ACI Structural Journal/September-October 2007 521

ACI Structural Journal, V. 104, No. 5, September-October 2007.MS No. S-2006-036.R2 received October 11, 2006, and reviewed under Institute

publication policies. Copyright © 2007, American Concrete Institute. All rights reserved,including the making of copies unless permission is obtained from the copyrightproprietors. Pertinent discussion including author’s closure, if any, will be published in theJuly-August 2008 ACI Structural Journal if the discussion is received by March 1, 2008.

ACI STRUCTURAL JOURNAL TECHNICAL PAPER

Columns with plain vertical bars and no detailing for earthquakeresistance are rehabilitated with shotcrete jackets connected to theold column through various means and subjected to cyclic uniaxiallateral loading up to ultimate conditions to investigate the effect ofdifferent means of connection at the interface on the effectivenessof the jacket. These test results are supplemented with data fromthe literature for different means of connection of the jacket to theold column to develop simple rules and expressions for thecalculation of the yield moment, the drift at yielding, the secant-to-yield stiffness, and the ultimate drift in cyclic loading of jacketedcolumns with or without lap splicing at the base of the old column.These rules give the yield moment, the drift at yielding, the secant-to-yield stiffness, and the ultimate drift in cyclic loading ofjacketed columns as a ratio of the corresponding quantities of anequivalent monolithic member.

Keywords: column(s); concrete jacket; cyclic loading; deformation; seis-mic; stiffness; strength; test.

INTRODUCTIONConcrete jacketing1 of members is widely used in seismic

rehabilitation of old buildings.2 Past experimental work onthe cyclic behavior of reinforced concrete (RC)-jacketedmembers is limited.3-21 With the exception of a recentcomparative monotonic testing program of jacketed columnswith different levels and means of connection at the interface,12

the influence of positive measures of connection between theold concrete and the jacket has not been systematically studied.

Guidance documents for concrete jackets, as well asconstruction practice, include measures for shear connectionof the old and the new concrete at the interface. It is oftenrecommended to expose the corner bars of the old column andconnect them to the corner bars of the jacket through U-shapedsteel inserts, lap-welded to both bars. A positive connection ofsteels of different grades (and hence different chemicalcompositions), however, is considered conducive to reinforcingbar corrosion, not to mention that reinforcing bars of highcarbon equivalent are normally nonweldable. Other commonconnection measures include roughening of the surface of theold column and/or provision of dowels over this surface.There is some concern, though, that dowels or welded steelinserts in the plastic hinge zone may trigger crushing anddisintegration of the jacket concrete around them.

Welded steel inserts between the new and the old cornerbars, or dowels at the interface, were considered essential inthe 1970s and 1980s to transfer, through shear, part of thecolumn gravity load from the old column to the jacket. Theconcern with gravity load capacity was partly due to the factthat concrete jackets were then mainly, if not exclusively,applied to heavily damaged columns, often with a partiallydisintegrated core. This concern is also reflected in pastrecommendations to partially unload the column beforejacketing, by propping up beams framing into the column via

wedges or even jacks. For undamaged or moderatelydamaged columns, however, concern about sharing the axialload between the jacket and the old column does not seemwarranted. Moreover, recent experimental evidence22 hasshown that concrete columns subjected to large post-ultimate drifts that heavily damage the concrete core retain alarge part of their gravity load capacity. So, given thecollateral negative effects of welded steel inserts or thepreviously mentioned dowels and the direct or indirect costsof such measures, it is worthwhile to investigate theeffectiveness of friction alone for the shear transfer at theinterface. Friction is enhanced by compressive stresses thatbuild up normal to the interface, as the old member restrainsshrinkage of the concrete jacket in the radial andcircumferential directions. Such restraint induces confinementstresses in the old column, along with radial compression inthe jacket and circumferential tension both in the jacket andin its perimeter tie. If, indeed, there is significant shrinkage-induced friction at the interface and confinementenhancement in the old column, shotcrete or normal cast-in-place concrete has an advantage in this respect as jacketmaterials, compared with more expensive and difficult toapply nonshrinking mortars.

Friction is enhanced if the interface is rough. The 2:3scale, two-story, one-bay frame tested in,13 with sizeableconcrete jackets constructed around the heavily damagedcolumns of the original frame with the interface justroughened, sustained story drifts of 1.25% (translated to adrift ratio of the clear column length over 4%, if drifts aredue to the columns alone) without loss of column lateralforce resistance and with apparent monolithic behavior ofthe jacketed column. It is not clear, however, whether thecostly artificial roughening of the interface is essential. Forexample, monolithic behavior of the jacketed columns andbeams was apparently achieved in the three-dimensional(3D) subassemblages tested in References 14 and 15 withoutany positive measures to enhance shear transfer at theinterface. The recent experimental and analytical study of sixjacketed concrete columns in Reference 10 concluded thatdifferent means and levels of shear connection at theinterface do not play an important role for the monotonicbehavior up to yielding and beyond.

More experimental work is needed to clarify the effect ofthe shear connection at the interface on the cyclic behavior andespecially on the cyclic deformation capacity. So, this papersupplements the available test results through an experimental

Title no. 104-S49

Strength, Stiffness, and Cyclic Deformation Capacity of Concrete Jacketed Membersby Stathis N. Bousias, Dionysis Biskinis, Michael N. Fardis, and Alexis-Loukas Spathis

ACI Structural Journal/September-October 2007522

program on rectangular RC-jacketed columns, focused on theimpact of measures taken to enhance the shear transfer at theinterface of the old and the new concrete. As the present testsare limited, they are supplemented with monotonic or cyclicexperimental data from the literature3-12,16-18 with variousbonding measures between the old and the new concrete(usually not applied in the same testing program), to drawmore robust conclusions about the effect of bonding at theinterface. This is done by expressing key properties of thejacketed column (notably, the yield moment, the yield drift,the secant stiffness at incipient yielding, and the ultimate drift)as ratios to the corresponding quantities of an equivalentmonolithic member. The conclusion is that this ratio does notsystematically depend on the type of measures taken toenhance the shear transfer at the interface of the old and thenew concrete. This allows development of simple rules andexpressions for the calculation of key properties of columnswith or without lap splices at floor level, after theirrehabilitation via concrete jacketing. The rules—publishedherein for the first time with full documentation—have beenadopted in the part of Eurocode 8 on seismic assessment andrehabilitation of buildings.23

