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1/20 6PCI Journal|Summer 2012

Fiber-reinforced-polymer (FRP) composite materi-

als have been used as construction materials overthe past two decades. FRP composites can replace

traditional steel reinforcement in new construction. Ex-

ternally bonded composite plates or sheets are being used

for the repair, strengthening, and rehabilitation of aging

and deteriorated structures. It is the latter application that

has shown the greatest promise as a cost-effective solution

to the growing problem of structurally deficient concrete

structures in the United States and worldwide.

Composite alternatives have advantages over traditional

strengthening methods in terms of strength-to-weight ratio,

corrosion resistance in the case of carbon and aramid fi-bers, speed and simplicity of application, and versatility in

conforming to various cross-sectional shapes. Many exper-

imental investigations have shown externally bonded FRP

composites to be effective in increasing the load-carrying

capacity of concrete members such as columns and gird-

ers. Most of these studies have focused on axial (confine-

ment) or flexural strengthening. While flexure is typically

the limiting mode of failure in bridge girder design, shear

failure may dominate in cases where the original trans-

verse reinforcement has severely corroded or the flexural

strength has been increased. In such cases, increasing the

shear capacity can prevent catastrophic shear failure.

This paper reports an investigation of the failure modes and

ultimate bearing capacity of 16 full-scale prestressed concrete

girders strengthened in shear with externally bonded carbon-fiber-reinforced polymer (CFRP) sheets.

Test parameters include the cross-sectional shape, effects of

preexisting damage, CFRP strengthening scheme, different

anchorage systems, and transverse steel reinforcement ratio.

The test results show that the failure modes are complex and

can vary considerably with respect to the test parameters.

The test results also show that the application of externally

bonded CFRP shear reinforcement might not increase the load-

carrying capacity of a prestressed concrete girder.

Behavior of prestressed

concrete I-girders

strengthened in shear

with externally bonded

fiber-reinforced-polymer sheets

Michael Murphy, Abdeldjelil Belarbi, and Sang-Wook Bae

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2/20Summer 2012 |PCI Journal4

Experimental studies on shear strengthening with FRP

are limited and have mostly considered only small-scale

reinforced concrete beams.1However, analytical models

proposed in the literature for shear strengthening with FRP

are numerous and in most cases contradictory in their pre-

dictions of the FRP shear contribution.2Even in traditional

reinforced concrete members without externally bonded

FRP reinforcement, shear design presents a complexchallenge that relies more on semiempirical methods in

contrast to design for axial load or flexure.

Accounting for externally bonded FRP shear reinforcement

with its specific characteristics adds to the complexity of

shear design. The difficulty in defining the shear contribu-

tion of FRP arises from its anisotropy combined with a

wide variety of possible reinforcement configurations. FRP

reinforcement configurations in shear include the selection

of surfaces to be bonded (side bonding, U wrap, complete

wrap), continuous reinforcement or a series of discrete

strips, and orientation of the primary direction of fibers.

The bond characteristics between the FRP and concrete

substrate are an additional complexity in understanding the

contribution of FRP to shear strength. The effectiveness of

the strengthening method has also been found to depend

on the mode of failure that has been experimentally shown

to vary between tensile rupture of the FRP and sequential

debonding of the FRP, depending on the anchorage condi-

tions.

Although design standards for FRP shear strengthening are

still under development, FRPs have already been used as

external strengthening in a number of field applications.3,4

Current guidelines for FRP shear strengthening have beenbased on modifications to existing shear provisions. The

applicability and accuracy of such design methods have

been validated through experimental testing, mostly on

small-scale reinforced concrete beams. This study was

designed to extend current knowledge by investigating the

effectiveness of using externally bonded FRPs for increas-

ing the shear strength of full-scale prestressed concrete

girders.

Research significance

Most of the present research has investigated the behaviorof traditional mild steel-reinforced concrete structures

strengthened in shear with FRP, with only limited studies

on prestressed concrete structures. In the present study, the

behavior of full-scale American Association of State High-

way and Transportation Officials (AASHTO)-type gird-

ers, particularly the failure modes, was investigated in a

comprehensive experimental program. The study revealed

that FRP strengthening for shear was not as effective for

prestressed concrete structures as for reinforced concrete

structures and the effectiveness varied with the shape of the

cross section.

Experimental program

A total of 8 full-scaleMoDOT (Missouri Department of

Transportation)LRFD Bridge Design Guidelines5Type 3

and Type 4 precast, prestressed concrete girders were con-

structed, with each girder designed to provide two distinct

test regions for a total of 16 test specimens. The depths of

the Type 3 and Type 4 girders are 39 in. (990 mm) and 45 in.(1140 mm), respectively. The test girders were designed

such that shear would be the governing failure mode with

consideration for both the American Concrete Institutes

(ACIs)Building Code Requirements for Structural Con-

crete (ACI 318-08) and Commentary (ACI 318R-08)6and

AASHTOs Standard Specifications for Highway Bridges

orAASHTO LRFD Bridge Design Specifications.7Table1

summarizes the test parameters, material properties, and test

results. The nomenclature of the specimen indicates the test

parameters considered in this experimental study:

size of girders (T4 = Type 4; T3 = Type 3)

stirrup spacing (12 in. or 18 in. [300 mm or 460 mm])

carbon-fiber-reinforced polymer (CFRP) strengthening

configuration (S90 = strips at 90 degrees, S45 = strips at

45 degrees)

presence and type of mechanical anchorage

presence of preexisting damage/cracks (PC)

Four different cross-section types were used for the test

girders (Fig.1). The differences in the cross-section typesinvestigated include the size of the girder according to the

MoDOT LRFD guidelines (Type 3 and Type 4), the pres-

ence or absence of a deck slab, the shape of the deck slab,

and the flexural reinforcement scheme.

