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copyNear-surface mounted FRP reinforcement: An emerging techniquefor strengthening structures

L. De Lorenzis a,*, J.G. Teng b

a Department of Innovation Engineering, University of Lecce, via per Monteroni, 73100 Lecce, Italyb Department of Civil and Structural Engineering, The Hong Kong Polytechnic University, Hong Kong, China

Received 27 January 2006; accepted 17 August 2006Available online 18 October 2006

Abstract

Near-surface mounted (NSM) fiber-reinforced polymer (FRP) reinforcement is one of the latest and most promising strengtheningtechniques for reinforced concrete (RC) structures. Research on this topic started only a few years ago but has by now attracted world-wide attention. Issues raised by the use of NSM FRP reinforcement include the optimization of construction details, models for the bondbehaviour between NSM FRP and concrete, reliable design methods for flexural and shear strengthening, and the maximization of theadvantages of this technique. This paper provides a critical review of existing research in this area, identifies gaps of knowledge, andoutlines directions for further research. 2006 Elsevier Ltd. All rights reserved.

Keywords: A. Polymer matrix composites (PMCs); B. Debonding; C. Analytical modeling; D. Mechanical testing; Near-surface mounted reinforcement

1. Introduction

Over the past decade, extensive research has beenconducted on the strengthening of reinforced concrete(RC) structures using externally bonded fiber-reinforcedpolymer (FRP) laminates; the technology has also beenimplemented in a large number of practical projects world-wide. More recently, near-surface mounted (NSM) FRPreinforcement has attracted an increasing amount of

research as well as practical application. In the NSMmethod, grooves are first cut into the concrete cover ofan RC element and the FRP reinforcement is bondedtherein with an appropriate groove filler (typically epoxypaste or cement grout). What is herein called NSM rein-forcement was previously given other names such asgrouted reinforcement[1], or embedded reinforcement[2,3].

Examples of the use of NSM steel rebars in Europe forthe strengthening of RC structures date back to the early1950s[1]. More recent applications of NSM stainless steelbars for the strengthening of masonry buildings and archbridges have also been documented (e.g. [4]). The advanta-ges of FRP versus steel as NSM reinforcement are betterresistance to corrosion, increased ease and speed of instal-lation due to its lightweight, and a reduced groove size dueto the higher tensile strength and better corrosion resis-

tance of FRP.Compared to externally bonded FRP reinforcement, theNSM system has a number of advantages: (a) the amountof site installation work may be reduced, as surface prepa-ration other than grooving is no longer required (e.g., plas-ter removal is not necessary; irregularities of the concretesurface can be more easily accommodated; removal ofthe weak laitance layer on the concrete surface is no longerneeded); (b) NSM reinforcement is less prone to debondingfrom the concrete substrate; (c) NSM bars can be more eas-ily anchored into adjacent members to prevent debondingfailures; this feature is particularly attractive in the flexuralstrengthening of beams and columns in rigidly-jointed

1359-8368/$ - see front matter 2006 Elsevier Ltd. All rights reserved.

doi:10.1016/j.compositesb.2006.08.003

* Corresponding author. Tel.: +39 0832 297241; fax: +39 0832 297279.E-mail address: [email protected](L. De Lorenzis).

www.elsevier.com/locate/compositesb

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frames, where the maximum moments typically occurat the ends of the member; (d) NSM reinforcementcan be more easily pre-stressed; (e) NSM bars are pro-tected by the concrete cover and so are less exposed to acci-

dental impact and mechanical damage, fire, and vandalism;this aspect makes this technology particularly suitable forthe strengthening of negative moment regions of beams/slabs; (f) the aesthetic of the strengthened structure isvirtually unchanged. Due to the above advantages, theNSM FRP method is in many cases superior to theexternally bonded FRP method or can be used in combina-tion with it, provided that the cover of the member is suf-ficiently thick for grooves of a desirable size to beaccommodated.

The existing knowledge on the NSM FRP method ismuch more limited than that on the externally bondedFRP method, as reflected by the absence of relevant provi-

sions in the existing guidelines on the FRP strengthening ofconcrete structures issued by fib [5] and ACI-440[6]. How-ever, the international engineering community has becomeincreasingly aware of the practical advantages of thismethod, which has led to accelerations of research andpractical applications worldwide. Both ACI-440 and fibare currently considering revisions to their documents toinclude NSM-related provisions. Against this background,this paper provides a critical review of existing research inthis area, identifies gaps of knowledge, and outlines direc-tions for further research.

This paper focuses on research work on the structural

aspects of NSM strengthening of concrete structures. Dis-cussions on some significant practical applications of theNSM method, on NSM strengthening of masonry and tim-ber structures, and on durability-related aspects are givenin Ref.[7]. Space limitation does not allow them to be dis-cussed in this paper.

2. Materials and systems

2.1. FRP reinforcement

In most existing studies, carbon FRP (CFRP) NSMreinforcement has been used to strengthen concrete struc-tures. Glass FRP (GFRP) has been used in most applica-tions of the NSM method to masonry and timberstructures. The present authors are not aware of any studyor practical application in which aramid FRP (AFRP) wasused. The tensile strength and elastic modulus of CFRP aremuch higher than those of GFRP, so for the same tensilecapacity, a CFRP bar has a smaller cross-sectional areathan a GFRP bar and requires a smaller groove. This inturn leads to easier installation, with less risks of interferingwith the internal steel reinforcement, and with savings inthe groove-filling material.

FRP bars can be manufactured in a virtually endless

variety of shapes. Hence, the NSM FRP reinforcementmay be round, square, rectangular and oval bars, as wellas strips (Figs. 1 and 2). For brevity, the term bars is

used herein as a generic term encompassing all cross-sectional shapes, while the term strips is reserved forthin narrow strips. Different cross-sectional shapes havedifferent advantages, and offer different choices for practi-cal applications. For example, square bars maximize the

bar sectional area for a given size of square groove whileround bars are more readily available and can be moreeasily anchored in pre-stressing operations. Narrow strips

Fig. 1. Types of FRP bars for NSM applications.

Fig. 2. Different NSM systems and nomenclature.

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maximize the surface area-to-sectional area ratio for agiven volume and thus minimize the risk of debonding,but require a thicker cover for a given cross-sectional area.In practical applications, the choice depends strongly on

the constraints of a specific situation, such as the depthof the cover, and the availability and cost of a particulartype of FRP bar.

FRP bars are also manufactured with a variety of sur-face textures, which strongly affect their bond behaviouras NSM reinforcement. Their surface can be smooth,sand-blasted, sand-coated, or roughened with a peel-plysurface treatment. Round bars can also be spirally woundwith a fiber tow, or ribbed[8].

2.2. Groove filler

The groove filler is the medium for the transfer of stres-

ses between the FRP bar and the concrete. In terms ofstructural behaviour, its most relevant mechanical proper-ties are the tensile and shear strengths. The tensile strengthis especially important when the embedded bars have adeformed surface, which produces high circumferentialtensile stresses in the cover formed by the groove filler(simply referred to as the cover or the epoxy coverhereafter) as a result of the bond action. In addition, theshear strength is important when the bond capacity ofthe NSM reinforcement is controlled by cohesive shear fail-ure of the groove filler. The effect of the modulus of elastic-ity of the groove filler has never been experimentally

investigated.The most common and best performing groove filler

is a two-component epoxy. Low-viscosity epoxy can beselected for strengthening in negative moment regions asthe epoxy can be poured into the grooves. For othercases, a high-viscosity epoxy is needed to avoid drippingor flowing-away. The addition of sand to epoxy canincrease the volume, control the viscosity, lower the coeffi-cient of thermal expansion, and raise the glass transitiontemperature. A drawback of this addition seems to bereduced adhesion at the barepoxy interface for a smoothbar surface[2].

The use of cement paste or mortar in place of epoxy as agroove filler has recently been explored in an attempt tolower the material cost, reduce the hazard to workers, min-imize the environmental impact, allow effective bonding towet substrates, and achieve better resistance to high tem-peratures and improved thermal compatibility with theconcrete substrate. However, cement mortar has inferiormechanical properties and durability, with a tensilestrength an order of magnitude smaller than that of com-mon epoxies. Results of bond tests and flexural tests[9,10]have identified some significant limitations of cementmortar as a groove filler. Given these limitations and thevery limited data available, the rest of this paper is focused

on epoxy-bonded NSM FRP reinforcement only, exceptwhen future research needs are discussed. Nevertheless,tests on NSM FRP reinforcement using cement grout as

the groove filler, if available, are included in Tables 14for completeness.

2.3. Groove dimensions

Fig. 2shows several configurations of NSM FRP rein-forcement, where db is the nominal diameter of a roundbar, and tf and hf are the thickness/width and the heightof an FRP strip or rectangular bar respectively. The groovewidthbg, the groove depth hg, the net distance between twoadjacent groovesag, and the net distance between a grooveand the beam edge aeare all relevant construction param-eters, which can influence the bond performance and hencethe structural behaviour.