RESEARCH SIGNIFICANCEConcrete jacketing is widely used in seismic rehabilitation

of old buildings. Despite the benefits and cost-effectiveness ofRC-jacketing, cyclic test results on RC-jacketed members arelimited. In this paper some commonly applied, labor-intensivemeasures to enhance the connection between the old and thenew concrete are experimentally studied. The present testresults are supplemented with data from the literature todevelop simple rules for the calculation of the moment

capacity, the effective stiffness at incipient yielding, and theultimate drift of members, with or without lap splices in theplastic hinge region, rehabilitated through concrete jackets.

EXPERIMENTAL PROGRAM: SPECIMENS, MATERIALS, AND TEST SETUP

The experimental program comprises five jacketedspecimens with a 250 mm (10 in.) square original column(Fig. 1) and a companion monolithic column. The continuousvertical bars in the original column are plain (used inpractically all buildings in the seismic regions of Europe untilthe 1970s), have a diameter of dbL = 14 mm (0.55 in.), yieldstress fy = 313 MPa (45.5 ksi), tensile strength of ft = 442 MPa(64 ksi) (average from three coupons), and 180-degree hooksat both ends. The internal diameter of the hooks is 2.5 bardiameters. The bar extends straight past the end of the curvedpart of the hook by 40 mm (1.6 in.). The stirrups of the originalcolumn are also plain, with a diameter of dbh = 8 mm (0.3 in.) atsh = 200 mm (7.9 in.) centers and have a 135-degree hook at oneend and a 90-degree hook at the other. Their yield and ultimatestress are fyw = 425 MPa (61.5 ksi) and ftw = 596 MPa (86.5 ksi)(average from three coupons). The concrete cylindrical strengthof the original column at testing fc′ is listed in Table 1.

The columns are tested as simple cantilevers, fixed into aheavily reinforced 0.6 m (23.5 in.) deep footing (Fig. 1). Thelateral load is applied 1.6 m (63 in.) above the top of thefooting. The cyclic behavior of the columns is dominated byflexure, as the ratio of shear span Ls to section depth in thedirection of testing h is equal to 4 or 6.4, with or withoutjacketing, respectively.

The jacket is 75 mm (3 in.) thick, which is almost theminimum that can provide proper cover to the perimeter tieand allow a 135-degree hook at its ends. It consists ofshotcrete and stops approximately 250 mm (10 in.) below thepoint where the lateral load is applied. The jacket isreinforced longitudinally with four 20 mm (0.8 in.) deformedbars having yield stress fy = 487 MPa (70.5 ksi), which wereembedded in the footing when the original column was cast(Fig. 2), anchored with 90-degree hooks at the bottom.Jacket transverse reinforcement consists of a single 10 mm(0.4 in.) perimeter tie with a yield stress fyw = 599 MPa(87 ksi) at 100 mm (4 in.) centers. The measured cylindricalcompressive strength of the jacket and the effective depth dof the jacketed section are listed in Table 1.

The jacketed columns are tested to study the effect ofmeasures that enhance the shear transfer at the interface ofthe old and the new concrete. Their identification in Table 1starts with Q-RC.

ACI member Stathis N. Bousias is an Assistant Professor in the Department of CivilEngineering, University of Patras, Patras, Greece. He received his MSc from CaseWestern Reserve University, Cleveland, Ohio, and his PhD from the University ofPatras. His research interests include the design and experimental response of newand rehabilitated concrete structures and structural testing techniques.

Dionysis Biskinis is a PhD Candidate in the Department of Civil Engineering,University of Patras. He received a diploma in civil engineering and an MS from theUniversity of Patras.

ACI member Michael N. Fardis is a Professor of design of concrete structures at theUniversity of Patras. He received his MSc and PhD from the Massachusetts Instituteof Technology, Cambridge, Mass. He received the ACI Wason Medal for MaterialsResearch in 1993.

Alexis-Loukas Spathis received a diploma and an MSc in civil engineering from theUniversity of Patras. His research interests include the experimental response ofconcrete structures.

Fig. 1—Cross section of original column and of jacketedcolumn. (1 in. = 25.4 mm)

Fig. 2—Connection of jacket to old member: (a) newreinforcement connected to old corner bars via welded U-bars; (b) dowels attached to old column; and (c) oldconcrete surface roughening.

ACI Structural Journal/September-October 2007 523

In Column Q-RCW each new corner bar was connected tothe corresponding old one by welding both of them to 16 mm(0.65 in.) diameter, 400 mm (16 in.) long deformed reinforcingbars bent into a U-shape (Fig. 2(a)). The bent-up legs of thesebars ended with a flat 70 mm (2.75 in.) long part. The cornerbars of the old column were exposed and welded to the 70 mm(2.75 in.) long flat central part of the U-bar. The two 70 mm(2.75 in.) long ends of the U-bar were welded to the corner barof the jacket. (Note that 70 mm [2.75 in.] is five diameters ofthe smaller of the two connected bars and is according tonormal practice.) Two U-bars were connected to each cornerbar, with their center 350 mm (14 in.) or 1050 mm (41 in.)above the top of the footing. Note that welding of the cornerjacket bars to the lower-most U-bar provides certain anti-buckling restraint in the plastic hinge.

