Behavior of Corroded Bar Anchorages

8/7/2019 Behavior of Corroded Bar Anchorages

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756 ACI Structural Journal/November-December 2007

ACI Structural Journal, V. 104, No. 6, November-December 2007.MS No. S-2006-344 received August 21, 2006, and reviewed under Institute publication

policies. Copyright 2007, American Concrete Institute. All rights reserved, includingthe making of copies unless permission is obtained from the copyright proprietors. Pertinentdiscussion including authors closure, if any, will be published in the September-October 2008 ACI Structural Journal if the discussion is received by May 1, 2008.

ACI STRUCTURAL JOURNAL TECHNICAL PAPER

An analytical model is developed to describe the mechanics ofcorrosion-induced bond strength degradation and its implicationson development capacity of bar anchorages. The model is africtional construct whereby bond strength is estimated fromthe coefficient of friction and the normal confining pressure alongthe anchorage. Both variables are evaluated considering the relevantdesign parameters (cover, shrinkage, and transverse reinforcement)and the effects of iron depletion. The model is used to interpret thebehavior of corroded anchorages as documented in publishedexperiments. To supplement model calibration with data representativeof long anchorages, two series of flexural specimens designed tofail in anchorage after yielding are tested after being conditioned inaccelerated corrosion to a predefined damage level in the anchoragezones. The effects of corrosion on bond strength are also considered inCFRP-patch repaired anchorages. The model correlates the exper-imental evidence obtained from the various alternative testarrangements and successfully reproduces the magnitude andparametric sensitivity of corrosion-induced bond degradation.

Keywords: anchorage; assessment; bond; corrosion; cover; cracking; FRPjacket; modeling.

INTRODUCTIONCorrosion of steel reinforcement has a detrimental influence

on stiffness, ductility, and deformation capacity of exposedreinforced concrete (RC) members and it may significantlycompromise their dependable strength.1 For this reason,corrosion is a critical parameter in assessment of residualstrength and service-life of old RC construction. In quantifying

the residual strength and deformation capacity of an affectedstructural member, three primary effects need be considered:1) reduction in bar diameter (bar section loss) owing to irondepletion; 2) embrittlement of steel and the ensuing loss ofdependable deformation capacity; and 3) the manifoldimplications on bond and the development capacity of thereinforcement, which include:

An effective reduction of the coefficient of friction onthe bar surface, as the gradually increasing rust layerpromotes separation and facilitates slippage betweenbar and concrete2;

Interruption of the chemical adhesion of concrete onthe bar surface by the interpolated rust layer. This isparticularly evident after the bar is stressed, thereforecontracting laterally due to Poissons effect. Confiningpressure due to drying shrinkage of concrete may beenhanced by a small amount of rust deposited on thebar. Therefore, for low corrosion level, a slight increaseof bond may be seen before bond failure,3,4 a phenomenonthat is more pronounced for smooth bars5;

Reduction in rib height of ribbed bars, thereby looseningmechanical interlock with concrete; and

Cracking or even spalling of the cover, owing to theexpansive tendency of iron upon oxidation. Cover damagecorrupts bond along the bar, thereby reducing or eveneliminating the so-called beam action mechanism of

behavior6 and weakening the flexural stiffness and momentresistance of the member. Once bond along the shear spanis degraded, reliance on the arching action to transfer theloads to the supports increases. This has severalimplications for safety, as development of the yield forceof the reinforcement (the tie of the arch) depends greatly onthe anchorage details of the bar near the support (usually acut-off point for flexural reinforcement).

The effect of corrosion on bond mechanics is the subjectof the present investigation. To evaluate the influence of rustaccumulation on bond strength, analytical modeling andcorrelation with experimental testing have been pursued.The proposed model refers to the basic frictional construct

that underlies the ACI 3187

guidelines and accounts for theinfluences of corrosion penetration and rust buildup onthe coefficient of friction and the normal pressure actingon the bar.

The predictive capability of the model is evaluatedthrough comparison with published test results from theliterature. Reported bond strengths vary substantiallybetween different investigations. The large range of values isinterpreted with reference to the different test setups and theinfluence these have on the mechanics of bond. To analyticallyreproduce this variability, it is necessary to properly representthe actual state of stress generated by the support and loadingconditions in the concrete cover and along the anchorage inthe analysis.8 For model calibration with data concerning

longer anchorages that are more representative of the field,two series of RC specimens, cantilever beams, and simplysupported slabs are included in the experimental part of thepresent study. Before mechanical load testing, specimens aresubjected to accelerated electrochemical corrosion up to apredefined level of mass loss. Parameters of the experimentalstudy are the post-corrosion residual bond strength anddeformation capacity defined as functions of corrosion massloss. The efficacy of a repair scheme that included replacementof the spalled concrete cover with a cement-based grout andsubsequent jacketing with carbon fiber-reinforced polymer(CFRP) wraps in the anchorage zones is assessed by reloadingthe specimens after repair.

RESEARCH SIGNIFICANCEQuantifying the dependable bond strength of corroded

reinforcement is an essential step in the assessment ofaffected RC structures exposed to aggressive environments.The primary contribution of the paper in this direction isformulation and calibration of a simple analytical model forbond strength of corroded reinforcement. The model is based

Title no. 104-S72

Behavior of Corroded Bar Anchoragesby S. P. Tastani and S. J. Pantazopoulou


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757ACI Structural Journal/November-December 2007

on the frictional concept for interface action and maintainsthe format of the code expressions for bond and developmentcapacity while accounting for the primary variables knownto affect behavior of corroded anchorages. The paperincludes experimental corroboration and calibration of theprimary behavioral aspects of the model. Due to itssimplicity and familiar format, the model can be easilyimplemented within a practical framework of assessmentwhere the residual strength, deformation capacity, andfailure hierarchy of individual structural members need to beestimated from basic principles.

CORROSION-BOND INTERACTIONAND BASIS OF ANALYTICAL PROCEDURE

Corrosion products have several times the volume of theparent metal (two to six times as large9). Oxides are depositedon the bar surface generating bursting pressure on the perimeterof the hole occupied by the bar, similar to the radial componentof bond in an uncorroded ribbed stressed bar. As in the caseof bond, pressure owing to rust build-up is resisted by hooptension in the cover. Bursting pressures owing to either bondor rust build-up compete for the same strength reserve,namely the splitting capacity of the concrete surrounding thebar. In the absence of transverse reinforcement, the tensilestrength of concrete sets an upper threshold in the magnitudeof attainable expansive pressure, marked by progressivecracking of the cover. The process of crack penetrationthrough the cover concrete has been modeled numerically.10

It was shown that the resistance curve of the ring of theconcrete surrounding the bar, given in terms of pressureversus the radial displacement of interior boundary, is acharacteristic property that depends primarily on theexternal-to-internal-ring-diameter ratio and the tensilestrength of concrete (Fig. 1). In this context, any displacementof the interior boundary owing to rust accumulation exhaustspart of the available cover resistance. Thus, in corroded baranchorages, only the residual part of the resistance curve isavailable for the bond mechanism.

