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The composite flexural action of prismatic reinforced concrete �RC� members repaired/strengthened by RC jacketing was modeled with a dual-section approach. The model considers the relative slip at the interface between the existing member and the jacket and establishes the mechanisms that are mobilized to resist this action, thereby supporting composite behavior. An iterative step-by-step incremental algorithm was developed for calculating the overall flexural response curve. Consideration of frictional interlock and dowel action associated with sliding at the interfaces as well as the spacing and penetration of flexure-shear cracks are key aspects of the algorithm. The proposed procedure was verified through comparison with published experimental data on RC jacketed members. The sensitivity of the upgraded member’s flexural response to jacket design variables was investigated parametrically. Monolithic response modification factors related to strength and deformation indices were evaluated and the sensitivity of the model was assessed.
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Flexural Behavior of Brittle RC Members Rehabilitated
with Concrete Jacketing
G. E. Thermou, Ph.D.1; S. J. Pantazopoulou, M.ASCE2; and A. S. Elnashai, F.ASCE3
Abstract: The composite flexural action of prismatic reinforced concrete RCmembers repaired/strengthened by RC jacketing was
modeled with a dual-section approach. The model considers the relative slip at the interface between the existing member and the jacket
and establishes the mechanisms that are mobilized to resist this action, thereby supporting composite behavior. An iterative step-by-step
incremental algorithm was developed for calculating the overall flexural response curve. Consideration of frictional interlock and dowel
action associated with sliding at the interfaces as well as the spacing and penetration of flexure-shear cracks are key aspects of the
algorithm. The proposed procedure was verified through comparison with published experimental data on RC jacketed members. The
sensitivity of the upgraded member’s flexural response to jacket design variables was investigated parametrically. Monolithic response
modification factors related to strength and deformation indices were evaluated and the sensitivity of the model was assessed.
DOI: 10.1061/ASCE0733-94452007133:101373
CE Database subject headings: Concrete, reinforced; Rehabilitation; Seismic design; Inelasticity; Flexure.
Reinforced concrete jacketing is a traditional method for seismic
upgrading of damaged or poorly detailed reinforced concrete con-
struction. In applying this technique, the objective is to suppress
alternative premature modes of failure that would otherwise pre-
vail in the structural members under reversed cyclic loading,
thereby promoting flexural yielding of primary reinforcement.
Through reinforced concrete RCjacketing stiffness and strength
are increased, whereas dependable deformation quantities may or
may not be enhanced, depending on the aspect ratio of the up-
graded element and the factors limiting deformation capacity in
the initial state of the element. For practical purposes, response
indices of the jacketed members such as resistance and deforma-
tion measures at yielding and ultimate are routinely obtained by
applying pertinent multipliers on the respective properties of
monolithic members with identical geometry. The multipliers are
referred to in the literature as monolithic factors,Ki.
Depending on the member property being scaled strength or
stiffness, the method of load application and the jacket function,
various values have been reported for Ki, ranging from 0.7 up to
1. Eurocode 8 CEN 1996recommends KR=0.8 for strength and
KK=0.7 for stiffness provided that: 1Loose concrete and buck-
led reinforcement in the damaged area have been repaired or re-
placed before jacketing; 2all new reinforcement is anchored
into the beams and slabs; and 3the additional concrete cross
section is not larger than twice the cross section of the existing
column. Based on the results of a recent experimental study con-
ducted by Vandoros and Dritsos 2006a,b, the monolithic factors
associated with strength, stiffness, and deformation vary greatly
depending on the techniques followed in constructing the jacket.
For example, it was shown that dowels improve the ductility ca-
pacity of the jacketed member, roughening of the interface in-
creases the energy absorption capacity, and a combination of the
two procedures improves stiffness.
Monolithic factors are used by codes of practice for con-
venience, as the mechanics of composite action of jacketed re-
inforced concrete members under cyclic shear reversals is too
complicated for practical calculations. So far the focus has been
on stiffness and strength, whereas no specific reference has been
made for monolithic factors related to deformation indices. A de-
tailed method for calculating these factors would be required in
order to assess their parametric sensitivity to the relevant design
variables. From the available experimental evidence it appears
that slip and shear stress transfer at the interface between the
outside jacket layer and the original member that serves as the
core of the upgraded element are controlling factors CEN 1996;
KANEPE 2004. Indeed, sliding failure at the interface limits the
strength and affects the rotation capacity of the entire member.
This paper presents a detailed procedure for estimating the
behavior of concrete members jacketed with an outer RC shell.
The composite action that jacketed reinforced concrete members
develop in flexure greatly depends on the force transfer that oc-
curs between the core and the jacket. Estimating strength and
deformation capacity of such members is a complex mechanics
problem that is hampered by the limited understanding of the
interfacial resistance mechanisms such as friction, interlock, and
dowel action. To calculate the monolithic factors and to establish
1Dept. of Civil Engineering, Laboratory of Reinforced Concrete,
Demokritus Univ. of Thrace, Vas. Sofias 12, Xanthi 67100, Greece.
2Professor, Dept. of Civil Engineering, Laboratory of Reinforced
Concrete, Demokritus Univ. of Thrace, Vas. Sofias 12, Xanthi 67100,
Greece corresponding author. E-mail:
3William J. and Elaine F. Hall Professor, Dept. of Civil and
Environmental Engineering, Univ. of Illinois at Urbana-Champaign,
2129 Newmark CE Lab., 205 North Mathews Ave., Urbana, IL 61801.
Note. Associate Editor: Dat Duthinh. Discussion open until March 1,
2008. Separate discussions must be submitted for individual papers. To
extend the closing date by one month, a written request must be filed with
the ASCE Managing Editor. The manuscript for this paper was submitted
for review and possible publication on October 27, 2005; approved on
March 9, 2007. This paper is part of the Journal of Structural Engineer-
ing, Vol. 133, No. 10, October 1, 2007. ©ASCE, ISSN 0733-9445/2007/
their dependence on critical design variables an analytical model
is developed in this paper from first principles. The significance
of jacket detailing on the resulting response and the associated
values of the monolithic factors for strength and deformation ca-
pacity is demonstrated and quantified through parametric studies
and correlation of analytical estimates with test results.
Analytical Model for RC Jacketed Members
It is assumed that the existing member core is partially connected
with the external jacket layer, so that the mechanisms of force
transfer at the interface are mobilized by relative slip of the two
bodies. In analyzing the flexural behavior, the cross section of the
upgraded member is divided into three layers. The two external
ones represent the contribution of the jacket, whereas the middle
layer represents both the core existing cross sectionand the web
of the jacket shell Fig. 1a兲兴. For reference in the remainder of
this derivation, the Cartesian coordinate system is oriented so that
the xaxis is parallel to the longitudinal member axis, the yaxis is
along the cross-sectional depth, whereas the zaxis is oriented
along the cross-sectional breadth Fig. 1d兲兴. The difference in
normal strain at the interface between layers accounts for the
corresponding slip in the longitudinal direction; thus, only the
implications of slip along horizontal planes are considered in the
model. The inaccuracy associated with neglecting shear transfer
along the vertical contact faces i.e., on faces normal to the zaxis
is small if jacket longitudinal tension reinforcement is evenly dis-
tributed in the perimeter Fig. 1a兲兴. Note that in that case, a
vertical slice of the jacketed cross section is self-equilibrating
consider for example the rectangular portion of the cross section,
to the left of line A-A’ in Fig. 1a兲兴. This means that the total
stress resultant is zero since compression and tension forces over
the height of the segment are in equilibrium; hence the shear
stress xz Fig. 1d兲兴 acting in a plane normal to the zaxis and
oriented in the longitudinal direction is also zero. As usually done
in flexural analysis of layered composite beams, it is assumed that
the three layers deform by the same curvature, Fig. 1a兲兴. From
free body equilibrium of any of the two exterior layers the shear
flow at the interface is calculated as the difference in the stress
resultant between two adjacent cross sections. The procedure is
implemented in an iterative algorithm that employs dual-section
analysis. A key element of the algorithm is the shear stress slip
relationship used to describe the behavior of the interface between
Interface Shear Behavior
Slip at the interface between the existing member and the jacket is
explicitly modeled. Mechanisms that resist sliding are: 1Aggre-
gate interlock between contact surfaces, including any initial ad-
hesion of the jacket concrete on the substrate; 2friction owing
to clamping action normal to the interface; and 3dowel action
of any pertinently anchored reinforcement crossing the sliding
plane. Thus, in stress terms, the shear resistance, vn, against slid-
ing at the contact surface, is
In Eq. 1vashear resistance of the aggregate interlock
mechanism; ␮⫽interface shear friction coefficient; Nnormal
clamping stress acting on the interface; and vDshear stress re-
sisted by dowel action in cracked reinforced concrete. The first
two terms collectively represent the contribution of concrete as
they depend on the frictional resistance of the interface planes.
