BOND OF FRP LAMINATES TO CONCRETE - Missouri …transportation.mst.edu/media/research/transportation/documents/...De Loren zis, L., B. Miller, And A. Nanni, "Bond of FRP Laminates

De Lorenzis, L., B. Miller, And A. Nanni, "Bond of FRP Laminates to Concrete", ACI Materials Journal, Vol. 98, No. 3, May -June 2001, pp. 256-264

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BOND OF FRP LAMINATES TO CONCRETE

Laura De Lorenzis, Brian Miller, Antonio Nanni

KEYWORDS: Bond, Concrete, External Reinforcement, Fiber Reinforced Polymers, Slip

ABSTRACT

Fiber reinforced polymer (FRP) laminates are being successfully used for strengthening of

existing reinforced concrete structures. Bond of the FRP reinforcement to the concrete

substrate is of critical importance for the effectiveness of the technique. In this project,

flexural test specimens were prepared to address some of the factors expected to affect bond,

namely, bonded length, concrete strength, number of plies (stiffness), sheet width and, to a

limited extent, surface preparation. Results are presented and discussed in this paper. A shear

lag approach, along with a simple shear model for the evaluation of the slip modulus, is used

to model the strain distribution at moderate load levels. Finally, expressions of the peeling

load and the effective bond length are presented. A design equation is proposed to calculate

the effective FRP ultimate strain to be used in design to account for bond-controlled failure.

INTRODUCTION

Fiber reinforced polymer (FRP) laminates are being successfully used for strengthening of

existing reinforced concrete (RC) and prestressed concrete (PC) structures. Bond of the

external FRP reinforcement to the concrete substrate is of critical importance for the

effectiveness of the technique, since it is the means for the transfer of stresses between

concrete and FRP in order to develop composite action. Some research efforts carried out so

far on this topic are outlined in the following.

Chajes et al. (1996) studied the bond and force transfer mechanism in composite material


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plates bonded to concrete, using a single -lap shear test specimen. A first set of tests was

performed to investigate the effect of surface preparation, type of adhesive and concrete

strength on the average bond strength. A second set of tests was used to study the force

transfer from the composite plates to the concrete. Two failure mechanisms were observed:

direct concrete shearing beneath the concrete surface and cohesive-type failure, depending on

the type of adhesive. Test results showed that (1) surface preparation of the concrete can

influence the bond strength, (2) if the failure mode of the joint is governed by shearing of the

concrete, the ultimate bond strength will be proportional to the square root of f'c, (3) there is

an effective bond length for a joint beyond which no further increase in failure load can be

achieved.

A group of researchers conducted a study on the effect of the type of concrete surface

preparation on the bond of carbon FRP (CFRP) sheets (Yoshizawa et al, 1996). The

specimen used in these tests was a concrete prism with CFRP sheets applied to two opposite

sides. The specimen was tested in tension, causing direct shear to be placed on the sheets.

The concrete surface of the specimens was prepared by either water jet or sandblasting. It

was found that the water jet doubled the capacity of the specimen as compared to

sandblasting. The bonded length of the CFRP sheet was determined to have little effect on

the ultimate load of the specimen.

Another group of researchers studied the effect of test method and quality of concrete on the

bond of CFRP sheets (Horiguchi and Saeki, 1997). Three different test methods, namely,

shear test, flexural test, and direct tensile test were investigated. The tensile test produced the

largest average bond strength, followed by the bending test. The lowest average bond

strengths were found in the shear test. Three failure modes were observed: shearing of the

concrete, delamination and FRP rupture. When the compressive strength of the concrete was

low, less than 3600 psi (25.3 MPa), then failure occurred in the concrete. Delamination


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occurred when the compressive strength was high or when the shear-type test was conducted.

FRP fracture was observed in the bending tests with higher strength of concrete. Bond

strength increased as the concrete compressive strength increased. The bonded length of the

CFRP sheet had minimal effect on the ultimate load.

Brosens and Van Gemert (1997) performed direct shear tests on two concrete prisms

connected with 3 layers of CFRP on two opposite sides. Their findings showed that an

increase in bonded length increases the failure load. This is contrary to findings of other

researchers. However, they did find that the influence of bonded length decreases at longer

lengths. They concluded that for computational purposes a linear bond stress distribution in

the FRP sheet may be assumed.