As corrosion progresses, a fraction of the rust is depositedin the layer of concrete around the bar filling the pores andthe cover cracks, whereas any soluble corrosion productsmay migrate away from the interface layer staining theexposed surface.9,10 Radial cracks initiate in the interiorboundary of the cover and propagate outwards with rustbuild-up.11 Radial pressures that are sustained by hoopstresses in the exterior uncracked part of the cover act toconfine the rust layer that is trapped between the bar and theconcrete cover. This confining action effectively compactsthe rust layer. Compacted rust has mechanical behaviorsimilar to that of cohesionless granular materials.2 Uponthrough splitting of the cover, hoop stresses and radial

confining pressure are released. Beyond that point, furtheringress of the aggressive agents necessary to sustain corrosion,such as oxygen, vapor, and chlorides, is greatly facilitated,leading to widening of cracks and eventual spalling of thecover concrete.

The simplest model representing the stress transferbetween steel and concrete is the so-called frictionalconcept. This model also forms the basis of the ACI 3187

requirements for bond and anchorage. Bond strength fb,quantified herein as an average shear stress acting on the

lateral surface of the bar is, apart from initial adhesion fadh,proportional to the normal confining pressure mobilized onthe bar over the anchorage. From free body equilibrium of adiametric slice (Fig. 2(a)) of the bar of length Lb, theresultant lateral shear force on the bar surface F equals0.5DbfbLb. The normal force N acting on the diametricplane of the bar equals nDbLb, where the normal confiningpressure n comprises contributions from the hoop stressesof the concrete cover c, the reaction of stirrups as they crossthe splitting plane st (calculated as the average normalcompressive stress in reaction to the stirrup tensile forces),and any transverse compressive stress field confexisting inthe anchorage zone. Considering that F = N +0.5Db fadhLb, where is the coefficient of friction, it

follows that

(1a)

The common expression for the familiar frictional equation indesign codes7 has the form

fb = n + fadh (1b)

Thus, the values of the term taken from the internationalliterature correspond to the product 2/. Generally,

fb2

------n fad h2

------ c st conf+ +( ) fad h+=+=

S. P. Tastani is a Research Engineer and an Adjunct Lecturer in the Department of

Architecture at Demokritus University (DUTh), Thrace, Greece. She received her Civil

Engineering Diploma, MSc, and PhD from DUTh. Her research interests include bond of

reinforcement (steel and FRP bars), reinforcement corrosion, and use of FRPs in seismic

repair/upgrading of reinforced concrete structures.

S. J. Pantazopoulou, FACI, is a Professor in the Department of Civil Engineering at

Demokritus University. She received her PhD and MSc in structural engineering from

the University of California at Berkeley, Berkeley, CA, and an Engineering Diploma

from the National Technical University of Athens, Athens, Greece. She is a member of

ACI Committees 341, Earthquake-Resistant Concrete Bridges; 374, Performance-Based

Seismic Design of Concrete Buildings; 408, Bond and Development of Reinforcement;

and Joint ACI-ASCE Committees 352, Joints and Connections in Monolithic Concrete

Structures, and 445, Shear and Torsion. Her research interests include the mechanics ofreinforced concrete and earthquake design of reinforced concrete structures.

Fig. 1Response of concrete cover under radial displacementimposed by corrosion.

Fig. 2(a) Schematic depiction of frictional model; and (b)modeling stirrup pressure along spacing s.


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decays with slip or crack opening. Stirrups produce a passiveconfining pressure over the anchorage, but unless they havea rather small diameter, they rarely reach yielding. Note thatin unconfined anchorages, slip is approximately twice theradial displacement12 of the internal bar boundary imposedby the displacing ribs. Thus, upon attainment of bondstrength at a nominal slip value of 0.1 mm (0.0039 in.), thecorresponding radial displacement is 0.05 mm (0.002 in.).(This result was obtained from local bond tests13 with anembedded length Lb = 5Db, cover c = 1.5Db and bar diameter

Db = 19 mm [0.748 in.].) The corresponding hoop strainequals the radial displacement divided by the radius to thepoint considered; thus, at the internal and external boundaries(free surface) the strain is 0.05 mm/9.5 mm = 0.0053 and0.05 mm/38 mm = 0.0013. In the presence of confiningpressure, radial displacement is greatly restrained. Forpressures in the order of 0.25fc , cover splitting is entirelysuppressed and failure occurs by pullout.13

Recently, a new technique for upgrading inadequateanchorages/splices by externally bonding FRP sheets (EB-FRP)orthogonal to the direction of reinforcement has been developed(FRP patch repairs). The technique has been shown toperform well, successfully enhancing anchorage strength andcontrolling bond degradation beyond attainment of strength,both for common and for corroded anchorages.14 In detailingthis repair method, the effective transverse strain f,effon thenearest surface of the cover is used to evaluate the sheetstress. For the example considered, f,eff= 0.0013 > cr, thatis, the cover is split through. This value may be considered asthe effective strain of EB-FRP jacket restraining theanchorage when a moderately high level of frictional resistancemay be relied upon ( 0.7 1).7 Due to the kinematicrelat ionship between slip, splitting crack width, and friction,higher values for f,effwould correspond to lower values.Using the format of the code expression7 (in metric units) theaverage bond strength is calculated from Eq. (1a) as follows

(2)

where contributions from concrete cover, stirrups, and EB-FRPjackets have been considered. In Eq. (2), the implicit valuesfor the parameters are = 0.9 (that is, in Eq. (1b), 0.6), andft = 0.5 ; c is the concrete cover; Nb is the number of barsrestrained by the stirrup legs included in Ast(Astis the cross-sectional area of stirrups crossing the splitting plane); and tfand Ef are the EB-FRP jacket thickness and modulus of

elasticity, respectively. The coefficient averages stirrup-induced confining pressures over the spacing s; note thatconfining pressure exerted on the bar by a single stirrup layeris maximum at the stirrup location and decays with transversedistance (Fig. 2(b)). Whereas the core of the cross sectionmay be effectively confined over the entire stirrup spacing,most of the bar length is outside the stirrup influence (dashedline). Herein, a parabolic distribution is assumed for theconfining pressure over the bar. This corresponds to auniform effective confining pressure acting on the longitudinalbar over the entire spacing of successive stirrups s approximatelyequal to 1/3 of the peak value (hence, = 0.33 in Eq. (2)).

fb2

------ c st f+ +( )2ftDb------------ c

Ast fst y,ftNbs

---------------------2 tfEff eff,

ftNb------------------------+ +

==

fb

0.29 cDb------ fc 0.19

Ast fst y,

DbNbs--------------- 1.15

tfEff e f f ,DbNb

------------------++=

fc

Equation (2) refers to anchorages that usually fail by splittingwhere the confining contribution of stirrups is limited due tosparse spacing. For this, it is necessary to impose an upperlimit of c + st 0.25fc .

13 Although this is a cruderepresentation of the local stress concentrations around theribs that engage in concrete, in a general sense, the simplefrictional model properly identifies the significance of manyimportant design parameters for bond, and may be easilyadapted to the conditions of a corroding reinforcing bar.