The clamping stress represents any normal pressure, p, externally
applied on the interface, but also the clamping action of reinforce-
ment crossing the contact plane as illustrated in Fig. 1b. From
equilibrium requirements it is shown
where pnormal pressure externally applied on the contact plane;
fsaxial stress of the bars crossing the interface; and
scorresponding reinforcement area ratio.
Shear transfer is affected by the roughness of the sliding plane,
by the characteristics of the reinforcement, by the compliance
of concrete, and by the state of stress in the interface zone. Dowel
action develops by three alternative mechanisms, namely, by di-
rect shear and by kinking and flexure of the bars crossing the
contact plane. A variety of models are available for modeling
the interface phenomena. In this study, the model developed by
Tassios and Vintzēleou 1987and Vintzēleou and Tassios 1986,
1987as modified by Vassilopoulou and Tassios 2003was used
due to its simplicity and robustness. The model estimates the
combined dowel and shear friction resistances for a given slip
value at the interface, as follows:
1. Frictional resistance: The concrete contribution term in
Eq. 1,vcs, is described by the following set of equations:
= 1.14
for s
0.5 3a
= 0.81 + 0.19
for s
0.5 3b
where sc,uultimate slip value beyond which the frictional
mechanisms break down sc,uis taken approximately equal
to 2 mm兲共CEB-FIP 1993. The normalizing term,
vc,uultimate frictional resistance of the interface, given by
2N1/3 4
where ␮⫽ultimate interface shear friction coefficient taken
equal to 0.4 and fc
concrete cylinder uniaxial compressive
strength Fig. 1b兲兴. To calculate the axial stress of the bars
crossing the interface, fs, the separation wbetween contact
surfaces as they slide overriding one another is considered
Fig. 1c兲兴. According to Tassios and Vintzēleou 1987the
Fig. 1. aStrain profiles; bnormal stresses at interface; cpull-out
displacement of bars crossing interface; and dstate of stress acting
on infinitesimal element in initial coordinate system
separation wand lateral slip, s, are related by: w=0.6·s2/3.To
account for w, it is assumed that the bars pull out by w/2
from each side of the contact surface. Considering uniform
bond stresses along the embedment length, the axial bar
stress, fs, at the contact plane is estimated from
In Eq. 5,Eselastic modulus of steel; and Dbdiameter
of the bars clamping the interface here, the stirrup legs of
the jacket.
2. Dowel resistance: In the dowel model the bar behaves as a
horizontally loaded free-headed pile embedded in cohesive
soil. Yielding of the dowel and crushing of concrete are
assumed to occur simultaneously. Dowel force the resultant
of term vDin Eq. 1兲兴 is obtained from the relative interface
slip sas follows Fig. 1c兲兴
= 0.5 2
for ssd,el = 0.006Db6a
0.5 s= 0.006Db
+ 1.76sd,u
− 0.5
where sd,elelastic slip value; sd,uultimate slip value;
VD,uultimate dowel force; and Dbdiameter of the bars
offering dowel resistance here, legs of the jacket transverse
In Eq. 6bthe dowel force, VDs, is estimated iteratively
given the slip magnitude, s. The ultimate dowel strength and as-
sociated interface slip are given by
VD,u= 1.3Db
fsy1−21/2;sd,u= 0.05Db7
where ␣⫽bar axial stress normalized with respect to its yield
value and fsyyield strength of steel.
The total shear resistance of an interface with contact area Aint
crossed by kdowels is
Vs=vcsAint +kVDs兲共8
where vcsand VDsare calculated from Eqs. 3and 6, re-
spectively, for a given amount of interface slip.
Estimation of Crack Spacing
Similar to conventional bond analysis, shear transfer at the inter-
face between the existing member and the jacket is carried out
between half crack intervals along the length of the jacketed
member Fig. 2a兲兴. To evaluate the crack spacing the stress state
at the crack is compared with that at the midspan between adja-
cent cracks Fig. 2b兲兴. It is assumed that at the initial stages of
loading cracks form only at the external layers jacketincreasing
in number with increasing load, up to crack stabilization. This
occurs when the jacket steel stress at the crack, fs,cr exceeds the
limit CEB-FIP 1993
fs,cr fctm
where fctmtensile strength of concrete; =Es/Ecmmodular
ratio; and s,effeffective reinforcement ratio defined as the total
steel area divided by the area of mobilized concrete in tension,
usually taken as a circular domain with a radius of 2.5Dbaround
the bar CEB-FIP 1993. Using the same considerations in the
combined section it may be shown that a number of the external
cracks penetrate the second layer coreof the jacketed member
Fig. 2a兲兴. From the free body diagram shown in Fig. 2bthe
shear flow, qs, at the contact between the bottom layer and the
core is estimated as
where NJnumber of bars in the tension steel layer of the jacket;
Db,Jbar diameter of the jacket longitudinal reinforcement;
fb,Javerage bond stress of the jacket reinforcement layer; and
bJwidth of the jacketed cross section. The crack spacing is es-
timated from free body equilibrium in the tension zone of the core
of the composite section Fig. 2b兲兴. Assuming that the neutral
axis depth is about constant in adjacent cross sections after stabi-
lization of cracking, the crack spacing is defined as follows:
where Ncnumber of bars in the tension steel layer of the
core; Db,cbar diameter of the core longitudinal reinforcement;
fb,c= average bond stress of the core reinforcement layer;
lcheight of the tension zone in the core component of the com-
posite cross section; and fctm,ctensile strength of concrete core.
Fig. 2. a,bDefinition of crack spacing; cestimation of vertical shear stress xy, denoted here as vdi; and drotation of jacketed cross section
due to slip
Shear Stress Distribution on Cross Section
of Jacketed Member
To analyze jacketed members in flexure the composite cross
section is assumed to deform in its plane of symmetry with a
curvature ; relative slip occurs in the horizontal contact planes
between the top and bottom jacket layers and the core Fig. 1a兲兴.
Section equilibrium is established and the normal stress resultant
of each layer, Fi, is estimated Fig. 2c兲兴. Using dual-section
analysis Vecchio and Collins 1988, and considering that the
shear force at any section equals the moment gradient along the
member length, the layer stress resultant Fi, is used to calculate
the vertical shear stress demand of the member i.e., stress xy,
oriented in the yaxis in Fig. 1d兲兴, at layer ith, denoted here by
the term vd,i, from
where cis obtained from Eq. 11兲关Fig. 2a兲兴. From basic mechan-
ics the vertical shear stress, vd,i, is taken equal to the horizontal
shear stress xy =yxmobilized along the interface for a given
slip magnitude, si.
Deformation Estimates at Yield and Ultimate
The cross section is considered to have attained a state of flexural
yielding when the extreme layer of tensile reinforcement reaches
first its yield strain syor alternatively when the concrete
strain at the extreme compression fiber exceeds the limit value of
c=1.5% FIB 2003, Chap. 4. Definition of an ultimate state is
also adopted so as to allow comparisons between the monolithic
and the detailed analytical approach. To this purpose an equiva-
lent monolithic curvature,u,M
eq , is estimated from the analysis
corresponding to a specified target drift at ultimate. The total
inelastic displacement comprises the elastic displacement at yield
y, the plastic displacement p,u, and the displacement owing to
interface slip slip,u
2;p,u=uylpLs− 0.5lp;
In Eq. 13b,ycurvature at yield of the composite section;
lplength of the plastic hinge region taken here as 0.08Ls
+0.022Dbfsy according to Paulay and Priestley 1992; and
ucurvature at ultimate. Terms in Eq. 13are calculated using
the proposed model and represent the tip displacements of a can-
tilever having a length Lsequal to the shear span of the member
in seismic loading the cantilever considered represents approxi-
mately half the member length under lateral sway.slip,uis
calculated at the ultimate from the slip values at the upper and
bottom interfaces, s1uand s2u, as shown in Fig. 2d. Owing to
interface slip the cross section rotates by slip,u=s1,u+s2,u/jd,
where jdrepresents the distance between the upper and the bot-
tom interfaces i.e., jdequals the core height which is usually the
cross-sectional height of the old member. Clearly, the end slip
rotation slip,uis greater for smoother interfaces, and therefore
deformation indices of jacketed members are expected to be
higher for lower interface friction properties.
