Corrosion-Damaged Reinforced Concrete Beams Repaired with
Fabric-Reinforced Cementitious Matrix (FRCM)
Mohammed Elghazy1*, Ahmed El Refai 2, Usama Ebead 3, and Antonio Nanni 4
The structural performance of corrosion-damaged reinforced concrete (RC) beams repaired with
fabric-reinforced cementitious matrix (FRCM) was investigated. Eleven large-scale RC beams
were constructed and tested in flexure under four-point load configuration. Nine beams were
subjected to an accelerated corrosion process for 70 days to obtain an average mass loss of 12.6%
in the tensile steel reinforcing bars while two other beams were tested as controls. One corroded
beam was repaired with carbon fiber-reinforced polymer (CFRP) before testing for comparison.
The test parameters included the number of fabric plies (1, 2, 3, and 4), the FRCM repair scheme
(end-anchored and continuous U-wrapped strips), and FRCM materials (carbon and
polyparaphenylene benzobisoxazole (PBO)). Test results showed that corrosion slightly reduced
the yield and ultimate strengths of the beams. The use of FRCM increased the ultimate capacity of
corroded beams between 5% and 52% and their yield strength between 6% and 22% of those of
the uncorroded virgin beam. Beams repaired with continuous U-wrapped FRCM strips showed
higher capacity and higher ductility than those repaired with the end-anchored bottom strips having
similar number of layers. A high gain in the flexural capacity and a low ductility index were
Ph.D. Candidate, Department of Civil and Water Engineering, Laval University, Quebec City, Quebec, G1V 0A6, Canada. Email: [email protected]
2 Associate Professor, Department of Civil and Water Engineering, Laval University, Quebec City, Quebec, G1V 0A6, Canada. Email: [email protected]
3 Associate Professor, Department of Civil and Architectural Engineering, College of Engineering, Qatar University P. O. Box 2713, Doha, Qatar. Email: [email protected]
4 Inaugural Senior Scholar, Professor and Chair, Department of Civil, Architectural & Environmental Engineering, University of Miami, 1251 Memorial Drive, Coral Gables, FL 33146-0630, USA. Email: [email protected]
reported for specimens with high amount of FRCM layers. A new factor was incorporated in the
design equations of the ACI-549.4R-13 to account for the FRCM scheme.
Authors’ keywords: Corrosion; Fabric-reinforced cementitious mortars; Flexure; Reinforced
concrete; Repair; Strengthening.
23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38
Introduction and background
Corrosion of steel reinforcement is one of the main causes of the deterioration of reinforced
concrete (RC) structures. Corroded structures suffer from loss of cross section of the bars, bond
deterioration, and concrete spalling, which can jeopardize the structure’s safety (Torres-Acosta et
al. 2007; Vidal et al. 2007; Xia et al. 2012). Several techniques had been adopted to repair and
strengthen corroded structures, with the use of externally-bonded steel plates and more recently
the epoxy-bonded fiber-reinforced polymers (FRP) being the most common techniques.
Numerous studies have documented the advantages of using FRP as repair materials for
corrosion-damaged structures (Masoud and Soudki 2006; Al-Saidy and Al-Jabri 2011; Malumbela
and Alexander 2011). However, concerns about the poor fire resistance of epoxy (Hashemi and
Al-Mahaidi 2012a), the incompatibility with the concrete substrate (Al-Salloum et al. 2012), and
the loss of ductility of the repaired structures (Hashemi and Al-Mahaidi 2012b) have been widely
reported. In a desire to overcome these drawbacks, the fabric-reinforced cementitious matrix
(FRCM) systems, with their cement-based adhesives, have been introduced as a promising
alternative to the FRP systems (Triantafillou and Papanicolaou 2005; Brückner et al. 2006;
Täljsten and Blanksvärd 2007; Schladitz et al. 2012; Tetta et al. 2015). FRCM systems are
characterized by their lightweight, high tensile strength, corrosion resistance, and ease of
application. More importantly, the mortars in the FRCM composites act as barriers against chloride
ions penetration thus protecting the reinforcing bars from corrosion. Their mechanical properties
are strongly influenced by the fabric’s material and geometry and the ability of the cementitious
matrix to impregnate the fabric. The bond strength at the fabric/matrix interface and at the FRCM
composite/concrete interface greatly affect the performance of the repaired element (Banholzer et
Several studies have reported on the use of FRCM in repairing RC flexural elements
(Triantafillou and Papanicolaou 2005; Brückner et al. 2006; Blanksvärd 2007; Hashemi and Al-
Mahaidi 2012a, b; Al-Salloum et al. 2012; Wang et al. 2013). D’Ambrisi and Focacci (2011)
reported 6 to 46% gain in the load-carrying capacity of RC beams repaired with carbon and PBO-
FRCM systems. Beams repaired with PBO-FRCM performed better than those repaired with
carbon-FRCM due to the superior bond characteristics of the former system at the fabric/matrix
interface. The use of polymer-modified cementitious matrix in the PBO-FRCM improved the bond
of the fabric to the matrix and consequently increased the ultimate capacity of the repaired beams.
In another study, Loreto et al. (2013) reported an increase between 35 and 112% in the flexural
capacity of RC slabs repaired with PBO-FRCM depending on the volume fraction of the fabric
used. The authors reported that increasing the number of FRCM layers reduced the ductility of the
repaired slabs. Slabs repaired with one ply failed due to slippage of the fabric within the matrix
whereas those repaired with four plies failed by fabric delamination at the fabric/matrix interface.
These results were also confirmed by Babaeidarabad et al. (2014).
In a comparison between the performance of FRCM- and FRP-repaired beams, Elsanadedy et al.
(2013) reported that basalt-FRCM systems were less effective than carbon-FRP systems in
enhancing the flexural strength of the beams, yet the FRCM-repaired beams showed more ductility
On the other hand, the feasibility of using FRCM systems to strengthen corrosion-damaged
concrete structures have received little attention. The challenges in repairing corroded RC elements
are two-fold, namely the absence of a sound concrete substrate due to corrosion and the durability
of the repair system should corrosion resumes. To the authors’ knowledge, only one study (El-
Maaddawy and Elrefai 2016) has documented the effectiveness of using basalt and carbon FRCM
systems to restore the ultimate capacity and serviceability of T-beams after a mass loss of 22% in
their tensile reinforcement due to corrosion. It was reported that the basalt-FRCM system could
not restore the original flexural capacity of the beam whereas the carbon-FRCM system restored
109% of the capacity. The authors reported that the use of a combination of internally-embedded
and externally-bonded carbon-FRCM layers was more effective in improving the strength and
ductility of the beams than the use of the same amount of FRCM layers internally embedded within
the corroded-repaired region.