PROPOSED ANALYTICAL MODELThe simplest method to introduce the effect of corrosionon the bond capacity fb is to apply a reduction coefficient(X) on Eq. (2) with an increasing level of corrosion such asfb

cor= (X)fb; Xis the percent loss of bar radius (referred toas depth of corrosion penetration). A number of models arelisted in the literature to represent this trend3,15-17 obtainedby fitting experimental bond data. In all cases, bond decayswith increasing X; however, the validity of the models islimited when tested with experimental measurements thathave not been used in their initial derivation. The discrepancybetween models is mostly owing to the great variety of specimenform and test setup used by the individual investigators. Notethat bond is usually measured indirectly, that is, it is reduced

from the global response of the specimen and testing hardwarebehaving as a complete structure. Unaccounted for frictionbetween specimen and test hardware, as well as the presenceof diagonal compression stress fields in the vicinity of theanchorage are known to superficially enhance the bondcapacity.18 Auyeung et al.3 and Hussein et al.4 collected datafrom direct tension pullout tests (that is, most adverse bondconditions [Fig. 3(a)]). Cabrera16 used conventionalpullout tests (very favorable bond conditions [Fig. 3(c)]).Rodriguez et al.15 examined short beam-end tests, and Stanishet al.17 tested simply supported slabs (bond is favored by thepressure at the support in the anchorage region [Fig. 3(b)]).

A common drawback in all available empirical models isthat a single variable, namely (X) is used to account for

different sources of bond degradation, that is, the reducedcapacity of the cracked cover and the attenuation of the frictionalproperties of the bar-concrete interface due to corrosion. Analternative approach consistent with the original frictionalbond model (Eq. (1a)) is proposed herein, where degradationof all relevant mechanisms is explicitly addressed

Fig. 3Test setups used in bond testing. Concrete is: (a) indirect tension; and (b) and (c) in compression.


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ACI Structural Journal/November-December 2007 759

(3)

In the proposed model, the coefficient of friction as wellas the various terms in Eq. (3) that represent normal pressures onthe bar surface are functions of corrosion intensity measuredby X. The term shris the shrinkage stress that is modeled forsimplicity as a compressive radial stress. The component of

bond attributed to initial adhesion fadh degrades rapidly witheither slip or corrosion. For smooth bars, it may represent asizeable fraction of the total resistance, but for ribbed bars, it isconsidered insignificant and is only included in Eq. (3) forcompleteness. Terms entering Eq. (3) are defined in thefollowing sections with regard to the principal design variablesas well as with reference to X using simplified mechanisticconstructs and physical argument.

Calculating crack front in cover Rcrdueto rust build-up

Radial cracks run through part of the cover of corrodedbars, owing to accumulation of the expansive rust products.An important parameter controlling the normal stress

components in Eq. (3) is the residual uncracked coverthickness, which can be relied upon to support bond action.The radius of the crack front Rcr is related to the radialdisplacement ur,o of the internal boundary of the cover ring(around the bar hole) imposed by rust build-up. The variableur,o is calculated from the depth of corrosion penetration X,and the volumetric ratio of rust to parent iron rs. For thedefinition of the associated boundary-value problem, thefollowing considerations are made.

Strain-displacement relationships for cover ringInaxisymmetric problems of continuum mechanics, the strain-displacement relations are expressed in polar coordinates: r=dur/drand = ur/rwhere uris the radial displacement and risthe radius from the line of axisymmetry (Fig. 4(a)). If smeared

strain definitions are adopted, these relations hold for thestate of strain in the cracked cover concrete. For a knowndisplacement of the interior boundary of the cover ring, theradius of the crack front may be evaluated from these relationsby setting the hoop strain equal to the cracking strain of concrete,that is, = cr= ft/Ec and hence, Rcr= ur,cr/cr.

The exact variation of the radial displacement from the valueur,o at the internal boundary Rb, to ur,crat the crack front Rcr, isgenerally unknown and may only be obtained from a completesolution of the associated nonlinear boundary-value problem.10

Herein, it is assumed that the radial displacement decays fromthe peakur,o value according to

(4)

where r is the radial compressive strain resulting from theresidual radial stress n

res that acts on the internal boundary(thus, for simplicity, ris taken constant in the cracked partof the ring). As corrosion proceeds, the radial stress attenuatesfrom its peak value (which represents the stress capacity ofthe ring n) following an inelastic path with a slope equal tothe initial concrete elastic modulus Ec (Fig. 4(b)). The radialstrain is then

(5)

b

co r 2

--- X( )n

re sfad h X( )

2

--- X( )=+=

c X( ) sh r X( ) st X( ) conf+ + +[ ] fad h X( )+

ur ur o, r r Rb( ) Rb r Rcr ;=

r rma x n n

re s( ) Ec=

Hoop strains, being tensile and attaining their peak valueat the internal boundary, result in gradual propagation ofradial cracks outwards through the cover. Thus, theuniaxial concrete compression stress-strain law sufficesto relate the radial strain r

max at the internal boundarywith the corresponding stress n (for example,Hognestads parabola) considering that Poisson effects arediminished after cracking.

Translation of internal boundary due to rust accumulationur,oThe volume of rust Vr is related to the volume of

depleted metal Vs and to the depth of corrosion penetrationX(X= Db/Db) through Vr= rs Vs. The volume ratiors is usually taken as 2, but higher values are possible if therust is fully hydrated.10 Assuming uniform corrosion on thebar surface (as opposed to pitting), deposition of the rustvolume Vraround the bar surface and in the radial cracksof the cover requires a radial displacement of the internalconcrete boundary by

(6)

For a given X, the radius of the crack front Rcris evaluatedby substituting Eq. (6) into Eq. (4); setting the radial

displacement ur,cr= crRcr, it may be shown

(7)

An initial estimate for the depth of corrosion penetrationXcrit associated with cracking is obtained from Eq. (6) bysetting ur,o= crRb and Rcr= Rb: thus, Xcrit= 1 [1 2cr/(1 rs)]

0.5.Second-order translation of internal boundary due to

compaction of accumulated rustun,oThe real increase inthe radial displacement of the internal concrete boundary

un,o not only depends on the rust volume stored in the poresand in the space within cracks, but it is also affected by thedegree of compaction experienced by the rust layer as aresult of the normal pressure generated by the severalmechanisms described previously. Rust behaves as acohesionless granular material. Its stress-strain law has theform shown in Fig. 4(c) (unloading of rust due to degradationof the n

res is determined by the tangent slope at n). A perti-nent mathematical model of this behavior has been formulatedby Lundgren2 where the relationship between the actual radialdisplacement un,o, the rust strain cor(radial compaction of therust layer), and the radial displacement due to free (uncom-pacted) rust deposition ur,o is as follows (Fig. 4(a))

ur o, 1 rs( )Rb2

2X X2

( ) Rb Rcr+( )=

Rcr

Rb

cr2

4 cr

r

+( ) 1 rs

( ) 2X X2

r++

cr

2 cr

r

+( )--------------------------------------------------------------------------------------------------------------------------------=

Fig. 4(a) Thick-walled cylinder model of cover and corrosionindexes; (b) stress-strain law of concrete cover under radialcompression and unloading path when radial pressure isreduced; and (c) stress-strain response of rust.


