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FRP for seismic strengthening of shear controlled RC columns: Experience from earthquakes and experimental analysis
Accepted Manuscript FRP for seismic strengthening of shear controlled RC columns: Experience from earthquakes and experimental analysis Marta Del Zoppo, Marco Di Ludovico, Alberto Balsamo, Andrea Prota, Gaetano Manfredi PII:
S1359-8368(17)30949-6
DOI:
10.1016/j.compositesb.2017.07.028
Reference:
JCOMB 5164
To appear in:
Composites Part B
Received Date: 16 March 2017 Revised Date:
19 June 2017
Accepted Date: 25 July 2017
Please cite this article as: Del Zoppo M, Di Ludovico M, Balsamo A, Prota A, Manfredi G, FRP for seismic strengthening of shear controlled RC columns: Experience from earthquakes and experimental analysis, Composites Part B (2017), doi: 10.1016/j.compositesb.2017.07.028. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
ACCEPTED MANUSCRIPT FRP for seismic strengthening of shear controlled RC columns: experience from earthquakes and experimental analysis Authors: Marta Del Zoppoa*, Marco Di Ludovicob, Alberto Balsamob, Andrea Protab, Gaetano Manfredib Department of Engineering, University of Naples “Parthenope”, Italy
b
Department of Structures for Engineering and Architecture, University of Naples “Federico II”, Italy
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a
*Corresponding Author: Marta Del Zoppo; Email: [email protected] Abstract
The high vulnerability of existing Reinforced Concrete (RC) structures, even to moderate seismic events,
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has been confirmed from recent post-earthquake surveys. Short and wall-like RC columns are particularly prone to brittle failures, governed by concrete crushing. To reduce the vulnerability of existing RC
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columns, the use of externally bonded Fibre Reinforced Polymer (FRP) reinforcement has been recognized as an effective method for preventing the aforementioned brittle failure and, hence, increasing members’ lateral capacity and ductility. In the first part of this study, the results of an observational analysis on columns shear failures in RC buildings severely damaged after the L’Aquila earthquake are presented. The second part of the study presents and discusses the results of an experimental program
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carried out on seven short RC columns governed by shear failure under load reversal and compressive axial load. Both columns in “as built” configuration and strengthened in shear with discontinuous carbon FRP (CFRP) strips have been tested. Two classes of concrete have been used, in order to simulate
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structures with medium or poor material quality, and different external reinforcement ratios have been investigated. The specimens’ responses have been analysed in terms of failure modes, strength/deformation capacity and strain distribution in CFRP strips.
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Keywords: seismic retrofit; columns; FRP strengthening; shear failure. 1. Introduction
Recent post-earthquake surveys [1-3] revealed the high vulnerability of existing Reinforced Concrete (RC) structures, often designed for gravity loads only, even to moderate seismic events. Indeed, the lack of a proper seismic detailing and a wrong shear-flexure hierarchy often lead to columns brittle failures due to shear before attaining the flexural yielding. Furthermore, short and wall-like RC columns are commonly subjected to such a brittle failure, governed by concrete diagonal compression failure [4-7]. Semi-basements, band-type windows, knee beam stairs and masonry parapets can turn the behavior of
ACCEPTED MANUSCRIPT slender columns into short ones during seismic events. Since short columns are subjected to higher shear forces than slender ones, their capacity cannot satisfy the shear demand and brittle failures occur in such members. Moreover, structures designed for gravity loads only or according to “first-generation” seismic codes can respond in a completely different way to a seismic event, especially in terms of shear failures.
exhibited shear failures also in slender columns [8].
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After the L’Aquila earthquake (2009), structures designed according to a “first-generation” seismic code
To prevent premature brittle failures of existing RC columns, the use of Externally Bonded
Reinforcement (EBR) made of Fibre Reinforced Polymer (FRP) strips has been recognized as a reliable
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and cost saving strategy for increasing members lateral strength capacity [9-11]. Plenty of researchers
focused the attention on the shear strengthening of RC beams, mainly due to gravity loads deficiencies
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[12-17, among many others]. Even though post-earthquake observed shear failures in columns are more diffused than shear failures of beams [18], only a few experimental studies have been developed on RC columns strengthened with FRP [19-24], demonstrating the effectiveness of this technique for such members.
The present work is focused on the behavior of shear critical short RC columns strengthened with CFRP.
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The results of an experimental program carried out on short RC columns strengthened with CFRP are herein presented and discussed. In particular, seven RC columns with square cross-section and governed by shear mode were analysed: two reference columns, termed “as built” in the following, and five
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columns externally reinforced with discontinuous CFRP strips along their shear length. Two classes of concrete have been used, in order to simulate structures with medium or poor quality concrete, and different external reinforcement ratios have been investigated. The specimens were cantilevers loaded
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with a low compressive axial load (i.e. axial load ratio 0.1) and subjected to horizontal cyclic displacements. The responses of columns have been analysed in terms of failure modes and increased strength and deformation capacity as a function of the external reinforcement ratio. Furthermore, to investigate on the effectiveness of CFRP external reinforcement, a detailed analysis related to the axial strains distribution along the fibres strips is reported. 2. Occurrence of shear failure in RC columns: the L’Aquila earthquake experience 2.1 Observed damage
ACCEPTED MANUSCRIPT After the L’Aquila earthquake (6th April 2009, magnitude 6.3), thousands of RC structures outside the historical centres were damaged. From the analysis of the post-earthquake reconstruction process [25, 26], a total of 2,799 RC buildings resulted affected by the seismic event: 1,738 RC buildings (i.e. 62%) were lightly damaged, with local damage to non-structural components but few or no damage to the
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structural ones, whereas the other 1,061 buildings (i.e. 38%) were severely damaged to both structural and non-structural components. Among the 1,061 severely damaged RC buildings, 284 structures (i.e. 27%) with severe damage to structural members were demolished, since repair and strengthening interventions resulted not viable due to economic reasons. From an extensive analysis carried out on the 284
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demolished RC buildings, resulted that 56 structures (i.e. 20%) exhibited brittle failures of columns due to shear actions. These 56 structures have been selected and analysed for investigating the reasons of shear
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failure occurrence.
The selected buildings were designed between ‘60-‘90 with three up to seven storeys, regular in both plan and elevation. The concrete mechanical properties, derived from in-situ material tests, were found in the range 8-30 MPa. Thus, in several cases, the poor quality of concrete contributed to a reduced resisting capacity to shear actions. Columns that achieved a shear failure were generally located at the ground floor
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or at the semi-basement of the selected buildings, where the shear demand attains maximum values. Columns affected by shear failure were characterized by a rectangular/square cross-section with dimensions variable between 300x350 and 300x600 mm2 in case of rectangular, and 300x300 or 400x400
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mm2 in case of square members. For almost all buildings, columns had constant cross-section dimensions for each storey level, thus they were all subjected to a similar shear demand. In 59% of selected buildings (i.e. 33 out of 56), the columns exhibited diagonal shear cracks, but they
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were still able to bear axial loads. Conversely, heavy damaged columns were detected in remaining 41% (i.e. 23) with large diagonal crack openings and no residual axial load capacity. The columns weakness to shear actions is mainly related to the low amount of transverse reinforcement found in typical existing members. Ties of 6 mm diameter spaced at 200-300 mm closed with 90° hooks were usually found in columns that failed for shear (see Fig. 1a). The large spacing between consecutive ties allows large cracks openings and longitudinal bars buckling, leading to a premature shear failure.