This paper reports the results of the flexural tests conducted on corrosion-damaged RC beams
repaired with FRCM systems. The test program included the type of the FRCM used (carbon and
PBO), the FRCM reinforcement ratios (represented by the number of fabric plies bonded to the
concrete substrate, namely 1, 2, 3, and 4 plies), and the repair scheme (end-anchored and U-
wrapped strips). The paper also reports on the failure modes, the load-carrying capacities, the
ductility, and the straining actions at different stages of loading of the tested beams. Theoretical
formulations are also presented to predict the flexural response of the beams.
Eleven large-scale RC beams were constructed and tested as follows: two specimens were neither
corroded nor repaired (UU), one specimen was corroded but not repaired (CU), seven specimens
were corroded then repaired with different FRCM systems, and one specimen was corroded and
repaired with carbon-FRP (CFRP) sheets. The test matrix is shown in Table 1.
The test specimen was 2800 mm long with a 150 × 250 mm rectangular cross section. All beams
were reinforced with 2-15M deformed bars at the bottom (As = 400 mm2) and 2-8M deformed bars 5
at the top (As' = 100 mm2). The tensile reinforcement ratio was 1.067%. All of the specimens had
a moment span of 800 mm and a shear span of 880 mm. The shear spans were reinforced with
10M deformed stirrups spaced at 100 mm to avoid a premature shear failure. A hollow stainless
steel tube with external and internal diameters of 9.5 mm and 7 mm, respectively, was placed at
80 mm from the specimen tension face to act as cathode during the accelerated corrosion process.
Typical dimensions and reinforcement details of the test specimen are shown in Figure 1.
Accelerated corrosion aging
Salt (NaCl) measured as 5% of the cement weight was added to the concrete mix used to cast the
middle-bottom of the corroded specimens (Figure 1). Corrosion of the main reinforcement was
localized in the middle 1200 mm of the beam’s span with a height of 100 mm. A power supply
was used to impress a constant electrical current of 380 mA to obtain a current density of 180
µA/cm2 in the reinforcing bars. This current density was chosen to obtain corrosion products
similar to those obtained in natural conditions (El Maaddawy and Soudki, 2003). During the
accelerated corrosion process, the bottom reinforcement acted as anode whereas the stainless steel
tube acted as cathode and the salted concrete acted as electrolyte. The test specimens were
electrically connected in series to ensure that the induced current was uniform in all specimens
(Figure 2). All specimens were subjected to wet-dry cycles that consisted of 3 days wet followed
by 3 days dry in a large environmental chamber. The wet-dry cycles provided water and oxygen
necessary for the corrosion process. In this study, a 10% mass loss in the reinforcing bars was
anticipated to represent moderate corrosion damage. According to Faraday’s law, the duration of
corrosion exposure required to achieve this mass loss was 70 days.
Two ready concrete mixes (normal and salted) with similar water/cement ratio were used to cast
the beams. Standard concrete cylinders (150 x 300 mm) were prepared from each concrete batch
and were tested in compression after 28 days and on the day of testing. Table 2 lists the
compressive strengths of both mixes. Prior to fiber application, the corroded beams were repaired
using local commercial cementitious repair mortar (Sikacrete-08SCC) having a compressive
strength of 55.36 MPa (standard deviation of 4.97 MPa) and flexural strength of 3.36 MPa
(standard deviation of 0.26 MPa) as determined by the authors. The nominal yield strength of the
reinforcement steel bars was 400 MPa with elastic modulus of 200 GPa.
Two commercial FRCM systems (PBO and carbon) in addition to carbon-FRP composites were
used to strengthen the corroded specimens (Figure 3). The fabric properties in the primary direction
as reported in the manufacturers’ data sheet are shown in Table 3. The PBO fabric consists of an
unbalanced net of spaced fiber rovings organized along two orthogonal directions as shown in
Figure 3a. The associated inorganic cementitious matrix had a compressive strength of 43.86 MPa
(standard deviation of 0.41 MPa) and a flexural strength of 3.01 MPa (standard deviation of 0.32
MPa) after 28 days as determined by the authors. On the other hand, the carbon-FRCM composite
consists of unidirectional net made of carbon-fiber strands oriented in one direction (Figure 3b)
and impregnated in an inorganic cementitious matrix of compressive strength of 42.11 MPa
(standard deviation of 4.27 MPa) and flexural strength of 3.26 MPa (standard deviation of 0.30
MPa) after 28 days as determined by the authors. Finally, the carbon-FRP composite consists of
unidirectional carbon fiber sheet (Figure 3c) and an epoxy resin. According to the manufacturer’s
data sheet, the composite has a tensile strength of 0.894 GPa, a tensile modulus of 65.4 GPa, and
an ultimate elongation of 1.33%. Table 4 lists the properties of the FRCM composite systems as
reported by Usama et al. (2016).
FRCM equivalent axial stiffness
According to the (ACI-549, 2013) provisions, the tensile stress-strain curve of the FRCM coupon
can be represented by a simple bilinear curve as shown in Figure 4. The first linear segment
represents the behavior of the FRCM composite prior to cracking and is characterized by the
uncracked modulus of elasticity, 𝐸𝐸𝑓𝑓∗ . The second linear segment represents the cracked behavior
with a reduced cracked modulus of elasticity, 𝐸𝐸𝑓𝑓 . An equivalent axial stiffness, Kf, was utilized to compare between the two FRCM systems used in this study based on their cracked elastic modulus
and the cross-sectional area of the fabric as given by Equation 1:
𝐾𝐾𝑓𝑓 = 𝜌𝜌𝑓𝑓 𝐸𝐸𝑓𝑓 = [(𝑁𝑁𝐴𝐴𝑓𝑓 )/𝑑𝑑𝑓𝑓 ]𝐸𝐸𝑓𝑓 𝜌𝜌𝑓𝑓 =
𝜌𝜌𝑓𝑓 , Af, and 𝐸𝐸𝑓𝑓 are listed in Table 1, Table 3, and Table 4, respectively. The equivalent axial
FRCM repair schemes
stiffness, Kf, of each repaired specimen is shown in Table 1.
