012. ACI 440.1R-06. 2006. “Guide for the Design and Construction of Structural Concrete Reinforced with FRP Bars”. American Concrete Institute, Farmington Hills MI, USA. CSA. 2006a. “Canadian Highway Bridge Design Code. CAN/CSA-S6-06”. Canadian Standard Association, Toronto, Ontario, Canada, 2006. CSA. 2006b. Commentary on CAN/CSA-S6-06, “Canadian Highway Bridge Design Code”. Canadian Standard Association, Toronto, Ontario, Canada, 2006. CSA-S806-12. “Design and construction of building structures with fibre reinforced polymers”. Canadian Standards Association, Rexdale, Ontario, 2012. El-Gamal, S., El-Salakawy, E., and Benmokrane, B. 2005. “A New Punching Shear Equation for Two-Way Concrete Slabs Reinforced with FRP Bars”. ACI Special Publication, SP-230-50, pp. 877-894. El-Ghandour, A., Pilakoutas, K., and Waldron, P. 1999. “New Approach for Punching Shear Capacity Prediction of Fiber Reinforced Polymer Reinforced Concrete Flat Slabs”. ACI journal, 135-144. El-Ghandour, A., Pilakoutas, K., and Waldron, P. 2000. “Punching Shear Behavior and Design of FRP RC Flat Slabs”. Proceedings of the international workshop on punching shear capacity of RC slab, Stockholm: TRITABKN Bulletin 57, p. 359– 66. El-Salakawy, E., Benmokrane, B., Masmoudi, R., Brière, F., and Breaumier, E. 2003. “Concrete Bridge Barriers Reinforced with Glass Fiber-Reinforced Polymer Composite Bars”. ACI Structural Journal, 100(6): 815–824. El-Salakawy, E., Masmoudi, R., Benmokrane, B., Briére, F., and Desgagné, G. 2004. “Pendulum Impacts into Concrete Bridge Barriers Reinforced with GFRP Composite Bars.” Canadian Journal of Civil Engineering, 31(4): 539-552. Jacobson, D., Bank, L., Oliva, M., Russell, J. 2005. “Punching Shear Capacity of Double Layer FRP Grid Reinforced Slabs”. ACI Special Publication, SP- 230-49, pp. 857-875. Japan Society of Civil Engineers (JSCE). 1997. “Recommendation for Design and Construction of Concrete Structures Using Continuous Fiber Reinforcing Materials”. Concrete Engineering Series, No. 23, 325 pp. Maheu, J., and B. Bakht. 1994. "A New Connection Between Concrete Barrier Walls and Bridge Decks", Proceedings of the 22nd CSCE Annual Conference, Winnipeg, Manitoba, Vol. II, pp 224-229. Matthys, S., and Taerwe, L. 2000. “Concrete Slabs Reinforced with FRP Grids, Part II: Punching Resistance”. ASCE Journal of Composites for Constructions, 4(3): 154-161 Ospina, C., Alexander, S., and Cheng, R. 2003. “Punching of Two-Way Concrete Slabs with Fiber-Reinforced Polymer Reinforcing Bars or Grids”. ACI structural Journal, 100(5): 589-598. Sennah, K. and Khederzadeh, H. 2012. “Crashworthiness of PL-3 Concrete Bridge Barrier Reinforced with SandCoated GFRP Bars.” Proceedings of 6th International Conference on Advanced Composite Materials in Bridges and Structures, Kingston, Ontario, pp. 1-10. Khederzadeh, H.R., Sennah, K. 2014. Development of Cost-Effective Concrete Bridge Barrier Reinforced with GFRP Bars. Static Load Test, Canadian Journal of Civil Engineering,41(4) pp. 368-379.
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