In Column Q-RCD, three dowels were driven into each sideof the old column, at distances of 200, 650, and 1100 mm (8, 25,and 43 in.) from the top of the footing (Fig. 2(b)). The dowelwas a 16 mm (0.65 in.) diameter deformed bar, epoxy-groutedinto a 100 mm (4 in.) deep hole drilled in the old concrete. Itprotruded into the 75 mm (3 in.) thick jacket for a straightlength of 50 mm (2 in.) and was then bent as a 90-degree 100 mm(4 in.) long horizontal hook.

In Column Q-RCR, the full lateral surface of the oldcolumn was roughened using a pneumatic chipping

device (Fig. 2(c)), until the hardened cement paste and thefine aggregates were removed and the coarse aggregateswere exposed. In Column Q-RCRD, roughening of theinterface, as with Column Q-RCR, is combined withdowels as with Column Q-RCD. In Column Q-RC, nospecial measures were taken to connect the jacket to theold concrete.

A monolithic column, Q-RCM, was also used as controlspecimen. It had the same external dimensions (400 mm[16 in.] square) and the same reinforcement as the jacket(four 20 mm [0.8 in.] vertical bars, 10 mm [0.4 in.] closedties at 100 mm [4 in.] centers), but it did not have the four14 mm [0.55 in.] plain vertical bars and the correspondingstirrups of the original 250 mm [10 in.] square column ofthe five jacketed specimens.

Unidirectional deflection cycles were applied withamplitudes increasing by 5 mm (0.2 in.) from one cycle tothe next. Testing continued (with increasing deflectionamplitude) until and beyond the column ultimate deformation,conventionally defined as that where peak resistance in acycle drops below 80% of the maximum recorded lateralresistance of the column. Often the conventionally definedultimate deformation was reached when one or more verticalbars fractured.

Table 1—Jacketed columns: test parameters and key test results

Specimen and connection of jacket

with old columnd,

mm (in.)

fc′, MPa (ksi)

ν = N/bhfc′ jacketed column

Yield moment

My ,kNm

(ft-kip)

Calculated yield moment My,calc , kNm

(ft-kip)

Drift at yielding θy, %

Yield curvature ϕy , 1/m

(1/ft)

Fixed-end rotation at yielding

(rad)

Ultimate drift θu,

%

Ultimate curvature ϕu , 1/m

(1/ft)

Fixed-end rotation at ultimate,

rad

Main features of the behavior and of failure mode

Original column Jacket

Q-RCW WeldedU-bars

355 (14)

22.9 (3.32)

28.7 (4.16) 0.130 210

(154.9)209.5

(154.9) 1.15 0.013 (0.0039) 0.0028 5.65 —* —*

All four jacket bars buckled and two ruptured

Q-RCD Dowels 355 (14)

27.4 (3.97)

55.3 (8.02) 0.085 240

(177.0)231.8

(171.0) 1.15 0.012 (0.0036) 0.0025 6.25 0.17

(0.052) 0.017

All four jacket bars buckled and one ruptured; minor splitting cracks along corner bars

Q-RCR Roughened 355 (14)

27.7 (4.0)

55.3 (8.02) 0.090 260

(191.8)238.0

(175.5) 1.20 0.015 (0.0046) 0.0025 5.65 0.13

(0.039) 0.027

Full disintegration near base; buckling of four jacket bars (one broke) and of interior bars (in old column); ties opened; partial height bond splitting/spalling along two corner bars

Q-RCRD Roughened + dowels

355 (14)

26.3 (3.81)

53.2 (7.71) 0.094 245

(180.7)238.5

(175.9) 1.00 0.015 (0.0046) 0.0025 5.30 0.14

(0.043) 0.020

All four jacket bars buckled and one ruptured

Q-RC Notreatment

355 (14)

26.3 (3.81)

55.3 (8.02) 0.080 235

(173.3)227.3

(167.6) 1.10 0.014 (0.0043) 0.002 5.30 0.11

(0.033) 0.017

Serious disintegrationnear base; lower-most tie opened; one interior bar (old column) and two jacket bars buckled; one bar ruptured

Q-RCM Monolithic 350 (13.8)

30.6 (4.44) — 0.180 251

(185.1)218.5

(161.2) 1.00 0.013 (0.0039) 0.0018 5.30 —* —*

Concrete crushed and all four bars buckled at base; one bar ruptured

*Measurements insufficient for estimation of curvature and fixed-end rotation.

524 ACI Structural Journal/September-October 2007

A jack at the top of the column maintained the axial loadat approximately 800 kN (180 kips), representing the axialforce in an internal column of a typical five-story building.Table 1 gives the mean value of the axial load during the test,normalized to the product of the external dimensions b and hof the jacketed section and to the fc′ value of the jacket, as ν= N/bhfc′ . The jack applying the axial load was manuallycontrolled and readjusted as the test progressed to keep theload constant. It acted against vertical rods connected to thelaboratory strong floor through a hinge. Bending momentsreported in Table 1 include the P-Δ′ moment, which is equalto the axial load times the ratio of the distance of the hingefrom the top of the footing (0.425 m [16.73 in.]) to thedistance of the hinge from the point where the lateral loadwas applied (2.025 m [79.72 in.]).

Linear variable differential transformers placed betweenthe top of the footing and two levels of the column above thebase (at 50 or 100% of the section depth of the originalcolumn, that is, at 125 mm [5 in.] or 250 mm [10 in.]) at oppositesides of the column in the loading direction (Fig. 1(d))provide the relative rotation and the mean axial strain of thecolumn between the top of the footing and each one of thetwo instrumented sections, including the effect of slippageand pull-out of vertical column bars from the base.

