ACI Structural Journal/July-August 2008 1

ACI Structural Journal, V. 105, No. 4, July-August 2008.MS No. S-2007-051.R1 received February 8, 2007, and reviewed under Institute

publication policies. Copyright © 2008, American Concrete Institute. All rights reserved,including the making of copies unless permission is obtained from the copyright proprietors.Pertinent discussion including author’s closure, if any, will be published in the May-June 2009 ACI Structural Journal if the discussion is received by January 1, 2009.

ACI STRUCTURAL JOURNAL TECHNICAL PAPER

Following previous studies, the authors describe an experimentalstudy performed to analyze the influence of the interface treatmenton the seismic behavior of columns strengthened by reinforcedconcrete (RC) jacketing to increase their ultimate bendingmoment. A numerical study, subsequently conducted to furtherinvestigate this subject, is also presented. It has been concludedthat, for undamaged columns with a bending moment/shear forceratio greater than 1.0, it is not necessary to consider any type ofinterface treatment before casting a RC jacket with a thickness lessthan 17.5% of the column width to obtain a monolithic behavior ofthe composite element.

Keywords: cyclic loading; interface; jacketing; seismic response;strengthening; surface preparation.

INTRODUCTIONReinforced concrete (RC) jacketing is a strengthening

technique most frequently used in seismic retrofitting.1 It hasbeen widely used after earthquakes in Mexico, Japan, theBalkans, and the U.S.2-4 To optimize the structural performanceof the composite element, it is essential to ensure its monolithicbehavior, which implies ensuring total adherence betweenthe original column and the added jacket. To fulfill thisobjective, the current practice consists of increasing theroughness of the interface surface, applying a bonding agentand, in some cases, steel connectors. Due to the reducedthickness of the jacket, a self-consolidating high-strengthgrout is usually adopted.

In all published experimental studies on this subject, thepreparation of the column surface before jacketing is alwaysreferred to.3,5-9 Nevertheless, a quantitative analysis of itsinfluence is never reported.

The authors decided to perform experimental studies toquantify the influence of four parameters on the bondstrength between concretes with different ages and differentcharacteristics. The parameters considered were: 1) theroughness of the interface surface; 2) the use of a bondingagent; 3) the added concrete mixture; and 4) the applicationof steel connectors. Slant shear tests and pushoff tests wereadopted to determine bond strength in shear and pulloff testswere used to assess bond strength in tension. It wasconcluded that: 1) between those adopted, sandblasting is thebest roughness treatment10; 2) the use of epoxy resins doesnot improve the interface strength if sandblasting is used11;3) adding a high-strength concrete (HSC) increases the inter-face strength12; and 4) the use of steel connectors doesnot significantly increase the interface debonding stress,although, after that, the shear stress is highly dependent onthe relation between the cross section area of steel connectorsand the area of the interface.13

Afterward, the authors conducted monotonic tests onseven column-footing models.14 These columns were

strengthened by RC jacketing after the interface surface hadbeen prepared according to the conclusions drawn from theresults of the experimental studies previously referred to. Itwas concluded that, for current undamaged columnssubjected to bending moment/shear force ratios greater than1.0 m (3.281 ft), a monolithic behavior of the compositeelement can be achieved, even without increasing theirsurface roughness or using bonding agents or applying steelconnectors, before strengthening by adding an RC jacketwith a thickness lesser than 17.5% of the column width. Inspite of that, it should be noted that, for other conditions suchas cyclic loading, RC short columns, or thicker jackets, theseconclusions may not apply.

The experimental study presented in this paper adds relevantinformation to the previous research conducted by theauthors, defining the response of the considered strengthenedcolumns to cyclic loading. In fact, this is particularly importantto predict if, for seismic loading, the jacketed column stillshows a monolithic behavior. The numerical simulation,conducted after the experimental study and also presented inthis paper, tests the hypothesis that, for short columns,debonding of the jacket may occur.