Another study on the bond mechanism of CFRP sheets was conducted by Maeda et al. (1997).

The variables in this testing were bonded length, number of layers of FRP, and type of FRP

sheet. Results of the tests showed that, as the stiffness of the fiber sheet increases, the

ultimate load increases. For bonded lengths above approximately 4 inches (100 mm), the

ultimate load did not change, implying the existence of an effective bond length.

Xie and Kharbari (1997) used a specially designed peel test to characterize the bond strength

between a carbon fiber/epoxy composite and a concrete substrate. The interfacial fracture

energy was measured for various peel angles and peel rates. When the peel angle increased,

or the peel rate decreased, the fracture energy at the interface was found to increase. Most

failures happened as cohesive failures in the adhesive. A finite-element analysis was

performed to evaluate the strain energy release rate in the peel test specimen for various crack

propagation paths: (1) cohesive in concrete (2) adhesive-concrete interface (3) cohesive in

adhesive (4) composite-adhesive interface. It was found that path 4 yielded the lowest strain

energy release rate, therefore, it is the most possible failure path. However, the difference

between paths 2, 3 and 4 is not significant. Cohesive debonding in the concrete substrate is


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the least possible failure mode. However, for a small peel angle the difference is small and

cohesive failure in the concrete is more possible than for a larger peel angle.

Täljsten (1994, 1997) carried out shear tests on concrete prisms with steel and CFRP bonded

plates. Results were compared with Volkersen's theory for lap joints and the comparison

showed that the theory can predict the shear stress in the joint fairly well for moderate load

levels. Failure in the joint always occurred by shearing of the concrete and test results seemed

to indicate that there exists a critical strain level at which the concrete starts to fracture. It was

also found that a critical anchor length exists above which longer lengths do not contribute to

the ultimate load.

Bizindavyi and Neale (1999) performed an experimental and analytical investigation on

transfer lengths and bond strengths of composite laminates bonded to concrete. The observed

modes of failure were shearing of the concrete beneath the glue line and rupture of the

composite coupon. An analytical model based on a shear lag approach was developed. A

comparison of analytical to experimental results was reasonably good for loads less than the

initial cracking load, especially for specimens with 1-ply bonded joints.

Volnyy et al. (1999) investigated bond between concrete and CFRP plates to be used as

connections in precast concrete walls. Failure occurred in all specimens due to one, or a

combination, of the following modes: CFRP rupture, delamination of the CFRP from the

concrete, and concrete surface shear failure. The stress distribution along the composite

appeared to be approximately linear, and to go asymptotically to zero after a certain distance,

the effective bond length of the connection.

Although a considerable amount of research has been devoted to bond of FRP laminates to

concrete, to date there is no acceptable prediction model of the bond failure load and the

effective bond length. In this project, flexural test specimens were prepared to address some

of the factors expected to affect bond, namely, bonded length, concrete strength, number of


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plies (stiffness), sheet width and, to a limited extent, surface preparation. Results are

presented and discussed in this paper. A shear lag approach, along with a simple shear model

for the evaluation of the slip modulus, is then used to model the strain distribution at moderate

load levels. Finally, expressions of the peeling load and the effective bond length are

presented. A design equation is proposed to calculate the effective FRP ultimate strain to be

used in design to account for bond-controlled failure.

EXPERIMENTS

Specimens

The specimen was a plain concrete beam with an inverted-T shape (Figure 1). The beam was

simply supported, with a span of 42 in (1067 mm) and a total length of 48 in (1219 mm). A

steel hinge at the top and a saw cut at the bottom, both located at mid-span, were used to

control the distribution of the internal forces. During loading, the saw cut caused a crack to

develop at the center of the beam and extend up to the hinge. Therefore, the compressive

force in the beam at mid-span was located at the center of the hinge and the internal moment

arm was known and constant for any given load level above the cracking load. This allowed

to compute with accuracy the tensile stress in the FRP.

A 2-in. (51-mm) wide CFRP strip was bonded to the tension face of the beam. A transverse

sheet was placed on one side to force failure to occur at the other end. Also, the sheet was left

unbonded approximately 2 in (51 mm) on each side of midspan. The design choices were

made to ensure that no cracking would occur within the bonded area.