(b)
(c)
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(a)
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In addition, for columns designed for seismic action in one direction only, the absence of longitudinal reinforcement in addition to corner rebars along the column secondary direction, which can help in shear
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resisting mechanism, was also observed (see Fig. 1b). In addition to the poor material quality or the absence of seismic detailing, the shear failure observed in columns often resulted as a consequence of the infills actions. In 23% of buildings under investigation
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(i.e. 13 buildings), the shear failure of columns was due to unexpected interaction effects with masonry infills (see Fig. 1c and d). Indeed, stiff infills generate significant shear actions concentrated near the top
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end section of columns, which can lead to premature brittle failures.
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(f)
Fig. 1. Columns shear failure after L’Aquila earthquake (2009).
Finally, semi-basements (see Fig. 1e) and irregularly distributed infills, like in case of band-type openings
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(see Fig. 1f), reduce the actual shear length of columns, causing local shear failures. Short columns are subjected to a high shear demand and, hence, are more prone to shear failures with respect to slender
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members.
Even though the reasons of columns shear failure occurrence during seismic events are quite well known [5], it is still difficult to understand why in a sample structure only a few columns failed for shear among a large number of columns with same geometrical and mechanical properties and under similar load conditions (i.e. similar axial load condition and lateral stiffness). Indeed, in 62% of selected structures (35
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buildings), the shear failure occurred only for 1-2 columns. Conversely, shear failures on 3-5 columns were attained in 27% of selected buildings (15 buildings). Only in 6 buildings (i.e. 11%), the shear failure occurred in more than five columns. This post-earthquake survey founding can be related to local
project.
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deficiencies due to non-uniform material quality or reinforcement details non-conforming to the design
2.2 Strengthening solutions
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In case of repair and strengthening of L’Aquila damaged RC structures, several techniques were used. From the analysis of the reconstruction process, it was found that the most widely employed strengthening technique was the use of FRP composite systems [27, 28]. Indeed, 475 (i.e. 45%) heavily damaged RC structures adopted FRPs as the main retrofit solution for beams, columns and beam-column joints.
ACCEPTED MANUSCRIPT Fig. 2a shows the local strengthening solution on a short RC column of a school building lightly damaged by 2009 L’Aquila earthquake [28]. The effectiveness of using FRPs as a strengthening technique has been also proven by the recent earthquakes (i.e. Central Italy earthquake 2016): Fig. 2b shows a school complex consisting of two jointed buildings where the RC building retrofitted with FRP is mostly
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undamaged and the masonry structure collapsed during the 2016 Central Italy earthquake. The post-earthquakes experience raised, among others, these two issues: the assessment of existing
members’ shear capacity and the design of strengthening interventions with FRP composite systems.
International codes provide refined formulations for evaluating shear capacity in existing RC columns and
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practice-oriented tools for assessing the ductile/brittle behaviour of columns are presented in the literature [29]. However, accuracy and reliability of the available shear capacity models in predicting the column
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shear capacity before and after the strengthening is still a challenging task [8, 30], especially because theoretical formulations for designing FRP-based interventions are mainly derived for beams. Indeed, for the selected 56 structures, all shear capacity models considered by the Authors underestimated the effective members shear capacity and predicted shear failures also for many un-damaged columns. An inadequate accuracy of capacity models for predicting the members original shear capacity and the
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contribution of the FRP strengthening systems to the total capacity leads to the design of costly and often unnecessary strengthening interventions. Thus, the following experimental analysis is specifically targeted at investigating shear behavior of columns in both “as built” and FRP strengthened
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configurations, with the aim to point out the effective contribution of FRP external reinforcement to the capacity and the failure mode of shear critical RC columns. 3. Experimental program
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3.1 Specimens geometrical and mechanical properties
(a) (b) Fig. 2. FRP interventions on RC columns in L’Aquila (a) and Amatrice (b).
ACCEPTED MANUSCRIPT Seven short RC columns were designed with the same geometrical properties and reinforcement detailing as depicted in Figure 3a. The short columns have been over-designed for flexure, with the aim to result in a brittle failure governed by shear. Each specimen had a square cross-section 300 x 300 mm2 reinforced with ten 22 mm diameter deformed rebars (longitudinal geometric reinforcement ratio, ρl = As/bh = 4.2%
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with As total area of longitudinal steel reinforcement and b,h, cross section dimensions). The longitudinal reinforcement was placed only along the loading direction, no reinforcement was provided along the
secondary direction as previously observed for existing columns. The concrete cover thickness was 20
mm for each specimen. Steel transverse reinforcement was made of 8 mm diameter ties, spaced at 300
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mm apart (transverse geometrical reinforcement ratio, ρw = 0.11%) and with hooked ends. A reduced
1 CFRP ply D A
300 mm
C B
C
2 CFRP plies D
B
260 mm
As= 10Ø22
C
A 165
300 mm
A
100
Ø8/300
900 mm
D
B
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600 mm
1250 mm
Ø8/50
300 mm
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spacing has been adopted in the zone of load application to avoid localized damage.
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1200 mm
(a)
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Fig. 3. Specimens’ geometry: “as built” (a) and CFRP strengthened (b).
The specimens were cantilevers, made by a column and a foundation block: the column height was 1,250 mm; the foundation block was 1,200 x 1,200 x 600 mm3. The foundation block and the columns of each specimen were cast in two different phases to reproduce the discontinuity of concrete at the columnfoundation interface, as common practice in concrete casting procedures. The lateral load has been applied at a distance of 900 mm from the foundation block, in order to simulate the behaviour of short columns (i.e. LS/h = 3 with shear length LS = 900 mm). Two classes of concrete have been herein adopted: •
low compressive strength for four specimens (termed S1, S2, S3 and S4);
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medium compressive strength for the remaining three specimens (termed S5, S6 and S7).
One specimen for each concrete class was unstrengthened and used as control specimen (“as built” specimens, named S1 and S5 for low and medium compressive strength, respectively) whereas the other five specimens were strengthened with discontinuous uniaxial CFRP strips with fibres perpendicular to
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the column longitudinal axis. The CFRP strips, width wf = 100 mm, were spaced at sf = 165 mm for the total height of the column. Before applying the external reinforcement, the columns corners were rounded with radius R = 20 mm, in order to avoid the fibres premature rupture. The strips overlap was more than 250 mm for each layer along the column sides orthogonal to the transverse load action direction. Further
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details about the CFRP application are reported in Figure 3b. Specimens characterized by low concrete strength were reinforced with one ply of CFRP strips; conversely, two plies of CFRP strips were used for
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medium concrete specimens. Different axial rigidities of CFRP bonded strips were investigated, where the axial rigidity is expressed by the product Ef ρf, where Ef the elastic modulus of carbon fibres and ρf = 2tf wf/bsf is the FRP reinforcement ratio, with tf the thickness of dry fibres (i.e. Ef ρf = 0.16, 0.31, 0.34, 0.61 and 0.67GPa for specimens S2, S3, S4, S6 and S7 respectively). A summary of the specimens’ mechanical properties and external reinforcement configurations is reported in Table 1.