The test girders were constructed in a local precast con-

crete plant and delivered to the testing laboratory. Deck

slabs were constructed for all girders except for Type I

girders (Fig. 1). CFRP sheets were used for strengthen-

ing. Twelve-inch-wide (300 mm), single-ply CFRP strips

were applied using the wet lay-up technique in a U-wrap

configuration. The fibers were oriented at either 90 or45 degrees relative to the longitudinal axis of the girders

(Table 1). The CFRP strips were spaced to provide a 6 in.

(150 mm) gap.

Table 1 summarizes the concrete compressive strengths

measured for each specimen at the time of testing. Me-

chanical properties for the transverse and flexural steel

reinforcement were also evaluated (Table2). The CFRP

used for strengthening consisted of unidirectional carbon-

fiber sheets. Table 2 also provides the mechanical proper-

ties for the CFRP as provided by the manufacturer and

validated by coupon testing.

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Four different anchorage systems were investigated for

their abilities to prevent or delay the debonding associated

with FRP shear strengthening (Fig. 2):

continuous mechanical anchorage system (CMA)

discontinuous mechanical anchorage system (DMA)

Table 1.Summary of test parameters and test results

Specimen*Cross-

section

typev f

Fiber

direction,

degrees

Anchorage

typesa/d f'c g, psi f

'c s, psi

Vcr ,

kip

Vu ,

kipFailure mode

T4-12-Control I 0.0031 0 None None 2.9 9970 n/a 139 202 Top flange

T4-18-Control I 0.0020 0 None None 2.9 9930 n/a 140 206 Top flange

T4-18-S90-NA I 0.0020 0.0014 90 None 2.9 10,020 n/a 149 193 Top flange

T4-18-S90-CMA II 0.0020 0.0014 90 CMA 2.9 10,120 5240 136 229 Top flange

T4-18-S90-DMA II 0.0020 0.0014 90 DMA 2.9 10,160 7370 161 244 Top flange

T4-18-S45-DMA II 0.0020 0.0010 45 DMA 2.9 10,190 7840 161 255 Top flange

T4-12-Control-Deck II 0.0031 0 None None 2.9 10,660 10,730 141 245 Top flange

T4-12-S90-SDMA II 0.0031 0.0014 90 SDMA 2.9 10,330 10,810 108 258 Top flange

T3-12-Control III 0.0031 0 None None 3.4 8890 8520 126 253Stress

concentration

T3-12-S90-NA III 0.0031 0.0014 90 None 3.4 8910 8760 130 271 Web crushing||

T3-12-S90-NA-PC# III 0.0031 0.0014 90 None 3.4 9470 8670 n/a 239 Web crushing||

T3-12-S90-DMA III 0.0031 0.0014 90 DMA 3.4 10,380 9700 115 249Stress

concentration

T3-18-Control IV 0.0020 0 None None 3.4 9590 9820 120 252Diagonal

tension**

T3-18-S90-NA IV 0.0020 0.0014 90 None 3.4 10,120 10,030 153 216Diagonal

tension**

T3-18-S90-HS IV 0.0020 0.0014 90 HS 3.4 10,190 10,900 133 221Diagonal

tension**

T3-18-S90-SDMA IV 0.0020 0.0014 90 SDMA 3.4 10,430 11,280 141 235Diagonal

tension**

*Specimen nomenclature: size of girders (T3 = Type 3; T4 = Type 4); stirrup spacing in inches (12; 18); carbon-fiber-reinforced polymer strengthening

configuration (S45 = strips at 45 degrees; S90 = strips at 90 degrees); presence and type of mechanical anchorage (CMA = continuous mechanical

anchorage plates; DMA = discontinuous mechanical anchorage plates; HS = horizontal carbon-fiber-reinforced polymer strips; NA = no anchorage;

SDMA = sandwiched discontinuous mechanical anchorage plates); presence of preexisting damage/cracks (PC).

Fiber direction with respect to the longitudinal direction of girders

Failure in the top flange near the reaction point

Stress concentration in the top flange near the reaction point

||Web crushing of concrete strut

#Specimen tested to investigate effects of preexisting cracks prior to FRP strengthening

**Diagonal tension cracking in the web

Note: a/d= shear spantodepth ratio; CFRP = carbon-fiber-reinforced polymer; f 'c g= concrete strength of girders measured at the time of girder

testing; f 'c s = concrete strength of deck slab measured at the time of girder testing; n/a = not applicable; Vcr= measured shear force at initiation of

cracking; Vu= measured maximum shear force at failure; f= CFRP reinforcement ratio; v= transverse steel (stirrup) reinforcement ratio.

1 kip = 4.448 kN; 1 psi = 6.895 kPa.

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4/20Summer 2012 |PCI Journal

Figure 1.Cross-sectional configurations of experimental precast concrete girders. Note: MoDOT = Missouri Department of Transportation. no. 3 = 10M; no. 5 = 16M;no. 6 = 19M; no. 8 = 25M; 1 in. = 25.4 mm.

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rizes the test results for measured shear force at initiation

of cracking Vcr, measured maximum shear force at failure

Vu, and corresponding failure modes. The measured shear

force at initiation of cracking Vcrcorresponds to the first

observation of shear web cracks and also correlates with

the point at which steel stirrups and FRP begin to take

load. The measured maximum shear force at failure Vu

corresponds to the maximum shear force measured during

each test.

Because the cross-sectional type varies among the test

girders, direct comparison of Vuand corresponding failure

modes should be limited to those specimens sharing a

common cross-sectional geometry. Such comparison

shows that the application of externally bonded CFRP

sheets for shear strengthening did not always yield an

increase in the maximum shear capacity. This is because

the shear capacity of the test girders was affected by the

failure modes. However, these results are not enough to

suggest that shear strengthening of prestressed concrete

girders using externally bonded FRPs is ineffective. Theresults presented here suggest that the effectiveness of

external strengthening using CFRP sheets is influenced by

the failure mode. Therefore, this paper discusses the failure

modes of test girders in each cross-sectional type and the

contribution of CFRP sheets.