Three series of specimens were tested. Each series consisted of six specimens with three

different bonded lengths. Either concrete strength or number of plies of CFRP was varied

between each series. Description of the specimens is reported in Table 1.


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After the beams had cured for 10-14 days, the concrete surface on which the CFRP had to be

applied was sandblasted to remove the top layer of mortar, just until the aggregate was

visible. The approximate depth of sandblasting was 0.06 in. (1.5 mm).

A composite system including a unidirectional fiber tow CFRP sheet, primer and saturant was

used for this experimental program. The sheet had an ultimate tensile strength of 620 ksi

(4360 MPa), a design tensile strength of 550 ksi (3870 MPa), a modulus of elasticity of 33000

ksi (232 GPa) and a fiber thickness of 0.0065 in (0.165 mm), as indicated by the manufacturer

(MBrace, 1998). The tensile properties of primer and saturant are reported in Table 2. The

resin was allowed to cure for at least 7 days prior to testing of the beams. The thickness of

each layer of the composite system was determined in previous work (Tumialan, 1998) by

using a Scanning Electron Microscope (SEM). The resulting thicknesses were: 0.017 in.

(0.432 mm) primer, 0.037 in. (0.940 mm) first resin layer, 0.0065 in. (0.165 mm) CFRP sheet.

After testing of the three series of specimens was complete, three more specimens were

tested. Two of them were made with 4-in. (102-mm) wide CFRP sheets instead of 2-in. (51-

mm) sheets. The purpose was to determine whether the width of the sheet had an effect on

the strain distribution. The bonded lengths of the specimens were 8 and 12 in. (203 and 305

mm), while the nominal concrete strength was 6000 psi (42.2 MPa). These specimens were

instrumented similarly to those of the first three series. Finally, one specimen was tested in

which the surface preparation was changed. The surface was roughened by adding notches

using a hammer and chisel. This specimen had 12-in. (305 mm) bonded length and 6000-psi

(42.2 MPa) nominal concrete strength. The purpose of this test was to determine if surface

preparation affected the average bond strength. No strain gauges were used in this test

because the main interest was to see if the ultimate load was increased by the different surface

preparation.


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Data collection was accomplished by a series of strain gauges located along the length of

CFRP sheet. One strain gauge was placed at the center of the unbonded region. For the 4-in.

(102-mm), 8-in. (203-mm) and 12-in. (305-mm) bonded lengths, there were 3, 4 and 6 strain

gauges placed along the centerline of the CFRP sheet in the bonded region, respectively. An

LVDT was located at mid-span to monitor the beam deflection.

Testing was performed on a universal testing machine. The beam was first loaded with 1500

lbs (6.67 kN) and then unloaded to 500 lbs (2.22 kN) to ensure that all data was being

recorded properly. Next, the beam was loaded until a crack formed at midspan of the beam,

and then unloaded to 500 lbs (2.22 kN). Load was then applied until failure.

Results

Test results in terms of ultimate load of the specimens are reported in Table 1. The data

collected from the strain gauges was used to develop strain-location curves. These curves

illustrate the strain vs. the distance the strain gauge is located from mid-span of the beam.

Each curve is plotted for a given load level. A typical strain-location graph for the 4-in. (102-

mm) bonded length can be seen in Figure 2, while Figure 3 shows a typical graph for the

longer bonded lengths. By comparing Figures 2 and 3, it can be seen that, at early stages of

loading, the curves show the same behavior. They both have a non linear shape, and the

strain gauges far from the center do not read strain. Also, as the load increases, the curves

tend to attain a linear shape. It can be assumed that joint failure begins immediately after the

point when the curve becomes linear. This corresponds to the attainment of a uniform bond

stress along the portion of laminate that is taking the load. The length of this portion is what

has been previously indicated as effective bond length. Once failure begins, the behavior in

the two cases is different. The specimens with the 4-in. (102-mm) bonded length fail

suddenly. For the 8 and 12-in. (203 and 305-mm) bonded lengths, the longer bonded length


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causes failure to occur in stages. This progressive failure is indicated by the strain

distribution. The strain becomes horizontal at the beginning of the bonded length, which

means that no load is transferred into the concrete in that portion of the bond because the joint

has begun to fail. In other words, the effective length of the CFRP sheet takes the entire load

to a certain point at which localized joint failure occurs causing the effective bond length to

shift. This shifting continues until the CFRP sheet has completely peeled from the concrete.