For round bars, De Lorenzis [11], based on results ofbond tests with square grooves (bg= hg) and defining

k=bg/db, proposed a minimum value of 1.5 for k forsmooth or lightly sand-blasted bars and a minimum valueof 2.0 forkfor deformed bars. Parretti and Nanni[12]sug-gested that bothbgand hgshould be no less than 1.5db. ForNSM strips, Blaschko [13] suggested that the depth andwidth of the cut groove should be about 3 mm larger thanthe height and thickness of the corresponding FRP striprespectively, so to obtain an adhesive layer thickness ofabout 12 mm. Also for NSM strips, Parretti and Nanni[12] recommended that the minimum width of a groovebe no less than 3tfand the minimum depth be no less than1.5hf. For a more detailed discussion, see the section onbond behaviour.

In the existing studies, NSM strips were bonded usingepoxy either along all four sides of the strip surface[14,15], or along three sides of the strip surface only[16,17] (Fig. 2). Due to the large width to thickness ratioof the strips, the reduction in the bond surface in the lattercase is negligible. In existing tests on NSM square bars [18],only three sides of the bar surface were bonded to the con-crete member.

2.4. Groove position

If a single NSM bar is to be provided to the tension side

of an RC member, it should naturally be centrally locatedover the beam width. When two or more NSM bars need tobe provided, then the distance between two adjacent NSMbars and the distance between the edge of the member andthe adjacent bar become important design parameters. Theeffect of these parameters is discussed in the section onbond behaviour.

2.5. Constructional aspects

Compared with the use of externally bonded FRP lam-inates, the need to cut grooves into the concrete member in

the construction process of the NSM method is the key dif-ference. A detailed discussion of the construction processcan be found in Ref. [7].

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Table 2Local bond strength of NSM systems (saw-cut grooves)

Type of bar (material,cross-sectional shape)

Surfaceconfiguration

Nominaldbortf hf (mm)

kor bg hg(mm)

Groovefiller

Direct tensile strength ofgroove filler (MPa)

f0c (MPa) Local bondstrength (MPa)

Mainb

failure modeReference

CFRP, round Sand-blasted 9.5 1.342.67 Epoxy 13.8 From manufacturer 28 8.6 BE-I [21]12.7 1.25 9.7

CFRP, round Ribbed 9.5 1.33 27.4 From testing 22 11.2 SP-C1 [26]1.59 15.42.12 16.6

GFRP, round Ribbed 9.5 1.36 9.11.64 10.02.18 12.5

CFRP, round Spiral ly wound andsand-coated

7.5 1.50 18.02.00 20.82.50 21.9

CFRP, square Smooth 10 10 N/Aa 31 From manufacturer N/Aa 9.0c N/Aa [9]CFRP, strip N/Aa 1.5 9.6 3.3 15 1622 From testing 35, 45, 70 19.8 BE [14]

Roughened (1.22) 20 Average3.3 22

33.3 From testing 32 and 46 20.0d BE-C [16]

Roughened 5 16 9 22 42.6 From testing 29 Cubestrength

1012 BE-I [20]

CFRP, round Ribbed 9.5 1.59 Cementpaste

6.3, Bending tensile strengthfrom testing

22 9.7 SP-C1, BE-C [26]2.21 6.4

GFRP, round Ribbed 9.5 1.64 8.02.27 8.3

CFRP, round Spiral ly wound andsand-coated

7.5 1.50 6.7

CFRP, square Sand-coated 10 10 N/Aa 9, Bending tensile strengthfrom manufacturer

N/Aa 4.3c N/Aa [9]

a N/A = not available.b In some cases more than one failure mode were concurrent.c Taken from the specimen with the shortest bonded length (100 mm).d With no influence of edge distance (i.e. a 0e P 150 mm).

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Table 3

Summary of existing experimental work on flexural strengthening with NSM FRP reinforcement beams with limited bonded lengths, slabs, and columns

Beams with limited bonded lengths Slabs Columns

Reference [17] [19] [20] [2] [42] [43] [42]

Test method 3-Point bending 4-Point bending Patch loading at centre 3-Point bending Uniform line load

at mid-span or at 330

from free end of

cantilever slab

Horizontal load

at top

Cross-sectional

shape anddimensions

(mm)

T, total height = 300,

web height = 250,flange width = 300,

web width = 150

Rectangular,

150 300

5400 229 RC, simply

supported

1800 229 PC, simply

supportedc

7600 460 1180 406 PC 600 600

Net span (m) 2.5 3 3 7.9 4.9 (mid-span), 1.8

(cantilever)

Height

1.83.4

Shear span (m)

f0c (MPa) 48 35 28 41 56 4550 56

Groov e fi ller Ep oxy

Direct tensile

strength of

groove filler

(MPa)

N/Aa 42.6 from testing N/Aa N/Aa N/Aa N/Aa N/Aa

Type of FRP/cross-

sectional shape/

surface

configuration

CFRP/strip/

N/AaCFRP/

round/RB

CFRP/

strip/PP

C FR P/ round/ SM C FR P/ round/ SM C FR P/ round/ SB C FR P/ round/ N/ Aa CFRP/round/

SBCFRP/strip/N/Aa CFRP/strip/N/Aa

CFRP/round/RB

FRP nominal db or

tf hf(mm)

1.2 25 9.5 4 16 3 9.5, 1.2 25 11 10; 1.4 25; 9.5 11

Number of FRP

bars

1 1 1 4 Bars every 102 mm

o.c. (long.), 1 bar every

102 mm o.c. (transv.)

1 Bar every 102 mm

o.c. (long.), 1 strip

every 51 mm o.c.

(long.)

20 bars (one every

381 mm o.c.)

10; 18; (mid-span)

6; 6; 8 (cantilever)

3 on each face

or 7 on each

face

Elastic modulus of

FRP (GPa)

150 111 151 172 172 (bars),

130 (strips)

119 147; 150; 111 119

Tensile strength of

FRP (MPa)

2000 1918 2068 1550 1550 (bars),

1790 (strips)

1240 1970; 2000; 1918 1240

bg hg (mm) 5 25 18 30 8 22 9.5 13 16 25, 3.2 29 14 19 16 30; 5 25; 16 30 14 19

Steel tension

reinforcement

Two 10-mm bars Two 12-mm bars 16-mm bars @102 mm

o.c. (long.) and

204 mm o.c. (transv.)

None 25-mm bars @127 mm

o.c. (long.) and 13-mm

bars @457 mm o.c.

(transv.)

Five 16-mm mild bars

(mid-span), four 16-mm

mild bars (cantilever) +

twelve 15-mm 7-wire

pre-stressing strands

Four 19-mm

bars

Bar anchorage

lengthb (mm)

1501200 1501200 01150 Termination: 152 from

support, 762 from free

edge

N/Aa Bar length 6 m N/Aa 381 mm in the

footing

Test variables Bar anchorage length Type of bar Number of bars

Observed failure

modes

Debonding by EC-Cb at

cut-off and MMR, BR

Concrete

Splittingc

(MMR)

CCS (cut-off),

CCS (MMR)

Punching

shear + DBdCC + DBd (bars),

CC (strips)

BR CC BR, CC

Increase in ultimate

load (%)

054 041 0106 15 300 27 36; 43; 39 102, 177

Acronyms: BR = bar rupture, CC = concrete crushing, CCS = concrete cover separation; DB = debonding; MMR = maximum moment region; RC = reinforced concrete; PC = pre-stressed concrete; PP = roughened with peel-ply

surface treatment; RB = ribbed; SB = sand-blasted; SM = smooth.a N/A = not available.b SeeFig. 3.c Not clear whether it was CCS.d

Debonding mechanism not clear.

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3. Bond behaviour

3.1. Summary of existing work

The bond between an NSM bar and the substrate mate-rial plays a key role in ensuring the effectiveness of theNSM strengthening method. The performance of the bonddepends on a number of parameters: the groove and thebar dimensions, the tensile and shear strengths of the con-crete and the groove filler, the bar cross-sectional shapeand surface configuration, and the degree of roughness ofthe groove surface. This large number of parameters callsfor extensive laboratory characterization, as well as analyt-ical and numerical modelling.Table 1provides a summaryof the existing experimental data on bond behaviour. Notethat the tests in some existing studies [17,19,20], althoughaddressing bond-related aspects, were conducted on beams

under bending and are therefore discussed in a latersection.

3.2. Failure modes and mechanisms

The bond tests summarized in Table 1 identified differ-ent possible bond failure modes of NSM systems (Fig. 3).These modes are described in some detail below and theunderlying mechanisms examined.

3.2.1. Bond failure at the barepoxy interface

This mode may occur as either pure interfacial failure(BE-I), or as cohesive shear failure in the groovefiller (BE-C). The BE-I mode is critical for bars with asmooth or lightly sand-blasted surface, i.e. when thedegree of surface deformation is insufficient to providemechanical interlocking between the bar and the groovefiller and the bond resistance relies primarily on adhesionbetween the bar and the filler. For round bars, this modebecomes critical if the groove size is sufficiently large toavoid splitting failure of the groove filler. For epoxy andconcrete of moderate strengths, De Lorenzis and Nanni[21] estimated that for lightly sand-blasted round bars, akvalue of 1.5 was enough to prevent splitting failure ofthe epoxy cover. For round bars, cracking of the epoxycover (Fig. 3) produced by the radial components of thebond stresses can accelerate the occurrence of a BE-Ifailure.

