ADVANCED COMPOSITE MATERIALS IN BRIDGES AND STRUCTURES MATÉRIAUX COMPOSITES D'AVANT‐GARDE POUR PONTS ET CHARPENTES Winnipeg, Manitoba, Canada, September 22 – 24, 2008 / 22, 23 et 24 septembre 2008 

EXPERIMENTAL AND NUMERICAL INVESTIGATION OF RC BEAMS STRENGTHENED IN BENDING WITH NEAR SURFACE MOUNTED CFRP S. M. Soliman University of Sherbrooke, Canada

E. El-Salakawy University of Manitoba, Canada

B. Benmokrane University of Sherbrooke, Canada

H. Abdel Baky University of Sherbrooke, Canada

Abstract

The use of FRPs as Near Surface Mounted (NSM) reinforcement for existing RC structures is relatively new. It has been demonstrated that the common mode of failure for RC beam strengthened with NSM-FRP bars is governed by debonding of the FRP off the concrete surface. Nevertheless, a full understanding of the debonding phenomena is somewhat lacking. As a contribution to bridge the existing gaps in the development of design and analysis approaches of NSM-FRP-strengthened RC beams, eight rectangular RC beams with a clear span of 2.6 m were tested up to failure. These specimens include various steel reinforcement ratios with different bonded length of FRP bars and the same CFRP axial stiffness. It is intended that debonding of the CFRP strips from the surrounding concrete layer is the main mode of failure of most of the tested specimens. An incremental nonlinear displacement-controlled 3D finite element (FE) analysis was used to numerically simulate the behaviour of the test beams. An appropriate bond–slip relation was adopted in the numerical simulations to characterize the behaviour of the CFRP/epoxy and epoxy/concrete interfaces. Comparisons between the FE predictions and experimental results showed very good agreement in terms of the load−deflection and load−strain relationships, ultimate capacities, and modes of failure for the tested beams.