RESEARCH SIGNIFICANCEThe research presented in this paper has proven that, for

the conditions considered (undamaged reinforced concretecolumns with a bending moment/shear force ratio greaterthan 1.0), it is not necessary to previously prepare theinterface surface, namely, increasing its roughness, applyinga bonding agent, or eventually steel connectors, beforecasting RC jacketing with a thickness less than 17.5% of thecolumn width to achieve a monolithic behavior of thecomposite element subjected to cyclic loading. Thisachievement leads to: a) significant savings in expensivematerials, such as epoxy-based bonding agents; b) significantsavings in time-consuming operations (for example, theapplication of steel connectors); and c) avoidance of the useinadequate tools to increase the roughness of the columnsurface, such as jackhammers, that promote microcrackingof the concrete substrate.

EXPERIMENTAL INVESTIGATIONSeven column-footing models were built using concrete

with approximately 35 MPa (5076 psi) nominal compressivestrength (Table 1) measured with cubic specimens at 28 daysand steel with 520 MPa (75,420 psi) nominal yielding stress.

Title no. 105-S45

Reinforced Concrete Jacketing—Interface Influence on Cyclic Loading Responseby Eduardo N. B. S. Júlio and Fernando A. B. Branco

ACI Structural Journal/July-August 20082

The dimensions adopted for the column cross section andheight were 0.20 x 0.20 m2 (0.656 x 0.656 ft2) and 1.35 m(4.43 ft), respectively. The column was symmetricallyreinforced with three bars with 10 mm (0.394 in.) diametersat each face. The transverse reinforcement of the columnconsisted of 6 mm (0.236 in.) diameter stirrups spaced 150 mm(5.905 in). Strain gauges were bonded to each central bar,close to the footing, of the longitudinal reinforcement and onthe second stirrup from the bottom in opposite branches (Fig. 1).

Three models were considered to serve as reference: thefirst model (M1) was left unstrengthened; the second model(M2) was strengthened with a nonadherent jacket, materializedwith a thin, hard, greased layer placed on the interface, withthe objective of reaching the lower limit of the structuralbehavior of the composite model; and the third model (M3)was produced monolithically with the purpose of getting theupper limit of that behavior because debonding of the jacketwas expected to occur on the remaining models. Three othermodels were conceived with the interface surface prepared

according to the conclusions of the previous studiesperformed by the authors10-13: the fourth model (M4) wasstrengthened by jacketing without any interface treatment;the fifth model (M5) was strengthened by jacketing after itsinterface surface had been treated by sandblasting; and thesixth model (M6) was strengthened by jacketing after itsinterface surface had been prepared by sandblasting and steelconnectors had been applied. A seventh model (M7) wasconsidered, differing from the others by having beenstrengthened after application of the axial force, but identicalto M6 in respect to the preparation of the interface surface.With this model, the objective was to study the influence ofstrengthening columns with and without considering an activeshoring, the latter being the most frequent situation in practice.

Three months after casting the models, each column wasencased, using a commercial high-strength self-consolidatinggrout with approximately 80 MPa (11,603 psi) nominalcompressive strength (Table 1), measured with cubic specimensat 28 days and steel with 520 MPa (75,420 psi) nominalyielding stress. The dimensions adopted for the reinforcedconcrete jacket thickness and height were 35 mm (1.378 in.)and 0.90 m (2.953 ft), respectively. The longitudinalreinforcement of the jacket consisted of three bars with10 mm (0.394 in.) diameters at each face anchored to thefooting, with a commercial epoxy resin, in a predrilled hole of250 mm (9.842 in.) depth. The transverse reinforcement ofthe added jacket consisted of 6 mm (0.236 in.) diameterstirrups spaced 75 mm (2.953 in.) and out of phase withthose of the column because this is the most effectivegeometry to obtain a monolithic behavior of the strengthenedcolumn.15 Strain gauges were also bonded to each centralbar, close to the footing, of the longitudinal reinforcementand on the second stirrup from the bottom in oppositebranches (Fig. 1).