Two FRCM repair schemes were utilized in this study as shown in Figure 5. Scheme I consisted
of one or more FRCM flexure plies having 150 mm width (equal to the width of the beam) and
applied to the soffit of the beam over a length of 2400 mm. The fabrics were oriented so that their
primary direction was parallel to the longitudinal axis of the beam. The plies were anchored at
each end using one U-shaped transverse strip of 300 mm width and 200 mm height as shown in
Figure 5a. Scheme II consisted of bottom flexural strips similar to those of scheme I but wrapped
with an additional U-shaped continuous ply along the beam’s span (Figure 5b). The primary
direction of the U-wrapped PBO ply was oriented parallel to the longitudinal axis of the beams.
For instance, the beam CR-4P-II consisted of 3 bottom flexural strips plus one U-shaped layer,
with the primary fibers of all 4 layers running parallel to the longitudinal axis of the beam.
Therefore, the 4 layers of the PBO-fabric contributed to the flexural performance of the beam. On
the other hand, the carbon fabric is a unidirectional fabric. Therefore, the bottom strips of the
carbon-FRCM were oriented parallel to the longitudinal axis of the beams whereas the U-shaped
layer was oriented in the transverse direction and therefore did not contribute to the flexural
behavior of the beam (Figure 5b). For example, specimen CR-3C-II was repaired with 3 flexural
strips in the longitudinal direction plus one U-shaped layer in the transverse direction. Only 3
layers of the carbon-FRCM were considered later in the analysis of this beam.
Corroded specimens were repaired before applying the FRCM repair system. Figure 6 depicts
the repair procedure. The deteriorated concrete was first removed using a hydraulic hammer. The
corroded steel bars were then brushed and the beams were repaired using Sikacrete-08SCC mortar.
After 7 days of curing in ambient temperature, sandblasting was used to roughen the concrete
substrate. The beam’s substrate was damped in water for 2 hours before applying the first layer of
the cementitious matrix with a thickness of 3 to 4 mm. Then, the fabric was installed and coated
with a second layer of matrix of similar thickness. The procedure was then repeated until the
specified number of layers was attained.
Test setup and instrumentation
All beams were instrumented at mid-span with a 60 mm long strain gauge bonded to the top
surface of concrete and 5 mm strain gauges bonded to the tensile steel bars. The repaired specimens
were instrumented with 5 mm strain gauges installed directly on the outer fabric of the FRCM
composite and distributed along the beam span as shown in Figure 7. The beams were tested under
four-point loading configuration as shown in Figure 1. The load was applied in displacement
control at a rate of 2 mm per minute using a MTS actuator. Beam deflections were measured by
means of three linear variable differential transducers (LVDTs) located at mid-span and under the
point loads. A data acquisition captured the readings of strain gauges and LVDTs at all stages of
Corrosion cracks and mass loss
Due to corrosion, continuous longitudinal cracks parallel to the reinforcing bars were observed
as shown in Figure 8 for specimen CU. No concrete spalling was observed. All of the corroded
specimens did not meet the ACI 318-14 service requirements that limits the maximum crack width
in service to 0.40 mm (ACI, 2014). The average and maximum measured crack widths after
corrosion were determined as 0.7 and 1 mm, respectively.
Visual inspection of the corroded beams revealed the existence of several corrosion pits randomly
dispersed along the surface of the bars. Six steel coupons, 200 mm long each, were extracted from
each corroded bar after testing. The actual mass losses of the examined bars were determined
according to the ASTM G1-03 standards (ASTM, 2011). The average tensile steel mass loss for
each specimen are listed in Table 1. The average, minimum, and maximum steel mass loss
determined for all specimens were 12.6, 11.5, and 13.7%, respectively.
Modes of failure
The modes of failure of the tested specimens are summarized in Table 1 and shown in Figure 9
for the tested beams. Beams UU (control) and CU (corroded unrepaired) failed by yielding of the
steel bars followed by concrete crushing (SY-CC). A similar mode of failure was observed in
specimen CR-1P-I as shown in Figure 9a. No loss of bond was observed between the PBO-FRCM
and the concrete substrate while loading. The PBO fabric remained intact with its matrix until
crushing of concrete occurred at ultimate. For the other repaired specimens, four different modes
of failure were observed:
a) FRCM delamination (FD): this type of failure occurred at the fabric/matrix interface with
complete delamination between the fabric and the first layer of the matrix adjacent to the concrete
substrate (Figure 9b). The delamination was caused by the propagation of flexural cracks to this
thin layer of the matrix and the relative deformation between the fabric and the matrix. This mode
of failure was reported for specimens CR-2P-I, and CR-4P-I.
b) Fabric slippage (FS): slippage occurred within the cementitious matrix (Figure 9c). Cracks
were first observed in the matrix of the U-shaped FRCM layer followed by the gradual slippage
of the fabric until the FRCM strengthening action was lost. This mode of failure was observed in
specimens CR-2P-II and CR-4P-II. It should be noticed that the continuous PBO-U-shaped ply
mitigated the FRCM delamination. Therefore, specimens that failed in this category showed a
more ductile behaviour compared to that observed in specimens that failed due to FRCM
c) Matrix cracking and fabric separation from the matrix [MC-SFM)]: this type of failure was
reported for specimens with carbon-FRCM namely, CR-2C-II and CR-3C-II, as shown in Figure
9d. As the applied load increased, progressive cracking in the cementitious matrix associated with
the separation of the carbon fabric from the matrix was observed. Matrix cracking took a web
pattern as shown in Figure 9d for the bottom of specimen CR-3C-II. This mode of failure was
more brittle than that observed in the PBO-repaired specimens, which can be attributed to the
superior characteristics of the cementitious matrix of the PBO-FRCM compared to those of the
d) CFRP laminate rupture (LR): this mode of failure was reported for specimen CR-1FRP-I
(Figure 9e). A longitudinal crack initiated at mid span at the laminate/concrete interface followed
by the sudden rupture of the laminate. This mode of failure was consistent with the high strains
recorded in the laminate at ultimate.