TEST RESULTS AND DISCUSSIONYield moment and deformations of jacketed columns at yielding and ultimate

Columns 7, 9, 10, and 11 in Table 1 give key test results atyielding of the column: the moment (including P-Δ effects)My, the drift ratio θy, the curvature at the base φy, and therotation of the column end section with respect to the baseowing to bar pullout from it (fixed-end rotation). Columns12, 13, and 14 in Table 1 give the ultimate drift ratio θu, theultimate curvature at the base φu, and the fixed end rotationdue to bar pull-out from the base, again at ultimateconditions. The column drift ratio (or chord rotation over theshear span, Ls = 1.6 m [63 in.]) is the deflection at the pointof lateral loading, divided by Ls. Yielding is identified witha distinct reduction in the slope of the moment-deflectioncurve and is close to the corner point of a bilinear envelopethat can be fitted to the ascending part of the force-driftresponse, that is, up to peak resistance (the chosen yieldpoints for the six present tests are shown in Fig. 3 with anasterisk). The curvature at the base of the column and thefixed-end rotation due to bar pullout were estimated from therotations of the two instrumented sections, assuming that thecurvature is constant within the lowest 250 mm (10 in.) ofthe column. Note, though, that these values are subject tosignificant uncertainty due to the assumptions involved intheir estimation and noise in the measurements.

Force-deflection loops of jacketed columnsFigure 3 compares the force-drift loops of the five jacketed

columns with those of the monolithic column, Q-RCM, thathas the same external dimensions and the samereinforcement as the jacket alone. In view of the variation ofsome specimen parameters (notably the values of fc′ ), fewsystematic differences are observed. The monolithic columnhas higher yield moment and stiffness at incipient yieldingthan practically all jacketed columns, although it isreinforced with the vertical bars of the jacket alone. Theultimate deformation of the jacketed columns is not less thanthat of the monolithic column. In fact, it was found to be

larger, if the shear connection at the interface is enhanced viaU-bars welded between the corner bars of the old columnand the jacket, or dowels, or roughening of the old columnsurface. Note that when dowels and surface roughening wereapplied together, the beneficial effect that each oneseparately has on deformation capacity was found to be lost.Due to the very small number of the present tests, however,conclusions about the effect of the connection at theinterface cannot be drawn on the basis of their results alone.Later in this paper, these results are considered together withthose of past tests3-12,16-18 on RC-jacketed members withdifferent means of connection at the interface, to draw morefirm conclusions.

The behavior of the monolithic Column Q-RCM wasgoverned by flexure throughout the test: the width of a fewhairline diagonal cracks that developed up to maximum load,decreased after ultimate strength. Flexure also controlled thebehavior and failure of all jacketed columns. In some ofthem, however, diagonal cracking was more visible than inthe monolithic column. A notable feature of the behavior ofjacketed Columns Q-RCD and Q-RCR is the cracking andsplitting of the concrete cover along corner bars nearultimate strength, which spread up to approximately 300 mm(12 in.) from the base in Q-RCD, or 450 mm (18 in.) in Q-RCR.It is interesting that bond splitting along vertical bars was notprevented by the dowels connecting the jacket to the oldcolumn in Q-RCD. There was no bond splitting in ColumnQ-RCW, where each corner bar was connected to itscounterpart in the old column at two points along its length,or in Column Q-RC, which had no treatment of the interface.Such cracking and splitting of the concrete cover along thebars does not appear to have any adverse effect on the force

Fig. 3—Force-drift loops of jacketed columns for variousmeans of connection of jacket and old column.

ACI Structural Journal/September-October 2007 525

and deformation capacity of the column. As a matter of fact,Columns Q-RCD and Q-RCR had at least as large ultimatedeformation as the other jacketed columns, or as themonolithic one. It is notable that in Specimens Q-RCR andQ-RCD, the bars of the old column inside the jacket buckledas well, as shown in Fig. 4.

Mean axial strain at bottom of jacketed columnsFigure 5 shows how the axial strain at the center of the

section (average value within the lower-most 250 mm [10 in.]of the column) varies with the applied drift. These axial strainsare derived from the average of the LVDT measurements onopposite sides of the column near the base and reflect the axialdisplacement of the jacket at a section 250 mm (10 in.) abovethe base with respect to the footing. The change in axial strainat the center of the section during each cycle of loading inproportion to the lateral deflection is due to flexure accordingto the plane-sections hypothesis.

The most notable result in Fig. 5 is the axial extension atzero value of the deflection. If the normalized axial load, ν =N/bhfc′ , were high, axial shortening of the column wouldhave accumulated during the cycling of the load beyondyielding and towards failure of the member, due toaccumulation of permanent compressive strains in theconcrete. Because all present columns had fairly lownormalized axial load, there is no ratcheting axial displacementat zero deflection in the monolithic Column Q-RCM, and inall jacketed columns except Q-RCD and (mainly) Q-RC.The large ratcheting axial elongation of these two lattercolumns is due to slippage of the jacket with respect to theold column inside. Roughening of the old surface of the oldcolumn or U-bars welded to the corner bars of the oldcolumn seem to prevent such slippage, while dowels alone atthe interface are only partly effective in this respect.Although slippage may influence the compression block andreduce the internal lever arm, there is no clear effect of it onthe global cyclic behavior of the columns in Fig. 5 and on theparameters that summarize it in Table 1. The analysis of theentire database of test results on jacketed members in theliterature in the next section, however, suggests that, if thereare no positive measures of connection at the interface, or ifsuch measures are limited to dowels or welded U-bars, thesecant stiffness at incipient yielding may be slightly lowerthan in columns with roughened interface.

STRENGTH, STIFFNESS, AND DEFORMATION CAPACITY OF JACKETED MEMBERS

Database of tested jacketed membersTo develop and calibrate simple rules for the estimation of

the yield moment, of the secant stiffness at apparent yieldingand of the cyclic deformation capacity of jacketed members,

the results of the present five tests are supplemented with theresults of 49 more tests from the literature.3-12,16-18 That infor-mation is given in the Appendix* in the form of tables listing alltest variables and key test results. Most of the jacketed membersin the database did not have lap splicing of the longitudinal barsin the original member, but some of them did.17

Regarding the tests in the literature, it is noted that:1. The three columns in Reference 3, the single one from

Reference 11, and the two columns in Reference 18 werejacketed for strengthening in shear and stiffening of the freelength outside the footing. The vertical bars of the jacketwere not connected to the footing. (In Reference 11, asizeable gap was provided between the end of the jacket andthe footing.) All other specimens had jacket longitudinalbars anchored at the footing and were representative of aglobal intervention enhancing flexural capacity as well;

2. In Reference 4, the vertical bars of the jacket werethreaded into a steel plate fixed at the top of the footing,preventing fixed-end rotation of the column at the basesection. Moreover, the column axial load at ultimateconditions was different (significantly less) from its value atyielding;

3. The jacketed columns in Reference 5 were subjected toeccentric compression without transverse load. So, onlyyield moments and curvatures reported in Reference 5 can beused in the present work;

Fig. 4—Damage of retrofitted columns: (a) Q-RCD (notedowel and broken bar); and (b) Q-RCR.