Visual inspection of the specimens after the test revealed that failure occurred in the concrete-

adhesive interface, with very little or no sign of damage in the concrete surface.

When comparing results, it was found that the bonded length did not affect the bond failure

load. In fact, it was found that the bond strength decreases as the bonded length increases for

all three series. It was concluded that an effective length exists beyond which no stress is

transferred until peeling occurs. As already outlined in the introduction, other researchers

have also reported the existence of a measurable bonded length beyond which no further

increase in the transferred load can be achieved. While it was expected that the concrete

strength would have an effect on the bond strength, there was no evidence from this

investigation. Since failure occurred at the concrete-epoxy interface, the concrete strength did

not affect the ultimate load. The number of plies used to make the CFRP laminate affects the

bond failure load. In order for two plies of CFRP sheet to be as efficient as one ply, the

ultimate load would have to double. As expected, this does not occur. The average of the

ultimate loads of Series II is only 1.5 times that of Series I.

Two specimens with 4-in. (102-mm) sheet width were tested in order to verify that the width

of the CFRP sheet did not change the bond strength. The two specimens, with 8-in. and 12-

in. bonded length, failed at a load of 7890 lbs (35.09 kN) and 7990 lbs (35.54 kN),

respectively. These ultimate loads are approximately twice those of the specimens with same

bonded length and 2-in. wide sheets. Also, failure mode and strain distribution resulted the


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same. It was concluded that the width of the sheet did not influence the bond strength.

Finally, the performance of the specimen with roughened surface was much better than that of

the specimen with sandblasted surface. The former failed at a load level of 5590 lbs (24.86

kN), as opposed to 3830 lbs (17.04 kN) and 3390 lbs (15.08 kN) of specimens 6-1-12-1 and

6-1-12-2. CFRP rupture was attained with the roughened surface. The sheet began peeling

until it reached the location of the first set of notches. The notches seemed to anchor the sheet

to the concrete. It was concluded that the surface preparation of the concrete can significantly

affect the average bond strength.

ANALYSIS OF RESULTS

Linear Analysis of the Bonded Joint

The analysis of the bonded joint in the linear elastic range can be conducted by means of a

simple shear lag approach. The differential equation governing bond is as follows:

0))((1

2

2

=⋅⋅

− xsEtdx

sdτ (1)

where s is the slip, τ the bond stress, x the coordinate along the bonded length of the laminate,

t the thickness and E the elastic modulus of the FRP.

Equation (1) comes from equilibrium and compatibility relations on a finite element of sheet

of length dx, along with the assumptions that: (1) the FRP sheet is linear elastic, (2) the

concrete strain is negligible if compared to that of the FRP, (3) the adhesive is only exposed

to shear forces. At moderate load levels, a linear bond stress-slip behavior can be adopted:

sK ⋅=τ (2)

Solving (1) with τ(s) given by (2) yields the following functions:

xcoshCxsinhC)x(s αα ⋅+⋅= 21 (3a)

xsinhCxcoshC)x( ααααε ⋅⋅+⋅⋅= 21 (3b)


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)()( xsKx ⋅=τ (3c)

where:

EtK⋅

=α (4)

and C1 and C2 are to be determined from the boundary conditions. If the coordinate x is such

that x = 0 corresponds to the free end of the sheet, where no strain exists, and x = l (being l

the bonded length) corresponds to the end of the sheet where the external load is directly

applied, the boundary conditions are as follows:

0)0( =ε (5a)

El applσ

ε =)( (5b)

Equations (3a-c) become:

xcoshlsinhE

)x(s appl ααα

σ⋅

⋅⋅= (6a)

xsinhlsinhE

x appl αα

σε ⋅

⋅=)( (6b)

)()( xsKx ⋅=τ (6c)

The previous approach has been previously applied to the analysis of bonded FRP plates. In

this case K, the so called slip modulus, can be established by means of a simple shear model

as the ratio of the shear modulus of the adhesive and the thickness of the adhesive layer.

When FRP sheets rather than plates are used, such as in our case, the slip modulus can be

evaluated as follows:

int

int

tG

K = (7)

where Gint = shear modulus of the FRP sheet-concrete interface and tint = thickness of the

interface. Since the composite was formed using the wet lay-up technique, the resin system


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intrinsically serves both as the matrix for the composite and the interlayer between the

concrete and the composite, i.e., as the adhesive layer. This layer then is the main medium for

the transfer of shear stresses between the composite and the concrete (Kharbari et al., 1997).