Failure modes

A variety of failure modes was observed among the test

specimens, including failure along the top flange, debond-

ing of FRP, localized rupture of FRP, diagonal shear

tension, web crushing, mechanical anchorage failure, andstress concentration at reaction point. Table 1 presents the

ultimate contributing failure modes for each test girder,

and Fig.58show the test girders at failure.

The first series of tests was performed on the test girders

with a Type I cross section (T4-12-Control, T4-18-Control,

and T4-18-S90-NA). For the unstrengthened specimens,

T4-12-Control, and T4-18-Control, diagonal cracks formed

first in the web. The maximum shear force was measured

when the diagonal cracks propagated into the top flange

near the reaction point (Fig. 5). For the strengthened speci-

men, T4-18-S90-NA, diagonal cracks formed in the web

sandwich panel discontinuous anchorage system

(SDMA)

additional horizontal CFRP strips (HS)

The CMA system used continuous precured CFRP plates

anchored in place with concrete wedge anchors. The

DMA system used discontinuous precured CFRP plates

anchored in place with bolts running through the web. The

SDMA system used discontinuous precured CFRP plates

with sandwich wrapped ends anchored in place with bolts

running through the web. The HS system used 5 in. wide

(130 mm) strips of bidirectional ( 45 degrees) CFRP

strips applied parallel to the longitudinal axis of the beam

and covering all of the free edges of the vertical CFRP

strips as well as along the interface of the web and bottom

flange (locations where debonding is expected to initiate).

This anchorage system was installed immediately after

application of the vertical CFRP shear reinforcement to

ensure a better bond between the vertical and horizontal

CFRP sheets.

The test setup for all specimens consisted of a three-point

monotonic loading configuration (Fig.3). Load was ap-

plied using two hydraulic actuators operating in parallel at

the farthest support under deformation control. By using

this loading configuration, the load demand on the actua-

tors is minimized, thus allowing failure of the specimens

to be achieved without exceeding the load capacity of the

actuators. The test setup also consisted of an additional ex-

ternal strengthening system composed of a series of hollow

steel sections and no. 11 (36M) reinforcing bars. This sys-

tem was intended to prevent failure from occurring outsideof the designated test region and to protect the second test

region from premature damage during testing of the first

test region. Preexisting cracks were introduced to specimen

T3-12-S90-NA-PC prior to the application of CFRP sheets

by applying load equivalent to 60% of the anticipated load-

carrying capacity. Figure4shows the preexisting cracks of

girder T3-12-S90-NA-PC.

Test results

Failure testing of the girders was performed in order of

cross-section Type I, II, III and IV (Fig. 1). Table 1 summa-

Table 2.Mechanical properties of steel and CFRP reinforcement

Reinforcement type Grade, ksi Yield strength, ksiUltimate strength,

ksi

Modulus

of elasticity, ksi

No. 3 stirrups 60 65 99 26,000

No. 6 flexural tension steel 60 78 98 25,800

0.6 in. seven-wire strand 270 n/a 291 29,100

CFRP sheets n/a n/a 550 33,000

Note: CFRP = carbon-fiber-reinforced polymer; n/a = not appl icable. no. 3 = 10M; no. 6 = 19M; 1 in. = 25.4 mm; 1 ksi = 6.895 MPa.

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Figure 2.Anchorage systems. Note: CFRP = carbon-fiber-reinforced polymer; CMA = continuous mechanical anchorage plates; DMA = discontinuous mechanicalanchorage plates; HS = horizontal FRP strips; SDMA = sandwiched discontinuous mechanical anchorage plates. 1 in. = 25.4 mm.

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Figure 3.General test setup configuration.

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that shear strength will only increase if the test girders fail

by diagonal cracking through the bottom flange.

Based on the observations of the first test series, the second

series of tests (T4-18-S90-CMA, T4-18-S90-DMA, T4-18-

S45-DMA, T4-12-Control-Deck, and T4-12-S90-SDMA)

was designed so that the top flange had greater stiffness.

This was accomplished by adding a 12 in. deep (300 mm)

deck slab (Fig. 1). A complete CFRP wrap strengthen-

first, and the FRP strips started debonding. The maximum

shear force was measured when the diagonal cracks propa-

gated into the top flange near the reaction point (Fig. 5)

as also seen in the unstrengthened specimens. There was

no shear strength increase observed. The diagonal cracks

propagated into the top flange rather than the bottom flange

because the bottom flange is stiffer than the top flange. It

can be concluded that with this failure mode, shear strength

increase is not possible. Furthermore, it can be assumed

Figure 5.Type I test girders at failure.

T4-12-Control T4-18-Control T4-18-S90-NA

Figure 6.Type II test girders at failure.

T4-18-S90-CMA T4-18-S90-DMA T4-18-S45-DMA

Figure 4.Preexisting cracks in T3-12-S90-NA-PC.

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(Fig. 7). Because the only difference between the two

specimens (T3-12-S90-NA and T3-12-S90-NA-PC) was

the presence of preexisting cracks, it can be concluded thatthe preexisting cracks did not affect the failure mode. The

web crushing failure was believed to occur as a result of

spalling of the concrete cover rather than debonding at the

CFRP-concrete interface. Spalling of the cover resulted

in a significant reduction in the width of the concrete

compressive struts within the thin web of the prestressed

concrete girders, leaving them vulnerable to web crushing.

The strengthened specimen with the mechanical anchor-

age system T3-12-S90-DMA failed in the same manner as

observed in the control specimen T3-12-Control. Failure

was due to the stress concentration in the top flange near

the reaction point. A slight increase in shear strength wasobserved in T3-12-S90-NA, but not in the other cases.