The equivalent monolithic curvature, u,M
eq , is obtained by as-
suming equal displacements at ultimate for both the monolithic
and the composite members. Therefore
eq =
2+y,MlpLs− 0.5lp
lpLs− 0.5lp14
In Eq. 14,ototal tip displacement calculated from
Eq. 13a兲兴; and y,Mcurvature at yield of the monolithic cross
section obtained from conventional sectional analysis.
Calculation Algorithm
In the proposed model the interfaces between old concrete and the
jacket are treated as the weak link of the composite behavior;
thus, the shear force demand introduced in the contact surfaces
for any level of flexural curvature cannot exceed the associated
interface strength that corresponds to the level of slip already
attained. Calculations are performed for monotonically increasing
curvature using stepwise iteration. Initially, interface slip is taken
to be zero at both contact surfaces. Hence, in the first step of the
solution for very small strainsthe longitudinal strain profile is
identical to that of the monolithic approach Fig. 1a兲兴. The shear
flow is calculated from dual-section analysis, using the estimated
flexural stresses. Based on classical mechanics a longitudinal
shear flow i.e., shear stress xy in Fig. 1d兲兴 may be calculated at
any distance yiFig. 1a兲兴 from the neutral axis of a monolithic
elastic cross section, as
where Sifirst moment of area from yito the top of the cross
section; Imoment of inertia of the composite cross section; and
Vshear force on the member calculated from the estimated flex-
ural moment of the monolithic section divided by the shear span,
Ls. In the subsequent steps the longitudinal strain gradient is
modified by allowing for sequentially increasing discontinuities
of strain at the interface levelsas required to satisfy equilibrium.
The interface slip is related to the magnitude of strain discontinu-
ity at the upper and bottom interfaces, ⌬␧1, and ⌬␧2, as follows:
where variables c1,j2,j3, and c2normal strains in the section
layers above and below the contact surfaces Fig. 1a兲兴; and
caverage crack spacing Fig. 2a兲兴. Interface shear resistance is
mobilized depending on the slip magnitude: interface shear resis-
tances v1and v2Eq. 1兲兴 are obtained from the respective slip
values s1and s2using the constitutive relationships for interface
behavior Eqs. 38, illustrated in Fig. 2c兲兴. Shear demand
values vd,1 and vd,2, estimated from Eq. 12, are compared with
the dependable resistance values from Eq. 1兲兴 for equilibrium. If
equilibrium is not attained then the slip estimate is subsequently
revised and the above calculation repeated until convergence. The
final step in the algorithm involves establishing equilibrium of
forces over the composite member cross section. The strain pro-
file of the cross section is revised if there is a nonzero residual
section force resultant, i.e., if FiNext 0; the algorithm con-
verges to a final solution when both equilibrium requirements are
The algorithm is summarized in the flowchart presented in
Fig. 3. It comprises the following steps:
1. For a selected level of sectional curvature, n, it is required
to calculate the associated moment resultant, Mn.Note that
problem unknowns are: the normal strain at the top fiber of
the jacketed cross section, J1
n,m; the interface slip at the upper
n,rand bottom interfaces s2
n,r; and the associated moment
resultant, Mn. Therefore, start by setting the sectional curva-
ture equal to n;
2. Estimate normal strain at the top fiber of the cross section,
n,mFig. 1a兲兴;
3. Estimate the interface slip at the upper and bottom interfaces,
n,rand s2
n,rEq. 16兲兴. Crack spacing is calculated from
Eq. 11;
4. Calculate the shear stress at the upper and bottom interfaces,
n,rand v2
n,r, from the respective slip values, s1
n,rand s2
Eqs. 38兲兴;
5. Define shear stress demands, vd,1
n,rand vd,2
n,rEq. 12兲兴.If
both v1
n,rand v2
n,rproceed to Step 6, otherwise
return to Step 3 and set s1
n,r+1 =s1
n,r+1 =s2
the selected increment in the slip value;
6. Check cross-section equilibrium. If FiNexttolerance
go to Step 7. In any other case return to Step 2 and set
n,m+1 =J1
n,m+dJ.dJis the step increment in the top strain
of the jacketed cross section;
7. Set J1
n,rand store convergent values;
8. Estimate the moment resultant Mn. Repeat Steps 1–7 for
n=n+1. Calculations stop when the capacity of the shear
interface is exhausted.
Experimental Validation
Although RC jacketing is one of the most commonly applied
rehabilitation methods worldwide, a limited number of experi-
mental programs on RC jacketed subassemblages have been re-
ported Ersoy et al. 1993; Rodriguez and Park 1994; Bett et al.
1998; Gomes and Appleton 1998; Bousias et al. 2004; Vandoros
and Dritsos 2006a,b. The rather limited experimental database
compared to fiber reinforced polymer FRPjacketing, for ex-
ampleis a serious impediment in the development of design
expressions for this upgrading methodology.
In order to investigate the validity of the proposed analytical
model for RC jacketed members published experimental data
are used. From among the available tests those conducted by
Rodriguez and Park 1994, Gomes and Appleton 1998, Bousias
et al. 2004, and Vandoros and Dritsos 2006a,bsummarized in
Table 1 are used for model verification as they are considered
representative examples of columns under combined flexure and
shear. It is noted that reinforcement slip owing to bond was
included. Other relevant studies that were not included either
concerned short column specimens Bett et al. 1998, or tests
that had been conducted under constant moment no shear, Ersoy
et al. 1993.
Details of the experimental program are outlined in Table 1 for
all specimens considered such as geometric properties and rein-
forcement details of the original as well as the jacketed elements.
In the identification code adopted for the present comparative
study the first character is either S or M corresponding to
strengthened members with jacketing after cyclic loading or
specimens built monolithically with a composite section to be
used as controls, respectively. The second character represents the
treatment at the interface: rcorresponds to roughened interface
achieved by chipping or sandblasting or other such methods,
whereas srepresents a smooth interface. The third character Dor
Nidentifies specimens with dowels marked by Dor without
dowels marked by Ncrossing the interface between the interior
core and the jacket. The fourth character pd corresponds to pre-
damaged units. The numerals 15, 25, 30, and 45 stand for the
lap splice length of the existing unit corresponding to 15Db,
25Db,30Db, and 45Db, respectively. The character lcorresponds
to U-shaped steel links utilized to connect the longitudinal rein-
forcement of the jacket to the existing member coreand the
character wcorresponds to welding of stirrup ends of the first four
stirrups from the base of the jacketed member. The numeral in
the end is the specimen number considered in successive order
in Table 1. For easy reference, the original code names used
for the specimens by the original investigators are also listed in
Table 1 column “Specimens”.
The calculated lateral load versus lateral displacement curves
along with the curves obtained from standard sectional analysis of
the monolithic cross sections for the total number of tested units
are plotted in Figs. 4–8. The experimental curves plotted on the
same figures represent the envelope of the recorded lateral load
versus lateral displacement hysteretic loops.