Load-deflection relationships of the tested beams are shown in Figure 10 to Figure 12. The
flexural response of the virgin beam (UU), the corroded-unrepaired (CU), and the FRP-repaired
specimen (CR-1FRP-I) are also shown for reference. The load-deflection curve of specimen CU
indicated that corrosion slightly reduced the load-carrying capacity and stiffness of the beam. The
load-deflection curve of the repaired beams consisted of three segments with two turning points
indicating the concrete cracking and the yielding of the tensile steel. The flexural response of the
repaired beams was highly dependent on the FRCM repair scheme, its type, and the number of
FRCM layers used.
Figure 10 shows the load-deflection relationships of the beams repaired with PBO-FRCM using
scheme I. All of the beams showed similar stiffness prior to yielding of steel reinforcing bars
indicating the slight influence of the FRCM composite on the flexural response at this stage.
Increasing the number of the PBO plies increased the post-yielding stiffness of the repaired
specimens in comparison to the control ones. Specimen CR-1FRP-I (repaired with one layer of
CFRP fabric) showed higher post-yielding stiffness than that of specimen CR-1P-I (repaired with
one layer of PBO fabric). However, the later specimen showed a more ductile behavior than the
Figure 11 shows the effect of the FRCM scheme on the load-deflection response of the PBO-
repaired beams. Specimens repaired with two and four PBO plies in scheme II showed a slight
enhancement in the pre-yielding and post-yielding stiffness, which can be attributed to the
enlargement of the beam width and the effect of the continuous U-wrapped strips in delaying the
delamination of the FRCM.
Figure 12 compares the load-deflection responses of the Carbon- and PBO-FRCM repaired
beams using scheme II. It can be noticed that specimens repaired with carbon-FRCM showed
higher post-yielding stiffness than that of their PBO-repaired counterparts. The former specimens
exhibited a sudden drop after reaching the ultimate load whereas specimens repaired with PBO-
FRCM showed a gradual decreasing trend after reaching the ultimate. This can be related to the
brittle mode of failure reported for specimens repaired with carbon-FRCM.
Table 5 gives the strength results of the tested beams. The experimental yield, 𝑃𝑃𝑦𝑦𝑒𝑒𝑒𝑒𝑒𝑒 , and ultimate,
𝑃𝑃𝑢𝑢𝑒𝑒𝑒𝑒𝑒𝑒 , loads of all specimens were normalized to those of the virgin specimen. It can be noticed
that corrosion of the main reinforcement reduced the yield and ultimate loads by 8% and 5%,
respectively. The reduction in the load-carrying capacity due to corrosion was smaller than the
steel mass loss due to the good anchorage of the bars within the shear zone, which allowed a tied-
arch action to be developed when the specimen approached failure (Masoud et al. 2001).
Effect of number of FRCM plies on strength
The use of a single PBO-FRCM layer in specimen CR-1P-I restored 95 and 105% of the yield
and ultimate loads of the virgin beam, respectively. Increasing the number of PBO-FRCM layers
further increased the yield and ultimate loads (specimen CR-2P-I restored 106 and 108% and
specimen CR-4P-I restored 111 and 125% of the yield and ultimate loads, respectively). However,
the strength enhancement was not linearly proportional to the added number of FRCM layers. On
the other hand, the FRP-repaired specimen CR-1FRP-I restored 104 and 121% of the yield and
ultimate loads of the virgin beam, respectively, compared to 111 and 125%, respectively, for
specimen CR-4P-I repaired with four PBO-FRCM plies. It is important to note that the FRP and
FRCM systems in these beams had almost similar axial stiffness (99.7 and 95.3 MPa, respectively,
as shown in Table 1). However, the PBO-FRCM repaired specimen showed superior capacities
than its FRP counterpart.
A similar trend was encountered in specimens repaired with scheme II. Increasing the number of
FRCM layers enhanced the yield and ultimate strengths of the repaired beams. Specimen CR-4P-
II showed an increase of 22 and 44% of the yield and ultimate loads, respectively, compared to 14
and 28% for specimen CR-2P-II. Similarly, the use of two layers of carbon-FRCM in specimen
CR-2C-II increased the yield and ultimate strengths by 6 and 30%, respectively, compared to 16
and 52% for specimen CR-3C-II.
Effect of FRCM repair scheme on strength
Scheme II was more effective than scheme I in restoring the yield and load-carrying capacity of
the repaired beams. This was depicted from the results of the beams repaired with two and four
PBO-FRCM layers. The enhancement in yield load was 6 and 14% for specimens CR-2P-I and
CR-2P-II, respectively. Their corresponding ultimate strengths increased by 8 and 28%,
respectively. The use of four layers of PBO-FRCM with scheme II in specimen CR-4P-II increased
the yield and ultimate loads by 22 and 44%, respectively, in comparison to 11 and 25% for
specimen CR-4P-I having the same number of PBO-fabric layers.
Effect of axial stiffness on strength
Figure 13 shows the effect of changing the axial stiffness of the repair system, Kf, on the
normalized ultimate load of the tested specimens. Specimens with similar axial stiffness of their
repair system didn’t show similar ultimate capacities. This can be depicted from the results of
specimens CR-2P-I and CR-2P-II having the same axial stiffness of their FRCM system but with
different repair schemes. The former specimen showed a load-carrying capacity of 86.4 KN versus
102.2 KN for the later one. Similarly, specimens CR-4P-I and CR-4P-II, also having the same
axial stiffness, showed 99.6 KN and 114.4 KN, respectively. This finding was also demonstrated
in specimens CR-4P-II and CR-2C-II having almost similar axial stiffness but repaired with two
different FRCM systems. Specimen CR-4P-II showed a load carrying capacity of 114.4 KN
whereas the specimen CR-2C-II showed a load carrying capacity of 104 KN. This finding indicates
that the axial stiffness of repair system, Kf, should not be used solely to compare the strengthening
actions of different FRCM systems without taking into account the material properties and the
repair scheme used.
