Shear Strengthening of Full-Scale RC T-Beams with CFRP Sheets

Abdeldjelil Belarbi

Professor and Chair, Department of Civil and Environmental Engineering, University of Houston, Texas, USA

Michael Murphy

Structural Engineer, CTL Group, Washington, D.C., USA

Sang-Wook Bae

Assistant Professor, Department of Department of Civil and Environmental Engineering, Texas Tech University, Texas, USA

ABSTRACT

Many research studies have been conducted to investigate the behavior of reinforced concrete (RC) beams strengthened in shear with externally bonded fiber-reinforced polymer (FRP) composite materials over the past two decades. As a result, analytical models and design equations are available in the form of code/guidelines/specifications in the United States and other countries. However, most of the existing studies were conducted on small-scale specimens. Thus, analytical models and design equations developed from those studies may not be able to predict the shear capacity of large-scale RC beams strengthened in shear with externally bonded FRPs. This study was, therefore, performed to evaluate the performance of existing analytical models and design equations with the test results from eight full-scale RC T-beams as well as the experimental database consisting of 375 test results obtained from existing literature. The results of the full-scale tests conducted in this study indicated that FRP shear strengthening on full-scale RC beams is as effective as that observed on small-scale RC beams. The performance of the analytical models considering the effective strain concept appeared to be more reliable to predict the shear strength for both small- and large-scale beams than the ones considering non-uniform distribution of strains in FRP.

KEYWORD

shear, T-beam, FRP, strengthening

S1A02

1. INTRODUCTION Analytical models and design equations for shear strengthening with externally bonded fiber reinforced polymer (FRP) sheets are currently available around the world either in the form of code/specifications or guidelines. However, the existing models and equations have not considered all the parameters affecting the shear behavior of RC beams strengthened with FRPs; therefore, their predictions are sometimes controversial [1]. This could be primarily due to the fact that most have been developed based on a limited number of experimental results on small-scale, rectangular RC beams. A number of research studies have been conducted to evaluate the performance of reinforced concrete (RC) beams strengthened in shear with externally bonded fiber reinforced polymer (FRP) sheets over the last two decades. An experimental database consisting of 375 experimental results from 50 existing experimental studies available in literature has been reviewed by the authors [2]. The review revealed that about 97% of existing experimental studies used scaled beams with a depth of less than 508 mm. Furthermore, most studies focused on rectangular cross-sections, while actual RC bridge girders usually have a T-shaped cross section. In addition, not many tests have been conducted on members with a span length that is comparable to that of real bridge girders, and research studies that have investigated the influence of scale effect on the shear behavior of members strengthened with FRPs are very rare. The effects of pre-existing cracks were also not widely investigated while the primary use of FRP is to strengthen damaged structures. The objective of this study was, therefore, to validate and expand on the findings of previous experimental and analytical studies and to enrich the experimental database with results from experiments on full-scale bridge RC T-beams. This study was conducted as part of a National Cooperative Highway Research Program (NCHRP) Project 12-73 [3] to assess the shear strengthening of reinforced and prestressed concrete bridge girders with FRP. This paper reports only the performance of the effectiveness of shear strengthening with FRP on RC T-beams.

2. TEST PROGRAMS 2.1 Materials The concrete used was a ready-mix concrete with a 28-day target strength of 27.6 MPa. Transverse steel reinforcement (stirrups) consisted of #3 Grade 40 reinforcing bars (As = 71 mm2 and fy = 276 MPa) while all longitudinal reinforcement was Grade 60 reinforcing bars (fy = 414 MPa): #5 bars (As = 200 mm2) for compression reinforcement in the top flange, and #11 bars (As = 1006 mm2) for flexural tension reinforcement. Carbon fiber reinforced polymer (CFRP) sheets were used as external shear reinforcement and had a tensile strength of 3792 MPa, elastic modulus of 228 GPa, and ultimate strain of 0.017. The hybrid FRP plates used for mechanically anchored systems consisted of a glass and carbon hybrid pultruded strip embedded in a vinyl ester resin. These strips were 3.2 mm thick and 102 mm wide with mechanical properties provided in Table 1.