INTRODUCTION

In the last few years, near surface mounting (NSM) technique has received more attention as an alternative for external laminates in the flexural strengthening of concrete beams. Major innovative applications have been witnessed, at the beginning of this century, in this field to overcome the problems arising when externally FRP composites are subjected to a severe environmental condition or mechanical damage [1-5]. Both the FRP bars and strips have been applied in the NSM repairing system. In this strengthening technique, longitudinal grooves are first cut into the concrete cover of concrete beams/slabs, then the FRP bars/strips are inserted into these grooves and bonded with an appropriate binding agent; typically an epoxy paste or a cement grout. The first application of this promising technique was the implementation of steel bars into slots to strengthen a bridge deck slab in Lapland, Finland in 1940s [6]. Recently, due to their superior characteristics, FRP bars/strips are used as an alternative for the steel bars [7-11]. Hassan and Rizkalla [12] investigated concrete beams strengthened in flexural using NSM-CFRP bars with different bonded length ranging from 15d, 55d, 80d and 120d, where d identifies the diameter of the strengthening CFRP bar. The beam with the smallest bonded length (15d) provided an insignificant increase in the strength due to early debonding of the CFRP bar. While, beams with a bonded length of 55d, 80d or 120d provided 19.6, 30.4 and 41.1% increase in the ultimate capacity over that of the unstrengthened one, respectively. The maximum CFRP strain was 40% of the ultimate tensile strain of the CFRP composites (cases of 80d and 120d bonded length). From results of several applications, it has been concluded that the increase in the flexural capacity of strengthened beams applying NSM approach is greater than that using external bonded technique for the same area of the FRP laminates. This is attributed to the fact that the former approach delays the debonding of the FRP laminate, and thus increases the load carrying capacity and the FRP strength utilization ratio; i.e., the ratio of the strain in the FRP at failure to its ultimate strain [13]. Accordingly, NSM technique has gained a wide acceptance worldwide as a promising substitute for external bonded laminates in the rehabilitation of concrete beams. In contradict to the unstrengthened concrete beams; the failure of NSM-FRP-strengthened beams is generally governed by separation of the concrete cover either at the end of the FRP laminate or between two intermediate cracks. These premature failures caused by debonding often limit the effectiveness of this strengthening technique and prevent these beams from attaining their ultimate flexural capacities [7]. However, these debonding failures are less likely to occur with NSM FRP strips compared to that with externally bonded FRP sheets. To date, our knowledge of the various debonding phenomena and the associated mechanics of the bond between the NSM-FRP strips and concrete are somewhat limited. Further details on the advancements related to using NSM strengthening technique can be found in various state-of-the-art reports and review papers [7]. As far as numerical modelling is concerned, Hassan and Rizkalla [12] have presented a nonlinear finite element modelling of the beam cross section using 2D analyses. Their model intended to address the stress distributions and load transfer mechanism between NSM-FRP bars and concrete. Three-dimensional nonlinear analysis has been presented using the finite element package ABAQUS in order to simulate the load−displacement relationship of NSM-FRP-strengthened beams in the work of Kange et al. [14]. Although no failure criterion has been introduced in the literature to simulate the debonding phenomena of FRP strips from the surrounding concrete layer, these numerical simulations predict with a reasonable accuracy the load−deflection profiles as claimed by their proponents. This is probably because of the fact that these models simulate beams tested in particular experimental programs and assume the debonding strain limit in the numerical simulations from the experimental data. The limited number of the finite element simulations of NSM-FRP-strengthened concrete beams is due to the lack of our knowledge on the interfacial behaviour between the FRP and concrete and the bond mechanisms involved. A failure criterion that defines the termination of the analysis according to the strain limit in the FRP is still missing. Besides, the rare of theoretical investigations on this repairing technique is probably attributed to the fact that the experimental studies are not enough to interpret the behaviour of such beams. Therefore, this paper introduces an attempt to present 3D nonlinear finite element analysis for NSM-FRP-strengthened beams. Interface elements are used to simulate the interfacial behaviour between FRP bars and surrounding concrete beams rather than assuming full bond as that available in the literature. These interface elements are aligned between the NSM-FRP bars and concrete beams in direction parallel to the beam axis. Accordingly, this research is one of the first attempt to consider the relative slip between the NSM bars and concrete beams in the numerical simulations.

In this paper, the experimental program consisting of eight full-scale RC beams is first presented. Then, the finite element analysis using the finite element package ADINA [15] is introduced to simulate the aforementioned tested beams. Finally, a parametric study is conducted to address the effect of the steel reinforcement ratio on the debonding load and strain level at FRP bars. The general objectives of this research project that have been conducted at the University of Sherbrooke through the NSERC research Chair in Innovative FRP Composite Materials for Infrastructure are: (1) to use a NSM system composed of carbon FRP V-ROD bars manufactured by Pultrall Inc. [16] and adhesives manufactured by Hilti Inc. [17] as a compatible strengthening system; (2) to study the flexural behaviour of concrete beams strengthened with NSM-CFRP bars, in terms of cracking, deflection, carrying capacity, and mode of failure, having different bonded lengths; and (3) to investigate the effect of steel reinforcement ratio on such behaviour. EXPERIMENTAL PROGRAM In this study, experimental results of eight RC beams simply-supported over a clear span of 2600 mm were constructed and tested to failure. These beams are categorized in two series; namely, A and B. The tested beam specimens have rectangular cross-sections of 200 × 300 mm. The beams were strengthened with NSM-CFRP bars and tested in four-point bending with a shear span of 1000 mm, as shown in Figure 1.