Based on the results of the third series of tests, it was

concluded that failure due to stress concentrations could

be avoided if the deck slab size was increased and better

confined with reinforcement. The fourth series of tests

examined specimens T3-18-Control, T3-18-S90-NA,

T3-18-S90-HS, and T3-18-S90-SDMA. As an additional

measure, to prevent the failure mode that occurred in the

third series of tests, additional external strengthening in the

form of hollow steel sections and no. 11 (36M) reinforc-

ing bars was added near the reaction frame within the test

ing scheme for better confinement was also used in some

specimens (Fig. 6). Mechanical anchorage systems were

installed to avoid premature failure due to FRP debond-ing (Fig. 6). These modifications postponed the cracking

observed in the first series of tests. However, failure did

not occur due to either FRP debonding or rupture. Instead,

buckling of the longitudinal compression reinforcement

produced horizontal cracking along the midheight of the

top flange (Fig. 6). No increase in shear strength due to

FRP was observed.

The third series of tests (T3-12-Control, T3-12-S90-NA,

T3-12-S90-NA-PC, and T3-12-S90-DMA) was designed to

avoid the failure modes observed in the previous two test

series by eliminating the compression bars and adding deckslabs. The unstrengthened control specimen, T3-12-Con-

trol, ultimately failed at the top flange near the reaction

point. The failure was not because the diagonal cracking

penetrated into the top flange as seen in the previous series

of tests. Rather, it was a buildup of stress concentrations

due to a combination of diagonal compression stresses,

flexural compression stresses, and contact stresses induced

by the reaction frame.

For the strengthened specimens without mechanical an-

chorage systems (T3-12-S90-NA and T3-12-S90-NA-PC),

failure occurred because of crushing of the web concrete

Figure 7.Type III test girders at failure.

T3-12-S90-NA T3-12-S90-NA-PC T3-12-S90-DMA

Figure 8. Type IV test girders at failure.

T3-18-Control T3-18-S90-NA T3-18-S90-SDMA

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region. The failure of this series of tests was characterized

by diagonal shear-tension failure in the web preceded by

some level of debonding when CFRP shear reinforcement

was present. The mechanical anchorage system used for

specimen T3-18-S90-SDMA helped to limit such debond-

ing. However, the mechanical anchors created a plane of

weakness along which the critical cracks were observed

to propagate (Fig.9). This might explain the lower shear

strength of the strengthened specimen, T3-18-S90-SDMA,

compared with the control specimen, T3-18-Control.

Overall, there was no increase in shear strength due to FRP

strengthening. This is attributed mainly to the complex

failure modes of prestressed concrete girders.

Web crushing failure

To understand web crushing failure, it is first necessary

to understand the behavior of the bond between FRP

sheets and concrete. Numerous research studies have been

conducted on this topic. Most research has focused on a

simple shear test (Fig.10) in which the maximum inter-

facial shear stress maxand effective bond lengthLeareexperimentally determined. Debonding occurs first within

the effective bond lengthLe, resulting from debonding of a

thin layer of concrete rather than debonding at the interface

between the FRP and concrete. Equation (1) is the com-

mon form for expressing the maximum normal force at

debonding Puthat can be carried by the FRP sheet.

Pu= maxbfLe (1)

where

bf= width of FRP strips

Figure 9.Crack propagation along line of mechanical anchorage for specimen T3-18-S90-SDMA.

Figure 10.Shear test for effective bond length. Note: FRP = fiber-reinforced

polymer; Le= effective bond length; P= applied load.

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The knowledge accumulated from simple shear tests is not

directly applicable to beam shear problems and thus must

be modified. Figure11presents three different models

for debonding based on crack spacing. Unlike the simple

shear test used in previous research studies, a large piece of

concrete debonded in the prestressed concrete girder tests

(Fig.12). This was due to the damage from the compres-

sion forces and interlocking of aggregate near the cracks.

Thus the debonding lines (Fig. 11) can be assumed to

be parabolic. It can also be assumed that spalling of the

cover occurred within the effective bond length and that

the debonding strength depends on the concrete tensile

strength because the debonding did not occur at the inter-

face of the FRP and concrete.

Figure 11.Shear crack and debonding models. Note: FRP = fiber-reinforced polymer; Le= effective bond length; P= applied load; Se= crack spacing.

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12/20Summer 2012 |PCI Journal4

The MoDOT Type 3 girders with stirrups spaced at 18 in.

(460 mm) showed larger crack spacing compared with

the MoDOT Type 3 girders with stirrups spaced at 12 in.

(300 mm). This may explain why web crushing failure did

not occur.

Behavior of anchorage systems

Four different mechanical anchorage systems were tested

in this study (Fig. 2):

the CMA system

the DMA system

the SDMA system

the HS system

The CMA system performed poorly, showing premature

anchorage failure (Fig.13). This is mainly due to buck-

ling of the continuous FRP plate and short embedment

length of concrete anchor bolts. The DMA system did not

prevent FRP debonding, though it did delay the debond-

ing of FRP and performed better than the CMA system

(Fig.14). The HS system delayed FRP debonding but wasnot as effective as the DMA system. The SDMA system

performed best among the four anchorage systems. The

SDMA system prevented FRP debonding and led to FRP

rupture. However, in one specimen it created a weak plane

along the anchor bolts. The diagonal cracks eventually

propagated along the weak plane. As a result, premature

failure occurred along the weak plane without any increase

in shear strength (Fig. 9).

Shear contribution of CFRP sheets

Although an increase in the shear capacity was not always

With the assumptions for the suggested models, when the

crack spacing is greater than 2Le(Fig. 11), which was the

case for the MoDOT Type 4 girders, the area of debonded

concrete cover is small compared with the total area of the

compression strut. Therefore, web crushing failure did not

occur. In this case, Eq. (1) is still applicable and max can be

replaced with the direct cracking strength of concrete

(4 fcl for normalweight concrete,8where fclis the speci-

fied compressive strength of concrete in psi).

The MoDOT Type 3 girders with stirrups spaced at 12 in.

(300 mm) exhibited tighter crack spacing. The bond

behavior of FRP and concrete can be treated as the cases

in which crack spacing is less thanLeor less than 2Le(Fig. 11). In these cases, the loss of concrete cross sec-

tion due to FRP debonding in the concrete strut cannot be

ignored, and thus web crushing is likely to occur. Where

crack spacing is less than 2Le, Eq. (2) can determine the

maximum debonding force to account for the overlapped

effective bond length.