Test
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Table 1. Summary of the experimental program. fcm
fyl
fyw
ρf
Ef ρf
[MPa]
[MPa]
[MPa]
[GPa]
n° plies
CFRP type
-
-
-
14.3
S2
14.3
0.06
0.16
1
A
0.13
0.31
1
B
0.13
0.34
1
C
23.9
-
-
-
-
S6
32.3
0.27
0.61
2
B
S7
36.4
0.27
0.67
2
C
S3 S4
16.5 14.0
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S5
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S1
[%] -
531
525
The same class of steel has been used for all the columns: average yielding strength fyl = 531 MPa for longitudinal reinforcement and fyw = 525 MPa for transverse reinforcement were derived from tensile tests on coupon bars. Four concrete 150-mm side cubes were cast during each specimen preparation and were cured under the same environmental conditions. The mean cylindrical concrete strength, fc, derived for each specimen is reported in Table 1. Three types of CFRP composite material with different unit weight and elastic modulus were used, in order to obtain different axial rigidities of bonded strips. The
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∆ [%] 0.2 mm/s
1.0 mm/s
2.0 mm/s
10%
5%
-5%
-10%
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0%
t [s]
-15% 1000
2000
3000
4000
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0
mechanical properties of the dry fibre material types declared by the manufacturers (i.e. elastic modulus
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of dry fibres Ef and ultimate strain εfu) are reported in Table 2. Table 2. Dry carbon fibres material properties.
εfu [%] 1.9 1.3 1.9
tf [mm] 0.16 0.33 0.33
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CFRP Unit weight Ef type [g/m2] [GPa] A 300 252 B 600 230 C 600 252 3.2 Test set-up, instrumentation and loading
The short columns were subjected to a constant compressive axial load and horizontal cyclic displacements. The lateral load was applied at a distance of 900 mm from the base cross-section by using
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an electrohydraulic actuator (stroke ± 250 mm, maximum load +500 kN, -300 kN) integrated in a steel reaction wall and a 500 kN load cell was used to measure lateral loads. A detailed description of the test
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setup and relevant instrumentation is reported in [31] and shown in Figure 4a.
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(b) Fig. 4. Test set-up (a) and drift protocol (b).
All the specimens were subjected to a normalized compressive axial load (ν = N / Acfc, where N is the compressive axial load, Ac is the concrete gross area and fc is the mean cylindrical concrete strength) equal to 0.1 and a lateral cyclic loading under drift control (i.e. ∆ = d/ LS with d equal to the horizontal
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applied displacement and LS = 900 mm), see Figure 4b. A rate of 0.2 mm/s was adopted for initial four
cycles, then higher rate of 1.0 mm/s was adopted for the next five cycles and 2.0 mm/s for the last cycles. For each cycle, the target displacement was achieved three times, see Figure 4b. Only one specimen (S7)
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was subjected to cyclic lateral displacements for the first seven cycles and then to monotonic load up to a drift ratio ∆ = 11% (i.e. d = 100 mm), as the actuator achieved its maximum load capacity in the negative
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load action direction. Strain gauges were located at the mid-height and mid-span of each strip for all the column sides, for the first four CFRP strips from the base, with the aim to monitor axial strains in the external reinforcement. 4. Experimental results
4.1 Lateral load-drift relation and failure mode
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The experimental lateral load-drift relationships for the seven short RC columns are depicted in Fig. 5a-g. The envelope curves normalized with respect to the concrete contribution,
with d the section
effective depth, are reported in Fig. 5h. The relevant experimental outcomes for both positive and negative load-action directions along with the observed failure mode are reported in Table 3: peak forces
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and relative drift ratios, Fmax and ∆Fmax, drift ratios at conventional failure, ∆0.8Fmax, defined as the values corresponding to 80% of the peak force on the envelope curve, percentage of increased strength due to the
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FRP strengthening with respect to the “as built” specimens in the positive and negative load direction, respectively, δFmax, and ductility, μ∆, defined as the ratio between ultimate and yielding drift. The drift at the yielding is assumed as the intersection point between the secant at the 70% of the peak strength and the horizontal at the peak strength. The crack patterns and relevant damage at conventional failure are shown in Fig. 6. Table 3. Experimental outcomes and observed failure mode. Test S1
Fmax [kN] 159.6 -151.8
∆Fmax
∆0.8Fmax
δFmax
[%] 1.6 -1.6
[%] 2.9 -2.8
[%] -
μ∆ [-]
failure mode
FRP failure
-
shear
-
S4 S5 S6 S7
3.2 -3.2 3.2 -3.2 2.4 -2.4 1.6 -1.6 6.4 -6.4 9.1 -
4.3 -3.7 6.2 -6.0 7.5 -7.2 2.7 -2.5 8.5 -8.0 >11 -
+61 +56 +57 +67 +80 +89 +63 +57 +87 -
4.0 3.9 4.8 4.6 >6 -
shear
2nd strip rupture
shear
-
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S3
256.8 -236.3 249.9 -253.0 286.7 -286.5 188.9 -191.0 308.3 -300.6 354.0 -
flexure-shear
local rupture
shear
-
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S2
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flexure-shear
-
flexure
-
The two “as built” specimens, S1 and S5, exhibited a brittle failure governed by shear before the flexural yielding, see Fig. 5a and e. Diagonal cracks extending throughout the entire shear length of the column
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were developed along the two sides parallel to the load action direction (i.e. B and D sides), see Fig. 6a and e. The peak force was achieved for both specimens at a drift ratio of 1.6%, as reported in Table 3. After the peak cycle, a sudden drop of the lateral strength was observed and the tests were stopped. The
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failure was governed in both cases by concrete web crushing, so defined shear diagonal compression failure, typical of short columns. After the conventional failure, the columns were still able to carry the
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axial loads. Peak lateral strength of +159.6 kN (-151.8 kN) and +188.9 kN (-191.0 kN) were attained on S1 and S5 respectively, see Table 3. Thus, a difference of 60% in concrete compressive strength increased the shear capacity, governed by concrete crushing, about 18%.
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(g) (h) Fig. 5. Lateral load-drift relationships for the seven specimens.
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(c)
(e)
(f)
(g)
Fig. 6. Crack patterns and damage at failure.
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(a)
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The detailed behaviour of strengthened specimens is described as a function of the axial rigidity, Ef ρf, in the following: Specimen S2 (Ef ρf = 0.16GPa)
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•
Under lateral cyclic loading (Fig. 5b), the specimen exhibited a brittle failure governed by the fragile rupture of CFRP reinforcement before the flexural yielding, see Fig. 6b. The CFRP rupture, probably due to a strain concentration at corner, was limited to the second strip starting from the column base. Comparing the response of specimen S2 with the “as built” specimen S1, the external reinforcement
•
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increased the shear capacity up to +256.3 kN (-236.3 kN), which means 56-61% of strength enhancement. Specimen S3 (Ef ρf = 0.31GPa)
Specimen S3, exhibited a brittle behaviour governed by shear, see Fig. 5c. In this case, no rupture of the
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composite material was detected (Fig. 6c). Therefore, the shear failure in this case is controlled by the deterioration of unconfined concrete between consecutive strips, see Fig.6h, and by the loss of aggregate interlock inside the CFRP strips. Peak forces recorded for specimen S3 are close to those attained from
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specimen S2, about +249.9 kN (-253.0 kN), which means 57-67% of strength enhancement with respect to the “as built” specimen, see Table 3. •
Specimen S4 (Ef ρf = 0.34GPa)
Fig. 5d shows that specimen S4 exhibited a “quasi-ductile” behaviour, attaining the flexural yielding before a sudden drop of lateral capacity due to the CFRP rupture. In this case, a local rupture of the fibres is limited to the third strip starting from the column base, see Fig. 6d. Thus, an external reinforcement ratio of 0.40% changed the failure mode of the specimen from brittle to “quasi-ductile”, with ultimate drift ratios of +7.5% (-7.2%), see Table 3.