The ductility index, ΔI, defined as the ratio of the midspan deflection of the beam at ultimate, δu,
to its midspan deflection at yielding, δy, was used to quantify the ductility of the tested specimens.
In general, a higher ductility index means a higher ability of the beam to redistribute moment and
to exhibit large overall deformation and energy dissipation. Table 6 lists the deflections at yielding
and ultimate and the ductility indices for all of the tested beams normalized to that of the virgin
It can be noticed that corrosion of the steel bars increased the ductility index of the corroded
beam by 15%. All beams repaired with PBO in scheme I restored the ductility of the virgin beam
except beam CR-4P-I that showed a ductility index 13% less than that of the virgin beam. For this
set of beams, increasing the number of PBO plies decreased the ductility of the repaired beam. The
ductility indices of beams CR-4P-I, CR-2P-I, and CR-1P-I were 2.4, 2.8, and 3.0, respectively.
Comparing specimens CR-4P-I and CR-1P-I, quadrupling the number of PBO plies decreased the
ductility index by 20%. On the other hand, the CFRP-repaired specimen (CR-1FRP-I) did not
restore the ductility of the virgin beam and had a similar ductility index of its FRCM-repaired
counterpart (CR-4P-I) having the same axial stiffness.
The set of beams repaired with PBO in scheme II restored the ductility of the virgin beam. In
fact, these beams showed 2 to 13% increase in their ductility indices as compared to their
counterparts repaired with scheme I. However, increasing the number of the PBO plies in scheme
II had a less pronounced effect on the ductility index than in scheme I. Beams CR-4P-II and CR-
2P-II had ductility indices of 2.8 and 2.9, respectively, which indicates that doubling the number
of plies in scheme II resulted in only 3.5% reduction in the ductility index of the beam.
The ductility indices of the beams repaired with carbon-FRCM (CR-3C-II and CR-2C-II) were
lower than that of the beams repaired with PBO-FRCM having the same number of plies and the
same repair scheme. Both beams couldn’t restore the ductility of the virgin beam. Their ductility
index was 14 and 22% less than that of the virgin beam, respectively. This reduction in ductility
was attributed to their brittle mode of failure that was due to the rapid loss of the strengthening
action at the fabric/matrix interface. It is important to note that increasing the number of carbon
plies in this set of beams increased the ductility index of the beam, which is contrary to what has
been noticed in the PBO-repaired beams. This increase was attributed to the increase in the ultimate
load of the carbon-FRCM repaired beams with similar yielding deflections in comparison to their
Table 6 lists the strains measured at midspan in both concrete and the outer fabric at ultimate.
Figure 14 and 15 show the load-strain curves for specimens repaired with scheme I and scheme II,
respectively. Similar to the load-deflection responses of the repaired beams, the load-strain curves
consisted of three segments with two turning points that indicated the concrete cracking and the
yielding of the tensile steel.
Figure 14 shows that, prior to yielding, all repaired specimens showed a similar increase in
concrete strains as the applied load increased. This increase in concrete strains continued after
yielding but at different rates depending on the repair system used. Specimen CR-1FRP-I showed
the highest rate of increase in concrete strains when compared to the PBO-repaired ones. On the
other hand, specimen CR-1P-I recorded the maximum tensile strains in the outer fabric of FRCM
(14921 μɛ) as no fabric delamination was observed for this specimen until failure. Specimens CR-
2P-I and CR-4P-I, repaired with two and four plies, failed by FRCM delamination and therefore
recorded low tensile strains in the PBO fabric (8670 μɛ and 9541 μɛ, respectively).
As shown in Figure 15, the concrete strains measured in the PBO-repaired specimens were higher
than those recorded in their carbon-FRCM counterparts. For instance, specimens CR-2P-II and
CR-2C-II recorded concrete strains of 3491 and 2370 μɛ, respectively. It was observed that
concrete strains of the carbon-FRCM specimens increased at higher rate than that of strains of the
PBO-FRCM specimens. This can be depicted from the strains recorded for specimens CR-2C-II
and CR-3C-II in Figure 15. On the other hand, the tensile strains in the carbon-FRCM at failure
was lower than those in PBO-FRCM. Specimens CR-2C-II and CR-3C-II recorded tensile strains
in the outer fabric at failure of 5753 μɛ and 5991 μɛ, respectively, whereas their counterparts CR-
2P-II and CR-4P-II recorded tensile strains of 11262 μɛ and 9598 μɛ, respectively. These findings
were consistent with the mode of failure of the carbon-FRCM repaired specimens where premature
matrix cracking and fabric separation were encountered. They were also consistent with the
measured ductility indices for these beams.