Series A Series B

Figure 1 Dimension and reinforcement details of the test specimen of A, and B series

All tested specimens were constructed using a ready-mixed concrete with a targeted 28-day concrete compressive strength of 35 MPa. The actual concrete compressive and tensile strengths were determined based on standard cylinder tests (three cylinder specimens, 150 × 300 mm, for each concrete batch). The cylinders were tested at the same time of testing the specimen. The obtained concrete compressive strength ranged between 38 and 44 MPa. The average concrete tensile strength ranged between 2.9 and 3.6 MPa. Size No.10 (9.5 mm-diameter) CFRP sand-coated bars manufactured by Pultrall inc. [16], (Thetford Mines, Quebec) were used in this study. The CFRP bars were tested to obtain the tensile strength and modulus of elasticity according to the ACI 440.3R-04 [18]. Two deformed steel bars No.10M (11.3 mm-diameter) and No.15M (16 mm-diameter) were used in reinforcing the

concrete beams. The obtained properties of the steel and CFRP reinforcing bars used in this study are listed in Table 1. An epoxy adhesive, type HIT RE 500 produced by Hilti Inc. [17] was used in this study. The HIT RE 500 is a high strength two-part epoxy based adhesive. This type of adhesive, which can be applied on wet or dry surfaces, is specially designed for fastening into solid base materials in a wide range of material temperatures (49oC down to -5oC). The tensile strength and modulus of elasticity of the HIT RE 500 adhesive are 43.5 and 1493 MPa, respectively. Table 2 shows the details of the tested beams including the original steel and NSM-CFRP reinforcement ratios. The test beams are divided into two series, A and B, based on the internal steel reinforcement ratio.

Table 1: Mechanical properties of the reinforcing bars

Bar type Bar diameter, mm

Bar area, mm2

Modulus of elasticity, GPa

Tensile strength, MPa

Ultimate strain %

CFRP 9.5 71 122 1536 1.22

Steel (10M) 11.3 100 200 fy = 454

fu = 571 εy = 0.23

(15M) 16 200 200 fy = 460 fu = 658 εy = 0.23

Table 2: Description of test specimens

Beam ID Steel ratio Diameter of NSM bar Groove dimension Bonded length

Series A A0 Control A1

0.8% 9.5 2.00d

12d A2 24d A3 48d A4 60d

Series B B0 Control B1 0.40% 9.5 2.00d 24d B2 48d

Note: d is the diameter of the NSM bar

All the beams failed due to debonding of the NSM-CFRP strips at the vicinity of shear cracks with a separation of the concrete cover along the level of the tensile steel reinforcement bars. The debonding of the NSM bars was preceded by a typical vertical flexure cracks in the pure bending region and shear cracks in the shear zone. The concrete cover separation was initiated by internal cracking of the surrounding epoxy followed by formation of inclined and longitudinal cracks (parallel to the beam axis) in the concrete surrounding the slits. All the strengthened beams failed by debonding in the form of concrete cover splitting starting at the cut-off points of the FRP bar.

For series A, the control specimen, A0, failed due to steel yielding followed by concrete crushing at a load level of 130 kN. Beam A1 did not show any increase in the ultimate load carrying capacity but it showed about 16 % increase in the yielding load. While Beam A2 had an increase in the yield and ultimate load carrying capacities of 19.0 and 18.7%, respectively. Generally, it was observed that Beam A2 with a bonded length of 48d achieved a significant increase compared to other beams in terms of the ultimate carrying capacity. Following debonding of the CFRP bar, the load dropped to a load level similar to that of the unstrengthened beam.

For beams of Series B, the unstrengthened beam, B0, failed due to steel yielding followed by crushing of concrete at a load of 55.0 kN. Beam B1 with the shortest bonded length (12d) failed at a load level of 67 kN showing an increase of 22% in the ultimate carrying capacity compared to that of Beam B0. The other three beams (Beams B2, B3 and B4) failed at load levels of 73, 94 and 96 kN, respectively, showing an increase in the ultimate carrying