The layer of primer must be also considered. Therefore, Gint and tint can be computed from

the thicknesses obtained from the SEM and the elastic properties of resin and primer given in

Table 2. From the rule of mixture for the shear modulus:

resinprimerprimerresin

primerresin

GtGt

GGG

+⋅

⋅=int (8)

where:

)1(2 resin

resinresin

EG

ν+= (9)

and the same relation is valid for primer. Using equations (8) and (7), a shear modulus of

interface Gint = 74.86 ksi (526.3 MPa) and a slip modulus K = 1386 ksi/in. (383.6 MPa/mm)

were computed for the tested specimens. Figure 4 shows the comparison of experimental and

theoretical strain distributions for specimen 6-1-18-1 at moderate load levels. It can be seen

that reasonable agreement is achieved. The correlation is less accurate in the case of two FRP

layers, the experimental response being somewhat stiffer than the theoretical curve.

Local Bond Stress – Slip Relationship and Fracture Energy

The local τ-slip curves were obtained from the experimental data as follows. The bond stress,

τ, can be found by equilibrium of forces, considering also the linear elasticity of the FRP:

dxxd

Etx r )()(

ετ ⋅⋅= (10)

where ε r is the FRP strain. Therefore, the τ vs. location diagram can be obtained from the

first derivative of the strain vs. location diagram multiplied by the elastic modulus E and the


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thickness t of the FRP sheet. From the definition of slip, and assuming that the concrete

strain, ε c, may be considered negligible when compared to the FRP strain, ε r, it is:

rdxds ε= (11)

from which:

∫+=x

r dxxsxs0

)()0()( ε (12)

Assuming that s(0), slip at the free end of the sheet, can be considered negligible prior to

peeling, the slip vs. location diagram can be obtained by integrating the strain vs. location

curve. The local τ – slip rela tionship can be obtained by combining the two curves τ(x) and

s(x) expressed by equations (10) and (12). A τ vs. x curve and a s vs. x curve can be obtained

for each value of the applied load. Therefore, a τ vs. slip diagram corresponding to each

value of applied load can be drawn. Figure 5 illustrates the τ - slip curves obtained from some

of the tested specimens at a load level corresponding to imminent peeling. This was

identified as the load level at which the strain distribution becomes linear. A first steep

ascending branch is followed by a second region characterized by plastic or softening

behavior, until an ultimate value of slip is reached. The plastic branch of the curve explains

the attainment of a linear strain distribution close to peeling. The limited value of ultimate

slip explains the brittle behavior of the joint, that is, the localization of load transfer within a

short effective area even for long bonded lengths.

The area underneath the τ - slip curve, indicated as G f, is the fracture energy per unit area of

the bonded joint. The values of G f and τm, the latter being the maximum value of τ in the τ -

slip curve, were obtained from the τ - slip curves. The parameter τm was consistently higher

for the specimens of the second series (two plies of FRP) than for those of the first series (one


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ply of FRP). The average values were 0.51 ksi (3.59 MPa) for the first series and 0.74 ksi

(5.20 MPa) for the second one. The ratio between the two is:

I

II

mI

mII

tE

tE

)(

)(45.1 ≅=

ττ

(13)

Therefore, it can be assumed that τm is proportional to tE , that is:

tEkm 1=τ (14)

Computing k1 for each specimen and taking the average, a value of 0.0184 with a coefficient

of variation of 12.6% can be obtained, therefore:

tEm 0184.0=τ (15a)

with t in mm, E and τm in MPa, or:

tEm 035.0=τ (15b)

with t in in., E and τm in ksi.

The value of Gf was consistently the same for the 1-ply and 2-plies cases. Therefore, it can be

concluded that, as the stiffness of the FRP sheet increased, the local τ - slip relationship

experienced an increase in the “local bond strength”, τm, and a decrease in the “local bond

ductility” such that the area underneath the τ - slip curve (that is, the fracture energy)

remained the same. An average value of 6.05 lbs-in./in.2 (1.06 N-mm/mm2) was obtained for

Gf.

Peeling Load and Effective Bond Length

The issue of bond is of great importance in design of externally bonded FRP laminates.