The BE-C failure mode was observed for NSM stripswith a roughened surface [13,16]. This mode occurs whenthe shear strength of the epoxy is exceeded.

Inter-laminar shear failure within the bar, although the-oretically possible, has never been observed. Shearing-off ofribs in ribbed bars has never been reported as a failuremode itself, unlike in the case of FRP ribbed bars as inter-nal reinforcement for concrete[22]. However, in some tests

[21], the surface of the ribbed bars was found to have beendamaged after bond failure, indicating that this could be anupper-bound failure mode.

3.2.2. Bond failure at the epoxyconcrete interface

Bond failure at the epoxyconcrete interface mayoccur as pure interfacial failure (EC-I), or as cohesiveshear failure in the concrete (EC-C). The EC-I failure

mode was found to be critical for pre-cast grooves[23]. For spirally wound bars or ribbed bars with lowrib protrusions, this was found to be the critical fail-ure mode whenever the groove was preformed, indepen-dent of the value of k. For ribbed bars with highrib protrusions, this mode was found to be critical onlyfor k values larger than a minimum value (equal toapproximately 2.00 for ribbed bars in epoxy), and forlower k values, splitting failure of the epoxy coverdominated.

The EC-C failure mode has never been observed in bondtests, but it has been observed in bending tests on beams(Fig. 3, see also Fig. 7f) within the strengthened region

[11,20] or at the bar cut-off point [17]. The latter authorsconsidered this failure mode in their theoretical model fordebonding of NSM strips.

3.2.3. Splitting of the epoxy cover

Longitudinal cracking of the groove filler and/or frac-ture of the surrounding concrete along inclined planes isherein referred to as cover splitting. This was observed tobe the critical failure mode for deformed (i.e. ribbed andspirally wound) round bars.

The mechanics of cover splitting bond failure in anNSM system is similar to that of splitting bond failure of

steel deformed bars in concrete, on which a good under-standing has been developed from decades of research[24]. For an NSM FRP bar, the radial component of thebond stresses is balanced by circumferential tensile stressesin the epoxy cover which may lead to the formation of lon-gitudinal splitting cracks of the cover. The concrete sur-rounding the groove is also subjected to tensile stressesand may eventually fail when its tensile strength is reached,causing fracture along inclined planes. Whether fracture inthe concrete occurs before or after the appearance of split-ting cracks in the cover or even after the complete fractureof the cover, depends on the groove size and the tensilestrengths of the two materials.

The tensile strength of epoxy is one order of magnitudelarger than that of concrete, but the epoxy cover thicknessfor NSM FRP reinforcement is one order of magnitudesmaller than the thickness of concrete cover to internalsteel reinforcement in an ordinary RC member. Moreover,longitudinal steel reinforcement in RC beams benefits fromthe restraint of shear links, but this restraint is not availablefor NSM longitudinal reinforcement, unless externalrestraint of some form (e.g. FRP U jackets as shear rein-forcement) is provided. These factors explain why coversplitting is a likely bond failure mode for an NSM system.Figs. 4bc illustrate how the bond mechanism of an NSM

system can be modelled in the plane perpendicular to thebar axis, as further explained later. These figures also clar-ify the difference in bond mechanism between round bars



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and strips. In the latter case, the normal component of thebond stresses is transverse to the thick lateral sides ofthe groove [13,25] so that splitting failure is less likely tooccur.

The different patterns of cover splitting failure of NSM

systems are shown in Fig. 3. When the kratio is very low(e.g. specimens in Ref. [21] with k= 1.121.18), failure islimited to the epoxy cover and involves little damage in

the surrounding concrete (SP-E failure). For higher valuesofk, failure involves a combination of longitudinal crack-ing in the epoxy cover and fracture of the surrounding con-crete along inclined planes (SP-C-1 failure); concretefracture starts as soon as the epoxy cover cracks and the

tensile stresses are redistributed[26]. The inclined fractureplanes in the concrete have been observed to form an angle(bin Fig. 3) of approximately 30with the horizontal. For

Fig. 3. Bond failure modes of NSM systems observed in bond tests.



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large groove depths and/or when the tensile strength ratiobetween concrete and epoxy is small, fracture of concretemay occur before the epoxy crack has reached the externalsurface (SP-C-2 failure).

The bond failure modes discussed above are for anNSM bar located centrally in a wide member, whereedge effects are unimportant. When an NSM bar is closeto the edge of a concrete member, failure involves thesplitting of the edge concrete (SP-ED failure)[16]. In tests,this failure mode was found to occur when a0e < 20 mm[16], with the angle b 0 defined inFig. 3 ranging from 45to 70.

The bond strength associated with the SP-E mode isexpected to depend strongly on the tensile strength of theepoxy, whereas those associated with the SP-C-1 and SP-C-2 modes are expected to depend strongly on the concretetensile strength. In all cases, the bond strength is expectedto increase with the cover thickness of the NSM bar (i.e.the groove depth). However, the rate of strength increasewith the groove depth has been observed to reduce afterthe groove depth exceeds about 2 times the diameter,around which the SP-C-2 mode replaces the SP-C-1 modeas the critical bond failure mode. The bond strength ofthe SP-E mode also increases as the surface deformationsbecome less pronounced[26].

3.3. Effect of groove detailing on bond performance

The effect of the groove width-to-depth ratio on the

bond performance has not yet been investigated in detail.Through finite element modelling, the debonding load ofan NSM strip was found to increase with the groove width

[17]. As the groove depth was kept constant, an increasedgroove width implied a larger interfacial area betweenepoxy and concrete. This in turn implied a larger debond-ing load since debonding was assumed to occur by cohesive

shear failure in the concrete at the epoxyconcrete interface[17].

By simplified analytical modelling for deformed roundbars[25], the cracking load of the epoxy cover was foundto decrease with an increase in the groove width-to-depthratio (for a given depth) but the failure load was foundto remain substantially unchanged due to failure being inthe concrete along fracture surfaces nearly independent ofthe groove width-to-depth ratio. The first result was con-firmed by Hassan and Rizkalla[19]through finite elementmodelling. It was also found that the tensile stresses in theconcrete decrease with an increase in the groove width,which implies a larger concrete cracking load (but not nec-essarily a larger failure load). However, experimental evi-dence on the effect of the groove width-to-depth ratio onbond performance is still lacking.

From bond tests on NSM strips, Blaschko [13,16]indicated that a minimum a0e a

0e ae bg=2 of about

20 mm was required to avoid a splitting failure of the con-crete corner, and fora0evalues larger than 30 mm, no crackswere observed in the concrete at bond failure. He suggestedthat a0e be no less than 30 mm or the maximum aggregatesize, whichever is greater. The maximum aggregate sizewas suggested as a limit to avoid damaging the concreteduring the cutting of the groove. In his bond tests, a0e did

still influence the bond behaviour until the maximum inves-tigated value of 150 mm, beyond which no further influ-ence was assumed.

Fig. 4. Schematics of the bond behaviour of NSM FRP reinforcement: (a) bond stresses in the longitudinal plane; (b) normal stresses in the transverseplane generated by a round bar; (c) normal stresses in the transverse plane generated by a rectangular bar.



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Based on finite element modelling for round deformedbars, Hassan and Rizkalla [19] suggested a minimum agand a minimum ae of two and four bar diameters respec-tively, regardless of the groove width. However, one of the

beams tested by De Lorenzis[11], which was strengthenedwith NSM spirally wound round bars with ag= 30 mm(i.e. about 1.8 times the groove size and 3.6 times the bardiameter) andae= 69 mm (i.e. about 4.3 times the groovesize and 8.6 times the bar diameter), failed by debondingof the NSM bars involving the spalling of the concrete coverof the longitudinal steel reinforcement along the edges. Thistest thus suggested that the minimum values for agandaeasspecified in Ref. [19] are insufficient to eliminate interactionsbetween an NSM bar and the edge of a beam.

3.4. Local bond strength

3.4.1. Experimental resultsIn any type of bond test, the average bond strength usu-

ally decreases with increases in the bond length, as a resultof the non-uniform distribution of bond stresses (Fig. 4a).The local bond strength refers to the maximum value ofbond stress that the interface can resist, in contrast to theoverall bond strength (or simply bond strength as used inthis paper) which refers to the maximum transferable loadof the joint. The local bond strength must then be obtainedeither from very short specimens or from a long specimenby elaborative strain (and/or slip) measurements.

Several authors studied the local bond strengths of

NSM systems[13,27,16,20,21,25]. A summary of their val-ues, reported inTable 2, allows the following observationsto be made:

the local bond strength of the cover splitting mode(deformed bars), as expected, is higher if the groovedepth is larger and the bar surface deformation lesspronounced;

the local bond strength of the barepoxy interfacial (BE-I) failure mode, which was observed for sand-blastedbars, is not influenced by the groove size and is lowerthan that for deformed bars;

the local bond strengths of NSM strips from two testseries by different authors[14,16]are very close to eachother and are comparable to that of spirally wound bars[26]; by contrast, the local bond strength obtained froma third test series[20]is notably lower.