capacity of 32, 71 and 75%. Additional details concerning the experimental program can be found in Soliman et al [5, 10]. NONLINEAR FINITE ELEMENT ANALYSIS A displacement-controlled nonlinear load−deformation analysis of FRP-strengthened concrete beams is carried out using the finite element package ADINA [15]. The formulations for the concrete, steel, and FRP of this software package are employed in our analysis. They are described in detail in the ADINA theory and modelling guide [15], and are briefly summarized below. The material model for the interface truss elements consists of the bond−slip relationship obtained from the experimental work (Soliman et al. 2007). Geometrical modelling In the 3D simulations, only one quarter of the beam is modelled. Figure 2 depicts a typical 3D mesh and the type of elements used. Eight-node brick elements are used to model the concrete, CFRP and epoxy layers. The steel reinforcement is simulated using 2-node truss elements. Three translational degrees of freedom are considered at each node. The interface elements between the epoxy and concrete nodes are aligned in the longitudinal direction of the beam, (Details B & C in Figure 2) while a full strain compatibility is assumed in the other two directions and also between the epoxy and FRP elements. The element sizes of the concrete are selected to be 50 mm cube, except for the part of the concrete beam between the longitudinal tensile steel bars and the FRP laminates; where the element sizes are taken as small as 12.5 mm cube. Utilizing 3D modelling is proposed primarily to simulate the propagation of the cracks though the beam cross section.

The interface elements are assumed between the epoxy and concrete nodes assuming a full bond between the FRP and epoxy nodes. This is done since the observed failure mode of the tested specimens was a separation of the concrete cover along the concrete/epoxy interface. Besides no relative slips were observed between the CFRP bars and the epoxy in the aforementioned experimental program. In the analysis, the area of each interface element is equal to ff lb where fl is the length of the interface element (12.5 mm in the current analysis), fb is the perimeter of the epoxy layer corresponding to each concrete node.

Axis of Symmetry

Support

XY

Z

Axis of Symmetry Truss element (steel)

3-D Solid element (concrete)

3-D Solid element (Epoxy)

3-D Solid element (FRP)

3-D Interface elementA

Detail A

nodeEpoxy

Detail B

Concretenode

Z

B

C

Epoxy

Detail Cnode

Concretenode

Figure 2: Finite element models for the strengthened beams

Material models for concrete, epoxy, steel and FRP A hypo-elastic model is utilized to describe the nonlinear stress–strain relationship for concrete. The general multi-axial stress−strain relations are derived from the nonlinear uni-axial stress−strain relation depicted in Figure 3a. Figure 3b depicts the concrete biaxial failure envelope. A compressive uniaxial nonlinear relationship is used until the maximum concrete characteristic strength, cf ′ is reached beyond which the behaviour softens until the concrete crushes (Figure 3a). The ultimate uniaxial compressive stress, uσ , is taken as 0.85 cf ′ and the ultimate uniaxial compressive strain, uε , is assumed to be 0.0035. Poisson’s ratio, ν, is taken as 0.18 for the concrete. Failure envelopes are utilized to establish the uniaxial stress−strain law accounting for multiaxial stress conditions and to identify whether tensile or crushing failures of the concrete have occurred. The behaviour of the cracked concrete is described assuming a system of orthogonal cracks. Once a crack occurs in any direction, i , the material is considered orthotropic with the directions of orthotropy being defined by the principal stress directions. Cracking of the concrete occurs when the principal tensile stress lies outside the tensile failure envelope. The elastic modulus of the concrete is reduced to zero in the direction parallel to the principal tensile stress direction and then a redistribution of stresses takes place. Once cracking occurs, the shear reduction factor decreases linearly from 1.0 for the uncracked section to 0.5 for cracked sections at a strain level ranges from eight to twelve times the cracking strain, mε in Figure 3a, and then remains constant.