In general, the monolithic approach grossly overestimates the
actual response of the jacketed member; however it is successful
in reproducing the trends of member behavior even if interface
slip is neglected. The analytical model provides a lower bound of
the response of the jacketed members and it may generally be
considered conservative, while matching well the experimental
values. At low deformation levels response curves obtained by the
analytical approach and by the monolithic approach almost coin-
cide. This is expected as long as crack formation is at an early
In addition to these general observations, the following points
are noted: for the first group of specimens Rodriguez and Park
Fig. 3. Flowchart of proposed algorithm
Table 1. Summary of Test Units
Group Code nameaSpecimenb
Rodriguez and Park
SrNpd-1 SSI 350 350 20 2.05 6 0.16 29.5 325 550 550 16 0.89 10 0.36 32.9 502 10.0 1,425
SrN-2 SS2 350 350 20 2.05 6 0.16 29.5 325 550 550 16 0.89 10 0.36 34.0 502 10.0 1,425
SrN-3 SS3 350 350 20 2.05 6 0.16 29.5 325 550 550 12 0.75 10 0.94 19.4 491 10.0 1,425
SrNpd-4 SS4 350 350 20 2.05 6 0.16 25.9 325 550 550 12 0.75 10 0.94 25.2 491 10.0 1,425
Gomes and Appleton
SsNpd-5 P2R 200 200 12 1.13 6 0.22 53.2 480 260 260 12 1.64 6 0.33 58.2 480 6.0 1,000
SsNpd-6 P3R 200 200 12 1.13 6 0.66 58.2 480 260 260 12 1.64 6 0.49 49.6 480 7.1 1,000
MsN-7 P4 200 200 12 1.13 6 0.22 56.2 480 260 260 12 1.64 6 0.33 56.2 480 6.3 1,000
Bousias et al.
MsN-8 Q-RCL0M 250 250 14 0.98 8 0.24 30.6 313 400 400 20 1.29 10 0.44 30.6 500 18.0 1,600
SsN-9 Q-RCL0 250 250 14 0.98 8 0.24 26.3 313 400 400 20 1.29 10 0.44 55.8 500 7.9 1,600
SsN15-10 Q-RCL1 250 250 14 0.98 8 0.24 27.5 313 400 400 20 1.29 10 0.44 55.8 500 8.4 1,600
SsN25-11 Q-RCL2 250 250 14 0.98 8 0.24 25.6 313 400 400 20 1.29 10 0.44 55.8 500 8.4 1,600
SsNpd15-12 Q-RCL01pd 250 250 14 0.98 8 0.24 28.1 313 400 400 20 1.29 10 0.44 20.7 500 25.0 1,600
SsNpd25-13 Q-RCL02pd 250 250 14 0.98 8 0.24 28.6 313 400 400 20 1.29 10 0.44 20.7 500 27.0 1,600
Bousias et al.
SsN15-14 R-RCL1 250 500 18 0.81 8 0.24 36.7 514 400 650 18 1.13 10 0.44 55.8 500 6.6 1,600
SsN30-15 R-RCL3 250 500 18 0.81 8 0.24 36.8 514 400 650 18 1.13 10 0.44 55.8 500 6.6 1,600
SsN45-16 R-RCL4 250 500 18 0.81 8 0.24 36.3 514 400 650 18 1.13 10 0.44 55.8 500 5.2 1,600
Vandoros and Dritsos
MsN-17 Q-RCM 250 250 14 0.98 8 0.24 24.7 313 400 400 20 1.29 10 0.44 24.7 487 21.2 1,600
SsNl-18 Q-RCW 250 250 14 0.98 8 0.24 22.9 313 400 400 20 1.29 10 0.44 18.8 487 21.6 1,600
SsD-19 Q-RCD 250 250 14 0.98 8 0.24 27.0 313 400 400 20 1.29 10 0.44 55.8 487 8.9 1,600
SrN-20 Q-RCR 250 250 14 0.98 8 0.24 27.0 313 400 400 20 1.29 10 0.44 55.8 487 8.9 1,600
SrD-21 Q-RCRD 250 250 14 0.98 8 0.24 27.0 313 400 400 20 1.29 10 0.44 55.8 487 8.9 1,600
SsNw-22 Q-RCNT 250 250 14 0.98 8 0.24 27.0 313 400 400 20 1.29 10 0.44 17.8 487 25.6 1,600
SsN-23dQ-RCNTA 250 250 14 0.98 8 0.24 23.8 313 400 400 20 1.29 10 0.44 34.5 487 11.8 1,600
SsDw-24 Q-RCE 250 250 14 0.98 8 0.24 36.8 313 400 400 20 1.29 10 0.44 24.0 487 20.6 1,600
aSstrengthened members with jacketing, Mspecimens built monolithically, rroughened interface, ssmooth interface, D, Nspecimens with or without dowels, respectively, pdpredamaged units,
15, 25, 30, and 45: stand for the lap splice length corresponding to 15Db,25Db,30Dband 45Db;bU-shaped steel links utilized, wwelding of stirrup ends of the first four stirrups, the numeral in the
end is the specimen number considered in successive order.
bOriginal code names used for the specimens by the original investigators.
cAxial load ratio % calculated on the basis of concrete strength of the jacket.
dJacket constructed under axial load.
1994, Fig. 4the previous damage of units SrNpd-1 and SrNpd-4
had no significant influence on the response as compared to units
SrN-2 and SrN-3, which had not suffered any damage prior to
jacketing. Clearly, the analytical result is very close to its experi-
mental counterpart in the case of specimens MsN-7 and SsNpd-5
Gomes and Appleton 1998, Fig. 5. The experimental curve rep-
resenting specimen SsNpd-5 lies below that of specimen MsN-7
and this is attributed to the initial damage of unit SsNpd-5. In the
third group Bousias et al. 2004, Fig. 6the experimental response
of the units seems insensitive to the lap splice length of existing
reinforcement and to the degree of previous damage imparted to
units SsNpd15-12 and SsNpd25-13. This is also observed in the
case of the fourth group of units Bousias et al. 2004, Fig. 7.In
the last group of units Vandoros and Dritsos 2006a,b Fig. 8the
estimated strength of the monolithic unit MsN-17matches the
experimental evidence but the actual secant to yield stiffness is
lower. The response of unit SsNw-22 is very close to the response
of the monolithic approach, although slip at the interface modifies
the response somewhat, as shown by the analytical curve.
The response of the jacketed members is influenced greatly by
the interface model utilized. A more sensitive model that could
describe in more detail the interface shear behavior would provide
better results. In general, a softer response than the experimental
envelope implies too compliant an interface, whereas the opposite
trend implies the interface stiffness has been overestimated. This
is demonstrated in the following sections, where a parametric
investigation of the model’s sensitivity is explored. Interface be-
havior requires further calibration, and this would have been done
if a critical mass of experiments were available. However, even as
things stand, by explicitly accounting for this aspect in calculating
the flexural response of composite members, the model introduces
a degree of freedom that enables consideration of an important
response mechanism that was previously overlooked.
Parametric Investigation
A parametric investigation is conducted in the present section so
as to establish the sensitivity of the monolithic factors to the
important design and model variables. Note that these factors
are used to estimate the response indices of jacketed, composite
reinforced concrete members, from the corresponding response
variables of monolithic members with identical cross section, on
the premise that the latter quantities are easily established from
conventional flexural analysis. The magnitude of monolithic fac-
tors depends on the property considered strength, stiffness, or
deformation, on the jacket characteristics and on the interface
Parameters of Study
A sensitivity analysis of monolithic factors is conducted in this
section through a detailed evaluation of two reference cases. The
core of the composite member is the existing member, represen-
tative of former construction practices. In Case 1 the core used
had a 350 mm square cross section, reinforced longitudinally
with a steel area ratio, lc, equal to 1% and transverse confining
reinforcement ratio wc =0.13% perimeter stirrups Ø6 / 200 mm.
Concrete cylinder uniaxial compressive strength fc
was 16 MPa
and steel yield strength fsy was 300 MPa. In Case 2 the core had
a rectangular cross section of 250 by 500 mm, with a longitudinal
reinforcement ratio, lc =0.8%, transverse confining reinforcement
ratio wc =0.24% perimeter stirrups Ø8 mm/ 200 mm, concrete
uniaxial compressive strength fc
=16 MPa, and steel yield
strength fsy =300 MPa. In both cases the jacket considered was
Fig. 4. Lateral load versus drift for first group of units adapted from
Rodriguez and Park 1994
Fig. 5. Lateral load versus drift for second group of units adapted
from Gomes and Appleton 1998
Fig. 6. Lateral load versus drift for third group of units adapted from
Bousias et al. 2004
75 mm thick. After application of the jacket the shear span ratio
was reduced from 4.3 to 3 flexure dominatedand from 3 to 2.3
shear dominatedfor the two case studies, respectively.