Twenty-eight days after being strengthened, the modelswere submitted to cyclic loading (Fig. 2). The loadingsystem consisted of a horizontal force varying according to apredefined displacement histogram and a constant axialforce of 170 kN (38,218 lbf). The horizontal force wasapplied with a hydraulic jack, positioned horizontally at 1.0 m(3.281 ft) from the column footing, with both ends hinged.The measured load was obtained from the difference betweenthe values read in two load cells, placed on opposite sites ofthe column top. The imposed horizontal displacement wasmeasured by a displacement transducer. The axial force was

Eduardo N. B. S. Júlio is an Assistant Professor at the University of Coimbra, Coimbra,Portugal; Vice Chairman of the Scientific Committee of the Civil Engineering Department;and Director of the MSc course on rehabilitation of the built environment. Hisresearch interests include strengthening and structural rehabilitation.

ACI member Fernando A. B. Branco is a Full Professor at the Technical Universityof Lisbon, Lisbon, Portugal; Head of the Construction Sector; and a Consultant formajor public works in Portugal. He is a member of ACI Committee 342, Evaluation ofConcrete Bridges and Bridge Elements. His research interests include design, rehabilita-tion, and construction technology of concrete structures.

Table 1—Description of models and compressive strength of concrete

Models Description

Compressive strength ofconcrete, MPa (psi)

Original column Added jacket

M1 Nonstrengthened column 34.89 (5060) —

M2 Column with nonadherent jacket 35.51 (5150) 83.71 (12,141)

M3 Column with monolithic jacket (cast simultaneously) 35.02 (5079) 35.02 (5079)

M4 Column jacketed without surface preparation 34.95 (5069) 78.25 (11,349)

M5 Column jacketed after surface preparation with sandblasting 35.06 (5085) 76.01 (11,024)

M6Column jacketed after surface preparation with sandblasting

and application of steel connectors 35.17 (5101) 79.96 (11,597)

M7Column jacketed after surface preparation with sandblasting and after loading axial force

35.40 (5134) 80.87 (11,729)

Fig. 1—Testing installation, instrumentation, and cross section.

Fig. 2—Cyclic test of Model M4.

ACI Structural Journal/July-August 2008 3

applied with a tubular system of two sets of two welded Uprofiles, connected with two prestressing tendons, tensionedwith a hydraulic jack. The corresponding value was measuredwith a load cell, placed between the top set of the welded Uprofiles and the hydraulic jack used to apply the axial force.

Taking into account the variety of solutions adopted bydifferent researchers16-21 and in the absence of standards forcyclic testing of reinforced concrete structures, the history ofimposed displacements was determined based on a recom-mendation of the European Convention for ConstructionalSteelwork (ECCS).22 According to ECCS,22 four cycles,each of increasing amplitude, were defined as 0.25δy, 0.50δy,0.75δy, and 1.00δy, followed by additional sets of threecycles each (also of increasing amplitude) of 2δy, 4δy, 6δy,and 8δy, with δy being the yielding displacement determinedwith the monotonic tests formerly performed by the authorson seven models with the same interface treatments.14

Because of the symmetry of the models, the amplitudeadopted in each cycle was the same for both positive andnegative displacements. The displacements were slowlyimposed with a velocity of 0.1 mm/s (0.00394 in./s). Toillustrate the loading history defined according to ECCS,22

Fig. 3 shows the one applied to Model M1.

RESULTS AND DISCUSSIONThe results analyzed were the cracking pattern and other

parameters obtained, directly or indirectly, from the hystereticdiagrams (Fig. 4): maximum load, capacity for dissipatingenergy, and damage level.

Cracking patternExcept for the model with the nonadherent jacket, Model M2,

cracking was not observed at the jacket top, the only crosssection where the boundary line of the column/jacket interfacewas visible. A significant horizontal crack was registerednear the footing in all models excluding, again, Model M2(Fig. 5). In this model, several smaller cracks appeareddistributed over a height approximately equal to the crosssection width (Fig. 5). Also, the observed concrete crushinglevel in the models strengthened with a high-strength jacketwas lower for the nonstrengthened model, Model M1, andthe monolithic model, Model M3, both built with normal-strength concrete.