The distribution of the outer fabric tensile strains along the beam axis are plotted in Figure 16 to
Figure 18 for specimens CR-4P-I, CR-4P-II, and CR-3C-II, respectively, at a service load equal to
60% of ultimate (0.6 Pu), at the yielding load (Py), and at two post-yielding loads equal to 0.9 Pu,
It can be noticed that the strains in the fabric increased with the increase of the applied load until
yielding occurred. Post yielding, a significant increase in fabric strains were observed, with the
maximum increase occurring in the constant moment zone. This finding indicates that the FRCM
system became more effective in resisting the applied loads after yielding of the steel bars. The
repair scheme had marginal effect on the fabric strain profiles as can be depicted from Figures 16
The flexural behavior of the tested beams were predicted according to the provisions of the ACI
318-14 standard (ACI, 2014) and the ACI 549.4R-13 committee (ACI, 2013). Perfect bond was
assumed between the fabric and the cementitious matrix and between the FRCM and the concrete
substrate. A bilinear-elastic behavior of the FRCM repair system was presumed up to failure. The
cracked tensile modulus of elasticity, Ef, of the FRCM system was used after cracking,
396 397 398
The FRCM effective tensile strain at failure, 𝜀𝜀𝑓𝑓𝑓𝑓 , was limited to the FRCM design tensile
strain, 𝜀𝜀𝑓𝑓𝑓𝑓 , as given in Equation 2a (ACI 549.4R, ACI, 2013). The effective tensile stress in the FRCM at failure, 𝑓𝑓𝑓𝑓𝑓𝑓 , was calculated in accordance with Equation 2b. 𝜀𝜀𝑓𝑓𝑓𝑓 = 𝜀𝜀𝑓𝑓𝑓𝑓 ≤ 0.012
𝑓𝑓𝑓𝑓𝑓𝑓 = 𝐸𝐸𝑓𝑓 𝜀𝜀𝑓𝑓𝑓𝑓
400 401 402
Equation 3 using the strain compatibility principle as shown in Figure 19. 𝜀𝜀𝑓𝑓𝑓𝑓 𝜀𝜀𝑠𝑠 𝜀𝜀𝑠𝑠′ 𝜀𝜀𝑐𝑐 = = = 𝑑𝑑𝑓𝑓 − 𝑐𝑐𝑢𝑢 𝑑𝑑 − 𝑐𝑐𝑢𝑢 𝑐𝑐𝑢𝑢 − 𝑑𝑑′ 𝑐𝑐𝑢𝑢
The nominal flexural strength, Mn, was calculated in accordance with Equation (4) as follows:
Strains in concrete, steel reinforcing bars, and FRCM systems were computed in accordance with
𝑀𝑀𝑛𝑛 = 𝑀𝑀𝑠𝑠 + 𝑀𝑀𝑓𝑓 + 𝑀𝑀𝑠𝑠′ 𝑀𝑀𝑠𝑠 = 𝑇𝑇𝑠𝑠 ( d − 𝑀𝑀𝑓𝑓 = 𝑇𝑇𝑓𝑓 ( d − 𝑀𝑀𝑠𝑠′ = 𝐶𝐶𝑠𝑠′ (
𝛽𝛽1 𝐶𝐶𝑢𝑢 ) 2
𝛽𝛽1 𝐶𝐶𝑢𝑢 ) 2
𝛽𝛽1 𝐶𝐶𝑢𝑢 − 𝑑𝑑′ ) 2
(4) (4a) (4b) (4c)
The concrete stress block factors, 𝛽𝛽1 and 𝛼𝛼1 , and the modulus of elasticity of concrete, Ec, were
calculated as follows (ACI-318, 2014):
4ε′ −ε (C )
β1 = �6ε′c−2εc (Cu )�
3ε′c εc (Cu )−[εc (Cu )]2
α1 = �
3β1 (Cu )ε′2 c
𝐸𝐸𝑐𝑐 = 4700�𝑓𝑓′c
The force equilibrium was satisfied in accordance with Equation (9) and as shown Figure 19:
𝜀𝜀𝑐𝑐′ = 1.7𝑓𝑓𝑐𝑐′ /𝐸𝐸𝑐𝑐
𝑇𝑇𝑠𝑠 = 𝑅𝑅𝑐𝑐𝑐𝑐𝑐𝑐 𝐴𝐴𝑠𝑠 𝑓𝑓𝑦𝑦
𝐶𝐶 = 𝛼𝛼1 𝑓𝑓𝑐𝑐′ 𝛽𝛽1 𝑐𝑐𝑢𝑢 𝑏𝑏
𝐶𝐶𝑠𝑠′ = 𝐴𝐴′𝑠𝑠 𝐸𝐸𝑠𝑠 𝜀𝜀𝑠𝑠′
𝑇𝑇𝑠𝑠 + 𝑇𝑇𝑓𝑓 = 𝐶𝐶 + 𝐶𝐶𝑠𝑠′
𝑇𝑇𝑓𝑓 = 𝑁𝑁𝑁𝑁𝑓𝑓 𝑏𝑏𝑓𝑓𝑓𝑓𝑓𝑓
where 𝑅𝑅𝑐𝑐𝑐𝑐𝑐𝑐 = 1 – average tensile steel mass loss and,
between the experimental and theoretical values was obtained especially for specimens repaired
in scheme I. However, the capacities of specimens repaired with scheme II were under-estimated.
The theoretical formulations adopted do not account for the effect of the U-shaped FRCM layers
on the flexural response of the beams. The obtained results suggested the increase of the nominal
capacity, Mn, of FRCM-repaired beams with U-wrapped layers by 10% to account for the scheme
of the FRCM used.
Table 5 lists the theoretical ultimate loads, 𝑃𝑃𝑢𝑢𝑡𝑡ℎ , for all of the tested specimens. Good agreement
According to the provisions of the ACI-549 committee (ACI, 2013), the design flexural strength,
𝑀𝑀𝐷𝐷 , is calculated in accordance with Equation 10. The strength reduction factor, 𝜙𝜙𝑚𝑚 , is given by 20
Equation 11. In addition, the ACI-549 committee limits the increase in the nominal flexural
strength provided by the FRCM reinforcement by 50% of the flexural capacity of the structure
prior to repair. Table 5 lists the theoretical design load ϕ𝑚𝑚 𝑃𝑃𝑢𝑢𝑡𝑡ℎ and the ratio 𝑃𝑃𝑢𝑢𝑒𝑒𝑒𝑒𝑒𝑒 /ϕ𝑚𝑚 𝑃𝑃𝑢𝑢𝑡𝑡ℎ . It can be
noticed that applying both the flexural strength reduction factor and the 50% increase limitation
makes the gap between the experimental and design values lager, especially for the specimens
repaired with scheme II.
𝑀𝑀𝐷𝐷 = 𝜙𝜙𝑚𝑚 𝑀𝑀𝑛𝑛
0.90 for ɛ𝑡𝑡 ≥ 0.005 0.25�ɛ𝑡𝑡 −ɛ𝑠𝑠𝑠𝑠 �
𝜙𝜙𝑚𝑚 = � 0.65 + 0.005−ɛ
0.65 for ɛ𝑡𝑡 ≤ ɛ𝑠𝑠𝑠𝑠
for ɛ𝑠𝑠𝑠𝑠 < ɛ𝑡𝑡 < 0.005
This study investigated experimentally and analytically the structural performance of corrosion-
damaged RC beams strengthened with PBO- and carbon-FRCM systems. The following
conclusions can be drawn:
• An average mass loss of 12.9% in the tensile steel reduced the yield and the ultimate loads of
the beam by 8% and 5%, respectively. The corroded-unrepaired specimens failed to meet the
provisions of the ACI-318 standards for crack width criteria.