Parameters of the investigation were the percentage of
the longitudinal reinforcement of the jacketed cross section
lJ =AJ/bJhJbchc兲兴 which varied between 1 and 3%, the trans-
verse confining reinforcement ratio of the jacket wJwhich var-
ied for the square cross section Case 1between 0.3 and 1.25%,
and for the rectangular cross section Case 2between 0.4 and
1.75% and the axial load Napplied on the jacketed cross section
expressed as a fraction of the theoretical crushing capacity Agfc
of the jacketed cross section which varied between 0 and 0.3. The
cylinder compressive strength of the jacket concrete was taken as
=20 MPa. Yield strength of both longitudinal and transverse
jacket reinforcements was taken as fsy = 500 MPa.
The results of the parametric study are presented in terms of
the monolithic factor values both for flexural strength and for
deformation capacities. In this regard, the following three defini-
tions are adopted for the objectives of the study
where KM,K, and Kmonolithic factors for flexural strength,
curvature, and ductility. Subscripts yand uyield and ultimate,
respectively; whereas and ⌬⫽curvature and displacement duc-
tilities. The moments at yield, My, and ultimate, Mu, of the RC
jacketed member are estimated by multiplying the corresponding
moments, My,M, and Mu,M, of the monolithic member with factors
Mand Ku
MEq. 17a兲兴. Pertinent monolithic factors Ky
and Ku
may be used in the same way in order to obtain the curvature
at yield, y, and ultimate, u, of the RC jacketed members
Eq. 17b兲兴. Similarly, by multiplying the curvature ductility ,M
and the displacement ductility ,M, of the monolithic cross sec-
tion with appropriate monolithic factors K,and K,, the curva-
ture ductility and the displacement ductility of the jacketed
member may be estimated Eq. 17c兲兴.
Role of Characteristics of Jacket
The direct effect induced by any change in the design character-
istics of the jacket is depicted for both the yield and the ultimate
stage in Fig. 9. The circular mark in Fig. 9 corresponds to the
reference case of the parametric study with lJ =1%, wJ= 0.3%,
and N=0 for the square section example, and lJ =1%,
wJ =0.4%, and N= 0 for the rectangular one. The arrows indicate
the influence on the monolithic factors plotted in the xand yaxes,
effected by a corresponding change in the parameter studied.
With reference to the square cross section Case 1, increasing
both the percentage of longitudinal reinforcement of the jacket
lJ and the applied axial load ratio N=N/Agfc
results in a
reduction of Ky
Mand an increase of Ky
Fig. 9a兲兴. This is also
observed in the shear dominated member Case 2, however, the
influence is less pronounced, especially on Ky
M. This indicates that
flexure-dominated members are more sensitive to changes of
axial load and longitudinal reinforcement compared to the shear-
dominated ones. As discussed earlier, the jacketed member
reaches a yield at lower strength but at increased curvature as
compared to its monolithic counterpart, owing to the increased
Fig. 7. Lateral load versus drift for fourth group of units adapted from Bousias et al. 2004
Fig. 8. Lateral load versus drift for fifth group of units adapted from
Vandoros and Dritsos 2006a,b
deformation due to interface slip. The opposite is observed when
confinement reinforcement of the jacket wJis increased since
interface slip is suppressed with confinement i.e., the cross sec-
tion approaches more toward the monolithic condition.
Results of the parametric investigation at a nominal ultimate
limit state for both reference Cases 1 and 2 are presented in
Fig. 9b. The nominal ultimate is taken here to correspond to a
lateral drift of 2% for both the analytical and monolithic model.
This level was selected as a performance limit state and a point of
reference as it corresponds to a displacement ductility in excess of
3 for regular frame members, which is considered an upper bound
for the acceptable level of ductility demand in a redesigned struc-
ture. Increasing the longitudinal jacket reinforcement ratio lJ
and applied axial load ratio Nproduce a simultaneous reduc-
tion in the monolithic factors for strength and deformation at
ultimate, whereas the reverse effect is obtained by increasing the
amount of jacket confinement reinforcement wJFig. 9bfor
Case 1. In the case of the shear-dominated member Case 2the
response is differentiated with regards to the axial load influence:
as the axial load ratio increases the monolithic factor for strength
at ultimate also increases, whereas the monolithic factor for de-
formation at ultimate decreases.
The influence that each of the parameters under investigation
has on the various monolithic factors is depicted in Figs. 10–12
and is discussed in detail in the following subsections for both
the square Case 1and rectangular Case 2cross sections,
Longitudinal Jacket Reinforcement Ratio lJ
,K,, and K,are all reduced with increasing value
of this variable Fig. 10. The reverse trend is observed for Ky
Axial Load N
Increasing the applied axial load ratio Nleads to a simulta-
neous reduction of Ky
M, and Ku
Fig. 11, but also of K,and
K,. The monolithic factor of curvature at yield Ky
for an axial load ratio up to 0.2, but the trend is not uniform for
both cases Case 1 and 2at higher axial loads.
Confining Reinforcement wJ
As illustrated in Fig. 12, Ky
Mand Ku
Mincrease mildly as the per-
centage of jacket confining reinforcement wJincreases. Be-
cause the dowel function of transverse reinforcement is mobilized
passively, Ky
is almost insensitive to wJ, whereas there is a
strong increase of Ky
with confinement. Similarly, K,and K,
both increase with wJ.
Fig. 9. Monolithic factors of strength and deformation at: ayield;
Fig. 10. Influence of jacket longitudinal reinforcement on monolithic
factors Case 1, Case 2
Fig. 11. Influence of axial load on monolithic factors Case 1,
Case 2
Fig. 12. Influence of jacket confinement reinforcement on monolithic
factors Case 1, Case 2
Discussion of Results of Parametric Study
The results of the current parametric study provide an insight
into the mechanical effect that jacket characteristics play on
the lateral load response of jacketed members. Monolithic factors
are sensitive to the design variables of the jacket and do
not generally assume an obvious fixed value. From the results of
the parametric study for both the square Case 1and the rectan-
gular Case 2cross section the values of the monolithic factors
range as follows: 1Ky
=0.340.90;5K,=0.15– 0.93;
and 6K,=0.390.94.
The above results are consistent with the values suggested by
EC8 CEN 1996for the monolithic factor of strength KR=0.8 no
differentiation is made by the code between yield and ultimate,
although the range of estimated values is larger for the ultimate
M. The estimated values for Ky
show that jacketed cross
sections reach yield at greater curvatures, owing to slip at the
interface between the existing member and the core. The Ku
less than 1.0, thus, in general the curvature at ultimate uesti-
mated from the analytical approach is smaller than the monolithic
estimate u,M. Considering that slip at the upper and bottom
interfaces contributes to lateral drift, the reduced value of curva-
ture at ultimate 2% driftdefined by the analytical approach is
justified. The monolithic factors of curvature and displacement
ductilities K,,K,are less than 1.0, which emphasizes that
analytical curvature and displacement ductilities are both lower
than the corresponding monolithic values.
Sensitivity of Analytical Model
The proposed analytical model is primarily sensitive to param-
eters that affect the estimation of crack spacing and the shear
strength of the contact interfaces. Each of these variables has a
distinct influence on the computational procedure; however,
selection of the shear interface model is fundamental. Variables of
the shear transfer model used herein Tassios and Vintzēleou
1987; Vintzēleou and Tassios 1986, 1987; Vassilopoulou and
Tassios 2003are the interface strength vc,uand the slope of the
postpeak branch Fig. 13a兲兴.
A brief parametric investigation was conducted in order to
explore the sensitivity of the model to the primary variables.
The square cross section used in the preceding as Case 1 is
used as a point of reference. Geometric characteristics and mate-
rial properties of the existing member core were already given
in earlier sections. Longitudinal jacket reinforcement ratio
was selected as lJ =1%, with transverse confining reinforce-
ment wJ =0.3% fsy =500 MPa. No axial load was applied on
the jacketed cross section, whereas fc
=20 MPa for the jacket
First, the influence of the interface shear friction on the re-
sponse of the jacketed member was studied. To model unfavor-
able conditions at the interface, the shear friction coefficient is
increased stepwise up to 0.65 values used are: 0.4, 0.55, 0.65
while keeping =1; this coefficient indirectly accounts for the
roughness of the interface. The steepness of the descending
branch was examined for =1, 1 / 2, and 1 / 3, whereas vc,u
was given by Eq. 4. The values selected for are based on
published experimental data Bass et al. 1989; Papanicolaou and
Triantafillou 2002. The results of the parametric investigation
are summarized in Fig. 13 in a moment versus curvature diagram.