Maximum loadAn analytical approach was performed to predict the

maximum load, assuming two hypotheses: total nonadherence

and perfect bonding of the jacket. For the first case, it wasassumed that the curvature radius of the original column andof the added jacket were the same at the support crosssection. For the second case, compatible strain diagrams ofthe original column and of the added jacket at the supportcross section were assumed. The ultimate concrete strainwas fixed at the extreme concrete fibers of the added jacket.The strain diagram was established iteratively until thecorresponding stress diagram presented a resultant force ofthe same value as the measured axial force. A parabola-rectangle stress diagram was adopted for concrete. With theresultant bending moment, the maximum force could beeasily determined. For Model M7, strengthened after theaxial force had been applied, the procedure adopted todetermine the theoretical maximum force was adapted totake into account an initial strain state due to that load.

The comparison between experimental data (Table 2) andanalytical values (Table 3) leads to the conclusion,confirmed by visual inspection, that there was no jacketdebonding in any model, except for Model M2. In fact,considering perfect bonding, the relative error between theexperimental and the theoretical value varied from –0.5% to+3.5% (Table 3) except for Model M2. Considering total

Fig. 3—Adopted loading history.

Table 2—Peaks of horizontal load (kN)in dark background

Cycle δ/δy*

M1 M2 M3 M4 M5 M6 M7

1+0.25 +17.80 +25.06 +35.92 +20.02 +33.11 +39.06 +25.26

–0.25 –14.46 –27.42 –34.03 –24.27 –22.18 –38.41 –32.32

2+0.50 +23.10 +40.24 +47.24 +31.67 +48.49 +53.33 +40.24

–0.50 –19.43 –39.33 –42.20 –40.63 –37.89 –53.85 –47.50

3+0.75 +27.48 +52.08 +56.01 +43.64 +65.50 +63.99 +51.23

–0.75 –22.84 –46.26 –54.37 –50.65 –48.03 –64.65 –59.48

4+1.00 +31.02 +59.54 +62.82 +57.38 +70.67 +72.83 +59.09

–1.00 –25.85 –51.23 –61.31 –60.07 –57.91 –73.74 –68.64

5+2 +34.02 +66.94 +72.70 +73.74 +77.41 +79.11 +69.49

–2 –31.28 –55.55 –66.55 –70.60 –75.05 –77.60 –79.37

6+2 +32.39 +62.68 +67.66 +70.99 +72.43 +74.14 +63.34

–2 –30.56 –52.28 –68.51 –68.44 –72.04 –74.40 –77.41

7+2 +32.39 +61.44 +65.89 +70.27 +73.15 +72.30 +61.96

–2 –30.03 –50.78 –67.59 –65.43 –67.33 –72.76 –75.77

8+4 +33.50 +68.64 +71.45 +80.29 +80.61 +79.96 +72.43

–4 –32.26 –60.26 –73.55 –75.25 –78.85 –76.88 –82.44

9+4 +32.06 +64.52 +66.54 +69.49 +74.53 +74.99 +64.39

–4 –28.73 –53.66 –70.80 –72.63 –75.31 –76.10 –78.98

10+4 +30.95 +58.43 +63.73 +70.54 +72.70 +71.65 +62.55

–4 –23.82 –49.47 –68.31 –71.06 –73.68 –72.89 –77.28

11+6 +30.49 +64.91 +65.50 +77.80 +77.60 +73.74 +68.64

–6 –19.76 –50.32 –65.63 –74.66 –76.29 –76.30 –80.35

12+6 +24.34 +60.85 +61.24 +70.80 +72.70 +68.64 +58.50

–6 –15.83 –41.22 –60.07 –71.71 –73.42 –69.29 –76.95

13+6 +20.87 +50.91 +57.06 +68.97 +68.44 +65.96 +56.27

–6 –13.15 –38.28 –53.13 –69.49 –70.80 –62.62 –73.68

14+8 – +62.16 +59.94 +72.89 +73.42 +49.93 +56.80

–8 – –39.06 –44.10 –72.17 –72.37 –48.81 –74.66

15+8 – +57.58 +49.99 +66.81 +68.77 +39.65 +49.47

–8 – –34.55 –36.45 –67.79 –69.55 –42.27 –52.48

16+8 – +51.43 +39.98 +62.95 +61.18 +38.15 +38.80

–8 – –31.67 –28.33 –65.83 –67.33 –36.19 –47.05*δ/δy = amplitude ratio between positive or negative cycle and yielding displacement.