When failure is bond-controlled, the maximum stress in the FRP to be considered for design

cannot be equal to the tensile strength of the FRP material. A possible approach, as


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discussion within the ACI committee 440 seems to indicate, would be taking a reduced

ultimate strain level in the FRP reinforcement (American Concrete Institute, 1999):

urub k εε ⋅= (16)

where εu is the FRP ultimate strain, and k r is the reduction factor. Since FRP materials are

linearly-elastic up to failure, limiting the strain level will dictate the limitation to the stress

level. The problem is, therefore, the determination of k r.

Täljsten developed the following equation for the bond failure load by using non linear

fracture mechanics:

fub EtGbP 2= (17)

where b, E and t are width, elastic modulus and thickness of the FRP sheet, respectively and

Gf is the fracture energy per unit area of the joint. However, equation (17) can also be

obtained simply by equilibrium of the element dx along with linear elasticity of the composite

sheet. This equation is valid in the case of brittle joints.

From equation (16), the tensile strain in the loaded end of the FRP sheet at bond failure results

as follows:

tE

G fub ⋅

⋅=

2ε (18)

In the experiments described in this paper, failure always occurred at the interface between

CFRP sheet and concrete. Therefore, the value of Gf reported in the previous section is the

fracture energy related to this particular crack propagation path. As outlined in the

introduction, in most of the experiments on bond conducted so far, failure occurred by

shearing of a thin concrete layer underneath the FRP laminate. The fracture energy related to

cracking within the concrete substrate depends on the concrete compressive strength and

surface preparation. If the value of Gf corresponding to cracking of the concrete could be

determined as a function of these two variables, it would be possible to predict the bond


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failure mode, once Gf related to delamination is known for the given type of adhesive. The

lower of the two values of G f would control. Therefore, equation (17) could be used to

determine the corresponding peeling load. However, Xie and Kharbari (1997) reported that

the difference of strain energy release rate between the possible crack paths is not significant

for small peel angles, and the peel angle in our case is equal to zero.

Table 3 summarizes some of the experimental data currently available in the literature. For

the values of bonded lengths that have been considered, the bond failure load does not depend

on the bonded length. Furthemore, the difference in failure mode does not affect considerably

the fracture energy, for what has been said in the previous paragraph. Therefore, all data can

be plotted in the same graph. This has been done in Figure 6, where k r is plotted versus the

product Et, which represents the axial stiffness for unit width of the laminate. Best fit

interpolation conducted to the following equation that gives k r as a function of Et:

Etkr

75.6≅ (19a)

where E is in ksi and t in in., or:

Etkr

51.0≅ (19b)

where E is in MPa and t in mm. This expression is coherent with equation (18) that follows

just by equilibrium considerations.

In order to quantify the influence of the concrete strength, the values of k r from the

experimental data of Table 3 were all reconducted to the same stiffness value by using

equation (19). When plotting the "corrected" values of k r versus the concrete compressive

strength, no consistent relationship could be found. It appears that the laminate stiffness is the

parameter that mostly influences the bond failure load. Equation (19) is suggested as a

possible design equation to calculate k r and then, through (16), the effective maximum strain


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in the FRP to be used for design. Further study is necessary to quantify the influence of

surface preparation on the bond performance.

The effective bond length can be expressed as follows:

peel

ubeff

dxd

lεε

= (20)

Unlike calculation of the ultimate load, calculation of the effective bond length requires an

assumption on the shape of the local τ-slip relationship of the joint. If the τ-slip curve can be

modeled as having an initial ascending branch followed by perfectly plastic behavior at a

value τm of bond stress, it is:

tEdxd m

peel

τε= (21)

Using (18) and (21), (20) becomes:

m

feff

EtGl

τ

2= (22)

From equation (15) and the value of Gf reported in the previous section, it follows:

mm) (80 in. 14.3=effl (23)

Other researchers indicated that the effective bond length was between 3 in. (76 mm) and 4 in.

(102 mm).

CONCLUSIONS

Flexural test specimens were prepared to address some of the factors expected to affect bond.

Failure occurred in the concrete-adhesive interface, with very little or no sign of damage in

the concrete surface. The bonded length did not affect the ultimate load, so confirming the

existence of an effective length beyond which no stress is transferred until peeling occurs.