Table 1 Mechanical Properties of Hybrid FRP Plates [4]

Mechanical properties Strength/Capacity

Stress at failure [MPa] 834 Modulus of elasticity [GPa] 62

Open hole tensile strength [MPa] 834 Unconstrained bearing capacity [kN] 236

Concrete anchor bolts with the embedment length of 152 mm and diameter of 12.7 mm were used for the mechanical anchorage systems. 2.2 Test Beam Details The test beams used in this study were designed to simulate a bridge located in Troy, NY, USA. The bridge was built in 1932 and had to be strengthened in shear and flexure in 1999 [5]. Some modifications were made, however, in the actual design of test beams mainly because of the different concrete and steel from those in the old bridge and the necessity to induce shear failure rather than flexural failure. The cross-section of the test beams are presented in Figure 1. 2.3 Test Matrix Table 2 summarizes the test matrix used in this study. The denomination of the specimens indicates the stirrup spacing in inches (8 or 12),

the strengthening configuration (S90 = FRP strips at 90 degrees), the presence of mechanical anchorage (NA = no anchorage, DMA = discontinuous mechanical anchorage, SDMA = sandwich panel mechanical anchorage, HS = additional horizontal strips), and the presence of existing cracks (PC). Two different stirrup spacings were chosen to simulate moderate and low amounts of transverse reinforcement. The design using moderate transverse reinforcement with 8 in. (203 mm) stirrup spacing met the minimum requirement for transverse reinforcement specified in AASHTO LRFD Bridge Design Specifications [6]. This design also corresponds to the stirrup spacing of the Troy Bridge. The design used low transverse reinforcement, with 12 in. (305 mm) stirrup spacing, which did not meet the AASHTO-LRFD requirement and thus simulated a reduced amount of steel due to corrosion or design and construction fault. Pre-existing cracks for RC-12-S90-HS and RC-12-S90-SDMA-PC were introduced by preloading the beams up to 60 % of the anticipated un-strengthened shear capacity.

1067

178

762

934

longitudinal reinforcement(9.5 mm-diameter)

Stirrups(9.5 mm-diameter)

457 Fig.1 Cross section of the test beams

(All dimensions are in mm) 2.4 FRP Strengthening and Anchorage Installation All test beams, except for the control beams, were strengthened with one-ply CFRP strips in the form of a U-wrap with 90 degrees fiber orientation to the horizontal axis, as shown in Figure 2. The anchorage systems used in this study include discontinuous mechanical anchorage (DMA system), sandwich panel mechanical anchorage (SDMA system), and additional horizontal FRP strips (HS system). HS system was the simplest of the three anchorage

systems in which additional FRP sheets were applied continuously along the test region at the critical bond locations (i.e., at the end of vertical FRP strips) as shown in Figure 3. This system consisted of bi-directional (+45º/-45º) CFRP sheets applied immediately following application of the vertical FRP strips to ensure better bond between the two sets of CFRP sheets. The width of the CFRP sheets for HS system was designed as 178 mm based on the estimated effective bond length of the vertical FRP strips.

Table 2 Test Matrix Test Parameters

Test I.D. Strengthening Scheme

Anchorage Type Stirrups

RC-8-Control #3@8 in.

RC-12-Control None N.A.

#3@12 in. RC-8-S90-NA N.A.

RC-8-S90-DMA Strip/90

DMA #3@8 in.

RC-12-S90-NA N.A RC-12-S90-DMA

Strip/90 DMA

#3@12 in.

RC-12-S90-SDMA-PC SDMA RC-12-S90-HS-PC

Strip/90 HS

#3@12 in.