Pu= maxbf(SeLe) = 4 fcl bf(SeLe) (2)

where

Se= crack spacing

Where the crack spacing is smaller than the effective bond

lengthLe, Eq. (3) can determine the maximum debonding

force using the crack spacing instead of the bond length.

Pu= maxbfSe= 4 fcl bfSe (3)

Based on the results, it can be concluded that the debond-

ing and spalling of concrete will occur at a lower level of

stress in FRP, and web crushing will be the failure mode

when stirrups are spaced such that crack spacing is less

than the effective bond lengthLe.

Figure 12.Close-up of debonded fiber-reinforced polymer sheet with spalled concrete cover attached.

T3-12-S90-NA T3-12-S90-NA-PC

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sistance must equal the applied shear force. The three

individual components (Vc, Vs, and Vf) of the internal shear

resistance were then evaluated from crack-based free body

diagrams of a portion of the test girders along the critical

shear cracks. Strain gauge measurements were used toquantify the shear resistance provided by the steel stirrups

Vsand CFRP strips Vfwithin the test regions. Only those

strain gauge measurements closest to the critical shear

crack were used for this analysis. To satisfy equilibrium,

the shear resistance provided by concrete Vcwas consid-

ered as the difference between the applied shear force and

the shear resistance contributions from the stirrups Vsand

FRP Vf.

The contributions from the stirrups and CFRP strips were

minimal before cracking initiated in the web (Fig. 15 and

16). The applied shear force was thus carried primarily

realized among the CFRP-strengthened specimens, CFRP

did contribute to the shear resistance of the girders as mea-

sured by strain gauges applied to the CFRP strips. To better

understand the shear resistance mechanisms and to better

quantify the CFRP contribution to shear capacity, it isnecessary to decouple the individual contributions to shear

resistance. In this case, the primary components contribut-

ing to shear resistance are the concrete Vc, steel stirrups

Vs, vertical component of prestressing force Vp(which is

zero in the specimens of this study), and externally bonded

CFRP strips Vf.

A shear component analysis was conducted on the experi-

mental data to identify the contribution of each component

throughout the loading history of the test girders.

Figures15 and 16show the results. The analysis was

based on the understanding that the internal shear re-

Figure 13.Mechanical anchorage failure of specimen T4-18-S90-CMA.

Figure 14.Mechanical anchorage failure of specimens T4-18-S90-DMA and T4-18-S45-DMA.

T4-18-S90-DMA T4-18-S45-DMA

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Figure 15.Shear component diagrams for Missouri Department of Transportation Type 4 specimens. Note: Vc= contribution of the concrete to the shear resistance of

a member; Vf= contribution of the externally bonded FRP to the shear resistance of a member; Vs= contribution of the transverse steel reinforcement (stirrups) to theshear resistance of a member. 1 kip = 4.448 kN.

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Figure 16.Shear component diagrams for Missouri Department of Transportation Type 3 specimens. Note: 1 kip = 4.448 kN.

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the FRP contribution to shear resistance. The 17 models

found in the literature926can be categorized into four gen-

eral groups based on similarities in their approach:

empirically determined value of stress/strain associ-

ated with failure of the member for which the FRP

contribution is determined

effective FRP strain concept that is generally derived

from the regression of experimental data

accounting for a nonuniform strain distribution in the

externally bonded FRP

theoretical approaches that are mechanics based and

do not rely on empirical regression or calibration

techniques

Table3compares the FRP shear contributions predicted

by these 17 models926with the experimentally measured

FRP shear contributions determined from a shear com-

ponent analysis. None of the models was consistently

accurate at predicting the FRP contributions for the experi-

mental prestressed concrete girders. Reasonable average

strength ratios Vf,exp/Vf,predwere obtained for the models of

Khalifa and Nanni,15Triantafillou and Antonopoulos,16and

Hsu et al.20However, the standard deviations and coef-

ficients of variation for these models are high, indicating

poor correlation with the test results and therefore unreli-

able precision. The models by Khalifa and Nanni15and

Triantafillou and Antonopoulos16have been used as the

basis for current design guidelines.27,28

Conclusion

In this study, the shear behavior and failure modes of full-

scale prestressed concrete I-girders strengthened in shear

with externally bonded CFRP were investigated. The effec-

tiveness of externally bonded FRP for shear strengthening

was found to be significantly affected by the cross-sectional

shape of the girders. The results show that the failure modes

vary depending on the cross-sectional shape and shear

reinforcement schemes. Debonding of the FRP is typically

accompanied by peeling of the concrete cover that reduces

the thickness of the web and can result in a web crushingfailure mode for these thin webbed girders. In such cases,

FRP shear strengthening can result in a reduction of the

ultimate shear strength of a prestressed concrete girder.

The use of an FRP anchorage system was found to delay

debonding of the FRP, resulting in greater shear resistance.

Horizontal strips of FRP used as mechanical anchor-

age provided little additional shear capacity. Continuous

CFRP plates with anchorage bolts were also ineffective in

anchoring the CFRP sheets due to buckling of the plates

and insufficient embedment length to prevent pullout of

the anchor bolts. The most effective FRP anchorage system

by the concrete contribution Vc. Upon onset of diagonal

cracking, a portion of the applied shear force was trans-

ferred to the steel stirrups and CFRP strips as shown by

the sudden jump in the shear contribution responses. The

stirrups and CFRP strips continued to take load, as shown

by the gradual increase in shear contribution response, until

yielding of the stirrups or debonding of the CFRP strips oc-

curred. Yielding of all stirrups along the critical shear crackis indicated as a plateau in the shear contribution response

for the stirrups Vs. Sudden or gradual drops in the CFRP

contribution responses signify debonding of the CFRP

strips. The severity of the drops in the shear contribution

of CFRP strips reflected the magnitude of debonding that

occurred during load history.