4 ACI Structural Journal/July-August 2008

nonadherence, the relative error between the experimentaland the theoretical value varied from –15.5% to –27.6%(Table 3) for the same models, confirming the conclusionthat no slippage between the original column and the addedjacket occurred in these models. For Model M2, the experi-mental and theoretical values confirmed that its behaviorwas closer to the theoretical behavior considering totalnonadherence of the jacket, although some friction waspresent. This was probably due to the fact that the axial loadwas applied after strengthening of the column, which caused anexpansion of the cross section due to Poisson’s effect, therebymobilizing some friction between the column and the jacket.

It was observed that strengthening the column with theaxial load already applied had no relevant influence on testresults and that the resistance of the strengthened models

Fig. 4—Hysteretic diagrams of Models M1 to M7.

Table 3—Experimental and theoretical values* of maximum load of each level

Experimental values

Theoretical values

Nonadherent jacket Monolithic cross section

Models Maximum load, kN (lbf)

Maximum load, kN (lbf)

Error, %

Maximum load, kN (lbf)

Error, %

M1 34.0 (7644) — — 33.0 (7419) —

M2 68.6 (15,422) 64.8 (14,568) –5.9 82.0 (18,434) 16.3

M3 73.6 (16,546) 63.7 (14,320) –15.5 74.9 (16,838) 1.7

M4 80.3 (18,052) 65.5 (14,725) –22.6 83.1 (18,682) 3.4

M5 80.6 (18,120) 65.5 (14,725) –23.1 83.0 (18,659) 2.9

M6 80.0 (17,985) 65.4 (14,703) –22.3 82.9 (18,637) 3.5

M7 82.4 (18,524) 64.6 (14,523) –27.6 82.0 (18,434) –0.5*Assuming nonadherent jacket and monolithic cross section.

ACI Structural Journal/July-August 2008 5

was considerably higher than that of the original columnand slightly higher than that of the monolithic model. Thelast remark can be explained taking into account thatdebonding of the jacket did not occur and that the compressivestrength of the added concrete was on the order of 80 MPa(11,603 psi) instead of approximately 35 MPa (5076 psi) ofthe original concrete.

Normalized dissipated energyTo account for the differences between the seven models

when comparing their behavior in terms of capacity to dissipateenergy, it was decided to normalize this parameter bydividing the energy dissipated in each cycle by the energytheoretically dissipated, in a cycle of equal amplitude,assuming an elastic-perfectly plastic behavior of models.The expression from ECCS22 for cyclic tests of steel structuralelements was adopted

(1)

where η is the normalized dissipated energy in a given cycle; is the energy dissipated in this cycle; Fyt

+ is thevalue of the horizontal force needed to yield all the reinforcingbars of the model, acting in the positive direction; Fyt

– is thevalue of the horizontal force needed to yield all the reinforcingbars of the model, acting in the negative direction; δ yt

+ and δ yt–

are the corresponding horizontal displacements of thesection where the force is applied; and δ+ and δ– are themaximum positive and negative displacements in this cycle.

To illustrate this computation, Fig. 6 shows the experimentalcurve and the corresponding theoretical curve assumingelasto-plastic behavior for the 15th cycle of Model M5. Thearea defined by the first curve represents the energy dissipatedin this cycle by the model and the area defined by the secondcurve represents the energy theoretically dissipated.

In Fig. 7, the normalized dissipated energy in each cycle ispresented for all models. Although quantitatively different,all strengthened models, including the nonadherent Model M2,displayed a qualitatively similar behavior. A decrease of thisparameter is observed from the fifth to the sixth cycle and asmaller decrease from this to the seventh cycle. This tendencyis observed in all sets of cycles of constant amplitude. The onlysignificant conclusion that can be drawn from this analysis isthat the normalized dissipated energy of the nonstrengthenedModel M1 was lower than the corresponding values of theother models.