Fig. 5—Evolution of mean vertical strain at center ofsection (over lower 250 mm [10 in.] from base) with drift.

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

526 ACI Structural Journal/September-October 2007

4. In Reference 7, the original column was prefabricated,with the base fixed into a 750 mm (30 in.) deep socket bygrouting. The jacket was connected to the footing via epoxy-grouted anchors that had strength sufficient to force columnyielding and plastic hinging to take place right above theanchors. So, the critical section was not at the base of thecolumn but at the section where the anchors were terminated;

5. The tests in Reference 8 were done on three-storycantilever walls with barbelled section. Jacketing was limitedto the first story. It comprised doubling the web thickness andadding new vertical reinforcement to the barbells, welded tostarter bars from the footing. Drifts were measured at the topof the third story, 3.01 m (118.5 in.) above the base. Togetherwith the lateral loads applied at story levels, a moment wasapplied at the top of the third story. Failure took place in thetwo upper (unstrengthened) stories;

6. In Reference 12, the vertical bars of the jacket wereanchored with epoxy at the top of the footing, preventingfixed-end rotation of the column at the base section; and

7. The beams specimens in Reference 16 weremonotonically loaded at third-points of their span andreported results do not include information from which thedrift ratio and the curvature at the section of maximummoment can be derived. So, only the yield moments reportedin Reference 16 can be used herein.

Strength, stiffness, and deformation capacity of monolithic members

The yield moment, the secant stiffness at apparent yielding,and the cyclic deformation capacity of the jacketed member arecalculated here, as a multiple or fraction of the correspondingquantities of a monolithic member, as appropriate. The yieldmoment My and the yield curvature φy of the end section of amonolithic member with longitudinal bars not lap-spliced in thevicinity of that section may be determined from firstprinciples.24 The effective stiffness to yielding at each end ofsuch a member is determined as EIeff = MyLs/3θy, where thechord rotation at yielding at that end θy may be determined asfollows23,25

In beams or columns

(1a)

(1b)

In rectangular, T-walls, or barbelled walls

(2a)

(2b)

θyφy Ls aVz+( )

3------------------------------ 0.0013 1 1.5 h

Ls

-----+⎝ ⎠⎛ ⎞ 0.13φydbL fy

fc′------------------------------+ +=

fy fc′in MPa,( )

θyφy Ls aVz+( )

3------------------------------ 0.0013 1 1.5 h

Ls

-----+⎝ ⎠⎛ ⎞ 0.34φydbL fy

fc′------------------------------+ +=

fy fc′in ksi,( )

θyφy Ls aVz+( )

3------------------------------ 0.002 1 0.125

Ls

h-----–⎝ ⎠

⎛ ⎞ 0.13φydbL fy

fc′------------------------------+ +=

fy fc′in MPa,( )

θyφy Ls aVz+( )

3------------------------------ 0.002 1 0.125

Ls

h-----–⎝ ⎠

⎛ ⎞ 0.34φydbL fy

fc′------------------------------+ +=

fy fc′in ksi,( )

where αV = 1 if shear cracking precedes flexural yielding at theend section (that is, if My > LsVR,c, with VR,c denoting the shearresistance without shear reinforcement); otherwise, αV = 0; zequals the length of internal lever arm, taken equal to the distancebetween tension and compression reinforcement in beams,columns, or walls with barbelled or T-section, or z = 0.8h inwalls with rectangular section; and dbL is the diameter of thetension reinforcement.

The third term is the fixed-end rotation due to pull-out of thetension reinforcement from its development beyond themember end and is omitted if such pull-out is not physicallypossible (for example, if the end of the longitudinal bars isfixed at the end section, as in References 4 and 12). The valueof fc′ along the development of the tension reinforcementbeyond the end section should be used in this term.

The flexure-controlled deformation capacity of themonolithic or of the jacketed member is expressed herein interms of the plastic part, θu

pl of the ultimate chord rotation(total ultimate chord rotation θu minus value at yielding θy)at the end of the member. An empirical expression has beendeveloped in Reference 25 (and adopted in Reference 23) forthe value of θu

pl of members with: a) rectangular crosssection; b) ductile deformed longitudinal bars; c) possibilityfor slip of the longitudinal bars from their developmentbeyond the section of maximum moment; and d) detailingfor earthquake resistance (including no lap-splicing in theplastic hinge region, but not including diagonal bars)

(3a)

(3b)

where αwall = 1.0 for beams or columns and αwall = 0.6 forwalls; ν = N/bhfc′ (with b being the width of the compressionzone and the axial force N considered positive forcompression); ω = ρfy/fc′ , mechanical reinforcement ratio oftension longitudinal reinforcement (including anylongitudinal reinforcement between the two flanges); ω′ =ρ′fy/fc′ , mechanical reinforcement ratio of compression,reinforcement; ρs = As/bwsh, transverse steel ratio parallel tothe direction of loading; and α equals confinementeffectiveness factor from Reference 26.

The numerous available monotonic data show that formonotonic loading, the right-hand side of Eq. (3) is dividedby 0.48. Also, in members with no possibility of physicalpull-out of tension reinforcement (for example, refer toReferences 2 and 10), the available data show that, if this isthe case, the plastic part of the ultimate chord rotation θu

pl,monotonic or cyclic, should be multiplied by 1.625 (= 1 + 5/8).