Available information on the bond behaviour of squarebars is still very limited. Nordin and Taljsten [9] reportedaverage bond strengths from specimens whose lengths wereat least 10 times the width of the square cross-section butno local bond strengths were reported. An unusual aspectof the results reported by these authors is that the averagebond strength increased with the bond length.

De Lorenzis[25]reported the local bond strengths at theepoxyconcrete interface for specimens with different typesof bars tested by De Lorenzis et al. [23]. As EC-I failure

was the critical mode, the surface configuration of thebar did not have a significant effect on the bond behaviourand the difference in local bond strength between ribbedand spirally wound bars was basically due to the different

diameters, and hence to the different groove size corre-sponding to a given kvalue. The local bond strength wasfound to decrease with an increasing groove size.

3.4.2. Theoretical models for NSM strips

It is interesting to compare the experimental local bondstrengths of NSM strips reported by Sena Cruz and Barros[27] with the predictions by the formula proposed byBlaschko[13], and with those given by Hassan and Rizka-llas theoretical model[17]. Blaschkos formula[13]is givenby:

smax 0:2 ffiffiffiffiffia0e4

p saf a0e 6 150 mm 1

where saf is the shear strength of the epoxy. Hassan andRizkallas formula[17]is given by:

smax f0cfct

f0c fct2

where f0c and fct are the (cylinder) compressive and tensilestrengths of concrete, respectively. The two formulae relatethe local bond strength to different parameters, consistentwith their own experimental observations: Blaschko ob-served cohesive shear failure in the epoxy and studied theeffect of a0e, whereas Hassan and Rizkalla observed cohe-sive shear failure in the concrete (hence, their value ofsmax

is the shear strength of concrete). The following differencesbetween the two formulae should also be noted: (a) Blas-chko performed pull-out bond tests to provide the experi-mental basis, while Hassan and Rizkalla conductedflexural tests on RC beams embedded with bars of varyinglengths; (b) Blaschkos formula was calibrated with bondtest results, while Hassan and Rizkallas formula was de-rived from Mohrs circle for the pure shear stress state,which, when used in finite element modelling, yielded pre-dictions of the debonding load in good agreement with testresults.

The 95 percentile characteristic value of saf was indi-cated by Blaschko [16] to vary between 20 and 25 MPafor common highly filled, two-component epoxies. Accord-ing to the tests by Blaschko, the ratio between the charac-teristic and the average values ofsafis about 0.89, hence theaverage value ofsafof common epoxies can be assumed tovary between 22.5 and 28.1 MPa. For a0e 150 mm (i.e.with no edge effect), Eq. (1) thus yields a local bondstrength ranging between 15.8 and 19.8 MPa, whose upperbound practically coincides with the values inTable 2fromtests in[14,16]. This is as expected as Eq. (1)was calibratedusing test results with an average value of 28.7 MPa for saf,very close to the upper bound of the average value rangementioned above.

For f0c ranging between 20 and 40 MPa and taking fctas 0:53

ffiffiffiffif0c

p [28], Eq. (2) predicts local bond strengths

between 2.1 and 3.1 MPa. The large difference between



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the predictions of Eqs.(1) and (2)is a result of the differentmaterials controlling the failure (epoxy for Eq.(1)and con-crete for Eq.(2)) and thus the different interfaces that thesetwo formulae correspond to; the concrete shear strength is

much smaller than that of the adhesive.

3.4.3. Theoretical models for NSM round deformed bars

The bond behaviour of the NSM round deformed barsis controlled by splitting tensile stresses in the epoxy coverand the surrounding concrete.Figs. 4a and b illustrate theapproach adopted in[11,17]to model this bond behaviourin the plane perpendicular to the bar axis. Two simplifyingassumptions are common to these two studies. First, thefrictional coefficient l (=1/tanc) relating bond shear stres-ses and internal splitting pressures is constant, although itis known to change during the loading process. Second,the distribution of the radial pressure is uniform, although

the pressures on the thicker concrete substrate are higherthan those on the thinner cover.

De Lorenzis [11] took c to be 45 (i.e. l= 1), whereasHassan and Rizkalla measured it according to ASTMG115-98[29]. However, it should be noted that the conceptof frictional coefficient used in this context is very differentfrom that defined by the ASTM specification. The latterpertains to pure frictional properties between materials,depending on material and surface characteristics, whereasthe frictional coefficient in the bond context is also influ-enced by structural variables such as the cover depthand the bar diameter in the case of internal reinforcing bars

(e.g. [30]). In the case of NSM reinforcement, even morevariables are involved. For steel bars in concrete, Tepfersand Olsson [31] performed ring pull-out tests in orderto estimate the angle c of bond stresses at different stagesof loading. They established curves relating the coefficientof friction to the slip for different types of bars. Similarmeasurements for NSM bars would require accurate mon-itoring of strains in epoxy and concrete transverse to thebar axis.

For round deformed bars, Hassan and Rizkalla[19]pro-posed a model for cover splitting bond failure, based onelastic finite element analysis. They provided two formulaeto predict the local bond strengths for the barepoxy andthe epoxyconcrete interfaces respectively, and the one cor-responding to the smaller local bond strength wouldcontrol.

Depending on how they are computed, these local bondstrengths correspond to either the first cracking of theepoxy cover or the first cracking of the concrete adjacentto the groove. The two formulae are:

smax barepoxyfatl

G23

smax epoxyconcretefctl

G14

wherefatis the tensile strength of epoxy, and G1and G2arecoefficients which were evaluated by finite element analysisand are dependent on the groove depth-to-bar diameter

and the groove width-to-bar diameter ratios. Graphs forG1 and G2 were developed. The range of the first ratioexamined by them is between 2.00 and 2.50, while that ofthe second ratio is between 1.50 and 2.50. G1and G2range

between 0.58 and 1.3 and between 0.5 and 0.72, respec-tively. Hence, Eq. (4) controls in all cases with values of0.771.72 times lfct. For a concrete cylinder compressivestrength of 2040 MPa, the local bond strength varies be-tween 1.4 and 5.2l, which are very low compared withthe test results in Table 2, already with l= 1 instead of0.5 as originally proposed in Ref.[19].

It should be noted that the so-called local bond strengthin this model is the bond shear stress corresponding to theinitiation of cracking in the epoxy or in the concrete,whereas the joint can still sustain significant incrementsof the applied load between first cracking and ultimate fail-ure. Also, the distinction between cracking of epoxy and

cracking of concrete as two independent modes of bondfailure contradicts the experimental evidence that failurein the concrete usually follows cracking of the epoxy.

An approximate two-dimensional elastic stress analysiswas conducted by De Lorenzis[25]to determine the bondshear stress corresponding to the cracking of the epoxycover. For computation of an upper and a lower boundto the ultimate radial pressure of the NSM system (andhence to the local bond strength if the frictional coefficientis known), different possible failure modes were analyzed.A uniform tensile stress distribution along the fracture lineswas assumed and justified on the basis of redistribution of

cohesive stresses between the crack faces. The experimentalvalues of local bond strength were shown to fall within thecomputed range. However, no equation for the local bondstrength was proposed.

3.4.4. Theoretical models for NSM bars in pre-formed

grooves

The local bond strength of the epoxy-to-concrete inter-face for pre-formed grooves was shown to decrease almostlinearly with increases in groove size. An equation is givenin Ref. [25] that may be used to compute the local bondstrength for different groove sizes for this case.

3.4.5. Comparison with internal FRP rebars and externally

bonded FRP laminates

Cosenza et al.[32]presented a summary of bond prop-erties of FRP bars used as internal reinforcement, obtain-ing an average local bond strength of 2.74 MPa forsand-blasted bars (with a COV of 52%) and of11.61 MPa for ribbed bars (with a COV of 34%). Thelocal bond strength value for sand-blasted bars is the aver-age of results from Ref.[33], conducted on only one type ofbars different from those used by De Lorenzis and Nanni[21], hence a direct comparison is not meaningful. The cat-egory of ribbed bars referred to by Ref.[32]encompasses

a wide variety of material and surface configurations as itcovers what are referred to as ribbed bars and spirallywound bars in the present paper, plus some cross-wound



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bars. The local bond strength values for spirally woundbars from Ref.[32]vary between 4.76 and 18.04 MPa, withan average of 11.9 MPa, which is lower than what has beenobtained for NSM systems (Table 2). In bond tests of spi-

rally wound bars as internal reinforcement, significantinterlocking was not evident, but when used as NSM rein-forcement, these bars normally produce cover splittingbond failure before the loss of bond at the barepoxy inter-face. The higher local bond strength for an NSM systemcan be attributed to a larger resistance at the barepoxyinterface than at the barconcrete interface.

Nanni et al.[34]conducted bond tests on GFRP ribbedbars of the same type as used by De Lorenzis et al. [26](although bars in this study had a larger diameter, equalto 12.7 mm), and found a local bond strength of 17 MPafor them as internal reinforcement, which is larger thanthe value obtained for an NSM configuration. The bond

failure of these rebars as internal reinforcement occurredby the shearing-off of the ribs, a mode which wasapproached but never attained in NSM bond tests as coversplitting bond failure always occurs first.