η

F t( ) δd0

t

∫Fyt

+δ

+δyt

+– δ–

δyt––+( ) Fyt

–δ

–δyt

–– δ+

δyt+–+( )+

--------------------------------------------------------------------------------------------------------------------=

F t( ) δd0

t

∫

Damage indexThere are several types of damage indexes that can generally

be divided into two groups: 1) damage indexes based on thestrength; and 2) damage indexes based on the response.23

The damage indexes of the first group are inconvenient asthey need to be calibrated based on the damages observed,using a large database. Regarding the damage indexes of thesecond group, several authors proposed damage indexesbased on: (a) maximum deformation; (b) cumulative damages;and (c) a combination of the two previous parameters.23-28

The damage index, selected for its simplicity to evaluatethe damages in the models, was defined as a ratio betweenthe initial stiffness, taken as the secant stiffness given bythe origin and the positive peak of the first cycle, and thestiffness in each cycle, taken as the secant stiffness givenby the origin and the positive peak of the respective cycle.In all models, a degradation of the secant stiffness fromcycle to cycle was observed (Fig. 8). Except for thenonstrengthened Model M1, which presented higher valuesthan the remaining models, no significant differences wereobserved considering this parameter.

Conclusions of experimental investigationThe analysis of the considered parameters indicates that all

models behaved monolithically independently of the adoptedinterface preparation method, with the exception of Model M2,in which the nonadherence of the jacket was intentional. Eventhis model presented a structural behavior between that of the

Fig. 5—Cracking pattern of Models M1 and M2.

Fig. 6—Hysteretic curves from fifth to 15th cycle of Model M5and of ideal elasto-plastic model.

Fig. 7—Normalized dissipation of energy per cycle by allmodels.

6 ACI Structural Journal/July-August 2008

theoretically perfect frictionless model and the theoreticallyperfect adherent model, although closer to the first. Theseconclusions are in agreement with those drawn in theprevious study conducted by the authors with identicalmodels that had been subjected to monotonic loading.14

The reason why debonding of the jacket was not observedin any of the models, apart from Model M2, is probablyrelated to the fact that the compressive strength of the addedconcrete was reached earlier than the bond strength of theinterface. This means that for shorter columns strengthenedby jacketing, that is, for columns with a lower bending

moment/shear force ratio, debonding of the jacket mayoccur. It was decided to conduct a numerical simulation todetermine the validity of this hypothesis.

NUMERICAL ANALYSISThe numerical analysis was performed using a finite

element program. Isoparametric finite elements, pentahedralwith six or 15 nodes and hexahedral with eight or 20 nodes,were used to simulate concrete, considering the Mohr-Coulombor Drucker-Prager failure criteria. Linear finite elementswith two or three nodes were used to simulate reinforcing steel,considering the Von Mises yield criterion. Interface finiteelements, triangular with six or 12 nodes and rectangularwith eight or 16 nodes, were used to simulate the interface,considering a delamination model. To solve the set ofnonlinear equations, a combined incremental-iterative Newton-Raphson method was used.

Pulloff tests performed previously10 were numericallyreproduced to calibrate the characteristics of the delaminationmodel adopted for the interface. Taking into account thesymmetry of the pulloff specimen, only one quarter of thespecimen was modeled. The support conditions consisted ofrestraining all degrees of freedom of the nodes correspondingto the zone of application of the support ring of the experi-mental device used. The loading consisted of an imposeddisplacement at the nodes corresponding to the top of thecore where, in the experimental test, a steel disc was epoxybonded to which the tension force was applied (Fig. 9). Slantshear tests also previously performed10 were also numericallyreproduced for the same purposes.

Afterward, monotonic tests performed with columnsstrengthened by jacketing14 were numerically reproduced.Due to the model symmetry, only one-half of the column wasmodeled. The support conditions consisted of restraining alldegrees of freedom of the base nodes and the displacementnormal to the bending plane of all the nodes in the symmetryplane. The loading, in the first 10 increments, consisted ofgradually applying the axial force, as a uniformly distributedload on the column top, as performed experimentally. The

Fig. 9—Adhesive failure of: (a) pulloff specimen; and (b)corresponding numerical model.

Fig. 8—Damage index for each cycle of each model.