For members with lap splices (of plain or of deformedbars) in the plastic hinge region, as well as for members withother details not appropriate for earthquake resistance (for

θupl 0.0145awall 0.25v( ) max 0.01 ω′,( )

max 0.01 ω,( )-----------------------------------

0.3fc′( )0.2=

LS

h-----⎝ ⎠⎛ ⎞

0.35

25

αρS

fyw

fc ′-------⎝ ⎠

⎛ ⎞

fy fc′ in MPa,( )

θupl 0.0212awall 0.25v( ) max 0.01 ω′,( )

max 0.01 ω,( )-----------------------------------

0.3fc′( )0.2=

LS

h-----⎝ ⎠⎛ ⎞

0.35

25

αρS

fyw

fc ′-------⎝ ⎠

⎛ ⎞

fy fc′ in ksi,( )

ACI Structural Journal/September-October 2007 527

example, not closed stirrups), modifications to Eq. (1) to (3)were proposed in Reference 25 and adopted in Reference 23.

Simple rules for strength, stiffness, and deformation capacity of jacketed members

The rules proposed herein on the basis of the tests in thedatabase use modification factors on the properties of anequivalent monolithic member. The strength, stiffness, anddeformation capacity of the equivalent monolithic memberare determined according to the previous rules and to the addi-tional considerations listed in Table 2. The concept behindassumptions A.3 and A.4 in Table 2 is that, for common ratiosof jacket thickness to depth of the jacketed section, it is mainlythe jacket that carries the full axial load at the critical endsection and in the plastic hinge of the column. Also, it is thejacket that mainly controls the shear resistance and the bondalong the longitudinal reinforcement of the jacket.

In the following, an asterisk is used to denote a calculatedvalue for the jacketed member; for example, in My*, θy*, andθu*, whereas no asterisk is used (for example, in My, θy, andθu

pl) for a value calculated for the monolithic memberaccording to the assumptions in Table 2 and Eq. (1) to (3). Theratios of the experimental values of My, θy, and θu for thejacketed members in the database to calculated values of My,θy, and θu for the monolithic member according to theassumptions in Table 2 and Eq. (1) to (3) were determined andare shown in Fig. 6. Note that θu,cal is taken in Fig. 6 (bottom)equal to θy* + θu

plEq.(3), with θy* = 1.05 θy,Eq.(1)&(2) being the

overall best estimate of the chord rotation at yielding for thejacketed member (with θy from Eq. (1) and (2)). With sodefined θy*, in Fig. 6 (third figure from the top): EIeff* =My,calLs/3θy*. The ratios My,exp/My,calc, θy,exp/θy,Eq.(1)&(2),EIexp/EIeff* and θu,exp/(θy* + θu

plEq.(3)) are given from top to

bottom of Fig. 6 separately for different ways of connectingthe jacket to the old member, and separately for thosemembers which were damaged by testing before jacketing.Specimens in which the longitudinal reinforcement of thejacket did not continue beyond the member end,1,9 orspecimens with lap-spliced reinforcement in the originalcolumn,15 are identified in Fig. 6, but otherwise lumpedtogether with the tests of members with continuous verticalbars in the original member. For those tests that did not reachultimate conditions and for the two walls in Reference 6 thatfailed in the unstrengthened stories, an upward arrowsignifies an experimental-to-predicted ratio greater than theplotted value.

The average value and plus or minus one standarddeviation estimates of the mean experimental-to-calculatedratios are shown in Fig. 6 separately for the various groupsof specimens that represent different types of jacket-to-old-member connection, with or without damage in the originalcolumn. The distance of the sample average from a certainreference value (for example, 1.0), normalized by thestandard-deviation of the mean, is a criterion on whether thejacketed member property of interest may be taken equal tothe value calculated for the monolithic member according tothe assumptions in Table 2 and Eq. (1) to (3), times thatreference value.

Figure 6 supports the following rules for the calculation ofthe yield moment, the chord rotation (drift ratio) at apparentyielding, and the ultimate chord rotation (drift ratio), My*, θy*,or θu*, respectively, of the jacketed member, in terms of thevalues My, θy, and θu

pl calculated for the monolithic memberaccording to the assumptions in Table 2 and Eq. (1) to (3)

1. For My*: My* = My,calc;

Table 2—Determination of properties of equivalent monolithic memberFlexural resistance and deformation capacity, deformations at flexural yielding

Case A: Jacket longitudinal reinforcement continues beyond member end section for development

A1 Dimensions External dimensions of section are those of jacket.

A2 Longitudinalreinforcement

Tension and compression reinforcement are those of jacket. Longitudinal bars of old member are considered at theiractual location between tension and compression reinforcement from jacket: they may supplement any longitudinal barsof jacket between tension and compression reinforcement and included in web reinforcement ratio considered asuniformly distributed between tension and compression reinforcement; in walls, old vertical bars near edges may beincluded in tension and compression reinforcement of jacketed member, as appropriate. Effect of lap splicing inintermediate old reinforcement may be neglected. Differences in yield stress between new and old vertical reinforcementshould be taken into account in all cases.

A3 Concrete strengthValue of f c′ of jacket applies over full section of monolithic member, except in third term of Eq. (1) and (2), where f c′ valueused is that of concrete where longitudinal reinforcement is developed beyond end section.

A4 Axial load Full axial load is taken to act on jacketed column as a whole, although it was originally applied to old column alone.*

A5 Transverse reinforcement Only transverse reinforcement in jacket is taken into account for confinement.†

Case B: Jacket longitudinal reinforcement stops at end section

B1Dimensions, longitudinal

reinforcement, andconcrete strength

My and φy (also in first and third term of Eq. (1) and (2)) are calculated for cross-sectional dimensions, longitudinalreinforcement, and fc′ value of old member, neglecting any contribution from jacket (even in compression zone). Effectof lap splicing of old reinforcement should be taken into account as in References 20 and 23. Section depth h in secondterm of Eq. (1) and (2) (that has to do with shear deformations) should be that of jacket.