For FRP laminates externally bonded to concrete, Luet al. [35] proposed a simple equation for the local bondstrength. Assuming an FRP-to-concrete width ratio of0.5, this equation yields a local bond strength of 1.5 timesthe tensile strength of concrete, which is significantly lowerthan the cover splitting local bond strength of NSM bars.This results from the fact that, in the cover splitting bond

failure mode of an NSM system, the fracture of concreterelates to the larger perimeter (see Fig. 3, mechanismsSP-C1 and SP-C2), while the nominal bond strength isdefined using the smaller bar perimeter. In the debonding

failure of an externally bonded laminate, the fracture planehas approximately the width of the laminate and the nom-inal bond strength is based on the same width.

3.5. Local bondslip behaviour

Local bondslip curves were deduced from test data byDe Lorenzis [25] for different types of NSM round bars,and by Sena Cruz and Barros [27] for NSM strips. DeLorenzis[25]reported three different types of local bondslip behaviour (types IIII), of which the first two areshown in Fig. 5. The third type (for sand-blasted roundbars) differs from the second in that the abrupt decay from

the maximum bond stress to the frictional bond stress levelis replaced by a linearly decreasing branch. This third typecould be seen as a special case of the second type. Theequation proposed by Sena Cruz and Barros [27] has thesame form as that of the type I curve shown in Fig. 5.

The type I equation seems to reproduce rather accu-rately experimental curves showing a gradual decrease oflocal bond stress after the peak. Such a gradual decreaseexists when bond failure is at an interface (the epoxycon-crete or the barepoxy interfaces) or by cover splitting gen-erated by ribbed bars with low rib protrusions. In both

b

Fig. 5. Typical bondslip curves of NSM FRP reinforcement.



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cases, a significant amount of post-peak friction develops,due to interfacial friction in the first case and to aggregateinterlocking of the cracked concrete in the second case.

Conversely, cover splitting failure generated by ribbed

bars with high rib protrusions and spirally wound barshas a more brittle nature (for a detailed description seeRef. [26]) with an abrupt decrease in bond stress uponthe attainment of the peak value. However, even after thecomplete loss of the cover, a small amount of residual fric-tion remains because half of the perimeter of the bar is stillin contact with the epoxy.

The type I equation, where the bond stress tends to zeroas the slip approaches infinity, cannot reproduce the resid-ual frictional branch. This is only a minor drawback ifa 0 islarger than 1, as in such cases the bond stress from thetype I equation reaches the frictional plateau only at verylarge values of slip and the area underneath the bondslip

curve is infinite. However, ifa 0 is smaller than1, this areabecomes finite, and the joint may be predicted to be unableto develop the full tensile capacity of the bar (see next sub-section), which contradicts the better behaviour of this typeof joint compared with type II joints. More work on localbondslip curves of NSM systems, both experimental andtheoretical, is still needed, in particular on the post-peakbehaviour which greatly influences the performance of barswith long embedment lengths such as those used inpractice.

De Lorenzis et al. [26]commented on the variations ofthe secant stiffness and the shape of the curve with test vari-

ables. In nearly all cases, the slip at peak stress was foundto be in the range of 0.10.3 mm. Sena Cruz and Barros[27]found a value of 0.25 mm for NSM strips.

Based on strain gage and free-end slip readings, Blas-chko[16] obtained local bondslip curves of NSM strips,very similar in shape to the type I curves discussed above.However, by plotting the local bondslip curves at differentmeasurement locations along the bond length, he notedthat the local bond strength tended to be larger close tothe free end and smaller close to the loaded end. Thiswas attributed to the influence of the transverse displace-ments of the concrete adjacent to the groove, which werealso measured in the same tests. In his analytical model,the author adopted a local bondslip curve consisting ofan ascending branch defined by a second-degree parabolaand a horizontal branch at the local bond strength. Thiscurve was assumed to represent the pure shear stressstrainbehaviour of the epoxy. The local bond strength was takenas the shear strength of the epoxy, multiplied by an empir-ical function of the transverse displacement of concrete.The transverse displacements were predicted with a sepa-rate elastic model. Mohrs failure criterion was used tomodel the effect of local normal stresses, associated withtransverse concrete displacements, on the local bondstrength. Regression analysis was used to maximize the

agreement between measured and predicted strain distribu-tions in the strip along the bond length. In summary, hislocal bondslip model considers the response of the NSM

bar and the concrete member as a system so that the influ-ence of edge distance can be effectively reflected. However,the model is rather complex to apply as it requires an iter-ative procedure. Also, it cannot accurately represent the

presence of a frictional asymptote in the local bondslipcurves obtained from tests and hence it overestimates thebond failure loads of specimens with long bond lengths.For this reason, the author relied on the direct calibrationof test results, rather than the output of the bondslipmodel, to obtain Eq.(1).

3.6. Effective bond length, development length

and anchorage length

The maximum stress that can be resisted by a bondedjoint between an NSM bar and the concrete substrate witha sufficiently long bond length is given by:

rmax

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi2E

R

AGf

r 5

with

Gf

Z 10

ssds 6

whereGf, being the area underneath the bondslip curve, isthe fracture energy of the bonded joint, R is the perimeterover which the bond stress acts, A is the cross-sectionalarea over which the tensile stress acts, and E is the elastic

modulus of the material on which the tensile stress is ap-plied[23]. For round bars bonded to epoxy (or to concretein the case of internal rebars), Eq. (5)reduces to

rmax

ffiffiffiffiffiffiffiffiffiffiffiffiffiffi8Ebdb

Gf

s 7

where Eb is the elastic modulus of the bar, and for lami-nates bonded to concrete (neglecting the thickness of theadhesive layer), Eq.(5) reduces to

rmax ffiffiffiffiffiffiffiffiffiffiffiffi2EfGft

fs 8

where Ef and tf are the elastic modulus and thicknessrespectively of the laminate. Ifrmax computed by Eq. (5)is below the tensile strength of the reinforcement, its fullcapacity cannot be developed no matter how long the bondlength is. In this case, a value of bond length exists (theeffective bond length) at whichrmaxis developed, and be-yond which a further increase in bond length does not pro-duce any benefit. The concept of effective bond length hasbeen well established for externally bonded FRP laminates[36,37]. Ifrmaxis larger than the tensile strength of the bar(and in particular for bondslip curves with infinite values

ofGf), the full capacity of the reinforcement can be devel-oped, and the corresponding value of bond length is usu-ally termed the development length.



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For type I bondslip curves, Gfhas an infinite or a finitevalue when a 0P 1 or a 0


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large, especially if long bond lengths are tested; it is diffi-cult to conduct the test in slip-control mode; and it isdifficult to visually inspect the behaviour of the joint duringloading, especially the initiation and propagation of

cracks.Direct pull-out tests overcome the drawbacks of beampull-out tests mentioned above. The simplest direct pull-out test specimen may be composed of a square/rectangu-lar concrete block embedded with an NSM bar on one ofthe sides, but in this set-up, the NSM bar leads to eccentricloading of the concrete block. The use of two bars on twoopposite sides[38]or even four bars on all four sides[3]hasbeen attempted to overcome this problem. The multiple-bar specimen has its own problem: any small deviationsof the groove/bar positions can induce flexural effects,significantly altering test results.

De Lorenzis et al. [23] introduced a C-shaped

block where a single NSM bar was placed at the centreof gravity of the block. The set-up performed well, butthe specimen dimensions had to be specifically designedfor each groove depth. This set-up is also not suitable forstudying edge effects due to the presence of two thickflanges.

Based on the above discussions, a direct shear test on asingle NSM FRP bar embedded in a concrete block, wherethe tensile force applied on the bar is reacted by compres-sive stresses on the concrete block at the loaded end, isprobably a good choice that combines simplicity with reli-ability. A similar test set-up has been popular in studies on

externally bonded laminates (e.g.[39]). Blaschko[16]usedsuch a set-up, in which a steel plate was used to providethe reaction to the concrete block. The steel plate had acentral hole of 80-mm diameter to avoid reactive stresseson the immediate vicinity of the groove.

To minimize the transverse friction generated by thebearing pressure, which could delay the initiation of split-ting cracks as generally observed in pull-out tests of steelrebars in concrete, layers of PTFE or similar materialscan be placed between the bearing plate and the concreteblock.

Three main methods are available for obtaining thelocal bondslip curve of a bonded joint:

Approximate it with the curve relating the average bondstress to the loaded-end slip (or the free-end slip, or theaverage of the two) from specimens with a short bondlength (SBL).

Obtain bond stresses and slips from free-end slip andstrain measurements at discrete points along the bondlength. This method is usually adopted when long bondlength (LBL) specimens are used, as strain gages on aSBL specimen are likely to significantly affect the bondperformance.

Calibrate the unknown parameters in the local bond

slip equation, whose form needs to be known orassumed in advance, from loaded-end slip and free-endslip measurements.