For lower values of , i.e., more gradual decay of the descending
branch of the shear stress strain curve Fig. 13a兲兴, higher levels
of curvature capacity are estimated Fig. 13b兲兴. Increasing
leads to higher shear capacity at the contact surface allowing
for the development of higher strength and curvature values
Fig. 13c兲兴.
Summary and Conclusions
An algorithm for calculating the monotonic response of rein-
forced concrete jacketed members is presented. The model intro-
duces a kinematic degree of freedom interface slipthat enables
consideration of an important mechanism of behavior that was
previously overlooked, namely the shear transfer mechanisms
mobilized due to sliding at the interface between existing and new
material. The weak link controlling deformations in this problem
is the interface. The capacity of the weakest link is evaluated and
checked in every step, to make sure it is not exceeded by the
demand. The shear demand at the interface is controlled by the
flexural stresses on the cross section and by the spacing of cracks
in the longitudinal direction, whereas the shear capacity is a func-
tion of slip. The shear stress slip relationship for the contact sur-
faces and the definition of crack spacing play a key role in the
algorithm. Analytical results show that the model can reproduce
successfully the observed response of jacketed members and cor-
relates well with experimental data. This analysis tool was used to
explore the difference between the ideal response of monolithic
members and the actual response of the RC jacketed members of
identical geometry with reference to the design variables. A para-
metric study was conducted and the dependence of various mono-
lithic factors on the characteristics of the jacket was investigated.
It was found that strength factors at yield Ky
Mrange between
0.63 and 0.95, whereas strength factors at ultimate Ku
between 0.49 and 0.97. Monolithic factors for deformation indi-
ces were found in the case of curvature at yield to range between
0.95 and 2.57, whereas in the case of curvature at ultimate be-
tween 0.34 and 0.90. The monolithic factors of curvature and
displacement ductilities K,,K,are both lower than the cor-
responding monolithic values with the former to range between
0.15 and 0.93 and the latter between 0.39 and 0.94.
Fig. 13. Role of shear friction interface model
The first writer was co-funded by the European Research Project
Seismic Performance Assessment and Rehabilitation” SPEAR,
through Imperial College of Science, Technology and Medicine
London, UKand by the Hellenic Ministry of Education and
Religion Affairs through the scholarship “HRAKLEITOS.” The
contribution of the second writer was funded by the Hellenic
Secretariat of Research and Technology GSRTthrough the
multi-Institutional Project “ARISTION.” The contribution of the
third author was funded by the US National Science Foundation
through the Mid-America Earthquake Center, Award No. EEC
The following symbols are used in this paper:
Aggross section area;
Aint contact area of interface;
bJwidth of jacketed cross section;
caverage crack spacing;
Dbdiameter of bars clamping interface;
Db,c,DbJ bar diameter of core and jacket longitudinal
reinforcement, respectively;
Eselastic modulus of steel;
Ecm elastic modulus of concrete;
fb,c,fb,Javerage bond stress of core and jacket
reinforcement layer, respectively;
concrete cylinder uniaxial compressive
fctm tensile strength of concrete;
fctm,ctensile strength of concrete core;
fsaxial stress of bars crossing interface;
fs,cr jacket steel stress at crack;
fsy yield strength of steel;
Imoment of inertia of composite cross
jddistance between upper and bottom
Kimonolithic factor, where subscript i=R,K
refers to strength and stiffness, respectively;
KM,K,Kmonolithic factors for flexural strength,
curvature, and ductility, respectively;
knumber of dowels;
Lsshear span of member;
lcheight of tension zone in core component of
composite cross section;
lplength of plastic hinge region;
Mmoment resultant;
Naxial load;
applied axial load ratio;
Nc,NJnumber of bars in tension steel layer of core
and jacket, respectively;
Next externally applied axial load;
pnormal pressure externally applied on
contact plane;
qsshear flow at contact between bottom layer
and core;
q0shear flow based on classical mechanics;
Sifirst moment of area;
slateral slip;
sc,uultimate slip value beyond which frictional
mechanisms break down;
Sd,el elastic slip value;
Sd,uultimate slip value;
s1u,s2uslip values at upper and bottom interfaces of
Vshear force on member;
VDsdowel force estimated for slip magnitude, s;
VD,uultimate dowel force;
vashear resistance of aggregate interlock
vcsfrictional resistance at slip, s;
vc,uultimate frictional resistance of interface;
vDshear stress resisted by dowel action in
cracked reinforced concrete;
vd,1,vd,2 shear demand values;
vnshear resistance;
v1,v2shear resistances at upper and bottom
wseparation between contact surfaces as they
slide overriding one another;
␣⫽bar axial stress normalized with respect to
its yield value;
p,uplastic displacement;
slip,udisplacement owing to interface slip;
yelastic displacement at yield;
0total tip displacement;
⌬␧1,⌬␧2magnitude of strain discontinuity at upper
and bottom interfaces, respectively;
cconcrete strain at extreme compression fiber;
c1,j2,j3,c2normal strains in section layers above and
below contact surfaces;
sy tensile reinforcement yield strain;
=Es/Ecmmodular ratio;
slip,urotation owing to interface slip;
␭⫽coefficient indirectly accounting for
roughness of interface;
␮⫽interface shear friction coefficient;
displacement ductility;
curvature ductility;
lc percentage of longitudinal reinforcement of
existing cross section core;
lJ percentage of longitudinal reinforcement of
jacketed cross section;
sreinforcement area ratio;
s,eff effective reinforcement ratio;
wc transverse confining reinforcement ratio of
existing cross section core;
wJ transverse confining reinforcement ratio of
jacketed cross section;
Finormal stress resultant of each layer, i;
Nnormal clamping stress acting on interface;
xy shear stress acting on plane with unit normal
parallel to x-axis, and oriented in y-axis
defined according to classical mechanics as
per Fig. 2;
xz shear stress acting on plane with unit normal
parallel to x-axis, and oriented in z-axis,
respectively Fig. 1d兲兴;
ucurvature at ultimate;
eq equivalent monolithic curvature;
ycurvature at yield of composite section; and
y,Mcurvature at yield of monolithic cross
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... Among the parameters for which the provisions have been revised are the monolithicity factors for jacketed members; currently these (unnamed) factors are just constant values (ranging between 0.9 and 1.05) that multiply the resistance parameter (strength or deformation) calculated assuming a monolithic section, while in the new Code some of them, become functions of the axial loading. This paper presents the background of these factors adopted in the new Eurocode 8-3 for jacketed reinforced concrete columns; specifically, we present the derivation of expressions by regression of experimental results for jacketed columns, as well as those resulting from the processing of the results of an extended parametric analysis of R/C columns retrofitted by the addition of R/C jackets using a model developed by Thermou et al. [6][7][8] that takes into account slip at the interface. It is noted that simpler analytical models have also been put forward [9] for jacketed members which do not account for slippage at the core-jacket interface. ...
... The analytical model for predicting the flexural response of existing R/C members strengthened with concrete jacketing under monotonic and cyclic loading conditions introduces a degree of freedom allowing the relative slip at the interface between the existing member and the jacket [6][7][8]. Slip along the member length is attributed to the difference in normal strains at the contact interfaces ( Figure 1). For flexural analysis, the crosssection is divided into three layers that bend with the same curvature, φ ( Figure 1). ...
... where ΣFi is the layer force resultant and bJ is the width of the jacketed cross section; sr is the crack spacing (distance between adjacent cracks assumed to be constant), defined as follows [7]: ...