ε εu c

c

εStrain

σu

εm

Stre

ss

ft

t

f

=8tε

f1c

2 cfσ

σ

Compressionfailure failure

Tensile

a. Uniaxial stress–strain relationship b. Biaxial concrete failure envelope

Figure 3: Concrete constitutive law

The steel reinforcement is modelled as a bilinear elastic-plastic material, with the tangent modulus in the strain-hardening regime taken to be one-hundredth of the elastic modulus as shown in Figure 4. Although a bond−slip relation is assumed for the FRP/concrete interface, as described below, full-bond is assumed between the concrete and steel reinforcement. A linear elastic relationship until rupture is assumed for the FRP composites. In case of the 3D analysis, the elastic modulus in the direction perpendicular to the fibres, tE , is assumed to be one-tenth of that in

the direction of the fibres, fE . This assumption has been generalized for 3D analysis since the experimental data do not address clearly the exact value of the transversal elastic modulus.

Figure 4: Stress-strain curve for steel

A linear elastic model is employed to constitute the mechanical characteristics of the adhesive layer. The Young’s modulus is taken as 4000 MPa with a Poisson’s ratio of 0.3.

FRP/concrete interface models

In this study a special technique was taken into account for the slip that occurs between the FRP bar and the concrete. In the tested beams, it was observed that the main slip occurred between the epoxy layer and surrounding concrete layer. This leads to no further attention to the relative slip between the CFRP bars and epoxy. In general, the interface model is meant for the overall interfacial behaviour of the layer beneath the longitudinal steel bars including the epoxy/concrete and epoxy/CFRP nodes. The bond–slip relationships employed in this simulation were driven from the experimental results of pullout tests of NSM-FRP embedded in concrete blocks [5, 10]. The bond–slip relationship correlates the local shear stress,τ , and the associated slip, S , between the FRP laminate and the concrete substrate. This bond–slip profile depends on the bonded length. For example, Figure 5 shows the bond–slip relationship profile for a bonded length of 48d and it is employed to constitute the interfacial behaviour between NSM system and the concrete layer for the Beam A3 and B2.

024

68

1012

1416

0 0.5 1 1.5 2 2.5Slip, mm

Bon

d st

ress

, MPa

Figure 5: Bond–slip profile

FINITE ELEMENT RESULTS AND DISCUSSION

The numerical results are presented in terms of the ultimate load carrying capacities, modes of failure and deformational characteristics. Special emphasis is given to the effect of the steel reinforcement ratio on the predicted debonding phenomena.

maxτ

maxS

Table 3 summarizes the predicted load capacities, .numP & .expP , and the corresponding deflections at ultimate loads for the various specimens. Comparisons between the numerical and experimental results are also given in this Table. The average numerical-to-experimental load capacity ratio and its standard deviation are 1.06 and 0.023, respectively, indicating an excellent agreement.

Table 3: Comparison between experimental and numerical results

Specimen Experimental Numerical

Pnum./Pexp. Failure load, kN

Deflection, mm

Failure load, kN

Deflection, mm

A0 130.14 62.43 138.45 65.41 1.06 A1 124.45 18.64 136.30 21.73 1.09 B2 153.77 31.34 160.09 34.05 1.04 B3 93.87 23.97 96.39 24.17 1.03 B4 96.37 26.80 102.63 26.67 1.06

The results, plotted in Figures 6a and 6b, represent numerical-to-experimental comparisons of the load–deflection relationships of selected specimens from both Series A and B, respectively. It can be seen that the proposed finite element model was capable of simulating the entire load–deflection relationships.

0

20

40

60

80

100

120

0 10 20 30 40Deflection, mm

Deb

ondi

ng lo

ad, k

N

B4, Lb=60d exp.B4, Lb=60d num.B3, Lb=48d exp.B3, Lb=48d num.

0

20

40

6080

100

120

140

160

180

0 20 40 60 80Deflection, mm

Deb

ondi

ng lo

ad, k

N

Control, A0 exp.Control,A0 num.A1, Lb=24d exp.A1, Lb=24d num.A2, Lb=48d exp.B2, Lb=48d num.