Each of the above methods has its advantages and dis-advantages. The first method does not require the use ofstrain gages, which simplifies specimen preparation anddoes not alter the bond properties between the bar and

the epoxy. However, the bond length still needs to be longenough for the specimen to behave as a representative sam-ple (e.g. for ribbed or spirally wound bars, a minimumnumber of deformations should be included) and to reducethe influence of end effects. The bondslip curve obtainedrepresents the average performance of the chosen bondlength of the specimen.

The second method, more onerous and altering tosome extent the barepoxy interface, has the advantagethat the bond performance over a longer and hence morerepresentative portion of the reinforcement can bestudied. A local bondslip curve can be retrieved at eachload level, or alternatively at each measurement point, so

that more data are obtained from such a test than fromone on a SBL specimen. It is worth noting that evenfor tests conducted with load control on LBLspecimens, strain measurements allow the descendingbranch of the local bondslip relationship to be obtainedas a result of stress redistribution over the bondlength, although these measurements tend to show a nota-ble scatter, due to damage to the strain gages resultingfrom slips of the reinforcement. Teng et al. [20] presentedtests where strain gauges were sandwiched between twoCFRP strips so that the strain gauges were well protectedand did not affect the interfacial properties. Wang et al.

[40] recently explored the use of fiber optic sensorsembedded inside FRP bars for the measurement ofstrains so that the use of strain gauges on the bar sur-face can be avoided. These and similar techniquesfor strain measurement are highly desirable for bondtests on LBL specimens to minimize damage to thestrain sensors and interference with the interfacialbehaviour.

The third method has the advantage that it allows thelocal bondslip curve to be obtained from LBL specimenswithout the need for strain measurements. The disadvan-tages are that the form of the bondslip equation mustbe known a priori, and the accuracy of the deduced equa-tion may be compromised if the assumed form is inappro-priate in some way. If many different forms of equationsare tested to find the most suitable form, this approachinvolves a much more onerous process than the othertwo approaches.

In all cases, the specimen must be carefully designedto ensure that the failure mode and load are not signifi-cantly influenced by the specimen dimensions. An exampleof inadequate specimen size can be found in Ref. [3],which reports shear fracture failure of the concrete involv-ing the entire specimen cross-section due to its limitedsection size. When the NSM bar is close to the edge of

the concrete block, test results have indicated a stronginfluence of the transverse stiffness of the concrete on themeasured bondslip curve [13], which then also varies



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along the bond length as a result of the variation of thebond stresses. This effect should either be a parameter tobe considered or be eliminated by the use of a sufficientlywide specimen.

In addition to section size, the bonded portion of the barshould start at a suitable distance from the loaded end ofthe concrete block. If this distance is insufficient, the behav-iour of a specimen with a short bond length may be similarto that of a fastener (see e.g. [38]), or the failure mode ofthe specimen is unduly influenced by the cracking of con-crete at the loaded end [16].

4. Flexural strengthening


A summary of the existing experimental data is reported

inTables 3 and 4.Some existing studies were conducted on beams

strengthened with NSM bars of limited embedment lengths(Table 3). Although such tests were intended to study bondfailure mechanisms, they are not pure bond tests as thebond performance is affected by flexural cracking. More-over, the NSM FRP bars in such tests generally extend intothe shear spans, where part of the interfacial shear stress isdirectly dependent on the transverse shear force in thebeam.

Hassan and Rizkalla[17,19]conducted flexural tests onRC beams with NSM CFRP round ribbed bars and strips

of varying embedment length. Failure of beams withNSM round ribbed bars occurred by splitting of the con-crete cover followed by the complete debonding of the barsin all cases. These authors concluded that the tensile ruptureof this type of bars is unlikely to occur, regardless of theembedment length, that the maximum usable strain of thesebars should be limited to 0.70.8%, and that the anchoragelength should not be shorter than 800 mm. In the case ofbeams with NSM strips, rupture of the strips occurred whenthe embedment length was larger than 850 mm.

Teng et al. [20] conducted flexural tests on RC beamswith NSM strips of varying embedment length. As theembedment length increased, the failure mode changedfrom concrete cover separation starting from the cut-offsection, to concrete crushing followed by secondary coverseparation close to the maximum moment region. In thebeams with the two longest embedment lengths, secondarydebonding mechanisms were also observed.

All existing test results of strengthened beams, slabs, andcolumns (Tables 3 and 4) indicate that the NSM reinforce-ment improved the ultimate load and the load at the yield-ing of steel reinforcement, as well as the post-crackingstiffness. Some test programs included identical beamsstrengthened with equivalent amounts of FRP providedas either externally bonded or NSM reinforcement. In all

cases, the NSM reinforcement performed more efficiently,as debonding of the NSM reinforcement occurred at ahigher strain or did not occur[17,4143].

One study [41] has compared equivalent amounts ofNSM reinforcement provided as round bars or strips. Asexpected, strips performed better and failed by tensile rup-ture as opposed to debonding of the round bars, as a result

of the higher local bond strength and larger lateral surfaceto cross-sectional area ratio of NSM strips.

4.2. Failure modes of flexurally-strengthened beams

The possible failure modes of beams flexurally-strength-ened with NSM FRP reinforcement are of two types: thoseof conventional RC beams, including concrete crushing orFRP rupture generally after the yielding of internal steelbars, for which the composite action between the originalbeam and the NSM FRP is practically maintained up tofailure, and premature debonding failure modes whichinvolve the loss of this composite action. Although deb-

onding failures are less likely a problem with NSM FRPcompared with externally bonded FRP, they may still sig-nificantly limit the efficiency of this technology.

The likeliness of a debonding failure depends on severalparameters, among which the internal steel reinforcementratio, the FRP reinforcement ratio, the cross-sectionalshape and the surface configuration of the NSM reinforce-ment, and the tensile strengths of both the epoxy and theconcrete. Some researchers [11,10] extended the NSMFRP reinforcement over the beam supports to simulateanchorage in adjacent members. Despite this anchorage,debonding failures can still occur [11]. The beam reported

in Ref.[10]failed by FRP rupture, as opposed to debond-ing observed in an identical beam with the NSM reinforce-ment terminated away from the supports. Blaschko [16]reported the results of two beam tests: the first one failedby concrete cover separation starting from the cut-off sec-tion but the second beam, which was provided with a steelU-jacket bonded to the cut-off section, failed by the ruptureof the FRP strips. In the same study it was observed thatfatigue loading to two million load cycles did not affectthe residual beam capacity.

There is still limited understanding of the mechanics ofdebonding in beams strengthened with NSM systems.Descriptions of failure modes in the existing literature areoften not sufficiently detailed to understand the progres-sion of the failure process. Based on the available experi-mental evidence, the possible failure modes of beamsflexurally-strengthened with NSM FRP reinforcement areclassified in Fig. 7 and described below. The interactionsbetween the main failure modes described below and thesecondary failure modes are still unclear and deserve fur-ther investigations.

4.2.1. Barepoxy interfacial debonding

This mode involves interfacial debonding between a barand the epoxy and has been observed for sand-blasted round

bars[44]. This mode correlates well with the failure modeobserved in bond tests on the same type of bars (see theprevious section). However, unlike in a bond specimen, the



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epoxy cover in the beam was intersected by flexuralcracks which facilitated the initiation of longitudinal split-

ting cracks and hence accelerated interfacial debonding(Fig. 7a).

Fig. 7. Debonding failure modes of NSM bars and strips observed in tests on flexurally-strengthened beams: (a) debonding at the barepoxy interface; (b)separation of concrete cover between two cracks in the maximum moment region; (c) separation of concrete cover over a large length of the beam; (d)separation of concrete cover starting from a cutoff section; (e) separation of concrete cover along the edge; (f) secondary loss of bond between epoxy andconcrete; (g) secondary splitting of the epoxy cover.



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4.2.2. Concrete cover separation

The formation of bond cracks on the soffit of thebeam has been observed in tests[11,16,20], and these bondcracks are inclined at approximately 45 [20] to the beam

axis. Upon reaching the edges of the beam soffit, thesecracks may propagate upwards on the beam sides main-taining a 45 inclination within the cover thickness, andthen propagate horizontally at the level of the steel tensionbars. Debonding may next occur in different forms,depending on the subsequent evolution of the crackpattern:

(a) Bar end cover separation.If the NSM FRP reinforce-ment is terminated at a significant distance from thesupports, separation of concrete cover typically startsfrom the cut-off section and propagates inwards[16,20] (Fig. 7d). This mode is similar to the cover

separation failure mode observed in RC beams withan externally bonded FRP laminate[37,45,46].

(b) Localized cover separation. Bond cracks within orclose to the maximum moment region, together withthe pre-existing flexural and flexural-shear cracks,may isolate triangular or trapezoidal concretewedges, of which one or more are eventually splitoff (Fig. 7b). This mode can be identified from photosof failed beams given in Refs. [11,15,20].

(c) Flexural crack-induced cover separation. Separationof the concrete cover occurs almost simultaneouslyover a long portion of the NSM reinforcement, often

involving one of the shear spans and the maximummoment region (Fig. 7c) [15,44]. This mode wasobserved by De Lorenzis et al.[44]to start from themaximum moment region, whereas the location ofinitiation was not made clear in Ref.[15]. This modeis similar to the intermediate crack-induced debond-ing failure mode observed in RC beams with an exter-nally bonded FRP laminate[37,47,48].