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(8)

An initial estimate ofur,o was obtained from Eq. (6) forX= Xcrit. Up to this value, the ring capacity n acting on theintact concrete cover is calculated using the basic mechanicsof the ring model as detailed in the following sections.Pressure n acts both on the cover and on the rust layer as thelatter builds up around the bar. Using the constitutive modelof compacted rust (Fig. 4(c)), the rust compaction strain coris evaluated from n. The actual (corrected) radial displace-ment of the internal boundary un,o is then estimated from Eq.(8). From this revised value of the internal boundary transla-tion, all relevant parameters of the problem are recalculated(for example, the crack front Rcrfrom Eq. (4) after settingur,cr= crRcrand r= r

max and the residual normal pressureon the bar).

In a stepwise calculation algorithm given in each step, thedisplacement of the internal boundary un,o and the radialstrain rfor a depth of corrosion penetration X, an incrementin penetration by Xwill lead to calculation of the newas follows

(9)

Equation (9) actually estimates the accumulated radialdisplacement considering the additional compaction of thenew rust layer associated with X. If the n

res(i) is less thanthe achieved maximum value n, then unloading bothcracked concrete and rust is assumed (along the unloadingpaths shown in Fig. 4(b) and (c), respectively).

Capacity of cover ring cThe concrete cover contribution c in Eq. (3) is the normal

pressure on the interior boundary of the hole occupied by thebar that would be required to split the cover (Fig. 1 and 4(a)).Assuming either fully elastic or fully plastic behavior for the

ring,19

its capacity to radial pressure is calculated as c = ft(c/Db) ( = 1, 2, respectively). If radial cracks due tocorrosion have propagated to the crack front Rcr, then theresidual capacity of the cracked ring is estimated accordingly as

(10)

The first term of Eq. (10) denotes the remaining pressurecapacity supported by the uncracked part of the cover ring.Quantity A(,o) is the residual pressure resistance supportedby the cracked part of the ring. Its magnitude depends on thepost-fracture characteristics of the stress-strain law ofconcrete in tension that defines its fracture energy.10 Thisterm is calculated when hoop strains at the internal boundary,o = un,o/Rb exceed the cracking strain cr.

Effect of drying shrinkage shrDrying shrinkage is accounted for as an isotropic (volume)

contraction with a magnitude that depends on the conditionsand duration of exposure. Accounting for that contraction asa uniform prestrain, the rust volume required to split the

un o, ur o, co r X Rb ur o,+( )=

un o,i( )

un o,i( )

ur o,i( ) co r

i( ) X Rb ur o,i( )

un o,i 1( )

( )+[ ]=

c X( ) cCc Rcr( )

c------------------------ A o,( );+=

A o,( )1

Db------ c

Rb

Rcr

o,( )dr=

cover is effectively increased, whereas shrinkage enablesincreased bond resistance for small amounts of rust deposition.

In the thick-walled cylinder analogy, shrinkage generatescompressive stresses in the radial direction shr as theembedded bar restrains the length change of concrete, actingas a rigid inclusion. In design codes, shrinkage is interpretedas an isotropic contraction by shr; based on CEB-MC1990,20 shr = 0.0003 3cr for humid atmosphericconditions. Thus, the radial stress shr is taken equal to themodulus of elasticity of concrete times the isotropic strain

shr: shr= Ec shr= 3ft. Consistent with the assumptionof isotropy of shrinkage, a compressive strain equal to shrisalso resolved in the hoop direction. This strain beingcompressive delays the onset of tensile hoop stresses due tocorrosion penetration up to an internal radial displacementequal to = shrRb ( is the necessary quantity forcomplete elimination of hoop compressive strains in thecover). After hoop strain exceeds the cracking limit, that is,(un,o/Rb) shrcr, initiation of cover cracking is inevitable.Beyond that stage, as radial cracks propagate through the coverthickness, shrinkage stress is assumed to reduce proportionatelywith cover resistance, that is, shr(X) = 3ft(Cc Rcr)/c.

The depth of corrosion penetration Xshrthat would cause a

radial displacement of the internal boundary by may becalculated from the model considering that no radial crackingmay occur up to that point (that is, /Rb = shr). Based onanalysis of several case studies,3,4,15,16 it was found thatXshr 0.001; this quantity is insensitive to cover thicknessand concrete quality.

The concrete contribution conc to bond strengthcomposed of the tensile strength of concrete cover c(X) andthe shrinkage-generated pressure shr(X) is a lower boundvalue because it is based on the assumption of symmetricconditions around the ring (central bar in the cover ring):conc(X) = c(X) + shr(X). In actual circumstances, the ringradius is defined by the smallest cover thickness to the freesurface of the member. Depending on the actual geometry of

the cover (whether the bar is centrally or eccentrically placedin the cross section), the estimated crack propagation representscracking in the thinner part of the cover only, whereas a largefraction of the thicker part of the cover may remainuncracked. Thus, the normal pressure in the inside boundaryof the ring, calculated from the force resultant of hoopstresses along an assumed diametric plane, would be morefavorable in an unsymmetric condition where part of theactual cover would exceed the minimum dimension used inthe ring model (Fig. 5(a)). To account for this increase in theestimated ring capacity, a geometric factor B 1 is introduced inthe terms representing the cover ring contribution; the factorB accounts for the area ratio of thin to thick cover areas(Fig. 5(a)). The resulting enhanced concrete contributionterm is estimated from

(11)

In Eq. (11), the term conc is the local pressure increase(over the conc value, which is the minimum value assuminga uniform cover equal to the minimum thickness as shown inFig. 5(a)) in the thicker part of the bar cover, that is, B conc = B(c + shr c(X) shr(X)).

un o,li m

un o,lim

un o,

li m

un o,li m

conctotal

X( ) conc X( ) B + conc=

1 B( ) c X( ) sh r X( )+( ) B c sh r+( )+=


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(72.5 ksi) were cast in each specimen. The bars were placedin the axis of symmetry of the beam cross section, with aclear cover ofc = 20 mm (0.79 in.). The required anchoragelength was estimated as Lb = fyDb/(4fb) = 415 mm (16.34 in.)(30Db, where the nominal bond strength was taken as perEC-222 as fb = 2.25 ft). The remaining 415 mm (16.34 in.) ofthe total embeddment length was provided with a bondbreaker. To preclude premature shear failure, auxiliarylongitudinal and transverse reinforcement was used (fourbars ofDb = 10 mm [0.39 in.] placed at the corners); only

vertical stirrup links with a diameter of 5 mm (0.20 in.) wereprovided to avoid any confining influence by closed stirrupsin the end regions of the test bars. The specimens wereloaded as a cantilever beam with one bar pulled out. The testbar was gripped externally using a tendon anchorage wedgethat was bearing under the seating plate; the latter wasadjusted on the loading frame. To preclude lateral slidingand out-of-plane rotation, two pin supports were fixed on theframe (one on each section side [Fig. 6(a)]). Tip displacementtip was obtained as the sum of deflection owing to flexuralcurvature and deflection owing to bar pullout from the fixedsupport sp (Fig. A1 in the Appendix*). These values wereobtained from horizontal and vertical DTs placed at the topof the specimen and near the support region, respectively

(Fig. 6(a)).The S-group consisted of four one-way slab-strips with a

cross-sectional height, width, and total length of 150 mm(5.91 in.), 330 mm (12.99 in.), and 1200 mm (47.24 in.),respectively (Fig. 6(b)). Concrete compressive strength was32 MPa (4.64 ksi) at 28 days (at the time of first testing, thestrength had increased to 40 MPa [5.80 ksi]). A single layerof two Db = 14 mm (0.55 in.) steel bars with fy = 500 MPa(72.5 ksi) was used as longitudinal reinforcement with aclear cover ofc = 30 mm (1.18 in.). Specimens were testedunder four-point loading. Reinforcement was covered with abond breaker in the central region of constant bendingmoment (400 mm [15.75 in.]) so as to control the bonddemand over the bar development length. The available

anchorage was 360 mm (14.17 in.) (that is, 26Db) andoccurred in the shear span; the bar was terminated 40 mm(1.57 in.) beyond the support.