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Specimen S6 (Ef ρf = 0.61GPa)
The response of S6 is “quasi-ductile” since the flexural yielding was attained before the sudden drop of lateral capacity caused by the shear action, see Fig. 5f. The specimen failure is governed by the concrete deterioration and loss of aggregate interlock inside and between the CFRP strips. No fibres rupture was
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detected (see Fig. 6f). Thus, the external reinforcement ratio of 0.70% changed the failure mode of the specimen from brittle to “quasi-ductile”, with ultimate drift ratios of +8.5% (-8.0%), see Table 3. •
Specimen S7 (Ef ρf = 0.67GPa)
As shown in Fig. 5g, specimen S7 manifested a pure flexural behaviour. As aforementioned, the test was
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carried out as cyclic for first cycles and then as monotonic up to a drift ratio of 11%. No strength
degradation was observed after the flexural yielding, probably due to absence of cyclic loading. At the
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end of the test, the specimen was completely undamaged, see Fig. 6g. 4.2 Stiffness degradation and energy dissipation
To investigate the effects of CFRP external reinforcement on the columns stiffness degradation caused by the damaging under cyclic loading, the experimental stiffness has been computed as the peak-to-peak secant stiffness of the first cycle of each imposed drift ratio. The ratio between the experimental secant
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stiffness, K, and the theoretical elastic stiffness, Kflex = 3EcIg/LS3 (where Ec is the concrete elastic modulus and Ig is the gross section inertia), is plotted for the seven short RC columns as a function of the drift ratio in Fig. 7a. In case of specimen S7, the secant stiffness is calculated only for the cyclic part of the load-
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drift ratio curve. At first cycle, the ratio K/ Kflex is in the range between 31% and 45%: specimens S4 and S7, strengthened with type C composite material characterised by a higher axial stiffness than other types,
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manifested a greater initial secant stiffness if compared with other specimens.
(a) Fig. 7. Secant stiffness degradation (a) and cumulative energy (b).
(b)
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S6 34.7 139.4 512.5 1146.6 2017.9 5693.1 12412.7 28290.2 51307.5 76494.2
S7* 33.0 152.5 572.5 1316.0 3487.5 8664.2 18268.3 25703.5
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Drift Cumulative Energy [kNmm] [%] S1 S2 S3 S4 S5 0.2 38.6 38.6 32.2 33.7 41.9 0.4 160.1 154.9 127.5 148.9 162.6 0.8 758.3 571.7 510.9 546.0 654.0 1.2 1615.3 1262.3 1070.7 1199.8 1504.0 1.6 2725.3 2179.2 1855.8 2082.6 2708.6 2.4 5222.0 4683.2 3946.5 4944.3 5721.1 3.2 7668.5 8421.3 7520.5 10687.7 8391.5 4.8 17540.0 15609.4 23557.6 6.4 38249.7 8 50847.9 * Energy calculated only for the cyclic part of the force-drift relationship.
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As built specimens, S1 and S5, experienced a rapid decrease of secant stiffness due to the propagation of concrete shear cracks. Stiffness degradation follows a similar parabolic trend for all the strengthened specimens, caused initially by the unconfined concrete deterioration between consecutive CFRP strips
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and then by the crushing of confined concrete. It should be noted that specimens reinforced with high axial stiffness composite materials manifested greater secant stiffness than other specimens, due to the confinement effect of concrete, which reduces the cyclic damage.
The cumulative dissipation energy due to the cyclic loading is reported as a function of the drift ratio for the seven specimens in Fig. 7b. The results of cumulative energy are also reported in Table 4. The
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cumulative dissipation energy is evaluated only for the cyclic part of specimen S7 load-drift ratio relationship. A slight difference in the cumulative dissipation energy is observed for the control specimens, S1 and S5, and the strengthened that failed due to shear before the flexural yielding, S2 and
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S3. Thus in specimens failed in shear the energy dissipation was not affected by the CFRP external confinement. For the members that exhibited a “quasi-ductile” behaviour, specimens reinforced with high
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axial stiffness material dissipated more energy with respect to other specimens of same concrete class. Table 4. Cumulative energy at different drift ratios. 5. Strain analysis
The effective strains on the CFRP strips due to combined shear action and concrete lateral dilatation is investigated. The CFRP strains were measured by means of strain gauges located at mid-span and midheight of each side of the first four strips starting from the base column. The layout of strain gauges on CFRP strips is shown in Fig. 8. The ratio between CFRP experimental strain recorded at the positive peak drift of each cycle, εf, and ultimate composite material strain reported in Table 2, εfu, is presented for different positive imposed drift ratios ∆+ up to conventional failure along the column shear span, LS (i.e.
ACCEPTED MANUSCRIPT 6% LS, 24% LS, 42% LS and 61% LS). The effective strains recorded for specimen S7 are plotted up to the last cycle achieved before the monotonic loading. The strain profiles for negative imposed drift ratios are similar to those obtained for positive drift ratios and are not herein reported for brevity. Carbon strips work simultaneously for shear action, when crossed by a concrete diagonal crack, and for confinement,
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contrasting the concrete lateral dilatation due to a non-uniform distribution of compressive stress given by the combined action of axial load and bending moment. Thus in the positive load action direction, column side C is subjected to compression whereas side A is in compression for small imposed drift ratios otherwise is in tension. Strips in sides B and D work mainly for the shear action.
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The carbon fibres effective strains recorded on the four sides of the same generic CFRP strip resulted quite uniformly distributed for almost all specimens. As expected, the most significant strains were
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recorded on sides B and D with respect to other sides, probably due to the formation of new diagonal crack, which causes stress/strain concentration.
The strains recorded on specimens reinforced with two plies of composite material are slightly lower with
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respect to those recorded for one ply reinforcement. This is particularly evident for specimen S7.
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Fig. 8. CFRP strain distribution along the column height.
ACCEPTED MANUSCRIPT The strain profiles along the member shear length for a generic imposed drift ratio follow a similar trend for all specimens: higher strains are achieved from the second (i.e. 24% LS) and third (i.e. 42% LS) strips starting from the base column if compared with strain developed in first and forth strip. This is reasonable, as large diagonal cracks developed in the concrete region between second and third strips can
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be observed for the “as built” specimens. Maximum CFRP strains recorded are generally lower or equal to 40% of the ultimate composite material strain εfu, except for specimens S3 and S6 that achieved maximum strains up to 75% εfu. However, it should be noted that S3 and S6 are strengthened with type B composite material, characterized by a lower εfu than other CFRP types.
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International codes provisions for shear capacity of RC columns strengthened with FRP follow the truss analogy method and just add the FRP contribution, Vf, to the original capacity of the member. The FRP
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contribution to shear capacity is commonly calculated as an equivalent transverse reinforcement, as reported in Eq. (1). =
,
(1)
where εf,eff is an effective fibres strain, which is usually lower than the ultimate strain εfu. It is defined as the strain to be multiplied for the FRP elastic modulus and the available FRP cross sectional area to
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provide the shear contribution carried by the FRP. Thus, it is not related to the experimental strain recorded during tests, but it is derived by the knowledge of members’ capacity before and after the strengthening. Current codes limit the effective fibres strain, εf,eff, due to debonding considerations [32,
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32, 35, 36] or for preserving the concrete integrity inside fibres and the aggregate interlock mechanism (i.e. εf,eff = min{0.4%; 75%εfu} [34]). However, debonding problems are rare in case of fully wrapped
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configurations, where the FRP tensile rupture usually governs the failure mode. Only the fib bulletin 14 [35] provides a specific formulation for evaluating the effective fibres strain in case of fully wrapped configurations and fracture controlled failure modes [14], even though it is developed using experimental tests on beams.