(d) Beam edge cover separation. NSM bars located nearthe edges may generate detachment of the concretecover along the edges (Fig. 7e).

4.2.3. Epoxyconcrete interfacial debonding

For beams with NSM strips of a limited embedmentlength, Hassan and Rizkalla [17] reported cohesive shearfailure in the concrete at the epoxyconcrete interface start-ing from the cut-off section. Unfortunately no picture ofthe failed specimens was provided. This mode is believedto be similar to the plate end interfacial debonding failureof RC beams with externally bonded FRP laminatedescribed in Refs. [37,45,46].

4.2.4. Secondary debonding failure mechanisms

Other debonding mechanisms have also been observed.They are herein classified as secondary failure modes

and the role they play in the context of debonding failuresis still unclear. It has been observed[20]that upon the for-mation of the bond cracks, the opening-up of these inclined

bond cracks was restrained by the dowel action of NSMreinforcement which in turn tended to cause the detach-ment of the NSM FRP reinforcement from the soffit ofthe beam. After failure, the prism formed by the CFRP

strip and surrounding epoxy was found to retain a thinconcrete layer of variable thickness on the sides (Fig. 7f),indicating that a strong epoxyconcrete bond existed.Moreover, localized splitting occurred in the epoxy cover,exposing the internal CFRP strip (Fig. 7g). Similar obser-vations had been reported previously[11].

4.3. Prediction of ultimate loads and loaddeflection

behaviour

For the safe design of an NSM FRP system for the flex-ural strengthening of an RC beam, the foremost issue is theprediction of the ultimate load. If the failure of a strength-

ened beam does not involve debonding, then the failureload can be easily predicted using equations developedfor externally bonded FRP based on the plane sectionassumption (e.g. [37]) and with the difference in positionbetween the two types of reinforcement duly taken intoaccount. Accurate predictions of debonding failure loadsare much more challenging.

In the past few years, numerous studies have examineddebonding failures of beams with externally bonded steelplates and FRP laminates [37,4549]. Research on RCbeams flexurally-strengthened with NSM FRP has beenmuch more limited. The few theoretical models developed

so far are extensions of approaches developed for exter-nally bonded FRP laminates. For instance, Hassan andRizkalla[17]proposed a theoretical model for NSM strips(valid only for concrete shear failure at the epoxyconcreteinterface at the cut-off section) which is an extension of theinterfacial stress-based approach proposed by Malek et al.[45]. The model has been compared with a very limiteddatabase. Moreover, the failure mode assumed in themodel of Hassan and Rizkalla [17] was only observed inthe tests by these authors but has not been observed inother tests.

Of the reported failure modes, the most critical appearsto be the mode of concrete cover separation starting fromthe maximum moment region. Provided that bars with areasonable degree of surface deformation are used, failureat the barepoxy interface is unlikely. Moreover, in caseswhere the NSM reinforcement is needed over the entirelength of a member, NSM bars can be easily anchoredto adjacent members so that debonding failure at a cut-off section can be prevented. For similar reasons, theintermediate crack-induced debonding failure mode hasbeen recognised as the most important mode for RCbeams strengthened with externally bonded FRP lami-nates [47].

For design purposes, the simplest approach to the pre-

diction of debonding is to establish a bond reduction factorfor the ultimate tensile strain of the reinforcement. Such anapproach is currently adopted by ACI-440[6] for debond-



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ing of externally bonded laminates. The predictive modelsin Refs. [47,48] also follow this approach. Alternativeapproaches [5,50] for this type of debonding failure ofexternally bonded FRP laminates consider stress gradients

in the FRP between two adjacent cracks instead of a simplebond reduction factor[51].Development of reliable predictive models for debond-

ing failures requires a thorough understanding of themechanics of debonding failures, and of the qualitativeand quantitative roles of relevant variables. The mostchallenging aspect in tackling this problem appears tobe the lack of a direct correlation between the bond fail-ure modes in bond specimens and the debonding failuremodes in flexurally-strengthened beams. The possiblereasons are the presence of flexural and flexural-shearcracks which alter the bond stress distribution, the curva-ture of the beam, the dowel action of the FRP bars

restraining the opening-up of inclined bond cracks, phe-nomena which are all absent in a bond specimen. Thesame problem has been encountered in predicting inter-mediate crack debonding failures in RC beams withexternally bonded FRP laminates, although to a lesserextent [48].

The loaddeflection behaviour of beams strengthenedwith NSM reinforcement can be predicted with reason-able accuracy by the conventional sectional approachneglecting tension stiffening and assuming a perfect bondfor both steel and NSM reinforcement [11,16]. Morerefined approaches where tension stiffening is taken into

account (with different laws for un-strengthened andstrengthened beams) [15] and where slips of steel andFRP reinforcement are modelled using experimentallydetermined bondslip equations [52] have delivered moreaccurate predictions of experimental loaddeflectioncurves.

4.4. Strengthening with pre-stressed NSM FRP

NSM FRP bars or strips can be more easily pre-stressedand anchored than externally bonded laminates, so flexuralstrengthening with pre-stressed NSM FRP is a promising

technique. Nordin and Taljsten[53]have explored the useof this technique by tensioning NSM bars to 20% of theirtensile strength, filling the grooves with epoxy, and releas-ing the pre-stressing force upon hardening of the epoxy.The expected gains in the cracking load and the stiffnessof the beam as a result of pre-stressing were achieved,and the failure mode was the tensile rupture of the FRPbars in all cases. The method used by these authors cannotyet be implemented in a real strengthening project as theirprocedure of tensioning and anchoring the bars requiresaccess to the ends of the beam, which is generally not pos-sible in reality. For this reason, a tensioninganchoringdevice for NSM bars was proposed by De Lorenzis et al.[54].

5. Shear strengthening


The use of NSM FRP reinforcement is also effective inenhancing the shear capacity of RC beams. For this pur-pose, the bars are embedded in grooves cut on the sidesof the member at a desired angle to the beam axis.

Only three studies appear to have been published on theuse of NSM FRP bars for shear strengthening of RCbeams. De Lorenzis and Nanni[55]carried out eight testson large size T-beams, of which six had no internal stirrups.CFRP ribbed round bars in epoxy-filled grooves were usedas NSM shear reinforcement. The test variables includedbar spacing, inclination angle and anchorage of the barsin the flange. The NSM reinforcement produced a shearstrength increase which is as high as 106% in the absence

of steel stirrups, and still significant in presence of a limitedamount of internal shear reinforcement.

Barros and Dias[56]tested beams of different sizes withno internal stirrups. Some of these beams were strength-ened with NSM CFRP strips of different inclinations, whilethe rest were strengthened with equivalent amounts ofexternally bonded FRP shear reinforcement. The reportedstrength increases ranged from 22% to 77%, and were in allcases larger than those obtained with externally bondedFRP. Although failure modes were not described, basedon the reported loaddeflection curves, at least some ofthe beams are believed to have failed in bending.

Nanni et al.[57]reported the test results of a single full-scale PC girder taken from a bridge and shear-strengthenedwith NSM CFRP strips. The beam failed in flexure at ashear force close to the shear resistance predicted by themodel given in Ref. [55].

5.2. Failure modes

Two different failure modes were identified by DeLorenzis and Nanni [55]. The first was debonding of theFRP bars by splitting of the epoxy cover and cracking ofthe surrounding concrete, associated with the diagonal ten-

sion failure of concrete (Fig. 8a). This failure mode may beprevented by providing better anchorage of the NSM barscrossing the critical shear crack, by either anchoring thebars in the beam flange or the use of inclined (e.g. 45) barsat a sufficiently close spacing to achieve a longer total bondlength. Once this mechanism was prevented, separation ofthe concrete cover of the steel longitudinal reinforcementbecame the controlling failure mode in the tests presentedin Ref. [55] (Fig. 8b). Unlike internal steel stirrups, NSMshear reinforcement does not exert a restraining action onthe longitudinal reinforcement subjected to dowel forces.These forces, in conjunction with the normal pressures gen-

erated by the bond action of the steel longitudinal rein-forcement, create considerable tensile stresses in the cover



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which may eventually lead to cover separation failure. Thissecond mode, however, may be attributed to the fact thatno or very limited steel stirrups were present in thesebeams, and is unlikely in beams with a significant amountof steel stirrups. The most important failure mode is thusdebonding of the FRP bars. Although it has not beenobserved so far, tensile rupture of the NSM reinforcementis another possible failure mode.

5.3. Prediction of ultimate loads

The truss analogy was used by De Lorenzis and Nanni[55] to compute the shear capacity of a member strength-ened in shear with NSM FRP reinforcement, and in partic-ular the load at diagonal tension failure of concreteinvolving debonding of the NSM bars. The basic assump-tion of their approach is that, at the instant of failure,the bond stresses are evenly distributed along the barscrossed by the critical shear crack, and are equal to thelocal bond strength. This assumption is acceptable if thebondslip behaviour is ductile enough to allow a substan-tial redistribution of bond stresses along the bars crossedby the shear crack. This basic assumption yields easily

the tensile stresses in the bars crossed by the shear crackand hence the corresponding shear force. This approachwas shown to compare favourably with test results[55,57]. Further research is obviously needed for the assess-ment and improvement of the model.