• Repairing corrosion-damaged RC beams with PBO- and carbon-FRCM restored 105 to 144%
and 130 to 152%, respectively, of the original load-carrying capacity of the virgin uncorroded
beam. The gain in capacity was highly dependent on the number of fabric layers, their material,
and the scheme used.
• Beams repaired with PBO-FRCM systems failed in a ductile mode due to either fabric
delamination (repair scheme I) or fabric slippage within the matrix (repair scheme II), whereas
beams repaired with U-wrapped carbon-FRCM systems showed a more brittle failure due to matrix
cracking and complete separation of the fabric.
• Beams repaired with carbon-FRCM showed higher post-yielding stiffness than that of their
PBO-repaired counterparts. The former beams exhibited a sudden drop after reaching the ultimate
load whereas the later beams showed a gradual decrease after reaching the ultimate.
• Increasing the number of FRCM layers increased the yielding and ultimate loads of the repaired
beams. However, specimens with similar axial stiffness didn’t show similar ultimate capacities.
Therefore, the FRCM material and the repair scheme used should be taken into account while
comparing the strengthening actions of different FRCM repair systems.
• U-wrapped FRCM scheme was more efficient than the bottom end-anchored scheme in
increasing the ultimate capacity of the repaired beams. The PBO-repaired beams with scheme II
showed ultimate strengths 15 to 18% more than those repaired with scheme I.
• Beams repaired with PBO-FRCM systems showed a more ductile behavior than their
counterparts repaired with carbon-FRCM or carbon-FRP systems. Most of the PBO-repaired
beams restored the original ductility whereas the carbon-FRCM and carbon-FRP repaired beams
showed lower ductility than that of the virgin beam.
• The theoretical formulations of the ACI-549.4R-13 committee reasonably predicted the
ultimate strengths of the FRCM-repaired beams with scheme I but underestimated those repaired
with scheme II. A scheme factor of 1.1 is then proposed while calculating the nominal strength of
beams repaired with U-shaped FRCM.
476 477 478 479 480 481 482 483
The authors would like to express their gratitude to the Qatar National Research Fund (a member
of Qatar Foundation) for funding this project under grant # NPRP 7-1720-2-641. The statements
made herein are solely the responsibility of the authors.
489 490 491 492 493 494 495 496 497 498 499 500 501 502 503 504 505 506 507 508 509 510 511 512 513 514 515 516 517 518 519 520 521 522 523 524 525 526 527 528 529 530 531
The following symbols are used in this paper:
Af = equivalent area of fabric per unit width (mm2/mm) As = cross-sectional area of tension steel reinforcement (mm2) 𝐴𝐴′𝑠𝑠 = cross-sectional area of compression steel reinforcement (mm2) b = width of the beam (mm) C = compression force provided by concrete (kN) Cs’ = compression force provided by the compression reinforcement (kN) cu = distance from extreme compression fiber to neutral axis (mm) d = distance from top of the beam to the centroid of tension steel (mm) d’ = distance from top of the beam to the centroid of compression steel (mm) df = distance from top of the beam to the centroid of fabric reinforcement (mm) Ef = cracked elastic modulus of the FRCM composite (GPa) Es = elastic modulus of steel reinforcement (GPa) Ec = elastic modulus of concrete (MPa) 𝑓𝑓𝑐𝑐′ = compressive strength of concrete (MPa) 𝑓𝑓𝑓𝑓𝑓𝑓 = effective tensile stress in FRCM composite at failure (MPa) ffu = ultimate tensile strength of FRCM composite (MPa) fy = yield strength of steel reinforcement (MPa) MD = design flexural strength (kN-m) Mf = moment contribution of FRCM reinforcement to flexural strength (kN-m) Ms = moment contribution of the tensile steel reinforcement to flexural strength (kN-m) Ms’ = moment contribution of the compression steel reinforcement to flexural strength (kN-m) Mn = nominal flexural strength (kN-m) N = number of fabric layers Rcor = corrosion reduction factor Ts = tension force in steel reinforcement (kN) Tf = tension force in FRCM reinforcement (kN) 𝜀𝜀𝑐𝑐 = compression strain in concrete (mm/mm) 𝜀𝜀𝑐𝑐′ = compression strain of unconfined concrete corresponding to 𝑓𝑓𝑐𝑐′ (mm/mm) εcu = concrete strain at ultimate (mm/mm) 𝜀𝜀𝑠𝑠′ = tensile strain in compression steel reinforcement (mm/mm) 𝜀𝜀𝑠𝑠𝑠𝑠 = tensile yield strain of steel reinforcement (mm/mm) 𝜀𝜀𝑡𝑡 = the net tensile strain in extreme tensile steel reinforcement at the nominal strength (mm/mm) 𝜀𝜀𝑓𝑓𝑓𝑓 = FRCM design tensile strain (mm/mm) 𝜀𝜀𝑓𝑓𝑓𝑓 = effective tensile strain in FRCM composite at failure (mm/mm) εfu = ultimate tensile strain of FRCM composite (mm/mm) 𝜌𝜌𝑓𝑓 = fabric reinforcement ratio 𝜅𝜅𝑓𝑓 = equivalent axial stiffness (MPa) ∆𝐼𝐼 = ductility index β1 = ratio of depth of equivalent rectangular stress block to depth to neutral axis α1 = multiplier of 𝑓𝑓𝑐𝑐′ to determine intensity of the equivalent block stress for concrete 𝜙𝜙𝑚𝑚 = strength reduction factor 24
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Table 1: Test matrix 603
Average mass loss (%)
𝐾𝐾𝑓𝑓 = 𝜌𝜌𝑓𝑓 𝐸𝐸𝑓𝑓 (MPa)
Mode of Failure**
UU, CU, and CR refer to Uncorroded-Unrepaired, Corroded-Unrepaired, and Corroded-Repaired specimens respectively. 1, 2, 3 and 4 in the specimen’s label refer to the number of FRCM or FRP plies. P, C, and FRP refer to the repair materials PBO-FRCM, Carbon-FRCM, and Carbon-FRP, respectively. I and II refer to the FRCM repair schemes. ** SY-CC = Steel Yielding followed by Concrete Crushing; FD = Fabric Delamination; FS = Fabric Slippage; MCSFM = Matrix Cracking with Separation of Fabric within the Matrix; LR = CFRP Laminate Rupture.