In the specimen identification code (Table A1 in theAppendix), the first character marks the specimen type (B-or S-group), the second is the specimen number in the group,while the last two characters refer to the conditioning (c forcorroded specimens and R for repairs with FRP jackets).

Corrosion conditioning of specimensBoth groups of specimens were connected to an electro-

chemical corrosion cell for a period of 2-1/2 to 3 months so as togenerate accelerated rust production in the laboratory. Onespecimen of each group was left uncorroded to be used as a

control (B0 and S0 [Table A1 in the Appendix]). Specimenswere immersed in a water solution containing 3% per weightNaCl with bars acting as the anode of the circuit. A steelmesh was used as a cathode, placed at the bottom of thecorrosion basin. Bars and steel mesh were connected inparallel to the power supply with the salt solution completingthe circuit; a power of 6 Volts was impressed at the circuitends. The electrical current was monitored at 12-hour intervals.To promote accumulation of expansive corrosion products,

that is, to achieve complete hydration of rust, a 3-day cycleof wetting/drying was used during conditioning (1/2 day ofwetting up to the middepth of the bars followed by 2-1/2 daysof drying, where the water level was lowered well under thespecimens bottom face).

Both bars in the S-group were subjected to electrochemicalcorrosion. In one specimen of the B-series, both longitudinalbars were connected to the power supply (B5cR and B6cR inthe Table A1 in the Appendix); this specimen was placedsideways in the corrosion basin so as to achieve simultaneous

corrosion of both bars. In all other specimens of the B-group,only one of the two cast bars was corroded. The mass of ironconsumed over the time period was estimated by the totalamount of current that flowed through the electrochemicalcorrosion cell using Faradays Law. Values for the measuredaccumulated current Icorrand the estimated depth of corrosionpenetration Xare given for all specimens in Table A1 (inthe Appendix).

The duration of corrosion conditioning was similar for thetwo groups of specimens, but the different concrete strengthsand cover-to-bar-diameter ratios (c/Db) produced differentdegrees of damage with regards to mass loss and cracking.For the B-group (c/Db = 1.43), cracks 2 mm (0.08 in.) widespread throughout the concrete cover and anchorage length;

in some cases, the network of cracks extended through theheight of the section (Fig. A2(a) in the Appendix). By theend of the conditioning period of the S-group with c/Db =2.14, visible longitudinal cracks 0.5 to 1 mm (0.02 to 0.04 in.)wide had developed directly over the bars in the anchoragezones only. Apparently, the bond breakers that were placedin the middle noncorrosive zone (Fig. A2(b) in the Appendix)inhibited the flow of oxygen starving the corrosion mechanism,effectively protecting the bar against rusting in that zone.

Mechanical load testingBoth groups of specimens were designed to fail in the

anchorage after flexural yielding. The test setup andinstrumentation are illustrated in Fig. 6. Load was applied

monotonically in load increments (up to yielding) of 1 kN(0.11 kips) with continuous monitoring of deformations;displacement control was used beyond yielding.

Beam-end specimensThe B-group was tested up to various levels of lateral load

with critical thresholds being: 1) yielding of longitudinal

Fig. 6Dimensions (in mm) and test setup of specimenstested: (a) beam-end (B-group); and (b) slab (S-group).(Note: 1 mm = 0.0394 in.)

*The Appendix is available at www.concrete.org in PDF format as an addendum tothe published paper. You may request a hard copy from ACI headquarters for a feeequal to the cost of reproduction plus handling at the time of the request.


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763ACI Structural Journal/November-December 2007

reinforcement; and 2) the extent of mechanical damage inthe anchorage zone at cover spalling; the aim was the reuseof specimens for testing various repairing techniques.

Figure A1 (in the Appendix) plots the load-displacementcurves obtained during the first loading stage before repair,that is, the c series marked by circles and triangles with noconnecting lines (also included in the same figure are theresults after the repair, as mentioned in the following).The term cR in the specimen label marks experimentalcurves obtained from the post-repair loading tests. Loading

of Specimen B1c (X= 10.9%) was stopped at 15 kN (3.37 kips)before yielding of the corroded bar, with limited cracking(the yield load of the control Specimen B0 was estimatedboth analytically and experimentally as 21 kN [4.72 kips]).Testing of Specimens B2c and B3c (X= 6.8 and 9.6%) wasstopped just near visible yielding (at approximately 17 kN[3.82 kips]). Nevertheless, a substantial level of damage,including spalling of the cover, was only seen in SpecimenB2c. Specimen B4c with a similar degree of corrosion asSpecimens B3c and B1c failed abruptly, developing only25% of the strength of Specimen B0; the embedded baryielded and subsequently fractured near the support region inthe specimen block (it sustained a low yield load due to localbar section loss) without widening of the existing longitudinal

cracks or formation of any new cracks associated with bondaction. For Specimens B5c and B6c with a low degree ofcorrosion penetration (X = 2.96 and 3.7%), testing wasterminated at higher load (up to 83% of the fai lure load ofSpecimen B0). At that stage, transverse cracks had developedin the lower half of the anchorage length whereas the mainlongitudinal crack owing to corrosion had widened (excessivedamage in the case of Specimen B5c).

In most specimens, corrosion impaired the flexuralstiffness of the members as a result of partial breakdown ofbond. Due to lack of bond, it is not possible to mobilizetension stiffening of concrete and crack widths becomeexcessive; in that case, pullout due to slip may govern thetotal lateral deformation.14 Exceptions to the experimental

trends are the responses of Specimen B5c and B6c that also hadthe lowest degree of corrosion; also, Specimen B3c had a verysimilar stiffness to Specimen B0 with lower yield point.

Comprehensive results of the preloading stage for theB-group are also given in Table A1 (in the Appendix).Average bond stress fb was obtained assuming uniformdistribution along the anchorage length and was evaluatedfrom statics of the cantilevers critical section using themeasured peak load P.