Unlike the case of flexurally-strengthened beams, exist-ing evidence indicates that the debonding failure mode ofNSM bars in shear-strengthened beams is similar to thebond failure mode of the same bars in bond specimens.Further work is needed to confirm this observation as itis important for the modelling of debonding failures ofshear-strengthened RC beams. If confirmed, this observa-

tion means that local bondslip relationships developedfrom bond tests can be directly used in predicting debond-

ing failures of RC beams shear-strengthened with NSMFRP bars. With such an approach, the truss model [55]can be easily generalized to the case of debonding failureby incorporating an appropriate bondslip curve insteadof the simple ideally plastic curve assumed in the originalmodel. Similarly, local bondslip relationships obtainedfrom bond tests can also be directly used in the numericalmodelling of debonding failures.

6. Strengthening of beamcolumn joints

Recently, Prota et al.[58]proposed the combined use ofFRP laminates and NSM bars for upgrading RC beamcol-umn connections. NSM bars were installed on the columnprior to wrapping and anchored through the beam (i.e.the bars passed through the beam). This provided addi-tional reinforcement which is fully anchored and effectivein the maximum moment region of the column. The pres-ence of FRP wraps prevented the NSM reinforcement frombecoming ineffective as a result of load reversals. The instal-lation of NSM bars enabled the transition of the failuremode from the column to the shear failure of the joint. Infurther specimens, additional strengthening was provided

to the joint to suppress joint shear failure. For this purpose,the FRP reinforcement was placed in two directions: (a)along the beam axis, NSM FRP bars were provided andtransversely confined by single-ply CFRP U-jackets bondedto the beams; (b) along the column axis, a one-ply FRP lam-inate was provided, and this laminate was terminated belowthe base of the upper column to simulate the field conditionof the presence of a slab. With this strengthening scheme,failure shifted to the columnjoint interface at the termina-tion of the FRP laminate. The upgrading of the joint zoneincreased its deformability and hence provided a significantcontribution to the ductility of the system.

This study [58] showed that the combination ofNSM bars with externally bonded laminates enabled the

Fig. 8. Debonding failure modes of NSM bars observed in tests on shear-strengthened beams: (a) debonding of NSM bars by splitting of epoxy cover; (b)local separation of concrete cover.



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advantages of both techniques to be exploited in a comple-mentary manner. This topic deserves further investigationas similar advantages may be realised by suitable combina-tions of the two techniques in solving other strengthening

problems.

7. Research needs

This paper has provided a comprehensive and criticalassessment of existing research on the structural behaviourof RC structures strengthened with NSM FRP reinforce-ment. On the basis of this review, the main research needsin this area for the immediate future are outlined asfollows.

7.1. Bond behaviour

There is a lack of experimental evidence on the effects ofthe many variables that are likely to have a significant effecton the bond behaviour of NSM FRP bars. The number ofvariables associated with NSM FRP systems is much largerthan that for externally bonded FRP systems, but theamount of work available on the former is much more lim-ited than that on the latter. Therefore, extensive furthertesting is obviously required.

Based on existing knowledge and experience, the pre-ferred bond test set-up is believed to be a simple pull-outtest where an NSM FRP bar subjected to tension is bondedto a concrete block which is supported at the loaded end.

Using such a set-up, effects of factors such as the distancebetween adjacent grooves (i.e. groove spacing), and the dis-tance between the member edge and the nearest groove (i.e.edge distance) should be examined in the near future. Insuch bond tests, specimens with either long bond lengthsor short bond lengths may be used, and the local bondslipcurves obtained from them need to be compared to under-stand the advantages and disadvantages of using short andlong bond lengths. To obtain reliable strain measurementswithout interference with the interfacial behaviour, thedevelopment and application of innovative strain sensorssuch as fiber optic sensors embedded in the FRP bar shouldalso be given due attention.

The most important outcome of the bond tests shouldbe local bondslip curves. Further tests are required toassess the existing local bondslip equations, and moreimportantly to explore the possibility of developing generallocal bondslip models with parameters expressed as func-tions of geometry- and material-related properties. Toachieve this, the tests should preferably be assisted bynumerical modelling of the bond behaviour. For externallybonded FRP laminates, an accurate meso-scale finite ele-ment method has recently been developed by Lu et al.[59] which produced numerical results for use with testresults in the development of a set of accurate bondslip

models [35]. A similar approach should be explored forNSM FRP reinforcement. The model by Lundgren [60]which successfully predicted the bond behaviour of steel

bars in concrete, coupled with appropriate modelling ofthe epoxyconcrete interface and constitutive modellingof the materials [26], also seems to be a promisingapproach.

Analytical modelling also has a significant role to play inthe modelling of the bond behaviour of NSM FRP bars. Inexisting bond tests, splitting of the epoxy cover of NSMFRP bars has been identified as an important failure mode.Here, the frictional coefficients of different deformed barswith different groove shapes and dimensions should bemeasured as a function of the slip in order to further clarifythe splitting bond failure mechanism and develop a split-ting bond strength model to be used in splitting-criticalcases. The approach in Ref. [25] can be considered a firststep towards the definition of a simple splitting bondstrength formula accounting for geometric parametersand material properties. For this purpose, the simplifying

assumptions of a friction coefficient equal to 1 and of a uni-form distribution of radial pressures need to be removed,on the basis of experimental measurements of the frictioncoefficient.

7.2. Flexural strengthening

Obviously, given the larger number of parameters thatcan affect the flexural behaviour of RC beams with NSMFRP reinforcement, a great deal of further experimentaland theoretical work is required. In particular, the debond-ing failure mechanisms in beams strengthened with NSMreinforcement need to be clarified through further testing.The relationship between concrete cover separation andother modes of debonding local to the NSM FRPcon-crete joint such as fracture at the epoxyconcrete interfaceand splitting of the epoxy cover needs further research.Furthermore, the behaviour of pre-damaged beamsstrengthened with NSM FRP is of significant practicalinterest, as cracking and damage to the cover of the steelreinforcement may have a significant effect on the debond-ing failure process.

The relationship between bond failure mechanisms inbond test specimens and debonding failure mechanisms

in flexurally-strengthened beams needs to be clarified bydetailed experimental studies as well as rigorous theoreticalmodelling. Here, the study of the interaction between flex-ural/flexural-shear cracking and bond stresses is of crucialimportance. Once this relationship is clarified, it will thenbe possible to develop numerical and analytical modelsfor predicting debonding failures.

7.3. Shear strengthening

More tests need to be conducted to further clarify thefailure modes of strengthened beams and to evaluate the

effects of various significant factors. More tests are alsoneeded to confirm the applicability of local bondslip mod-els from bond tests in predicting debonding failures of



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shear-strengthened beams. This confirmation will facilitatethe development of accurate numerical and analytical mod-els for RC beams shear-strengthened with NSM FRP.

7.4. Other issues

As mentioned earlier in the paper, the combined use ofNSM FRP reinforcement in conjunction with externallybonded FRP reinforcement has been found to be effectivein strengthening beamcolumn joints [58]. As externallybonded FRP reinforcement alone has met with only limitedsuccess in strengthening beamcolumn joints, this com-bined approach definitely deserves further work. This com-bined use to take advantages of both techniques shouldalso be explored in solving other strengthening problems.

It is widely accepted that pre-stressing the FRP refine-

ment before bonding it to concrete structures for strength-ening purposes is often desirable, both to improve theserviceability of the structure and to make more efficientuse of the FRP material. Pre-stressing externally bondedFRP reinforcement has had little success in practice sofar because it is difficult to tension and anchor FRP lami-nates on site, particularly when they are formed by thewet lay-up process. NSM FRP reinforcement has a muchbetter chance to succeed: NSM FRP bars can be tensionedmuch more easily, particularly when compared with dryfiber sheets, and are much better anchored than externallybonded laminates.

The use of cement grout to replace epoxy as the groovefiller has been explored by a limited amount of work[9,18,23,26]. There are benefits with the use of a cementi-tious groove filler as discussed earlier in the paper.Research is needed to provide a better understanding ofthe performance of cement grout as a groove filler and toformulate stronger cementitious groove fillers.

8. Concluding remarks

Strengthening of structures with NSM FRP reinforce-ment is a technique that has attracted a considerable atten-tion as a feasible and economic alternative to the techniqueof strengthening structures with externally bonded FRPreinforcement. The former technique offers some signifi-cant advantages over the latter, including the more efficientuse of the FRP material due to a reduced risk of debondingfailure and the better protection of the FRP material fromexternal sources of damage.

Research on the strengthening of structures using NSMFRP reinforcement started only a few years ago but has bynow attracted worldwide attention. A significant amountof research has been conducted on this emerging technique,particularly on the application of this technique in thestrengthening of concrete structures. This paper has pro-

vided a detailed and critical review

19 near-surface mounted

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