Unlike the specimens without preexisting cracks, the speci-

men with preexisting cracks (T3-12-S90-NA-PC) showed

a gradual increase in the shear resistance contributions of

the stirrups Vsand FRP Vffrom the early stage of loading

(Fig. 16).

Comparison of analyticalpredictions and experimentalresults

The shear resistance mechanism for reinforced concrete

and prestressed concrete members is complex; however,

the most commonly adopted analytical approach is a su-

perposition method that considers the shear resistance as a

summation of the concrete Vcand transverse steel Vscontri-

butions. The contribution of these two components is docu-

mented in the literature,6,7and procedures for determining

their magnitude have been adopted by most current designguidelines. In accordance with this design philosophy, the

contribution of externally bonded FRP shear reinforcement

Vfis accounted for by the addition of a third term. Equa-

tion (4) calculates the total shear resistance Vn.

Vn= Vc+ Vs+ Vf (4)

A total of 17 analytical models for predicting the shear

contribution of externally bonded FRP were found in the

literature.926An overview of these models is beyond the

scope of this paper. The FRP contribution is often as-

sumed analogous to that of the transverse steel reinforce-ment based on strut-and-tie methodology. Similar to the

steel stirrups, the FRP laminates are considered ties that

resist the tensile stresses along cracks between the con-

crete struts. However, the effectiveness of FRP in resisting

these tensile stresses is more complex than that of the steel

stirrups because it depends on the complex bond behavior

between concrete and FRP laminates, the material behavior

of the FRP laminate (that is, linear elastic up to failure as

opposed to the elastoplastic nature of steel), FRP laminate

geometry (that is, width and effective bond length), failure

mode, and anchorage. As a result of such complexities,

there is disagreement among current models for predicting

8/10/2019 JL-Summer-2012-9


contribution or the total shear resistance of the prestressed

concrete girders in this study because of the complexity of

the shear behavior and failure modes. Further research intothe influences of the cross-sectional geometry (particu-

larly in thin webbed members) and the interactions among

concrete, steel stirrups, and FRP strengthening will lead to

scientifically more rational and well-founded models for

defining the shear resistance of FRP shear-strengthened

girders.

was the discontinuous CFRP plates attached with epoxy

and anchored in place with bolts running through the web.

The effectiveness of this method was further improvedby the use of what has been termed a sandwich applica-

tion that helps to prevent slippage of the FRP sheet from

beneath the anchorage plate.

In the absence of an increase in the ultimate shear capacity

of a girder that is strengthened with FRP, a shear com-

ponent analysis can be used to understand the FRP shear

contribution to the total shear resistance. Such analysis em-

phasizes the interdependence that exists between the differ-

ent components of shear resistance (that is, concrete, steel

stirrups, and FRP). Existing analytical models proposed in

the literature were not able to accurately predict the FRP

Table 3.Performance of analytical models in terms of Vf,exp/Vf,pred

Analytical models for Vf

T4-1

8-

S90-NA

T4-1

8-

S90-

CMA

T4-1

8-

S90-DMA

T4-1

8-

S45-DMA

T4-1

2-

S90-

SDMA

T3-1

2-

S90-NA

T3-1

2-

S90-NA-PC

T3-1

2-

S90-DMA

T3-1

8-

S90-NA

T3-1

8-

S90-HS

T3-1

8-

S90-

SDMA

Average(mean)

Stan

darddeviation

Coefficientofvariation

Al-Sulaimani et al. 1994 0.40 0.39 0.20 n/a 0.30 0.23 0.24 0.19 0.11 0.19 0.68 0.29 0.16 0.53

Chajes et al. 1995 1.26 1.23 0.63 1.10 0.95 0.63 0.64 0.51 0.30 0.51 1.83 0.87 0.43 0.49

Triantafillou 1998 1.11 1.09 0.55 0.79 0.84 0.56 0.56 0.45 0.27 0.45 1.62 0.75 0.37 0.50

Khalifa et al. 1998 1.01 1.30 0.66 1.07 1.00 0.64 0.64 0.51 0.31 0.52 1.85 0.86 0.42 0.49

Malek and Saadatmanesh 1998 17.80 13.29 3.84 2.99 11.23 1.65 3.88 2.53 0.83 1.22 3.35 5.69 5.43 0.95

Hutchinson and Rizkalla 1999 0.82 1.16 0.56 1.20 1.10 0.50 0.48 0.46 0.16 0.49 2.33 0.84 0.57 0.68

Khalifa and Nanni 2000 1.25 1.62 0.82 1.44 1.24 0.78 0.79 0.63 0.38 0.64 2.30 1.08 0.53 0.49

Triantafillou and Antonopoulos

20001.40 1.37 0.70 1.22 1.05 0.70 0.71 0.56 0.34 0.57 2.04 0.97 0.48 0.49

Deniaud and Cheng 2001 2.60 3.37 1.72 1.77 3.89 2.79 2.83 2.26 0.89 1.49 5.33 2.63 1.19 0.45

Chaallal et al. 2002 6.68 8.65 4.40 9.90 7.74 4.86 4.93 3.94 2.04 3.42 12.27 6.26 2.95 0.47

Pellegrino and Modena 2002 3.99 5.13 2.60 4.56 3.89 2.72 2.65 1.99 1.22 2.04 7.20 3.45 1.64 0.48

Hsu et al. 2003 1.26 1.62 0.82 1.44 1.23 0.84 0.83 0.62 0.38 0.64 2.25 1.08 0.52 0.48

Chen and Teng 2003 2.11 2.85 1.45 2.51 2.18 1.41 1.41 1.10 0.67 1.12 3.99 1.89 0.92 0.48

Carolin and Taljsten 2005 2.46 4.12 1.97 5.24 4.29 1.64 1.56 1.61 0.49 1.73 9.49 3.15 2.43 0.77

Cao et al. 2005 0.85 1.12 0.57 1.40 0.86 0.51 0.51 0.40 0.24 0.40 1.44 0.75 0.39 0.52

Monti and Liotta 2005 0.63 0.74 0.36 0.54 0.70 0.35 0.33 0.30 0.11 0.32 1.52 0.53 0.36 0.68

Sim et al. 2005 4.79 6.03 3.06 7.40 5.64 3.88 3.79 2.87 1.43 2.38 8.44 4.52 2.07 0.46

Note: Vf= contribution of the externally bonded fiber-reinforced polymer to the shear resistance of a member; Vf,exp= measured fiber-reinforced

polymer contribution to shear resistance of a member as determined from a shear component analysis; Vf,pred= analytical prediction of the externally

bonded fiber-reinforced polymer contribution to the shear resistance of a member.