Due to the failure mode, the concrete strength did not affect the ultimate load. The number of


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plies used to make the CFRP laminate affected the bond failure load, but the average of the

ultimate loads of Series II (2 plies) was only 1.5 times that of Series I (1 ply). The width of

the sheet did not influence the bond strength. Finally, the performance of the specimen with

roughened surface was much better than that of the specimen with sandblasted surface.

Failure occurred in the former by rupture of the FRP sheet at a remarkably higher load. It was

concluded that the surface preparation of the concrete can significantly affect the average

bond strength.

A shear lag approach, along with a simple shear model for the evaluation of the slip modulus,

was used to model the strain distribution at moderate load levels. A reasonable agreement

with the experimental results was achieved, especially for specimens with 1-ply laminates.

Finally, expressions of the peeling load and the effective bond length were presented. A

design equation was proposed to calculate the effective maximum FRP strain to be used in

design to account for bond-controlled failure. It appears that the laminate stiffness

significantly influences the bond failure load. Further study is necessary to quantify the

influence of surface preparation on the bond performance.

ACKNOWLEDGEMENTS

The authors would like to acknowledge the University of Missouri - Rolla NSF

Industry/University Cooperative Research Center on Repair of Building and Bridges with

Composites (RB2C) for supporting this project.

REFERENCES

American Concrete Institute Committee 440 (1999). "Guide for the Design and Construction

of Externally Bonded FRP Systems for Strengthening Concrete Structures". ????


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Bizindavyi, L. and Neale, K.W. (1999). “Transfer Lengths and Bond Strengths for

Composites Bonded to Concrete,” Journal of Composites for Construction, ASCE,

vol. 3, No. 4, pp. 153-160.

Brosens, K. and Van Gemert, D. (1997). “Anchoring Stresses between Concrete and Carbon

Fibre Reinforced Laminates,” Non-Metallic (FRP) Reinforcement for Concrete

Structures, Vol. 1, Japan Concrete Institute, Japan, pp. 271-278.

Chajes, M.J.; Finch, W.W.Jr.; Januszka, T.F.; and Thomson, T.A. (1996). "Bond and Force

Transfer of Composite Material Plates Bonded to Concrete", ACI Structural Journal,

ACI, Vol. 93, No. 2, pp. 295-303.

Horiguchi, T. and Saeki, N. (1997). “Effect of Test Methods and Quality of Concrete on

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19

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20

Table 1: Description of Specimens and Test Results

Series Number

Specimen Code

Concrete Class [fc′]

(psi)*

Number of Plies

Bonded Length (in.)

Ultimate Load (lbs)

6-1-4-1 3720 6-1-4-2 4 3990 6-1-8-1 3560 6-1-8-2

8 3190

6-1-12-1 3830

I

6-1-12-2

6000 [6860] 1

12 3390 6-2-4-1 5930 6-2-4-2 4 5140 6-2-8-1 4630 6-2-8-2

8 6260

6-2-12-1 5590

II

6-2-12-2

6000 [5900]

2

12 5080 3-1-4-1 3300 3-1-4-2

4 3120

3-1-8-1 4450 3-1-8-2 8 2920 3-1-12-1 4770

III

3-1-12-2

3000 [3550] 1

12 3450 *Top number is class of concrete while 28-day strength is shown in brackets

Note: 1 in. = 25.4 mm; 1 psi = 7.03 kPa


21

Table 2: Neat Resin Properties in Tension (ASTM D-638)

Primer Saturant Maximum Stress psi 2500 8000 Stress at Yield psi 2100 7800

Stress at Rupture psi 2500 7900 Strain at Max. Stress 0.400 0.030

Strain at Yield 0.040 0.025 Strain at Rupture 0.400 0.035

Elastic Modulus psi 104,000 440,000 Poisson’s Ratio 0.48 0.40

Note: 1 psi = 7.03 kPa


22

Table 3: Literature Review

Bonded Length

(in.)

Concrete Strength

(psi)

tE (ksi-in)

Surface Preparation

Test Method

Resin Modulus

(ksi)

Ultimate Load/Width

(lbs/in)

Failure Mode Ref.