Full-text available
The paper presents the background to the expressions adopted in the new Eurocode 8—3 for jacketed reinforced concrete columns. These are based on the commonly adopted concept of monolithicity factors (ratios of resistance of the jacketed section to that of an identical monolithic one). These factors are derived here in two ways: (i) by fitting experimental results for jacketed columns and (ii) by an extended parametric study of substandard reinforced concrete (R/C) members that were retrofitted by adding R/C jackets, analysed using a model developed by the authors that takes into account slip at the interface. Apart from the cross-section geometry and the thickness of the jacket, parameters of the investigation were the material properties of the core cross-section and the jacket, as well as the percentage of longitudinal reinforcement of the jacket and the percentage of dowels placed to connect the existing member to the jacket. It was found that the parameter that had the most visible effect on these factors was the normalised axial load (ν). The finally adopted factors are either simple functions of ν or constant values.
... The limited displacement ductility of shear-deficient reinforced concrete (RC) columns could cause catastrophic building collapse under earthquakes [1]. Existing retrofitting techniques mainly include concrete jacketing (e.g., [2][3][4][5][6]), fiber-reinforced polymer wrapping (e.g., [7][8][9][10]), and metal jacketing (e.g., [11][12][13]). Their performance has been commonly assessed using columns with a low-to-medium axial load ratio of less than 0.3 (e.g., [13,14]), where the axial load ratio is defined as the ratio of the axial load demand to the axial load capacity of the column. ...
... To understand the influence of the ratio of the axial load taken by the UHPC jacket to the total axial load demand (R axial ) in the shear capacity of the retrofitted column, the shear capacity with different values of R axial is assessed using Eqs. (3)(4)(5). Fig. 16 shows the relationship of the calculated shear capacity of the retrofitted columns and R axial . It can be seen that the nominal shear capacity of the retrofitted column increases with R axial . ...
Full-text available
Ultra-high-performance concrete (UHPC) is a class of concrete materials that have high strength and ductility capacities. This study developed two UHPC jacketing methods to improve the seismic performance of shear-deficient RC columns in mid- to high-rise buildings without enlarging the cross-section of the columns. In the first method, the original concrete cover was replaced with a cast-in-place UHPC jacket that was reinforced by steel bar mesh. To simplify the on-site construction process, the second method utilized prefabricated UHPC panels incorporating steel bar mesh. The performance of the proposed strengthening methods was evaluated using cyclic loading tests on three full-scale RC columns with a high axial load demand. The test results showed that the developed UHPC jacketing methods improved the lateral strength and drift capacity of the shear-deficient column by at least 40% and 30%, respectively. They also reduced the residual displacement of the column by 65%-80%. Compared to the cast-in-place jacketing, the discontinuity in the prefabricated UHPC jackets inhibited full composite action of the jacketing, thus reducing the retrofitting efficiency. In addition to the experimental study, evaluation methods were proposed to evaluate the strength and behavior of the retrofitted columns.
... Although concrete jacketing method is effective and useful for strengthening purposes [4][5][6][7][8][9][10][11], however, there are few disadvantages of concrete jacketing method. For example, increase in weight and volume of the strengthened concrete column. ...
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The main objective of this study was to investigate the monotonic and cyclic axial stress versus strain response of the concrete confined with different types of fiber reinforced polymer (FRP) composites such as cotton fiber rope reinforced polymer (CFRRP), glass fiber reinforced polymer (GFRP) composites and carbon fiber reinforced polymer (CFRP) composites. Another objective was to investigate the applicability of existing ultimate strength models to predict the ultimate strength of concrete confined with CFRRP, GFRP and CFRP composites. For this, a total number of 28 concrete cylinders were cast, strengthened and tested under pure axial compression. The concrete cylinders were strengthened with two and four layers of FRPs. The experimental results indicate that all kinds of FRPs are very useful to alter the ultimate strength and strain of the confined concrete. The highest increase in ultimate strength was recorded for CFRP composite confined specimens and the highest increase in ultimate strain was observed for CFRRP composite confined specimens. Further, both ultimate strength and ultimate strains are found higher in the case of cyclic loads compared with the monotonic load. In general, the stress versus strain curves of cotton fiber rope and glass FRPs are found trilinear. Whereas, stress versus strain curves of CFRP confined specimens are observed to be bilinear. Existing ultimate strength models developed for CFRRP and CFRP composites are found well accurate to predict the ultimate strengths of CFRRP and CFRP confined concrete specimens tested in this study.
... Thermou et al. [12] studied the composite behaviour of RC members repaired/strengthened by RC jacketing through a dual-section analytical approach. The model included the relative slip at the interface between the old member and the new RC jacket. ...
The design of reinforced concrete (RC) jacketing for retrofitting substandard columns is a common practice when approaching the structural rehabilitation design of existing structures. In this framework, international technical codes suggest to assume the hypothesis of monolithic section and finally affecting the deduced performances by some empirical factors, known in the literature as monolithicity factors. The assessment of these coefficients is a difficult task, and only some rough assumptions can be found in the modern codes. This paper presents the development of a software package able to perform the static non-linear analysis of RC jacketed columns and evaluating the actual value of the monolithicity factor. The code is developed in an integrated environment, taking advantage of finite element (FE) framework OpenSees and adopting Matlab for developing an user-friendly Graphical User Interface (GUI). The fibre-section approach is automatically implemented for modelling the flexural behaviour of the column and the effective bond between the old and new column in considered through a distribution of non-linear springs. Finally, the parametric analysis performed by the proposed tool allows to present the monolithicity factors as a function of the key parameters.
... Various techniques and materials have been utilized in recent years for the rehabilitation of existing concrete structures. Common retrofitting techniques include concrete or steel jacketing [5,6], fiber-reinforced polymer (FRP) [7] wrapping, external prestressing [8], and near-surface mounted and fiber-reinforced composite materials [9][10][11][12][13]. Although these techniques can effectively achieve strengthening objectives and improve strength and durability, they have some disadvantages, such as increasing the weight of structures in concrete jacketing, corrosion, fire resistance, deboning in steel plate strengthening, and aging of adhesion materials at the interface in FRP strengthening [14]. ...
A quick increase in the traffic loads (repetitive) and aging deteriorates the bridge's life. To increase the life of the bridge, techniques like overlaying are developed & implemented. Concrete with higher compressive strength and appreciable interfacial bond, which can enhance the life of the deck slab, is used as an overlaying material. In this process, researchers developed a ground-breaking next-generation concrete named “Ultra-High-Performance Fiber Reinforced Concrete (UHPFRC).” UHPFRC gives a minimum 120 MPa compressive strength and 7 Mpa tensile strength, respectively, which is much higher than conventional concrete. The novelty of this paper is to highlight the constituents and exceptional properties of UHPFRC, like compressive strength, fatigue life, and interfacial bond with the already existing substrate, making out suitable material for overlaying. Further, the use of UHPFRC as an overlay in various structural elements has been studied. The applications of UHPFRC around the world for rehabilitation purposes are also highlighted by discussing various case studies that should be added.
... With all this knowledge, this new study has been carried out. Researchers have conducted studies [14][15][16][17][18][19][20][21][22][23][24] about jacketing of beams which also refer to this study. The aim of this study is to reveal the effect of jacketing of beams, which is a practical structural reinforcement method, on the dynamic behavior of a model reinforced concrete structure. ...
... Externally bonded fibre-reinforced polymers (FRP) have gained popularity due to good corrosion resistance, great durability, high strength to weight ratio, and adaptation of the technology to various structural elements (Triantafillou 1998;Fukuyama et al. 2000;GHOBARAH * and SAID † 2001;Stephen and P. 2001;Ghobarah and Said 2002;Júlio et al. 2003;E. et al. 2007;A. and H. 2007;Triantafillou and Papanicolaou 2013;Thermou et al. 2016). Such a technology allows for significant gains in deformation capacity, which is especially important in critical places where massive plastic deformations are predicted (Campione et al. 2007;Gonzalez-Libreros et al. 2019). Furthermore, confining RC sections with FR ...