Slab-strip specimensBond was measured indirectly in the S-group (the developed

bar force was obtained from the statics of the constantmoment region) and it was affected by the arching of thecompression strut near the support. From the three corrodedelements of this group (in all cases X 4%), Specimen S3c andthe control Specimen S0 were loaded to failure. Peak loadswere of similar magnitude (121 and 126.5 kN [27.2 and28.4 kips], respectively), but the mode of failure wasdifferent. A large flexural crack opened in the constantmoment region of Specimen S0, extending deep in thecompression zone where it curved and became almosthorizontal as the load increased (Fig. A3(a) in theAppendix). (Note that the single crack is owing to fact thatthe bar was unbonded in that region.) Near failure, the crackwas approximately 3 mm (0.12 in.) wide. A flexure-shear

crack also opened in the shear span near the point of loading;eventually failure was controlled by this crack that alsobranched towards the support (the photo in Fig. A3(a) in theAppendix was taken after load removal). Specimen S3cfailed abruptly in the anchorage zone due to corrosionprecracking with part of the concrete cover exploding away(Fig. A3(b) in the Appendix). Specimens S1c and S2c wereloaded up to 65% of the measured strength of Specimen S3so that they could be reused after repair (Fig. A3(c) in theAppendix). Peak load and the resulting bond stresses are

listed in Table A1 (in the Appendix).The influence of specimen type and particularly the

confining action of the supports over the bar is assessed bycomparison of the test results of the uncorroded specimens ofthe two series (the uncorroded Specimens B0 and S0 reachedaverage bond capacities of 5.61 MPa (0.81 ksi) and 7.07 MPa(1.03 ksi), respectively (Table A1 in the Appendix), that is, a20% difference in strength. This difference persisted system-atically in comparisons of similarly corroded specimens of thetwo groups as well (owing to the different specimen forms).

Influence of repair on anchorage capacityAfter the first loading phase, all damaged specimens were

repaired in the anchorage zones by CFRP patches externally

bonded over the anchorage length and transversely to thelongitudinal axis of the members. In all cases, the FRP jacketwas anchored laterally along the height of the cross sectionforming a U-shape (no mechanical anchorage was used).The intent in this repair alternative was to mobilize passiveconfining of the anchorage zone with consequently favorableinfluence on bond resistance. Nominal mechanical properties ofthe jacket materials were as follows. For the CFRP tensilestrength: 3500 MPa (508 ksi), strain at failure 1.5%, andsheet thickness of 0.13 mm (0.0051 in.); and for the resinmodulus of elasticity: 3800 MPa (552 ksi) and shear strengthof 30 MPa (4.35 ksi).

Repair was done as follows: in cases of excessive coverdamage (Specimens B2cR and B5cR), before application of

the CFRP jacket, the concrete fragments and the rust wereremoved and replaced by a low strength mortar (water-cement ratio of 0.7). For all other members with controlledcracking, CFRP plies were just glued to the free surfacewithout any particular preparation (that is, the corners werenot chamfered and cracks were not sealed) except for carefulcleaning. During repair, the S-type specimens were turnedupside down with respect to their testing position, so thatcracks were closed having a favorable effect on the anticipatedeffectiveness of the jacket.

The repaired specimens were reloaded monotonically tofailure. The contribution of CFRP jacketing on bond resistancein the case of the S-group was clearly favorable. Stiffnessincreased up to 20% of the initial value of the corrodedspecimens due to crack closure upon repair, whereas the modeof failure was ductile (in Fig. A3(c) in the Appendix, label Rin the plot corresponds to the post-repair response). Measuredbond strength fb = 6.7 MPa (0.97 ksi) was reduced by only 5%from the value of the control Specimen S0 (Table A1 inthe Appendix).

In the B-group, corrosion damage and preloading effectson bond resistance and on the stiffness of load-displacementresponse were almost nonrecoverable (post-repair responsecurves are plotted in Fig. A1 (in the Appendix) by the seriesof symbols with connecting lines). By combining coverreplacement with CFRP patching (Specimens B2cR and


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B5cR), it is possible to develop the available flexuralstrength (attained at the prerepair loading). Failure for allspecimens was marked by excessive slip of reinforcementowing to the preloading effects on the anchorage, theunrecoverable loss in rib height, and the fact that the rustlayer was not removed before casting the new cover. UsingCFRP patching alone was not as effective as in the case ofthe S-group, particularly in cases that had sustained a highlevel of preloading: Specimens B3cR and B6cR could notsustain more than 63% of the preload threshold and failed in

a similar manner as Specimen B2cR. Specimens B1cR,which had the lowest degree of mechanical damage (onlyfew narrow cracks near the support due to the low level ofprerepair loading, whereas the anchorage had experiencedthe least damage from bar slip) and was repaired by CFRPlayers without cover replacement demonstrated the mostfavorable performance. Its ultimate load reached 80% of theSpecimen B0 strength, with pronounced yielding beforefailure and unaffected initial stiffness. Measured bond strengthfb was lower in most cases (except for Specimen B1cR) than

the bond stress value attained during the first loading cycle(Table A1 in the Appendix).

VERIFICATION OF ANALYTICAL MODELFor a given depth of initial corrosion penetration, the

residual bond strength may be calculated from Eq. (3), withthe proposed values for the frictional coefficient and thecalculated values of the normal stress components. In thefollowing paragraphs, the performance of the analyticalmodel is explored through correlation with reported

measurements from relevant tests published in theliterature3,4,15-17 as well as with the results of the experimentalprogram conducted in the framework of the present study. Intotal, five different series of tests were considered for modelverification. Typical characteristics of the specimens andrelevant references are as follows:

1. Short embedment length, concentric bar arrangement,and tension pullout test setup3,4 (Fig. 3(a));

2. Short embedment length, eccentric bar arrangement,and beam tests15 (Fig. 3(b));

Fig. 7Model correlation. Data from: (a) conventional pullout tests; (b)direct tension; (c) beam-end ((a), (b), and (c) concern short anchorages);(d) slab strips; and (e) data from current test program ((d) and (e) concern longanchorages).


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ACI Structural Journal/November-December 2007 765

3. Short embedment length, concentric bar arrangement,and pullout tests16 (Fig. 3(c));

4. Long embedment length, eccentric bar arrangement,and slab-strip tests (data from present experimental program,as well as from Stanish et al.17) (Fig. 3(b)); and

5. Long embedment length, eccentric bar arrangement,beam-end tests (data from present experimental study[Fig. 3(b)]), and consideration of patch-FRP repairs.

Figures 7(a), (b), and (c) present the behavior of the modelwhen applied to Specimen Series 1 through 3.3,4,15,16 A

common characteristic of these cases is the short embeddedlength of the test bar (4 7Db). Stress conditions in thesurrounding concrete vary depending on the test setup andsupport hardware. Overall, the model is in good agreementwith the experimental data or with the empirical equationsincluded in Fig. 7 that summarize subgroups of test results.In the case of conventional pullout tests16 of Fig. 7(a), themodel is in good agreement with the experimental data in thelow range ofXvalues, whereas it becomes rather conservativefor higher levels of corrosion penetration. The influence ofthe reacting confining pressure confowing to the test setup(Fig. 3(b) and (c)) was considered in the correlation as specifiedby Eq. (3). Datapoints circled with the dashed line concern aspecial concrete mixture containing fly ash.