3 5235 None 750 1902 CS 3 6833 Grinding 750 2232 CS 3 6833 Mech. Abr. 750 2391 CS 3 6325 Mech. Abr. 320 2367 CS 3 6325 Mech. Abr. 34 2013 Resin 3 6325 Mech. Abr.+Prim. 34 2160 Resin 3 6325 Mech. Abr. 230 2364 CS 3 6325 Mech. Abr.+Prim. 230 2517 CS 4 5283 Mech. Abr.+Prim. 230 2879 CS 6 5283 Mech. Abr.+Prim. 230 2679 CS 8 5283

629.2

Mech. Abr.+Prim.

Direct Shear

230 2600 CS

[1]

3 5803 141.7 663 CS 6 5803 141.7 1051 CS

12 6159 141.7 1365 CS 3 6031 351.1 1142 CF 6 6031 351.1 834 FR 6 6074 283.4 1856 CS

28 6074 141.7

Sander+ Primer

Direct Shear N/A

1142 CS

[2]

4 12230 1976 8 13200 3140

12 14180 4008 16 14180

1142.5 Sandblasting Direct Shear 953

3072

CS [3]

4 6860 214.5 1860 4 6860 214.5 1995 8 6860 214.5 1780 8 6860 214.5 1595

12 6860 214.5 1915 12 6860 214.5 1695 4 5900 429 2965 4 5900 429 2570 8 5900 429 2315 8 5900 429 3130

12 5900 429 2795 12 5900 429 2540 4 3550 214.5 1650 4 3550 214.5 1560 8 3550 214.5 2225 8 3550 214.5 1460

12 3550 214.5 2385 12 3550 214.5

Sandblasting Beam Test 440

1725

Resin [4]

Note: [1] = Chajes et al.; [2] = Maeda et al.; [3] = Taljsten; [4] = present study.

Note: 1 in. = 25.4 mm; 1 psi = 7.03 kPa; 1 lb = 4.45 N


23

10”

2” (Both Sides)

UNBONDED

4”

4”

4”

BL10”

48”

8”

48”

21”

2”

10’4” Saw Cut

Hinge

MONITOREDSIDE

Figure 1: Test Specimen

Note: 1 in. = 25.4 mm


24

0

2000

4000

6000

8000

10000

2 3 4 5 6Distance From Center (in)

Stra

in in

the

FRP

( µε) 1600 lb

2000 lb2500 lb3000 lb3500 lb3700 lb

6-1-4-1

Figure 2: Strain vs. Location

Note: 1 in. = 25.4 mm; 1 lb = 4.45 N


25

6-1-12-1

0

2000

4000

6000

8000

10000

2 4 6 8 10 12 14Distance From Center (in)

Stra

in in

the

FRP

( µε)

1500 lb2000 lb2500 lb3000 lb3300 lb3700 lb

Figure 3: Strain vs. Location (1 in. = 25.4 mm; 1 lb = 4.45 N)

Note: 1 in. = 25.4 mm; 1 lb = 4.45 N


26

0

1000

2000

3000

4000

5000

6000

7000

2 4 6 8 10

Distance from Center (in.)

Stra

in in

the

FRP

( µε)

P = 600 lbs theo.P = 600 lbs exp.P = 2000 lbs theo.P = 2000 lbs exp.P = 2500 lbs theo.P = 2500 exp.

Figure 4: Theoretical and Experimental Strain Distribution for Specimen 6-1-18-1

Note: 1 in. = 25.4 mm; 1 lb = 4.45 N


27

0

0,1

0,2

0,3

0,4

0,5

0,6

0,7

0 0,005 0,01 0,015slip (in.)

τ (k

si)

6-1-4-16-1-8-16-1-8-26-1-12-16-1-12-2

Figure 5: Local τ - slip Relationships

Note: 1 in. = 25.4 mm; 1 psi = 7.03 kPa


28

y = 6,7526x-0,51

R2 = 0,7725

0

0,1

0,2

0,3

0,4

0,5

0,6

0,7

0,8

0,9

1

0 200 400 600 800 1000 1200tE (kips/in.)

kr

Chajes et al.Maeda et al.Taljstenpresent study

Figure 6: Experimental Results in Terms of k r

Note: 1 in. = 25.4 mm; 1 kip = 4.45 kN

Documents

BOND OF FRP LAMINATES TO CONCRETE - Missouri …transportation.mst.edu/media/research/transportation/documents/...De Loren zis, L., B. Miller, And A. Nanni, "Bond of FRP Laminates