Test Region (2,740)

254127

FRPStrips

#1#2#3#4#5#6

Fig. 2 Configuration of CFRP sheet strengthening

(All dimensions are in mm) The DMA system consisted of two hybrid FRP plates bonded together with epoxy then anchored firmly in place with epoxy and concrete wedge anchors as shown in Figure 4. The SDMA system consisted of a modification to DMA system in which the ends of the vertical FRP strips are wrapped around the first FRP plate and overlapped with a second FRP plate as shown in Figure 5. This detailing forms a layered connection which crimps the ends of the vertical FRP strips providing greater resistance to slippage of the

(35.8 mm-diameter)

strips from underneath the plates. The hybrid FRP plates were used for DMA and SDMA systems.

2,740

178

Fig. 3 Configuration of HS system

(All dimensions are in mm)

Mechanical Anchor

FRP Sheet

356

102

Fig. 4 Configuration of DMA system

(All dimensions are in mm)

FRP Sheet

Wedge Anchor

CFRP Plate

(a) DMA system (b) SDMA system

Fig. 5 Details of DMA and SDMA systems

(All dimensions are in mm) 2.5 Test Set-up All test beams were tested under three-point bending as shown in Figure 6. Two actuators at the loading point pulled the beam upward while the reaction force at the reaction point produced the shear force in the test region. In addition, external shear reinforcement was installed to protect the untested region from premature failure and insure the shear failure within the test region. For this external reinforcement, two HSS steel members were placed on top and bottom of the beam, then the steel members were connected by two #14 Dywidag bars (As=1452 mm2) with a yield strength of 517 MPa.

ExternalShear Strengthening

Load Cell

274345722743

LoadingFrame

Reaction Frame

CFRP Sheet

Fig. 6 Test set-up

(All dimensions are in mm) 3. RESULTS AND DISCUSSION 3.1 Overall Behavior The control beam without FRP strengthening, RC-8-Control, failed when the web diagonal cracks propagated into the flange. Another control beam, RC-12-Control, showed the identical failure mode. FRP strengthened beams without mechanical anchorage systems, RC-8-S90-NA and RC-12-S90-NA, failed due to FRP debonding. Beams with the DMA system, RC-8-S90-DMA and RC-12-S90-DMA, failed primarily due to FRP debonding; however, the debonding of FRP was delayed due to the mechanical anchorage. The DMA system eventually failed due to the bearing failure of the vertical FRP strips around the mechanical anchor bolts, eventually exhibiting a slippage between the hybrid FRP plates and vertical FRP strips. Because of the bearing failure and slippage, the FRP strips did not reach the ultimate strength (i.e. FRP rupture was not observed). As a result, the FRP strengthened beams with the DMA system failed due to FRP debonding although the DMA system delayed the initiation of FRP debonding and increased the shear contribution of FRP sheets. To improve the performance of the DMA system, it is necessary to increase the bond between the FRP plates and FRP sheets. The FRP strengthened beam with the HS system, RC-12-S90-HS-PC, failed primarily due to FRP debonding while the HS system could delay FRP debonding. More information about the failure mode can be found in Reference [2]. Figure 7 presents the shear force versus deflection at the loading points for all test beams. As shown in Figure 7, FRP shear strengthening increased the shear capacity, and the use of anchorage systems increased the shear capacity further: the SDMA system performed best followed by the DMA system, and HS system as shown in Figure 7(b).

It is interesting to note that FRP strengthening was more effective on RC-12 series beams than on RC-8 series beams. For example, the shear strengths of RC-8-S90-NA and RC-8-S90-DMA appeared to be 25% and 38% greater than that of RC-8-Control while the shear strengths of RC-12-S90-NA and RC-12-S90-DMA showed an increase of 38% and 64% as compared to that of RC-12-Control. This was primarily due to the fact that the crack angles for RC-8 series beams were steeper than that for RC-12 series beams, and as a result, the number of stirrups bridging over the critical shear cracks decreased as the amount of stirrups increased, eventually decreasing the shear contribution of stirrups. Therefore, it is recommended to consider this interaction between external FRP strips and stirrups. Further discussion on this interaction can be found in Reference [2], in which the contribution of each shear component was calculated using the measured strain readings.