B2 Transverse reinforcement

Deformation capacity θu is calculated on basis of old column alone, considering that old column is confined by jacketand its transverse steel. Value of ρs = As/shbw for Eq. (3) is determined from value of As/sh in jacket and width of oldcolumn as bw. Confinement effectiveness factor α may be taken α = 1.0 (jacket serves as cushion distributing confinementto full extent of old section).

Shear resistance

C

Regardless of whether longitudinal reinforcement of jacket continues beyond member end section for development or stops there, shear resistance(including that without shear reinforcement, VR,c , for determination of value of V in first term of Eq. (1) and (2)) and anything that has to do withshear should be calculated on basis of external dimensions and transverse reinforcement of jacket.† Contribution of old transverse reinforcementmay be considered only in walls, provided it is well developed into (new) boundary elements.

*Assumption is supported by conclusion in Reference 12 that application of full axial load on old column alone at time of jacketing had almost no effect on monotonic response upto yielding and beyond, compared with column with axial load applied just at time of testing.†Assumption is supported by strain measurements in Reference 12.

528 ACI Structural Journal/September-October 2007

2. For θy* (the main target being the effective stiffness atincipient yielding, computed as EIeff* = My*Ls/3θy* withMy* = My,calc), irrespectively of predamage in the originalcolumn: θy* = 1.05θy,Eq.(1)&(2).

(Note that this rule was adopted in Reference 23 only forroughening of the interface or for roughened interfacecombined with epoxy-grouted dowels, but that the moreconservative rule: θy* = 1.2θy,Eq.(1)&(2) was adopted inReference 23 for no treatment of the interface, or epoxy-

grouted dowels alone, or jacket bars connected to those ofthe old member via welded U-bars); and

3. For θu*: θu* = θy* + θupl

Eq.(3)Rules 1 to 3, supplemented with the relevant provisions of

Table 2 (notably, assumptions B1 and B2), also apply if thejacket longitudinal bars stop at the end section of the member.

If no differentiation is made depending on the type ofmeasures taken to enhance the shear transfer at the interfaceof the old and the new concrete, the ratio of the experimentalvalues to those predicted according to the previous Rules 1to 3 has overall mean value and coefficient of variation(COV) equal to 1.025 and 10.4%, 1.005 and 23.5%, 0.98 and31.3%, and 1.125 and 18.8% for My, θy, EIeff , and θu,respectively. Bonding measures at the interface of the jacketand the old member seem to have a statistically significanteffect only for the ultimate chord rotation θu. The proposedrules underestimate the measured ultimate chord rotationθu,exp by 27% on average for the few specimens employingeither dowels at the interface or U-bars welded to the newand the old longitudinal bars, but only by 8% for the morenumerous specimens with neither of these measures. WeldedU-bars seem to be slightly more beneficial for the ultimatechord rotation than dowels (possibly due to their anti-buckling action). Roughening of the old surface alone doesnot seem to positively affect the ultimate chord rotation.Rule 3, which neglects the favorable effect of any positiveconnection measure at the interface of the old and the newconcrete, is therefore safe-sided for the ultimate chordrotation θu underestimating its measured value by 12.5% onaverage. Interestingly, no systematic positive effect ofdowels or welded U-bars on the yield moment My and theeffective stiffness EIeff was found.

The COVs of the ratio of experimental My, θy, EIeff , andθu for monolithic members to the values calculated from thefirst principles or Eq. (1) to (3) has been quantified inReference 25 and exceed those of the data for jacketedmembers in Fig. 6.

None of the jacketed specimens in the database had anysign of shear distress. This is consistent with the observa-tion that in any test, the shear resistance (calculatedaccording to Reference 27 as a decreasing function ofdisplacement ductility demand) exceeded by more than30% the applied shear.

The experimental value of My, θu , and EIeff , divided by thecorresponding value of My*, θu*, and EIeff* = My*Ls/3θy*derived according to the previous Rules 1, 2, and 3 areplotted in Fig. 7 and 8 against:

a. The ratio of fc′ of the jacket fcj′ to that of the old member,fcc′ (compare assumption A.3 in Table 2);

b. The ratio of the cross-sectional area of the jacket Acj tothat of the old member Acc;

c. The ratio of the product of the yield stress times the ratioof longitudinal reinforcement of the jacket to the sameproduct for the old member;

d. The axial load, normalized to the full cross-sectionalarea of the jacketed section times the fc′ value of the jacketas ν′ = N/(Acj + Acc)fcj′ ;

e. The axial load, normalized to the actual compressivestrength of the jacketed section as ν′ =N/(Acj fcj′ + Acc fcc′ ); and

f. The ratio of the compression zone depth c at yielding tothe thickness of the jacket.

Specimens with lap-spliced reinforcement in the originalcolumn15 are appropriately identified. There is no clearupward or downward trend in any of the plots of the ratio of

Fig. 6—Test results on RC-jacketed members compared withprediction for monolithic member: (a) no treatment ofinterface; (b) no treatment, predamaged member; (c)welded U-bars; (d) dowels; (e) roughened interface; (f)roughened interface, predamaged member; (g) U-bars androughened interface; (h) U-bars and roughened interface,predamaged member; (i) roughened interface and dowels;(j) roughened interface and dowels, predamaged member;and (k) monolithic.

ACI Structural Journal/September-October 2007 529

the experimental values of My, θu, EIeff to My*, θu*, andEIeff *, respectively, against the aforementioned parameters(a) to (f) in Fig. 7 or 8. This shows that the predictions of My,EIeff , or θu as My*, θu*, EIeff* above, according toaforementioned Rules 1 to 3, are not systematically biasedwith respect to any of these parameters. Note that the data inFig. 7 and 8 support assumptions A3 and A4 in Table 2, evenwhen the compression zone extends beyond the jacket,within the section of the old column.