Figure 6: Comparisons between experimental and numerical load–deflection relationships

Effect of steel reinforcement ratio (ρs)

It has been demonstrated through the experimental program presented in the previous section that the steel reinforcement ratio has a significant effect on the debonding load in NSM-FRP bars. In this study, the effect of four levels of the steel reinforcement ratio on the predicted debonding load is investigated. Figure 7 shows the numerically predicted load–deflection relationships for four beams having different steel reinforcement ratio ranging from 0.20% to 1.61%. The four beams have the same area of the CFRP bars of 71 mm2. Increasing the steel reinforcement ratio increases the ultimate capacity. The mode of failure for the four beams shown in Figure 7 is debonding of FRP bars from the surrounding concrete beam. This is simulated numerically when an interface element reaches maxτ at a certain time step. In the subsequent time step, the slip value in this particular interface element tends to a huge value as a result of increasing the crack width passing through this element. Accordingly, the debonding numerically occurs at this particular element. At a steel reinforcement ratio of 1.6, the failure initiates with concrete crushing at a load level of 191 kN, then the final failure occurs when the FRP bars debond off the concrete surface at a load level of 200 kN.

020406080

100120140160180200220240

0 2000 4000 6000 8000 10000Strain, µε

Load

, kN

ρ = 0.20%ρ = 0.40%ρ = 0.80%ρ = 1.60%

Figure 7: Effect of the steel reinforcement ratio on load–strain relationships

Figure 8a shows the effect of the steel reinforcement ratio on the predicted debonding load. Increasing the steel reinforcement area increases the ultimate capacity of the beam. However, as shown in Figure 8b, the debonding strain level in FRP bars reduces with increasing the steel reinforcement are. From this figure, it is concluded that the high utilization ratio for FRP bars is obtained in a low steel ratio beam, however, the maximum ultimate capacity achieved for high steel ratio beam.

0

50

100

150

200

250

0 200 400 600 800 1000Reinforcement area, mm2

Deb

ondi

ng lo

ad, k

N

0

2000

4000

6000

8000

10000

0 200 400 600 800 1000Reinforcement area, mm2

Deb

ondi

ng s

train

µe,

Figure 8: Effect on reinforcement ratio (ρs) on debonding load and strain CONCLUSIONS Based on the findings of this experimental and numerical investigation, the following conclusions can be drawn.

• The utilized NSM-FRP/adhesive system is effective to increase both stiffness and flexural capacity of concrete beams. However the mode of failure for the strengthened beams was debonding in the form of concrete cover splitting at the level of the steel reinforcement bars.

• The proposed finite element model was capable of predicting the load deflection behaviour and mode of failure of the strengthened beams. In general, it was observed that the predicted strain in CFRP bars was lower than that measured from the experimental results. This is probably attributed to the full bond assumption between the CFRP bar and the surrounding epoxy layer. Although the numerical model successfully simulated the behaviour of strengthened beams with a relatively long bonded length, the model was unable to predict the load deflection behaviour for beams with small bonded length (12d).

0.20 0.40 0.80 1.60 3.2 Reinforcement ratio ρs

0.20 0.40 0.80 1.60 3.2 Reinforcement ratio ρs

ACKNOWLEDGEMENTS The authors would like to express their special thanks and gratitude to the Natural Science and Engineering Research Council of Canada (NSERC), FQRNT (Équipe de recherche), FQRNT-CRIB, FQRNT-CREPEC, ISIS Canada, Pultrall Inc. (Thetford Mines, Québec), Hilti Inc. (Montréal, Québec), and the technical staff of the structural lab in the Department of Civil Engineering at the University of Sherbrooke.

REFERENCES

1. Nanni A., “Flexural Behaviour and Design of RC Members using FRP Reinforcement”, Journal of Structural Engineering, 1993, 119(11), 44–59.

2. De Lorenzis L., Nanni A., “Bond Between NSM Fiber-Reinforced Polymer Rods and Concrete in Structural Strengthening”, ACI Structural Journal, 2002, 99(2), 123–132.