The most recent literature review in European seismic design provisions [38] reports another formulation, valid for both shear and confinement applications, for evaluating the effective strain of fully wrapped members, based on geometrical considerations and accounting for debonding of external layers. In Fig. 8, the ratios εf/εfu at peak load evaluated along the sides B and D of specimens S2 and S3, which achieved a brittle failure even after the strengthening, are reported together with some analytical previsions of fibres
ACCEPTED MANUSCRIPT effective strains. The experimental strains are reported for both the positive (i.e. full markers) and negative (i.e. empty markers) load action directions. It should be noted that the ultimate strains εfu are different for composite materials used for specimens S2 and S3, see Table 1. The following analytical formulations are investigated for comparison with experimental outcomes: the
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ACI 440 (2008) [34], which limits the effective strain to 0.004; the Eurocode 8 - Part 3 (2004) (i.e. EC83) formulation [32]; the fib bulletin 14 (2001) formulation [33] and the formulation suggested in Pantazopoulou et al. (2016) [38]. Other approaches based on debonding [33] usually give effective strain lower than that of ACI 440 (2008) which is the most conservative of these formulations; for this reason,
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they are not herein considered. In Fig. 8 is also reported the ratio εf,eff,exp/εfu, where εf,eff,exp is the experimentally derived effective strain and represents the strain that is needed to increase the shear
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capacity of the unstrengthened members, evaluated according to Eurocode 8-3 formulation [32], up to the experimental maximum loads achieved from specimens S2 and S3.
The comparison between the derived ratio εf,eff,exp/εfu and the analytical formulations results shows that the strain limitation to 0.004 gives effective strains compatible with the experimental recorded strains but conservative with respect to the derived ratio εf,eff,exp/εfu. Very similar results are obtained by using the
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approach based on debonding considerations suggested by EC8-3 (2004). Conversely, the formulation suggested in Pantazopoulou et al. (2016) overestimates the CFRP contribution so as to the approach indicated in fib bulletin 14 (2001), even if it is more close to the derived ratio εf,eff,exp/εfu. These results
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confirm the need of experimental research on RC columns governed by shear and strengthened with FRP,
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in order to improve the accuracy in the evaluation of the fibres effective strain.
(a) (b) Fig. 8. Effective strains at peak load for specimens S2 (a) and S3 (b).
ACCEPTED MANUSCRIPT Conclusions L’Aquila earthquake (2009) confirmed the high occurrence of columns brittle failure due to shear, especially in case of reduced shear length (i.e. band-type windows, semi-basement, etc.). Indeed, among 284 heavily damaged RC buildings that were demolished after the earthquake, 56 (i.e. 20%) achieved
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columns shear failure. The use of external reinforcement systems, such as CFRP, is herein experimentally investigated as a reliable method for increasing the capacity of short RC members and prevent the brittle failure caused by shear for developing plastic deformations.
•
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Based on the experimental evidences, the following conclusions are reported:
For specimens characterized by poor quality concrete, CFRP configurations with axial rigidity,
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Ef ρf, lower or equal to 0.31GPa increased from 56% to 67% the original shear capacity of the short RC columns. However, the external reinforcement was in both cases not enough to avoid the brittle failure of the member. •
An external reinforcement ratio of 0.34 was able to prevent the brittle failure of the poor quality concrete members, allowing the development of plastic deformations up to drift ratios of +7.5%
•
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(-7.2%) and ductility 4.0.
Specimens characterized by medium quality concrete and strengthened with two plies of composite material, fully developed plastic deformations, with ultimate drift ratios equal or
•
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greater than 8.5%.
High composite material axial stiffness and high external reinforcement ratios increased the
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peak-to-peak secant stiffness of short columns and allowed a greater energy dissipation with respect to other configurations.
•
The CFRP strain profiles along the column shear length underlined the non-uniform distribution of strains among the different strips, due to the formation of concrete diagonal cracks in the region between 24%LS and 42%LS.
•
The peak strains recorded were lower than 40% of ultimate CFRP strain for all specimens, except for specimens S3 and S6, where maximum strain of 75%εfu were achieved.
ACCEPTED MANUSCRIPT •
A big dispersion of analytical formulations results provided by international codes and literature reviews for evaluating the FRP effective strains, necessary to predict the FRP shear contribution, has been observed.
The experimental results presented in this work and the recent earthquakes (i.e. Central Italy 2016) have
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proven the effectiveness of FRP techniques for increasing the shear capacity of existing RC members and preventing their brittle failures. However, the comparison between the experimentally derived effective strains and those provided by available analytical formulations confirmed the need of more experimental research for calibrating specific formulations for fully wrapped RC columns, in order to optimize the
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design process of the strengthening interventions, especially in terms of cost-benefit ratio. Acknowledgement
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strength. Journal of structural engineering 1999; 125(3): 255-264.
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by Means of EBR FRP Systems. Design Procedures for the Use of Composites in Strengthening of Reinforced Concrete Structures. Springer 2016. p.97-130.
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Nonconforming Full-Scale RC Columns. Journal of Structural Engineering 2013; 140(5). [32] CEN. Design of structures for earthquake resistance - Part 3: Assessment and reofitting of buildings. EN-1998-3, Eurocode 8. Brussell: European Committee for Standardization, 2005.
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[33] CNR-DT 200 R1/2013. Guide for the design and construction of externally bonded FRP systems for strengthening existing structures e materials, RC and PC structures, masonry structures”. Italian National Research Council, 2013. [34]
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S1359-8368(17)30949-6
DOI:
10.1016/j.compositesb.2017.07.028
Reference:
JCOMB 5164
To appear in:
Composites Part B
Received Date: 16 March 2017 Revised Date:
19 June 2017
Accepted Date: 25 July 2017
Please cite this article as: Del Zoppo M, Di Ludovico M, Balsamo A, Prota A, Manfredi G, FRP for seismic strengthening of shear controlled RC columns: Experience from earthquakes and experimental analysis, Composites Part B (2017), doi: 10.1016/j.compositesb.2017.07.028. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
ACCEPTED MANUSCRIPT FRP for seismic strengthening of shear controlled RC columns: experience from earthquakes and experimental analysis Authors: Marta Del Zoppoa*, Marco Di Ludovicob, Alberto Balsamob, Andrea Protab, Gaetano Manfredib Department of Engineering, University of Naples “Parthenope”, Italy
b
Department of Structures for Engineering and Architecture, University of Naples “Federico II”, Italy
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a
*Corresponding Author: Marta Del Zoppo; Email: [email protected] Abstract
The high vulnerability of existing Reinforced Concrete (RC) structures, even to moderate seismic events,
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has been confirmed from recent post-earthquake surveys. Short and wall-like RC columns are particularly prone to brittle failures, governed by concrete crushing. To reduce the vulnerability of existing RC
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columns, the use of externally bonded Fibre Reinforced Polymer (FRP) reinforcement has been recognized as an effective method for preventing the aforementioned brittle failure and, hence, increasing members’ lateral capacity and ductility. In the first part of this study, the results of an observational analysis on columns shear failures in RC buildings severely damaged after the L’Aquila earthquake are presented. The second part of the study presents and discusses the results of an experimental program
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carried out on seven short RC columns governed by shear failure under load reversal and compressive axial load. Both columns in “as built” configuration and strengthened in shear with discontinuous carbon FRP (CFRP) strips have been tested. Two classes of concrete have been used, in order to simulate
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structures with medium or poor material quality, and different external reinforcement ratios have been investigated. The specimens’ responses have been analysed in terms of failure modes, strength/deformation capacity and strain distribution in CFRP strips.