The model best reproduces the data of the direct tensionpullout test3,4 as illustrated in Fig. 7(b). In this type of test,longitudinal concrete stresses are tensile in the vicinity of theanchorage, thus they are more consistent with the stress-statesurrounding the anchorage in the tension zone of an actualbeam. For the beam-end test series15 (Fig. 7(c)) with andwithout stirrups, the model produces a lower bound trend(according to Rodriguez et al.,15 stirrups were uncorroded).Parameter B was set equal to 0.75 taking into account theeccentric bar arrangement.

In the case of long anchorages (Fig. 7(d) and (e)), themodel correlates well with the values of bond strength (in thepresent study, bond strength was measured upon failure ofrepaired specimens in the second test phase [Fig. 7(e)]). In

the correlation, contribution of the FRP jacket (from Eq. (2))and influence of the inclined compression field due tosupport reaction were accounted for through Eq. (3) by meansof the reacting normal pressure conf + f (in the case ofdata17 of Fig. 7(d): f= 0). For Specimen Series 4, Parameter Bwas taken as 0.75 (side-to-bottom-cover ratio 2).

In the B-series of specimens (Series 5, Fig. 7(e)), thecomposite material did not mobilize sufficient restrainingpressure f, as this would require significant hoop strainsat the external boundary. Herein, replacement of corrosion-cracked cover as a repair option was analytically modeled byusing an intact cover. Thus, only the coefficient of friction was reduced as a function of X to account for corrosion-induced deterioration of the bar surface (herein,

conf= 0

because the inclined compression field affected the tail-end ofthe unbonded bar). In modeling the S-specimens without coverreplacement (Fig. 7(e), black line), the combination of confiningpressures in the anchorage zones exerted from the support andfrom the FRP jacket acted favorably to sustain bond strength upto severe values ofX, preventing widening of the existing covercracks (B = 0.75) in the analysis.

In all cases, the model consistently reproduced theexperimental trends and the varying degree of influenceof corrosion on bond strength owing to the stress-state of theconcrete in the anchorage zone and the boundary conditionsprovided by the test setup.

CONCLUSIONSTwo series of RC specimens (slabs and beam-end specimens)

were conditioned under accelerated corrosion and subsequentlytested under mechanical load so as to investigate bondperformance of corroded bar anchorages. Some of the damagedanchorage zones were repaired with partial cover replacement.All specimens were wrapped with a CFRP patch and retested tofailure. Test results showed that corrosion affects the mechanicsof bond and can result in explosive spalling failure at theanchorage. The efficiency of CFRP patching was controlled by

several parameters such as the extent of corrosion damage, theextent of cover damage due to preload, the state of stress in thevicinity of the anchorage, and by whether the damaged coverwas replaced before jacketing.

To interpret the test results, a mechanical model wasdeveloped from first principles based on the frictionalconcept, whereby bond strength was estimated from thecoefficient of friction and the normal confining pressurealong the anchorage. Both variables were evaluated consideringthe relevant design parameters (cover, shrinkage, and transversereinforcement) and the effects of iron depletion (bar diameterloss). The models ability to accurately predict deterioration ofbond strength with progressive corrosion penetration wasestablished through extensive correlation with data from

several published studies as well as with the data of thepresent experimental investigation. In all cases, the modelconsistently reproduced the experimental trends and thevarying degree of influence of corrosion on bond strengthowing to the stress-state of the concrete in the anchoragezone and the boundary conditions provided by the test setup.

ACKNOWLEDGMENTSThis research was conducted in the Laboratories of Demokritus University,

Greece. Funding was provided by GSRT (Hellenic General Secretary forResearch and Technology) through the PENED 2001 program. The compositematerials (SikaWrap 230C and Sikadur 330) were donated by SIKA Hellas.

NOTATIONAst/st

cor = cross sectional area of transverse reinforcement, reduced value

due to corrosionB = factor that accounts for area ratio of thin to thick cover areasc/Cc = clear concrete cover thickness and cover measured from bar centerDb/Rb = bar diameter and radiusEc/s/f = modulus of elasticity for concrete, steel, and FRPfadh = adhesionfb/b

cor = bond strength and reduced value due to corrosionfc/ft = compressive/tensile strength of concretefst,y/fy = yield stress of transverse and longitudinal reinforcement

Icorr = total current that flowed through electrochemical cell duringcorrosion

Lb = anchorage lengthNb = number of tension longitudinal bars enclosed by stirrupsRcr = radius of crack frontr = radius from line of axisymmetrys = stirrup spacingtf = FRP jacket thickness

ur,cr = radial displacement at crack frontur,o/n,o = radial displacement, value also affected by rust compaction of

internal boundaryX = depth of corrosion penetrationrs = volumetric ratio of rust to parent irontip/sp = total displacement and displacement owing to bar pullout at

top ofB specimensVr = volume of rust productsVs = volume of steel losscor = radial compressive strain of rustcr = cracking strain of concretef,eff = effective strain of the externally-bonded FRP sheetshr = shrinkage strainst,y = yield strain of stirrupsr = radial compressive strain of concrete


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= hoop tensile strain of concreteconf = hoop compressive strain due to confining pressure confr/sm = friction coefficient: for ribbed bar, for smooth barc = maximum normal pressure for cover splittingconf/f = confining pressure by transverse compressive stress field, by

FRP jacketn/n

res = maximum normal pressure due to frictional bond mechanismand residual value

shr = shrinkage stressst = confining pressure due to stirrups = factor that considers concrete ring response (1 for fully elastic

and 2 for fully plastic)

REFERENCES1. Li, C. Q., and Zheng, J. J., Propagation of Reinforcement Corrosion

in Concrete and its Effects on Structural Deterioration, Magazine of ConcreteResearch, V. 57, No. 5, 2005, pp. 261-271.

2. Lundgren, K., A Model for the Bond Between Corroded Reinforcementand Concrete, Proceedings from Bond in Concretefrom Research toStandards, Budapest, Hungry, 2002, pp. 35-42.

3. Auyeung, Y.; Balaguru, P.; and Chung, L., Bond Behavior of CorrodedReinforcement Bars, ACI Materials Journal, V. 97, No. 2, Mar.-Apr. 2000,pp. 214-220.

4. Hussein, N.; Yang, Y.; Kawai, K.; and Sato, R., Time DependentBond Behaviour of Corroded Bars, Proceedings from Bond in Concretefrom Research to Standards, Budapest, Hungry, 2002, pp. 166-173.

5. Cairns, J.; Du, Y.; and Johnston, M., Residual Bond Capacity ofCorroded Plain Surface Reinforcement, Proceedings from Bond in Concretefrom Research to Standards, Budapest, Hungry, 2002, pp. 129-136.

6. MacGregor, J., Reinforced Concrete: Mechanics and Design,Prentice-Hill Inc., 1997, 939 pp.

7. ACI Committee 318, Building Code Requirements for StructuralConcrete (ACI 318-02) and Commentary (318R-02), American ConcreteInstitute, Farmington Hills, MI, 2002, 443 pp.

8. Tastani, S., and Pantazopoulou, S., Experimental Evaluation of theDirect Tension-Pullout Bond Test, Proceedings from Bond in Concretefrom Research to Standards, Budapest, Hungary, 2002, pp. 268-276.