CONCLUSIONSThe experimental program on rectangular RC-jacketed

columns focused on the impact of measures taken to enhancethe shear transfer at the interface of the old and the newconcrete. The test results were then supplemented withmonotonic or cyclic experimental data from theliterature3-12,16-18 with various bonding measures betweenthe old and the new concrete (usually not applied in the sametesting program), to draw more robust conclusions about theeffect of bonding at the interface. This was done byexpressing key properties of the jacketed column (notably,the yield moment, the yield drift, the secant stiffness atincipient yielding, and the flexure-controlled ultimate drift)as ratios to the corresponding quantities of an equivalentmonolithic member. The conclusion was that such ratios donot systematically depend on the type of measures taken toenhance the shear transfer at the interface of the old and the

new concrete. There is an exception for the ultimate chordrotation, θu, which seems to be increased by approximately1/6, if dowels are employed at the interface, or if U-bars arewelded between the new and the old longitudinal bars.

On the basis of the increased database of tests on memberswith or without lap splices at the floor level, rehabilitated viaconcrete jacketing, simple rules and expressions weredeveloped for the calculation of key properties of columns.According to these rules, the jacketed member is consideredas monolithic, having:• External cross-sectional dimensions and fc′ value those

of the jacket;• The axial load applied on the full section of the

jacketed member;• Tension and compression reinforcement those of the

jacket (in walls, supplemented with old vertical barsnear the edges, taking into account differences in yieldstress with the new reinforcement and any lap splicingat the floor level),

• Longitudinal bars of the old member between the newtension and compression reinforcement included in aweb reinforcement ratio (possibly supplementing longi-tudinal bars of the jacket between the tension and com-pression reinforcement); and

• Confinement only owing to the transverse reinforcementin the jacket.

Fig. 7—Accuracy of predictions on basis of proposed rules for calculation of yield moment (top row), of secant stiffness atapparent yielding (middle row), and of ultimate chord rotation (bottom row) of RC-jacketed members, as function of: (a) ratio offc′ of jacket to that of old member; (b) ratio of cross-sectional area of jacket to that of old member; and (c) ratio of product of yieldstress times longitudinal reinforcement ratio of jacket, to same product of old member.

ACI Structural Journal/September-October 2007530

If the jacket longitudinal bars stop at the member endsections, the yield moment and the flexure-controlledcyclic deformation capacity of the jacketed member maybe estimated considering the jacketed member asmonolithic, with:• Cross-sectional dimensions, longitudinal reinforcement,

and fc′ value those of the old member;• The axial load applied on the section of the old

member; and• Confinement provided by the jacket and its

transverse steel (with the transverse reinforcementratio, ρs = As/shbw determined from the value of As/shin the jacket and the width of the old column as bwand the confinement effectiveness factor taken equalto 1.0, as if the jacket serves as a cushion distributingconfinement to the full extent of the old section).

With these rules, the yield moment of the jacketed memberis, on average, well estimated, while its drift ratio at incipientyielding is underestimated by approximately 5% on average.There is no systematic effect of any special connectionmeasures at the interface on the yield moment My and theeffective stiffness EIeff of the jacketed member. The ultimatechord rotation is underestimated by 8%, on average, if nomeasures are taken at the interface between the old and thenew concrete to improve the connection. It is underestimatedmore when dowels are employed at the interface, and even

more if U-bars are welded between the new and the oldlongitudinal bars. So, the rules proposed for the ultimatechord rotation are safe-sided in all cases.

Regardless of whether the longitudinal reinforcement ofthe jacket continues beyond the member end section fordevelopment or stops there, the shear resistance (as well asanything that has to do with shear, such as the part of the driftat yielding owing to shear deformations—compare secondterm in Eq. (1) and (2)) should be calculated on the basis ofthe external dimensions and the transverse reinforcement of thejacket. The contribution of the old transverse reinforcementmay be considered only in walls, provided it is well developedinto the (new) boundary elements.

ACKNOWLEDGMENTSThis study was funded partly by the European Commission (project

SPEAR of the GROWTH program, Contract No. G6RD-2001-00525) andpartly by NATO (via its Science for Peace Program).

NOTATIONb = width of compression zoneh = cross section depthLs = shear spanMy = yield momentMy* = yield moment of jacketed memberMy,exp = experimental value of yield momentMy,calc = value of yield moment calculated from first principlesN = axial force (positive for compression)

Fig. 8—Accuracy of predictions on basis of proposed rules for calculation of yield moment (top row), of secant stiffness atapparent yielding (middle row), and of ultimate chord rotation (bottom row) of RC-jacketed members, as function of: (a) axialload, normalized to full cross-sectional area of jacketed section times fc′ of jacket; (b) axial load, normalized to actualcompressive strength of jacketed section; and (c) ratio of compression zone depth to thickness of jacket.

ACI Structural Journal/September-October 2007 531

VR,c = shear resistance of member considered without shearreinforcement

z = length of internal lever armα = confinement effectiveness factor from Reference 26ν = normalized axial load N/bhfc′θy = chord rotation (drift ratio) at yieldingθy* = chord rotation (drift ratio) of jacketed member at yielding θy,exp = experimental value of chord rotation at yieldingθu = ultimate chord rotation (drift ratio)θu* = ultimate chord rotation (drift ratio) of jacketed memberθu,exp = experimental value of ultimate chord rotationθu

pl = plastic part of ultimate chord rotation ρ = tension reinforcement ratio, including bars distributed between

tension and compression flangesρ′ = compression reinforcement ratiosρs = transverse steel ratio parallel to direction of loadingφu = ultimate curvatureφy = curvature at base at yieldingω = mechanical reinforcement ratio of tension longitudinal

reinforcement ρfy/fc′ω′ = mechanical reinforcement ratio of compression reinforcement

ρ′fy /fc′

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