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Keywords: seismic retrofit; columns; FRP strengthening; shear failure. 1. Introduction
Recent post-earthquake surveys [1-3] revealed the high vulnerability of existing Reinforced Concrete (RC) structures, often designed for gravity loads only, even to moderate seismic events. Indeed, the lack of a proper seismic detailing and a wrong shear-flexure hierarchy often lead to columns brittle failures due to shear before attaining the flexural yielding. Furthermore, short and wall-like RC columns are commonly subjected to such a brittle failure, governed by concrete diagonal compression failure [4-7]. Semi-basements, band-type windows, knee beam stairs and masonry parapets can turn the behavior of
ACCEPTED MANUSCRIPT slender columns into short ones during seismic events. Since short columns are subjected to higher shear forces than slender ones, their capacity cannot satisfy the shear demand and brittle failures occur in such members. Moreover, structures designed for gravity loads only or according to “first-generation” seismic codes can respond in a completely different way to a seismic event, especially in terms of shear failures.
exhibited shear failures also in slender columns [8].
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After the L’Aquila earthquake (2009), structures designed according to a “first-generation” seismic code
To prevent premature brittle failures of existing RC columns, the use of Externally Bonded
Reinforcement (EBR) made of Fibre Reinforced Polymer (FRP) strips has been recognized as a reliable
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and cost saving strategy for increasing members lateral strength capacity [9-11]. Plenty of researchers
focused the attention on the shear strengthening of RC beams, mainly due to gravity loads deficiencies
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[12-17, among many others]. Even though post-earthquake observed shear failures in columns are more diffused than shear failures of beams [18], only a few experimental studies have been developed on RC columns strengthened with FRP [19-24], demonstrating the effectiveness of this technique for such members.
The present work is focused on the behavior of shear critical short RC columns strengthened with CFRP.
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The results of an experimental program carried out on short RC columns strengthened with CFRP are herein presented and discussed. In particular, seven RC columns with square cross-section and governed by shear mode were analysed: two reference columns, termed “as built” in the following, and five
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columns externally reinforced with discontinuous CFRP strips along their shear length. Two classes of concrete have been used, in order to simulate structures with medium or poor quality concrete, and different external reinforcement ratios have been investigated. The specimens were cantilevers loaded
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with a low compressive axial load (i.e. axial load ratio 0.1) and subjected to horizontal cyclic displacements. The responses of columns have been analysed in terms of failure modes and increased strength and deformation capacity as a function of the external reinforcement ratio. Furthermore, to investigate on the effectiveness of CFRP external reinforcement, a detailed analysis related to the axial strains distribution along the fibres strips is reported. 2. Occurrence of shear failure in RC columns: the L’Aquila earthquake experience 2.1 Observed damage
ACCEPTED MANUSCRIPT After the L’Aquila earthquake (6th April 2009, magnitude 6.3), thousands of RC structures outside the historical centres were damaged. From the analysis of the post-earthquake reconstruction process [25, 26], a total of 2,799 RC buildings resulted affected by the seismic event: 1,738 RC buildings (i.e. 62%) were lightly damaged, with local damage to non-structural components but few or no damage to the
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structural ones, whereas the other 1,061 buildings (i.e. 38%) were severely damaged to both structural and non-structural components. Among the 1,061 severely damaged RC buildings, 284 structures (i.e. 27%) with severe damage to structural members were demolished, since repair and strengthening interventions resulted not viable due to economic reasons. From an extensive analysis carried out on the 284
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demolished RC buildings, resulted that 56 structures (i.e. 20%) exhibited brittle failures of columns due to shear actions. These 56 structures have been selected and analysed for investigating the reasons of shear
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failure occurrence.
The selected buildings were designed between ‘60-‘90 with three up to seven storeys, regular in both plan and elevation. The concrete mechanical properties, derived from in-situ material tests, were found in the range 8-30 MPa. Thus, in several cases, the poor quality of concrete contributed to a reduced resisting capacity to shear actions. Columns that achieved a shear failure were generally located at the ground floor
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or at the semi-basement of the selected buildings, where the shear demand attains maximum values. Columns affected by shear failure were characterized by a rectangular/square cross-section with dimensions variable between 300x350 and 300x600 mm2 in case of rectangular, and 300x300 or 400x400
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mm2 in case of square members. For almost all buildings, columns had constant cross-section dimensions for each storey level, thus they were all subjected to a similar shear demand. In 59% of selected buildings (i.e. 33 out of 56), the columns exhibited diagonal shear cracks, but they
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were still able to bear axial loads. Conversely, heavy damaged columns were detected in remaining 41% (i.e. 23) with large diagonal crack openings and no residual axial load capacity. The columns weakness to shear actions is mainly related to the low amount of transverse reinforcement found in typical existing members. Ties of 6 mm diameter spaced at 200-300 mm closed with 90° hooks were usually found in columns that failed for shear (see Fig. 1a). The large spacing between consecutive ties allows large cracks openings and longitudinal bars buckling, leading to a premature shear failure.
(b)
(c)
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(a)
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In addition, for columns designed for seismic action in one direction only, the absence of longitudinal reinforcement in addition to corner rebars along the column secondary direction, which can help in shear
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resisting mechanism, was also observed (see Fig. 1b). In addition to the poor material quality or the absence of seismic detailing, the shear failure observed in columns often resulted as a consequence of the infills actions. In 23% of buildings under investigation
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(i.e. 13 buildings), the shear failure of columns was due to unexpected interaction effects with masonry infills (see Fig. 1c and d). Indeed, stiff infills generate significant shear actions concentrated near the top
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end section of columns, which can lead to premature brittle failures.
(d)
(e)
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ACCEPTED MANUSCRIPT
(f)
Fig. 1. Columns shear failure after L’Aquila earthquake (2009).
Finally, semi-basements (see Fig. 1e) and irregularly distributed infills, like in case of band-type openings
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(see Fig. 1f), reduce the actual shear length of columns, causing local shear failures. Short columns are subjected to a high shear demand and, hence, are more prone to shear failures with respect to slender
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members.
Even though the reasons of columns shear failure occurrence during seismic events are quite well known [5], it is still difficult to understand why in a sample structure only a few columns failed for shear among a large number of columns with same geometrical and mechanical properties and under similar load conditions (i.e. similar axial load condition and lateral stiffness). Indeed, in 62% of selected structures (35
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buildings), the shear failure occurred only for 1-2 columns. Conversely, shear failures on 3-5 columns were attained in 27% of selected buildings (15 buildings). Only in 6 buildings (i.e. 11%), the shear failure occurred in more than five columns. This post-earthquake survey founding can be related to local
project.
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deficiencies due to non-uniform material quality or reinforcement details non-conforming to the design
2.2 Strengthening solutions
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In case of repair and strengthening of L’Aquila damaged RC structures, several techniques were used. From the analysis of the reconstruction process, it was found that the most widely employed strengthening technique was the use of FRP composite systems [27, 28]. Indeed, 475 (i.e. 45%) heavily damaged RC structures adopted FRPs as the main retrofit solution for beams, columns and beam-column joints.
ACCEPTED MANUSCRIPT Fig. 2a shows the local strengthening solution on a short RC column of a school building lightly damaged by 2009 L’Aquila earthquake [28]. The effectiveness of using FRPs as a strengthening technique has been also proven by the recent earthquakes (i.e. Central Italy earthquake 2016): Fig. 2b shows a school complex consisting of two jointed buildings where the RC building retrofitted with FRP is mostly
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undamaged and the masonry structure collapsed during the 2016 Central Italy earthquake. The post-earthquakes experience raised, among others, these two issues: the assessment of existing
members’ shear capacity and the design of strengthening interventions with FRP composite systems.