0

200

400

600

800

1000

0 20 40 60 80 100

Shea

r For

ce (k

N)

Deflection at the Loading Point (mm)

RC-8-Control(681 kN)

RC-8-S90-NA(850 kN)

RC-8-S90-DMA(943 kN)

(a) RC-8-Series

0

200

400

600

800

1000

0 20 40 60 80 100

Shea

r For

ce (k

N)

Deflection at the Loading Point (mm)

RC-12-Control(553 kN)

RC-12-S90-SDMA-PC (948 kN)RC-12-S90-DMA (907 kN)

RC-12-S90-HS-PC(829 kN)

RC-12-S90-NA (765 kN)

(b) RC-12-Series

Fig. 7 Shear Force vs. Deflection at the Loading

Points The effects of pre-existing cracks were not fully investigated in this study mainly because there was no reference beam that could be compared to

RC-12-S90-HS-PC and RC-12-S90-SDMA-PC. However, from the test results obtained in this study as shown in Figure 7(b), it can be said that the pre-existing cracks would not significantly decrease the FRP strengthening effectiveness because the shear strengths of RC-12-S90-SDMA-PC and RC-12-S90-HS-PC were still greater than control beam RC-12-Control and even the FRP strengthened beam without pre-existing cracks, RC-12-S90-NA. 3.2 Effective strain of FRP Vertical FRP strains (i.e., strains in the fiber direction) in the test region were measured using strain gages, and the results from RC-8-S90-NA, plotted at multiple load levels, are presented in Figure 8.

-0.0005

0.0005

0.0015

0.0025

0.0035

0.0045

1 2 3 4 5 6 7

FRP strip

FRP

stra

in (i

n./in

.)V=100 kips

V=120 kips

V=140 kips

V=160 kips

V=170 kips

Fig. 8 Vertical FRP strain distribution over shear

span measured by strain gages The results clearly indicate that vertical strains are not uniform over the shear span even prior to debonding of FRP. In fact, the FRP strain distribution becomes more complex after debonding, showing irregular and more complex distribution over the shear span; however, Figure 8 only shows the distribution of FRP strains at loading steps prior to the initial debonding of FRP. Regardless of this complexity, some researchers [7-10] have attempted to model the non-uniform FRP strain distribution in order to predict the strain level of FRP and the shear forces carried by each FRP sheet in the shear span. However, based on the non-uniform strain distribution, the performance of their analytical models did not appear more accurate than other analytical models using simple equations based on effective strain (stress) concepts [11]. Because of the complexity of the FRP strain distribution, most currently available analytical models express the contribution of FRP to the

shear capacity as a function of the effective strain of FRP, feε [12-19]. The effective strain is typically expressed as a fraction of the ultimate strain and represents the average strain experienced by the FRP strips at ultimate shear capacity of the strengthened member. The effective strain, feε , of all beams tested in this study was calculated using Equation 1 shown below and assuming Vf as the actual shear gain experienced by the FRP strengthened beams:

( )sin cosf fe ff

f

A f dV

sβ β= +

( )sin cosf f fe f

f

A E dsε

β β= + (1)

where 2f f f fA n t w= is the area of transverse FRP reinforcement covering two sides of the beam, nf is number of FRP plies, tf is the FRP reinforcement thickness, wf is the width of the strip, sf is the spacing between the strips (defined as the distance from the centerline of one strip to the center line of an adjacent strip), and β is the angle between the orientation of the fibers and the horizontal axis of the beam. Effective strains were calculated not only for the test beams in this study, but also for the specimens included in the experimental database. Figure 9 plots the results against the rigidity of FRP over the square root of compressive strength of concrete ( '/f f cE t f ) since the effective strains of FRPs are believed to be affected by the tensile properties of concrete as well as the rigidity of FRP when they are used for bond critical applications. In addition, since the effective strain of FRP is largely dependent on the failure mode, the effective strains drawn from the database were grouped in this figure by the failure mode of the test specimen, i.e., either as debonding or rupturing of the FRP strips. According to Figure 9, the effective strain decreases as the rigidity of FRP increases. The effective strains of the full-scale beams tested in this study are similar to those drawn from the database with respect to FRP rigidity, indicating that the effectiveness of FRP shear strengthening on full-scale RC beams with the range of '/f f cE t f values used in this study can be regarded as much as that verified through small-scale experimental results.