Fig. 10—Load versus displacement of experimental andnumerical tests of Models M2 and M6.

Fig. 11—Distribution of vertical stresses in original modeland in model 50% shorter.

ACI Structural Journal/July-August 2008 7

axial force was kept constant in the following increments andan imposed displacement was applied to the points at 1.0 m(3.281 ft) above the base, as in the experimental tests, until avalue of 20 mm (0.787 in.) was reached to clearly include theplastic phase.

The monotonic tests14 of the nonstrengthened column, themonolithic column, the column strengthened withouttreating the interface, the column strengthened with theinterface prepared with sandblasting, and the column withthe nonadherent jacket were numerically modeled. Severalmeshes were defined and results compared. The best meshwas selected considering the accuracy of the numericalresults, by comparison with the corresponding experimentalresults, and also the time needed for the calculation. As anexample, in Fig. 10, the results of the experimental tests and ofthe corresponding numerical simulations of Models M2 and M6are plotted, that is, the nonadherent model and the model with theperfectly bonded jacket, respectively. It can be seen that thenumerical models predict higher strength and stiffness than thecorresponding experimental models, although, qualitatively, therelative difference in behavior is the same.

Based on the numerical model of the column strengthenedwithout treatment of the interface surface, similar numericalmodels were analyzed with 90, 80, 70, 60, and 50% of theheight of the experimental model. Results confirmed thehypothesis that debonding of the jacket may occur for shortercolumns strengthened without treatment of the interfacesurface. As an example, Fig. 11 shows the vertical stressdistribution in the original model and in the 50% shortermodel, corresponding to bending moment/shear force ratiosof 1 and 0.5 m (3.281 and 1.640 ft), respectively. In theneighborhood of the base cross section, in the first case, amonolithic distribution of vertical stresses is observed; andin the second case, a transition of vertical stresses fromcompression to tension in the jacket and, again, compressionin the column near the jacket.

CONCLUSIONSBased on the experimental program, it can be stated that,

for a sound column with a bending moment/shear force ratioof 1.0 m (3.281 ft) or greater subjected to cyclic loading, amonolithic behavior can be obtained without increasing theroughness of the interface surface or using bonding agents orapplying steel connectors before strengthening it with an RCjacket with a thickness less than 17.5% of the originalcolumn width, which is in agreement with the conclusionsdrawn previously from monotonic tests.14 From the numericalstudy performed subsequently, it can be concluded that, fora bending moment/shear force ratio lower than 1.0 m (3.281 ft),debonding of the jacket may occur without treatment of theinterface surface.

ACKOWLEDGMENTSWe are grateful to Sika, Hilti, Betão Liz, Dywidag, Pregaia, Cimpor, and

Secil for their collaboration in this research project.

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7. Alcocer, S., and Jirsa, J., “Assessment of the Response of ReinforcedConcrete Frame Connections Redesigned by Jacketing,” Proceedings of theFourth U.S. National Conference on Earthquake Engineering, V. 3, 1990,pp. 295-304.

8. Gomes, A., and Appleton, J., “Experimental Tests of StrengthenedReinforced Concrete Columns Subjected to Cyclic Loading,” Revista Portu-guesa de Engenharia de Estruturas, V. 38, 1994, pp. 19-29. (in Portuguese)

9. Stoppenhagen, D. R.; Jirsa, J. O.; and Wyllie, L. A., “Seismic Repairand Strengthening of a Severely Damaged Concrete Frame,” ACI StructuralJournal, V. 92, No. 2, Mar.-Apr. 1995, pp. 177-187.

10. Júlio, E. S.; Branco, F.; and Silva, V. D., “Concrete-to-ConcreteBond Strength: Influence of the Roughness of the Substrate Surface,”Construction and Building Materials, V. 18, No. 9, 2004, pp. 675-681.

11. Júlio, E. S.; Branco, F.; and Silva, V. D., “Concrete-to-ConcreteBond Strength: Influence of an Epoxy-Based Bonding Agent on aRoughened Substrate Surface,” Magazine of Concrete Research, V. 57,No. 8, 2005, pp. 463-468.

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