The compressive behavior of short concrete columns with square cross-sections confined with steel fiber reinforced polymer and grout composites, named SRP and SRG composites, was investigated experimentally in this paper. Comparing the loading capacity and ultimate strain of confined specimens to those of unconfined samples was used to assess the confining effect. The influential parameters such as the density of steel fibers, cross-section geometry at corners, surface preparation, and the number of confining layers were studied. The results showed that the general failure mode of SRP-and SRG-confined concrete columns were the detachment of overlapping faces of fibers, and separation between the core concrete and the composite, respectively. The average compressive strength and strain at maximum load, respectively, were 7.9% and 9.8% higher in sharp-cornered specimens compared to rounded-cornered specimens. The treatment of specimen surfaces caused a 5.1% and 13.4 % increase in compressive strength and strain at maximum load, respectively. The effect of fiber density on increasing strength and strains did not match with density increments. Furthermore, in SRG-confined concrete columns, increasing the number of overlapping faces improved axial strength.
... Bett et al. [8] observed flexural cracks in columns reinforced with concrete jackets at a relative lateral displacement level of less than one percent. Studies by Thermou et al. [9] showed that concrete jacketing increases the stiffness and strength of a structural member. The change in the deformable capacity of the member depends on factors such as the aspect ratio of the jacketed member and the limiting factors of deformable capacity in the initial conditions of the structure. ...
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Given the deterioration of civil infrastructure throughout the world, developing more efficient repair and strengthening is essential. Jacketing is one of the most common methods for retrofitting reinforced concrete (RC) columns. Notably, using engineered cementitious composite (ECC) within the jacketing area increases the bearing capacity and significantly enhances the ductility of the columns. The recent development of ECC concrete with suitable compressive strength and higher ductility of about 5 % can significantly enhance the performance of reinforced concrete structures. The behavior of retrofitted RC columns depends heavily on the cohesion between the jacket and the original column as well as the mechanical properties of the jacketing materials. This study investigates jacketed square and circular RC columns using ECC and conventional/normal concrete (NC) using different casing techniques to retrofit RC columns, namely galvanized mesh, U and L-shape joints, removing the cover, core drilling, and integrated models. All specimens were subjected to a compression test. The results indicate that in both square and circular specimens, the use of ECC as a super ductile material and vertical U-shaped elements to connect the longitudinal rebars of the casing and the core leads to much higher ductility and bearing capacity than in NC specimens. These elements also showed suitable ductility because of using ECC as a super ductile material. In order to optimize these methods, finite element analysis (FEA) was conducted using Abaqus software to verify experimental models, as well as a parametric study to achieve an optimum design of the jacketing.
This paper presents the structural performance of jacketed concrete columns subjected to cyclic lateral loading using ultrahigh-performance fiber-reinforced concrete (UHPFRC) with various advantages as construction materials of high strength and durability with self-compacting. The UHPFRC jacketing with/without textile reinforcement for concrete columns for the seismic retrofitting demonstrates its strengthening efficiency and reduction of jacket thickness through the experimental program in this study. Especially, textile-reinforced UHPFRC (TR-UHPFRC) jacketing has enhanced ductility compared with that of UHPFRC jacketing by changing critical failure crack patterns in jackets. The satisfactory interface shear strength between UHPFRC jackets and substrate concrete of columns was confirmed by a series of slant shear test programs provided that required roughness on substrate with moisture were prepared according to current design guidelines. This study also investigates the effects of discontinuity of tension forces in jackets at the floor level of columns on shear load paths, which are intended to alleviate jacketing construction for columns in existing buildings as a member intervention technique. Shears and axial forces induced in jacketed columns under lateral loading are mainly resisted by compression forces in diagonal struts through the jacket with substrate concrete columns and are explained by arch actions for enhancement of shear strength after jacketing.
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Highlights of the evolution over the past two decades of a seismic design strategy, used in New Zealand for reinforced concrete buildings, are reviewed. After a brief outline of some philosophical concepts of the capacity design methodology, the main features of its application with respect to ductile rigid jointed frames, structural walls and hybrid structural systems are sketched. Another aim of this strategy, complementary to ductility requirements, is to strive for high quality in detailing. Numerous examples are presented to illustrate how this can be achieved. A specific intent of this state of the art review is to report on features of design and detailing which are considered to have originated primarily in New Zealand.
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It is shown, from analysis of typical reinforced concrete beam sections, that current design practice, which assumes beam stiffness is independent of reinforcement ratio but equal to a constant fraction of gross section stiffness is inappropriate. The analyses indicate that effective beam yield curvature can be considered constant, when non-dimensionalized by beam depth and yield strain, indicating that beam stiffness is proportional to strength. Based on this observation, a simple expression for yield drift of frames is proposed and is calibrated by comparing with results of a large number of beam/column subassemblage experiments. Good agreement is obtained. It is pointed out that current estimates of frame stiffness are generally too high. A consequence is that simple calculations show that the vast majority of frame buildings will be unable to achieve code design ductility levels before exceeding code drift limitations.
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Moment-curvature analyses of cantilever shear walls are used to show that yield curvature, serviceability curvature, and ultimate (damage-control) curvature are insensitive to variations of axial load ratio, longitudinal reinforcement ratio, and distribution of longitudinal reinforcement. The results are used to determine available displacement ductility factors for walls of different aspect ratios and drift limits. It is shown that drift capacity will generally exceed code levels of permissible drift, and that code drift limits will normally restrict, sometimes severely, the design displacement ductility factor.
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A new spectral representation of seismic demand is described for use in the seismic design of new structures and in the evaluation and rehabilitation of existing structures. Yield Point Spectra (YPS) retain the intuitive appeal of the Capacity Spectrum Method (Freeman 1978) and join the Nonlinear Static Procedures of FEMA 273/274 (1997) and ATC 40 (1996) for use in estimating displacement demands. YPS also may be used to establish admissible combinations of strength and stiffness for the design of new structures to limit system ductility and drift to arbitrary values. Graphical procedures allow admissible design regions to be established to satisfy multiple performance objectives. YPS computed for 15 ground motions classified as Short Duration, Long Duration, or as containing near-fault Forward Directivity pulses are presented for bilinear and stiffness-degrading hysteretic models.
The effectiveness of three different repair and and/or strengthening techniques in enhancing the lateral load response of identical reinforced concrete short columns was studied. Based on an 18 in. square prototype, three column test specimens were constructed to two-thirds scale, using identical geometry and reinforcement. The test specimens representing existing columns intended to represent typical practice of column design in seismic regions of the U. S. in the 1950s and early 1960s. One of the specimens was tested, repaired by jacketing, and then retested. The remaining two specimens were strengthened by jacketing prior to testing. A single lateral displacement history and constant axial load were used for all tests. Both the strengthened and the repaired columns performed better than the original column. Columns strengthened by jacketing, both with and without supplementary crossties, were much stiffer and stronger laterally than the original, unstrengthened column.
An experimental program was designed to provide information on the interface shear capacities between new concrete cast against an existing concrete surface. Thirty-three full-scale, pushoff specimens were designed, constructed, and tested. Test variables included surface preparation, amount and depth of embedment of the interface reinforcement, reinforcement details in the new and existing concrete elements, and the compressive strength of both new and existing concrete elements. Testing consisted of repeated load cycles producing shear stresses along the plane of the concrete interface. Deflections along and across the concrete interface at several locations were measured. Peak strength, degradation of strength with repeated load cycles, and increasing deflections were observed. Special attention was given to determining the failure mechanism along the interface.
The behavior of reinforced concrete structures subjected to earthquakes may be greatly influenced by the behavior of cracked critical zones. This paper presents an experimental investigation of dowel-to-concrete shear transfer. Emphasis is given to identifying dowel mechanisms under cyclic actions. In addition, recommendations for the calculation of dowel strength are proposed, both under monotonic and cyclic loading conditions.
Two series of tests were made to study the behavior of jacketed columns. The first series consisted of uniaxially loaded specimens, and the behavior of jacketed columns was compared with a monolithic reference specimen. The main objective of this series was to study the effectiveness of repair and strengthening jackets and the differences between jackets made under load and after unloading. Specimens with repair and strengthening jackets behaved well when jacketing was introduced after unloading. Repair jacketing made load exhibited poor behavior. In the second series, jacketed columns were tested under combined axial load and bending (monotonic and reversed cyclic). Two monolithic specimens also were tested to serve as reference specimens. Effectiveness of repair and strengthening jackets was studied considering strength, stiffness, and energy dissipation. The influence of load history (monotonic and reversed cyclic) on the behavior of jacketed columns also was studied. Repair and strengthening jackets behaved well both under monotonic and reversed cyclic loadings.