3. Hassan T., Rizkalla S., “Investigation of Bond in Concrete Structures Strengthened with Near Surface Mounted Carbon Fibber Reinforced Polymer Strips”, Journal of Composites for Construction, ASCE, 2003, 7(3), 248–257.

4. Kotynia R., “Analysis of the Flexural Response of NSM FRP-Strengthened Concrete Beams”, In Proceedings of the 8th International Symposium of FRP Reinforcement for Concrete Structures, (FRPRCS-8), Fibre-Reinforced Polymer Reinforcement for Concrete Structures, Triantafillou T.C., editor, 2007, 10p on CD.

5. Soliman, S., M., El-Salakawy, E., Benmokrane, B. “Flexural Behaviour of Concrete Beams Strengthened with Near Surface Mounted FRP Bars.” Fourth International Conference on FRP Composites in Civil Engineering (CICE2008), 2008a, 22-24 July, Zurich, Switzerland, 8 p.

6. Asplund, S. “Strengthening Bridge Slabs with Grouted Reinforcement”, ACI Structural Journal, American Concrete Institute, 1949, 20(6), 397-406.

7. De Lorenzis L., Teng J.G., “Near-Surface Mounted FRP Reinforcement: An Emerging Technique for Strengthening Structures”, Composites, Part B: Engineering, 2007, 38,119–143.

8. Cruz JM S., Barros JAO. “Bond between Near Surface Mounted Carbon Fiber Reinforced Polymer Laminate strips and Concrete”, ASCE, Journal of Composites for Construction, 2006, 8(6), 519-527

9. Soliman, S., M., El-Salakawy E., Benmokrane B., ”Bond Properties of Near Surface Mounted (NSM) Carbon FRP Bars in Concrete”, Annual General Meeting & Conference Canadian Society of Civil Engineering, June 6-9, Yellowknife, Northwest Territories, Canada, 2007, 10p.

10. Soliman, S., M., El-Salakawy E., Benmokrane B. “Effectiveness of Using NMS CFRP Bars in Flexural Strengthening of Reinforced Concrete Beams”, Annual General Meeting & Conference Canadian Society of Civil engineering, June 10-13, Quebec city, Quebec, Canada, 2008, 9p

11. Nanni, A., Ludovico, M., Parretti R. “Shear Strengthening of a PC Bridge Girder with NSM CFRP Rectangular Bars”, Advances in Structural Engineering, 2004, 7(4), 97-109.

12. Hassan T., Rizkalla S., “Bond Mechanism of NSM FRP Bars for Flexural Strengthening of Concrete Structures”, ACI Structural Journal, 2004, 101(6), 830–839.

13. El-Hacha, R., Rizkalla, S., “Near Surface Mounted Fiber Reinforced Polymer Reinforcements for Flexural Strengthening of Concrete Structures”, ACI Structural Journal, American Concrete Institute, 2004, 101(5), 717-716.

14. Kang J.Y., Park Y.H., Park J.S, You Y.J., Jung W.T., “Analytical Evaluation of RC Beams Strengthened with Near Surface Mounted CFRP Laminates“, In Proceedings of the 7th International Symposium on Fibre-Reinforced Composite Reinforcement for Concrete Structures, (FRPRCS)-SP-230-45, volume 1, Shield C.K., Busel J., Walkup S., Gremel D., editors, Michigan, USA, ACI, 2005, 779–794.

15. ADINA, “Automatic Dynamic Incremental Nonlinear Analysis: Finite Element Software, Version 8.4”, ADINA R & D, Inc, Watertown, MA, USA, 2004.

16. Pultrall Inc. “Product Technical Specifications”, http://www.Pultrall.com, 2006. 17. Hilti Inc. Product Technical Guide, http://www.ca.hilti.com, 2006. 18. ACI Committee 440 “Guide Test Methods for Fiber-Reinforced Polymers (FRPs) for Reinforcing or

Strengthening Concrete Structures”, ACI 440.3R-04, American Concrete Institute, Farmington Hills Michigan, 2004, 40p.