International codes provide refined formulations for evaluating shear capacity in existing RC columns and
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practice-oriented tools for assessing the ductile/brittle behaviour of columns are presented in the literature [29]. However, accuracy and reliability of the available shear capacity models in predicting the column
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shear capacity before and after the strengthening is still a challenging task [8, 30], especially because theoretical formulations for designing FRP-based interventions are mainly derived for beams. Indeed, for the selected 56 structures, all shear capacity models considered by the Authors underestimated the effective members shear capacity and predicted shear failures also for many un-damaged columns. An inadequate accuracy of capacity models for predicting the members original shear capacity and the
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contribution of the FRP strengthening systems to the total capacity leads to the design of costly and often unnecessary strengthening interventions. Thus, the following experimental analysis is specifically targeted at investigating shear behavior of columns in both “as built” and FRP strengthened
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configurations, with the aim to point out the effective contribution of FRP external reinforcement to the capacity and the failure mode of shear critical RC columns. 3. Experimental program
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3.1 Specimens geometrical and mechanical properties
(a) (b) Fig. 2. FRP interventions on RC columns in L’Aquila (a) and Amatrice (b).
ACCEPTED MANUSCRIPT Seven short RC columns were designed with the same geometrical properties and reinforcement detailing as depicted in Figure 3a. The short columns have been over-designed for flexure, with the aim to result in a brittle failure governed by shear. Each specimen had a square cross-section 300 x 300 mm2 reinforced with ten 22 mm diameter deformed rebars (longitudinal geometric reinforcement ratio, ρl = As/bh = 4.2%
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with As total area of longitudinal steel reinforcement and b,h, cross section dimensions). The longitudinal reinforcement was placed only along the loading direction, no reinforcement was provided along the
secondary direction as previously observed for existing columns. The concrete cover thickness was 20
mm for each specimen. Steel transverse reinforcement was made of 8 mm diameter ties, spaced at 300
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mm apart (transverse geometrical reinforcement ratio, ρw = 0.11%) and with hooked ends. A reduced
1 CFRP ply D A
300 mm
C B
C
2 CFRP plies D
B
260 mm
As= 10Ø22
C
A 165
300 mm
A
100
Ø8/300
900 mm
D
B
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600 mm
1250 mm
Ø8/50
300 mm
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spacing has been adopted in the zone of load application to avoid localized damage.
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1200 mm
(a)
(b)
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Fig. 3. Specimens’ geometry: “as built” (a) and CFRP strengthened (b).
The specimens were cantilevers, made by a column and a foundation block: the column height was 1,250 mm; the foundation block was 1,200 x 1,200 x 600 mm3. The foundation block and the columns of each specimen were cast in two different phases to reproduce the discontinuity of concrete at the columnfoundation interface, as common practice in concrete casting procedures. The lateral load has been applied at a distance of 900 mm from the foundation block, in order to simulate the behaviour of short columns (i.e. LS/h = 3 with shear length LS = 900 mm). Two classes of concrete have been herein adopted: •
low compressive strength for four specimens (termed S1, S2, S3 and S4);
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medium compressive strength for the remaining three specimens (termed S5, S6 and S7).
One specimen for each concrete class was unstrengthened and used as control specimen (“as built” specimens, named S1 and S5 for low and medium compressive strength, respectively) whereas the other five specimens were strengthened with discontinuous uniaxial CFRP strips with fibres perpendicular to
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the column longitudinal axis. The CFRP strips, width wf = 100 mm, were spaced at sf = 165 mm for the total height of the column. Before applying the external reinforcement, the columns corners were rounded with radius R = 20 mm, in order to avoid the fibres premature rupture. The strips overlap was more than 250 mm for each layer along the column sides orthogonal to the transverse load action direction. Further
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details about the CFRP application are reported in Figure 3b. Specimens characterized by low concrete strength were reinforced with one ply of CFRP strips; conversely, two plies of CFRP strips were used for
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medium concrete specimens. Different axial rigidities of CFRP bonded strips were investigated, where the axial rigidity is expressed by the product Ef ρf, where Ef the elastic modulus of carbon fibres and ρf = 2tf wf/bsf is the FRP reinforcement ratio, with tf the thickness of dry fibres (i.e. Ef ρf = 0.16, 0.31, 0.34, 0.61 and 0.67GPa for specimens S2, S3, S4, S6 and S7 respectively). A summary of the specimens’ mechanical properties and external reinforcement configurations is reported in Table 1.
Test
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Table 1. Summary of the experimental program. fcm
fyl
fyw
ρf
Ef ρf
[MPa]
[MPa]
[MPa]
[GPa]
n° plies
CFRP type
-
-
-
14.3
S2
14.3
0.06
0.16
1
A
0.13
0.31
1
B
0.13
0.34
1
C
23.9
-
-
-
-
S6
32.3
0.27
0.61
2
B
S7
36.4
0.27
0.67
2
C
S3 S4
16.5 14.0
AC C
S5
EP
S1
[%] -
531
525
The same class of steel has been used for all the columns: average yielding strength fyl = 531 MPa for longitudinal reinforcement and fyw = 525 MPa for transverse reinforcement were derived from tensile tests on coupon bars. Four concrete 150-mm side cubes were cast during each specimen preparation and were cured under the same environmental conditions. The mean cylindrical concrete strength, fc, derived for each specimen is reported in Table 1. Three types of CFRP composite material with different unit weight and elastic modulus were used, in order to obtain different axial rigidities of bonded strips. The
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∆ [%] 0.2 mm/s
1.0 mm/s
2.0 mm/s
10%
5%
-5%
-10%
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0%
t [s]
-15% 1000
2000
3000
4000
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0
mechanical properties of the dry fibre material types declared by the manufacturers (i.e. elastic modulus
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of dry fibres Ef and ultimate strain εfu) are reported in Table 2. Table 2. Dry carbon fibres material properties.
εfu [%] 1.9 1.3 1.9
tf [mm] 0.16 0.33 0.33
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CFRP Unit weight Ef type [g/m2] [GPa] A 300 252 B 600 230 C 600 252 3.2 Test set-up, instrumentation and loading
The short columns were subjected to a constant compressive axial load and horizontal cyclic displacements. The lateral load was applied at a distance of 900 mm from the base cross-section by using
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an electrohydraulic actuator (stroke ± 250 mm, maximum load +500 kN, -300 kN) integrated in a steel reaction wall and a 500 kN load cell was used to measure lateral loads. A detailed description of the test
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setup and relevant instrumentation is reported in [31] and shown in Figure 4a.
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(b) Fig. 4. Test set-up (a) and drift protocol (b).