Table 2: Concrete compressive strengths
28-day Testing day
Compressive strength Standard deviation Coefficient of variation (MPa) (MPa) (%) Normal concrete 37.9 0.8 2 Salted concrete
613 614 615 616
Table 3: Fabric properties in the primary direction as given in the manufactures’ data sheet
Area per unit width (𝐴𝐴𝑓𝑓 ) (mm2/m) 50
Tensile strength (GPa) 5.8
Elastic modulus Ultimate strain (GPa) (%) 270 2.15
617 618 619 620 621 622
Table 4: Mechanical properties of FRCM systems (Usama et al. 2016)
Cracked tensile modulus of elasticity, Ef (GPa) 121
Ultimate tensile strength, ffu (GPa) 1.55
Ultimate strain, εfu (%) 1.4
623 624 625
Table 5: Strength results Specimen
627 628 629
ϕ𝑚𝑚 𝑃𝑃𝑢𝑢𝑡𝑡ℎ 𝑃𝑃𝑢𝑢 (KN) ϕ𝑚𝑚 𝑃𝑃𝑢𝑢𝑡𝑡ℎ
Average values reported Normalized with respect to the yield and ultimate loads of the virgin beam
630 631 632 633 634 635 636 637 638 639
Table 6: Ductility indices and strains at ultimate Midspan deflection (mm) δy δu
641 642 643 644 645 646 647 648 649 650 651 652 653 654 655 656 657 658 659 660 661 662 663 664 665 666
Concrete strains at ultimate (µ𝜖𝜖)
Fiber strains at ultimate (µ𝜖𝜖)
Average values reported Normalized with respect to the yield and ultimate loads of the virgin beam
667 668 669
Figure 1: Typical dimensions and reinforcement details of the test specimen
(all dimensions in mm)
671 672 673
Stainless steel bar (Cathode)
Steel bars (Anode)
Figure 2: Specimens connected in series inside the corrosion chamber
Primary direction Secondary direction
678 679 680
Figure 3: Repair materials: a) unbalanced PBO fabric, b) unidirectional carbon fabric, and c) unidirectional carbon fabric
681 682 683
Figure 4: Idealized tensile stress-strain curve of FRCM coupon specimen (ACI-549, 2013)
686 687 688 689 690 691 692 693 694 695 696 697 698
699 700 701
Figure 5: Repair schemes (a) scheme I and (b) scheme II
702 703 704 705 706 707 708 709 710 711 36
712 713 714 715 716
717 718 719
720 721 722 723
724 725 726 727
730 731 732
Figure 6: Repair procedure: a) removing the deteriorated concrete, b) patch repair, c) roughening the concrete surface with sandblasting, and d) FRCM application
735 736 737 738 739 740 741 742 743 744 745
746 747 748 749 750
Figure 7: Positions of the electrical strain gauges along the outer fabric
Figure 8: Corrosion cracks pattern for specimen CU
751 752 753 754 755 756 757 758 759 760 761 762 763 764 765 766 767 768 769 770 771 772 773 774 775 776 777 778 779 780 781 782 783 784 785 786 787 788 789 790 791 792 793 794
a) Beam CR-1P-I
b) Beam: CR-4P-I
Delamination at end of test
Side FRCM delamination
c) Beam: CR-2P-II
d) Beam CR-3C-II
Side Matrix crushing
e) Beam CR-1FRP-I
Figure 9: Typical modes of failure: (a) SY-CC in beam CR-1P-I, (b) FD in beam CR-2P-I, (c) FS in beam CR-2P-II, (d) MC-SFM in beam CR-3C-II, and (e) LR in beam CR-1FRP-I
80 CU 60
40 20 0 0
795 796 797 798
20 30 Deflection (mm)
Figure 10: Effect of number of PBO-FRCM plies on the load-deflection curves
799 800 801 802 140 CR-4P-I CR-1FRP-I CR-2P-I
40 20 0 0
20 30 Deflection (mm)
Figure 11: Effect of the repair scheme on the load-deflection curves
805 806 807 140 CR-3C-II
40 20 0 0
20 30 Deflection (mm)
Figure 12: Effect of FRCM materials on the load-deflection curves
Normalized ultimate load
1.5 CR-4P-II 1.4 1.3
CR-2C-II CR-4P-I CR-1FRP-I
810 811 812
40 60 80 100 120 Equivalent axial stiffness Kf (MPa)
Figure 13: Normalized ultimate load versus the equivalent stiffness
813 814 Load (KN)
CR-4P-I CR-1FRP-I CR-2P-I CR-1P-I UU
CU 60 40 20 0 -6000 -4000 -2000
815 816 817
4000 6000 Strain (με)
8000 10000 12000 14000 16000
Figure 14: Load-strain curves for specimens with repair scheme I
818 819 820 821
CU 60 40 20 0 -6000 -4000 -2000
822 823 824
4000 6000 Strain (με)
8000 10000 12000 14000 16000
Figure 15: Load-strain curves for specimens with repair scheme II
825 826 827 828 829 16000 Outer fabric strain (με)
PU= 99.6KN Py= 83.3KN
0.90Pu= 90KN 0.6Pu= 60KN
400 600 Distance from mid-span (mm)
Figure 16: Strain profile in the PBO fabric for specimen CR-4P-I
Outer fabric strain (με)
PU= 114.4KN Py= 91.3KN
0.90Pu= 103KN 0.6Pu= 68KN
832 833 834 835 836
400 600 Distance from mid-span (mm)
Figure 17: Strain profile in the PBO fabric for specimen CR-4P-II
837 838 839 16000 Outer fabric strain (με)
840 841 842 843 844 845
400 600 Distance from mid-span (mm)
Figure 18: Strain profile in the carbon fabric for specimen CR-3C-II
846 847 848 849 As’
Cs’ = As’ Es εs’ C= α1 fc β1 cu b
Ts = Rcor As fy
Tf = N Af b ffe
b 850 851 852 853 854 855 856 857 858 859 860 861 862 863 864 865 866
Figure 19: Stress and strain distribution at ultimate stage Screenshot should be taken without the red line under Af