9. Martin-Perez, B., Service Life Modelling of R.C. Highway StructuresExposed to Chlorides, PhD thesis, Department of Civil Engineering,University of Toronto, Toronto, ON, Canada, 1998, 164 pp.

10. Pantazopoulou, S., and Papoulia, K., Modelling of Cover-CrackingDue to Reinforcement Corrosion in R.C. Structures, Journal of EngineeringMechanics, ASCE, V. 127, No. 4, 2001, pp. 342-351.

11. Li, C. Q., Reliability Based Service Life Prediction of CorrosionAffected Concrete Structures, Journal of Structural Engineering, ASCE,V. 130, No. 10, 2004, pp. 1570-1577.

12. Lura, P.; Plizzari, G.; and Riva, P., 3D Finite-Element Modelling ofSplitting Crack Propagation, Magazine of Concrete Research, V. 54,No. 6, 2002, pp. 481-493.

13. Malvar, J., Bond of Reinforcement under Controlled Confinement,ACI Materials Journal, V. 89, No. 6, Nov.-Dec. 1992, pp. 593-601.

14. Harajli, M. H., Bond Strengthening of Steel Bars Using ExternalFRP Confinement: Implications on the Static and Cyclic Response of R/CMembers, Proceedings of the 7th International Symposium on Fiber-Reinforced Polymer Reinforcement for Concrete Structures, SP-230, C.Shield, J. Busel, S. Walkup, and D. Gremel, eds., American ConcreteInstitute, Farmington Hills, MI, 2005, pp. 579-596.

15. Rodriguez, J.; Ortega, L.; Casal, J.; and Diez, J., AssessingStructural Conditions of Concrete Structures with Corroded Reinforcement,Proceedings from Concrete in the Service of Mankind: Concrete Repair,Rehabilitation and Protection, R. K. Dhir and M. R. Jones, eds., E&FNSpon, 1996, pp. 65-78.

16. Cabrera, J. G., Deterioration of Concrete Due to ReinforcementSteel Corrosion, Elsevier Cement and Concrete Composites, V. 8, 1996,pp. 47-59.

17. Stanish, K.; Hooton, R.; and Pantazopoulou, S., Corrosion Effectson Bond Strength in Reinforced Concrete, ACI Structural Journal, V. 96,No. 6, Nov.-Dec. 1999, pp. 915-921.

18. CEB Task Group 2.5, Bond of Reinforcement in Concrete, FIB Bulletin

10, International Federation for Concrete, Lausanne, Switzerland, 2000, 427 pp.19. Tepfers, R., Cracking of Concrete Cover Along Anchored

Deformed Reinforcing Bars, Magazine of Concrete Research, V. 31,No. 106, 1979, pp. 3-12.

20. CEB-FIP Model Code 1990, Design Code, Comit Euro-Internationaldu Bton, Thomas Telford Publications, London, UK, 1993, 437 pp.

21. Tastani, S., and Pantazopoulou, S., Recovery of Seismic Resistancein Corrosion-Damaged R.C. through FRP Jacketing, International Journalof Materials and Product Technology, V. 23, No. 3/4, 2005, pp. 389-415.

22. Eurocode 2, Design of Concrete Structures (EC-2), EuropeanCommittee for Standardization, Brussels, Belgium, 2002, 227 pp.


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APPENDIX

b)

highlighted cracks

non-corrosive zone (bond breaker)

a)

Fig. A1 - Cracking pattern at the bottom face of corroded specimens: a) Beam-end and b) Slab.

1


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2

0

5

10

15

20

25

0 10 20 30 4

30

: tip : sp

B3: X =9.6

cR

c

0

5

10

15

20

25

0 10 20 30 40 50 60 70 80

3030

B2: X =6.8 (%)

B0

cR

c

0

5

10

15

20

25

0 10 20 30 40 50 60 70 80

B1: X =10.9 (%)

Load(kN)

c

B0

cR

0

5

10

15

20

25

30

0 10 20 30 4

0

5

10

15

20

25

30

0 10 20 30 40 50 60 70 80

0

5

10

15

20

25

30

0 10 20 30 40 50 60 70 80

c

Load

(kN)

c

B0

B4: X =10.2 (%)

displ. (mm)

B0

c

cR

B5: X=2.96(%)

displ. (mm) disp

Fig. A2 - Experimental load displacement curves (B-group). Circles mark total cantilever

Triangles mark cantilever tip displacement due to slip of the bar at the support. c labels pre-repair

(series without connecting lines), R labels post-repair response (series with connecting lines) [1

mm = 0.0394 in.].


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3

0

30

60

90

120

150

0 3 6 9 12 15 18

Deflection (mm)

Load(kN)

a)

21

S1c

S2c

S1cR

S2cR

S0

b)

S3cc)

highlighted flexural crack

Fig. A3 - Failure of the a) control (S0) and b) corroded (S3c) specimen. c)

Performance of repaired slabs upon reloading [1 kN = 0.225 kip ; 1 mm = 0.0394 in.].


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32

Table A1: Comprehensive results of the experimental program.

Spec.

ID

Corrosion

Icorr() / X(%)

Preloading stage

P(kN)/ fb(MPa)

Reloading stage

P(kN) / fb (MPa)

Deflection*

y /ult(mm)

B0 0 / 0 27.83f/5.61 27.83

f/5.61 29.74 / 104.5

B1cR 12.1 / 10.9 14.85 / 3.19 22.26 / 4.48 23.58 / 54.35

B2cR 7.73 / 6.81 16.34 / 3.28 15.18 / 3.05 33.54 / 33.54

B3cR 10.71 / 9.58 17.11 / 3.44 9.87 / 1.98 21.92 / 21.92

B4c 11.35 / 10.18 6.98f/ 1.40 6.98

f/ 1.40 23.30 / 23.30

B5cR 3.42 / 2.96 23.27 / 4.69 22.26 / 4.48 37.61 / 39.63Beam-endspecimens

(Bgroup)

B6cR 4.27 / 3.71 20.78 / 4.18 13.16 / 2.64 29.81 / 29.81

S0 0 / 0 126.47

f

/ 7.07 126.47

f

/ 7.07 7.5 / 9.5

S1cR 12.49 / 4.60 78.57 / 4.37 119.57 / 6.69 5.71 / 10.82

S2cR 9.88 / 3.61 76.79 / 4.27 120.0 / 6.71 5.28 / 9.08Slabtests

(Sgroup)

S3cR 10.99 / 4.03 121.36f/ 6.79 121.36 f/ 6.79 ---

Note: 1 MPa = 0.145 ksi; 1 mm = 0.0394 in. fb is an average measure of bond stress. For values

obtained prior to anchorage failure (as in the cases of preload) the actual bond stress distribution is

not uniform.: Duration of corrosion process: 75 and 85 days for the B and the S group respectively. Mass loss

Ms was defined from the Faradays law and corrosion penetration Xfrom Ms=X(2-X)Ms.f: Failure load.*: For the B-group are given values from the tip displ. whereas for the S-group the middle deflectionat the corresponding loading of the 5th col.

Behavior of Corroded Bar Anchorages

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