3.3 Comparison of Test Results with Predictions by Existing Analytical Model Existing analytical models [1, 7-10, 12-15, 18, 20-23] were evaluated for their accuracy in predicting FRP shear contribution for the beams tested in this study. Figure 10 shows the accuracy of each model by plotting the ratio of the shear contribution of FRP predicted by the models to the experimentally determined shear contribution of FRP, Vf , of each beam that was calculated based on the actual shear gain. According to the observation in Figure 10, the models based on the effective strain predicted the shear contribution of FRP within a relatively small range of error. Meanwhile, models that considered non-uniform distribution of FRP showed less accuracy; however, this result does not necessarily mean that their attempt to consider the non-uniform strain distribution is wrong. Instead, it implies that there are other parameters to be considered in the development of an improved analytical model. Overall, it can be said that the current analytical models based on the effective strain concept can be applicable to the prediction of shear contribution of FRP in full-scale RC T-beams within a reasonable error range for design purposes. 4. CONCLUSIONS The beams strengthened with FRP tested in this study showed an increase in shear strength, and the use of a mechanical anchorage system provided additional shear strength. The strain distribution on FRP over the test region was found to be non-linear, however for design purposes, an effective strain can be used to calculate the shear carried by the FRP sheets. The effective strains calculated for beams tested in this study were compared to the effective strains calculated in the same way from data in the existing experimental database. The comparison showed that the effective strains of full-scale beams tested in this study were similar in magnitude to those calculated from existing data with respect to FRP axial rigidity. Based on this similarity and the fact that the existing data reflects tests on small-scale beams, FRP shear strengthening on full-scale RC beams is as effective as that observed on small-scale beams.

0.000

0.005

0.010

0.015

0.020

0.025

0.030

0.035

0.040

0.045

0.000 0.050 0.100 0.150 0.200 0.250 0.300 0.350

Efr f/f'c1/2

e fe

FRP debonding

FRP fracture

RC-8-S90-NA

RC-8-S90-DMA

RC-12-S90-NA

RC-12-S90-DMA

RC-12-S90-SDMA-PC

RC-12-S90-HS-PC

0.004

0.006

0.004

0.007

0.011

0.008

Test Beams Effective Strain

Fig. 9 FRP effective strain versus FRP rigidity

Range of Predicted to Experimental Values for RC Beams

0.000.200.400.600.801.001.201.401.601.802.002.202.402.60

Trianta

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Malek a

nd Saa

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Khalifa

et. a

l. 199

8

Khalifa

and N

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Hutchin

son e

t. al.

Chaall

al et.

al.

Chen a

nd Ten

g

Pelleg

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nd M

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a

Zhang

and H

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Cao et

. al.

Trianta

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and A

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Monti a

nd Li

otta

Carolin

and T

aljste

n

Analytical Models

V(pr

edic

tion)

/V(e

xper

imen

tal)

Fig. 10 Range of predicted to experimental values for all the test specimens

(Note that Kalifa et al 1998 and Triantafillou and Antonpoulos are basis of the design equations in ACI 440.2R-08 [16] and fib-TG9.3 [19], respectively.)

ACKNOWLEDGEMENT The financial support from the National Cooperative Highway Research Program (NCHRP) and the National University Transportation Center (NUTC) at the Missouri University of Science and Technology is gratefully acknowledged.

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