8/10/2019 JL-Summer-2012-9


10. Chajes, M. J., T. F. Jansuska, D. R. Mertz, T. A.

Thomson, and W. W. Finch. 1995. Shear Strength

of RC Beams Using Externally Applied Composite

Fabrics.ACI Structural Journal92 (3): 295303.

11. Triantafillou, T. C. 1998. Shear Strengthening of Re-

inforced Concrete Beams Using Epoxy-Bonded FRP

Composites.ACI Structural Journal

95 (2): 107115.

12. Khalifa, A., W. Gold, A. Nanni, and M. I. Abdel Aziz.

1998. Contribution of Externally Bonded FRP to

Shear Capacity of RC Flexural Members.Journal of

Composites for Construction2 (4): 195202.

13. Malek, A. M., and H. Saadatmanesh. 1998. Ulti-

mate Shear Capacity of Reinforced Concrete Beams

Strengthened with Web-Bonded Fiber-Reinforced Plas-

tic Plates.ACI Structural Journal95 (4): 391399.

14. Hutchinson, R. L., and S. H. Rizkalla. 1999. Shear

Strengthening of AASHTO Bridge Girders Using

Carbon Fiber Reinforced Polymer Sheets. InACI SP-

188: 4th International SymposiumFiber Reinforced

Polymer Reinforcement for Reinforced Concrete

Structures, 945958. Farmington Hills, MI: ACI.

15. Khalifa, A., and A. Nanni. 2000. Improving Shear

Capacity of Existing RC T-Section Beams Using

CFRP Composites. Cement and Concrete Composites

22 (3): 165174.

16. Triantafillou, T. C., and C. P. Antonopoulos. 2000.

Design of Concrete Flexural Members Strengthenedin Shear with FRP.Journal of Composites for Con-

struction4 (4): 198205.

17. Deniaud, C., and J. J. R. Cheng. 2001. Shear Behav-

ior of Reinforced Concrete T-Beams with Externally

Bonded Fiber-Reinforced Polymer Sheets.ACI Struc-

tural Journal98 (3): 386394.

18. Chaallal, O., M. Shahawy, and M. Hassan. 2002. Per-

formance of Reinforced Concrete T-Girders Strength-

ened in Shear with Carbon Fiber Reinforced Polymer

Fabrics.ACI Structural Journal99 (3): 335343.

19. Pellegrino, C., and C. Modena. 2002. Fiber Rein-

forced Polymer Shear Strengthening of Reinforced

Concrete Beams with Transverse Steel Reinforce-

ment.Journal of Composites for Construction6 (2):

104111.

20. Hsu, C. T. T., W. Punurai, and Z. Zhang. 2003. Flex-

ural and Shear Strengthening of RC Beams Using Car-

bon Fiber Reinforced Polymer Laminate. InLarge-

Scale Structural Testing, SP-211, 89113. Farmington

Hills, MI: ACI.

Acknowledgments

The authors wish to express their gratitude and sincere ap-

preciation to the National Cooperative Highway Research

Program as well as the Intelligent Systems Center and

National University Transportation Center at the Missouri

University of Science and Technology in Rolla for financ-

ing this research work. The authors would also like to rec-ognize Egyptian Concrete Inc. for its support in construct-

ing the full-scale test specimens.

References

1. Belarbi, A., S.-W. Bae, and A. Brancaccio. 2012.

Behavior of Full-Scale RC T-Beams Strengthened in

Shear with Externally Bonded FRP Sheets. Construc-

tion and Building Materials32 (July): 2740.

2. NCHRP (National Cooperative Highway Research

Program). 2011.Design of FRP Systems for Strength-

ening Concrete Girders in Shear. NCHRP report 678.

Washington, DC: NCHRP.

3. Hag-Elsafi, O., J. Kunin, S. Alampalli, and T. Conway.

2001. Strengthening of Route 378 Bridge over Wyn-

antskill Creek in New York Using FRP Laminates. Spe-

cial report 135. Albany, NY: New York State Depart-

ment of Transportation.

4. Tljsten, B., A. Hejll, and G. James. 2007. Carbon

Fiber-Reinforced Polymer Strengthening and Moni-

toring of Grndals Bridge in Sweden.Journal of

Composites for Construction11 (2): 227235.

5. Missouri Department of Transportation (MoDOT).

2005.MoDOT LRFD Bridge Design Guidelines. Jef-

ferson City, MO: MoDOT.

6. ACI (American Concrete Institute) Committee 318.

2008.Building Code Requirements for Structural Con-

crete (ACI 318-08) and Commentary (ACI 318R-08).

Farmington Hills, MI: ACI.

7. AASHTO (American Association of State Highway

and Transportation Officials). 2008.AASHTO LRFDBridge Design Specifications. 4th ed. Washington, DC:

AASHTO.

8. Collins, M. P., and Mitchell, D. 1990. Prestressed

Concrete Structures. Englewood Cliffs, NJ: Prentice

Hall.

9. Al-Sulaimani, G. J., A. Shariff, I. A. Basanbul, M.

H. Baluch, and B. N. Ghaleb. 1994. Shear Repair of

Reinforced Concrete by Fiber Glass Plate Bonding.

ACI Structural Journal91 (4): 458464.

8/10/2019 JL-Summer-2012-9


fcsl = concrete strength of deck slab measured at the time

of girder testing

Le = effective bond length

P = applied load

Pu = debonding strength

Se = crack spacing

Vc = concrete contribution to shear resistance of a mem-

ber

Vcr = measured shear force at initiation of cracking

Vf = FRP contribution to shear resistance of a member

Vf,exp = measured FRP contribution to shear resistance of

a member as determined from a shear component

analysis

Vf,pred= analytical prediction of the externally bonded FRP

contribution to shear resistance of a member

Vn = total shear resistance of a member defined as the

sum of the concrete Vc, transverse steel reinforce-

ment Vs, and externally bonded FRP Vfcontribu-

tions

Vp = vertical component of prestressing force

Vs = transverse steel reinforcement (stirrups) contribu-tion to shear resistance of a member

Vu = measured maximum shear force at failure

f = shear reinforcement ratio for externally bonded

CFRP reinforcement

v = shear reinforcement ratio for steel stirrups

max = maximum interfacial shear stress

21. Chen, J. F., and J. G. Teng. 2003a. Shear Capacity

of FRP Strengthened RC Beams: FRP Debonding.

Construction Building Materials17 (1): 2741.

22. Chen, J. F., and J. G. Teng. 2003b. Shear Capacity of

Fiber-Reinforced Polymer-Strengthened Reinforced

Concrete Beams: Fiber Reinforced Polymer Rupture.

Journal of Structural Engineering129 (5): 615625.

23. Carolin, A., and B. Taljsten. 2005. Theoretical Study

of Strengthening for Increased Shear Bearing Capac-

ity.Journal of Composites for Construction9 (6):

497506.

24. Cao, S. Y., J. F. Chen, J. G. Teng, Z. Hao, and J. Chen.

2005. Debonding in RC Beams Strengthened with

Complete FRP Wraps.Journal of Composites for

Construction9 (5): 417428.

25. Monti, G., and M. A. Liotta. 2005. FRP-Strength-

ening in Shear: Tests and Design Equations. In 7th

International Symposium on Fiber-Reinforced Polymer

(FRP) Reinforcement for Concrete Structures, SP-230,

543562, Farmington Hills, MI: ACI.

26. Sim, J., G. Kim, C. Park, and M. Ju. 2005. Shear

Strengthening Effects with Varying Types of FRP

Materials and Strengthening Methods. In 7th Inter-

national Symposium on Fiber-Reinforced Polymer

(FRP) Reinforcement for Concrete Structures, SP-230,

16651680, Farmington Hills, MI: ACI.

27. ACI Committee 440. 2008. Guide for the Design andConstruction of Externally Bonded FRP Systems for

Strengthening Concrete Structures. ACI 440.2R-08.

Farmington Hills, MI: ACI.

28. fib(Fdration Internationale du Bton) Task Group

9.3. 2001.Externally Bonded FRP Reinforcement for

RC Structures. Bulletin 14. Lausanne, Switzerland:fib.

Notation

a/d = shear spantodepth ratio

bf = width of FRP strips

fcl

= specified compressive strength of concrete

fcgl = concrete strength of girders measured at the time of

girder testing

8/10/2019 JL-Summer-2012-9

20/20

About the authors

Michael Murphy, PhD, is astructural engineer at CTLGroup

in Washington, D.C. He is a

member of ACI. He received his

BS, MS, and PhD in civil engi-

neering from the Missouri

University of Science and

Technology. His research interests include perfor-

mance of externally bonded fiber-reinforced polymers

for strengthening reinforced and prestressed concrete

members.

Abdeldjelil Belarbi, PhD, PE, is

the chair and Hugh Roy and Lillie

Cranz Cullen Distinguished

Professor in the Department of

Civil and Environmental Engi-

neering at the University of

Houston in Houston, Tex. He is a

Fellow of ACI and ASCE and an active member of

several ACI and ASCE technical and educational

committees. His research interests include constitutive

modeling of reinforced and prestressed concrete

structures, the use of advanced materials in new

construction and strengthening of civil infrastructures,

structural health monitoring, smart sensors, andstructural performance and durability in natural

disasters.

Sang-Wook Bae, PhD, is an

assistant professor in the Depart-

ment of Civil and Environmental

Engineering at Texas Tech

University in Lubbock, Tex. He is

a member of ACI. His research

interests include analytical and

experimental investigation of concrete structures,

advanced repair methods for corrosion-damagedreinforced concrete structures, performance evaluation

of fiber-reinforced polymer composites for use as

internal reinforcement or external repair, and strength-

ening of concrete structures.

Abstract

This paper investigates the behavior of full-scaleprestressed concrete girders strengthened in shear with

externally bonded carbon-fiber-reinforced polymer

(CFRP) sheets. The study is aimed at identifying the

failure modes and effects on ultimate bearing capacity

associated with the application of CFRP laminates as

externally bonded shear reinforcement for prestressed

concrete I-girders. A total of 16 full-scale prestressed

concrete girder tests are reported. Test parameters

include the cross-sectional shape, effects of preexist-

ing damage, CFRP strengthening scheme, different

anchorage systems, and transverse steel reinforcement

ratio. The test results show that the failure modes are

complex and can vary considerably with respect to the

test parameters. The test results also show that the ap-

plication of externally bonded CFRP reinforcement for

shear may not yield an increase in the load-carrying

capacity of a girder compared with a reference mem-

ber that is not strengthened with CFRP.

Keywords

Anchorage, carbon-fiber-reinforced polymer, CFRP,

composite sheets, failure, shear.

Review policy

This paper was reviewed in accordance with the

Precast/Prestressed Concrete Institutes peer-review

process.

Reader comments

Please address any reader comments to journal@pci

.org or Precast/Prestressed Concrete Institute, c/o PCI

Journal, 200 W. Adams St., Suite 2100, Chicago, IL

60606. J

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