All the specimens were subjected to a normalized compressive axial load (ν = N / Acfc, where N is the compressive axial load, Ac is the concrete gross area and fc is the mean cylindrical concrete strength) equal to 0.1 and a lateral cyclic loading under drift control (i.e. ∆ = d/ LS with d equal to the horizontal
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applied displacement and LS = 900 mm), see Figure 4b. A rate of 0.2 mm/s was adopted for initial four
cycles, then higher rate of 1.0 mm/s was adopted for the next five cycles and 2.0 mm/s for the last cycles. For each cycle, the target displacement was achieved three times, see Figure 4b. Only one specimen (S7)
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was subjected to cyclic lateral displacements for the first seven cycles and then to monotonic load up to a drift ratio ∆ = 11% (i.e. d = 100 mm), as the actuator achieved its maximum load capacity in the negative
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load action direction. Strain gauges were located at the mid-height and mid-span of each strip for all the column sides, for the first four CFRP strips from the base, with the aim to monitor axial strains in the external reinforcement. 4. Experimental results
4.1 Lateral load-drift relation and failure mode
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The experimental lateral load-drift relationships for the seven short RC columns are depicted in Fig. 5a-g. The envelope curves normalized with respect to the concrete contribution,
with d the section
effective depth, are reported in Fig. 5h. The relevant experimental outcomes for both positive and negative load-action directions along with the observed failure mode are reported in Table 3: peak forces
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and relative drift ratios, Fmax and ∆Fmax, drift ratios at conventional failure, ∆0.8Fmax, defined as the values corresponding to 80% of the peak force on the envelope curve, percentage of increased strength due to the
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FRP strengthening with respect to the “as built” specimens in the positive and negative load direction, respectively, δFmax, and ductility, μ∆, defined as the ratio between ultimate and yielding drift. The drift at the yielding is assumed as the intersection point between the secant at the 70% of the peak strength and the horizontal at the peak strength. The crack patterns and relevant damage at conventional failure are shown in Fig. 6. Table 3. Experimental outcomes and observed failure mode. Test S1
Fmax [kN] 159.6 -151.8
∆Fmax
∆0.8Fmax
δFmax
[%] 1.6 -1.6
[%] 2.9 -2.8
[%] -
μ∆ [-]
failure mode
FRP failure
-
shear
-
S4 S5 S6 S7
3.2 -3.2 3.2 -3.2 2.4 -2.4 1.6 -1.6 6.4 -6.4 9.1 -
4.3 -3.7 6.2 -6.0 7.5 -7.2 2.7 -2.5 8.5 -8.0 >11 -
+61 +56 +57 +67 +80 +89 +63 +57 +87 -
4.0 3.9 4.8 4.6 >6 -
shear
2nd strip rupture
shear
-
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S3
256.8 -236.3 249.9 -253.0 286.7 -286.5 188.9 -191.0 308.3 -300.6 354.0 -
flexure-shear
local rupture
shear
-
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S2
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flexure-shear
-
flexure
-
The two “as built” specimens, S1 and S5, exhibited a brittle failure governed by shear before the flexural yielding, see Fig. 5a and e. Diagonal cracks extending throughout the entire shear length of the column
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were developed along the two sides parallel to the load action direction (i.e. B and D sides), see Fig. 6a and e. The peak force was achieved for both specimens at a drift ratio of 1.6%, as reported in Table 3. After the peak cycle, a sudden drop of the lateral strength was observed and the tests were stopped. The
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failure was governed in both cases by concrete web crushing, so defined shear diagonal compression failure, typical of short columns. After the conventional failure, the columns were still able to carry the
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axial loads. Peak lateral strength of +159.6 kN (-151.8 kN) and +188.9 kN (-191.0 kN) were attained on S1 and S5 respectively, see Table 3. Thus, a difference of 60% in concrete compressive strength increased the shear capacity, governed by concrete crushing, about 18%.
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(a)
(d)
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(c)
(f)
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(e)
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(g) (h) Fig. 5. Lateral load-drift relationships for the seven specimens.
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(c)
(e)
(f)
(g)
Fig. 6. Crack patterns and damage at failure.
(d)
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(a)
(h)
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The detailed behaviour of strengthened specimens is described as a function of the axial rigidity, Ef ρf, in the following: Specimen S2 (Ef ρf = 0.16GPa)
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•
Under lateral cyclic loading (Fig. 5b), the specimen exhibited a brittle failure governed by the fragile rupture of CFRP reinforcement before the flexural yielding, see Fig. 6b. The CFRP rupture, probably due to a strain concentration at corner, was limited to the second strip starting from the column base. Comparing the response of specimen S2 with the “as built” specimen S1, the external reinforcement
•
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increased the shear capacity up to +256.3 kN (-236.3 kN), which means 56-61% of strength enhancement. Specimen S3 (Ef ρf = 0.31GPa)
Specimen S3, exhibited a brittle behaviour governed by shear, see Fig. 5c. In this case, no rupture of the
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composite material was detected (Fig. 6c). Therefore, the shear failure in this case is controlled by the deterioration of unconfined concrete between consecutive strips, see Fig.6h, and by the loss of aggregate interlock inside the CFRP strips. Peak forces recorded for specimen S3 are close to those attained from
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specimen S2, about +249.9 kN (-253.0 kN), which means 57-67% of strength enhancement with respect to the “as built” specimen, see Table 3. •
Specimen S4 (Ef ρf = 0.34GPa)
Fig. 5d shows that specimen S4 exhibited a “quasi-ductile” behaviour, attaining the flexural yielding before a sudden drop of lateral capacity due to the CFRP rupture. In this case, a local rupture of the fibres is limited to the third strip starting from the column base, see Fig. 6d. Thus, an external reinforcement ratio of 0.40% changed the failure mode of the specimen from brittle to “quasi-ductile”, with ultimate drift ratios of +7.5% (-7.2%), see Table 3.
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Specimen S6 (Ef ρf = 0.61GPa)
The response of S6 is “quasi-ductile” since the flexural yielding was attained before the sudden drop of lateral capacity caused by the shear action, see Fig. 5f. The specimen failure is governed by the concrete deterioration and loss of aggregate interlock inside and between the CFRP strips. No fibres rupture was
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detected (see Fig. 6f). Thus, the external reinforcement ratio of 0.70% changed the failure mode of the specimen from brittle to “quasi-ductile”, with ultimate drift ratios of +8.5% (-8.0%), see Table 3. •
Specimen S7 (Ef ρf = 0.67GPa)
As shown in Fig. 5g, specimen S7 manifested a pure flexural behaviour. As aforementioned, the test was
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carried out as cyclic for first cycles and then as monotonic up to a drift ratio of 11%. No strength
degradation was observed after the flexural yielding, probably due to absence of cyclic loading. At the
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end of the test, the specimen was completely undamaged, see Fig. 6g. 4.2 Stiffness degradation and energy dissipation
To investigate the effects of CFRP external reinforcement on the columns stiffness degradation caused by the damaging under cyclic loading, the experimental stiffness has been computed as the peak-to-peak secant stiffness of the first cycle of each imposed drift ratio. The ratio between the experimental secant
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stiffness, K, and the theoretical elastic stiffness, Kflex = 3EcIg/LS3 (where Ec is the concrete elastic modulus and Ig is the gross section inertia), is plotted for the seven short RC columns as a function of the drift ratio in Fig. 7a. In case of specimen S7, the secant stiffness is calculated only for the cyclic part of the load-
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drift ratio curve. At first cycle, the ratio K/ Kflex is in the range between 31% and 45%: specimens S4 and S7, strengthened with type C composite material characterised by a higher axial stiffness than other types,
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manifested a greater initial secant stiffness if compared with other specimens.
(a) Fig. 7. Secant stiffness degradation (a) and cumulative energy (b).
(b)
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S6 34.7 139.4 512.5 1146.6 2017.9 5693.1 12412.7 28290.2 51307.5 76494.2
S7* 33.0 152.5 572.5 1316.0 3487.5 8664.2 18268.3 25703.5
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Drift Cumulative Energy [kNmm] [%] S1 S2 S3 S4 S5 0.2 38.6 38.6 32.2 33.7 41.9 0.4 160.1 154.9 127.5 148.9 162.6 0.8 758.3 571.7 510.9 546.0 654.0 1.2 1615.3 1262.3 1070.7 1199.8 1504.0 1.6 2725.3 2179.2 1855.8 2082.6 2708.6 2.4 5222.0 4683.2 3946.5 4944.3 5721.1 3.2 7668.5 8421.3 7520.5 10687.7 8391.5 4.8 17540.0 15609.4 23557.6 6.4 38249.7 8 50847.9 * Energy calculated only for the cyclic part of the force-drift relationship.
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As built specimens, S1 and S5, experienced a rapid decrease of secant stiffness due to the propagation of concrete shear cracks. Stiffness degradation follows a similar parabolic trend for all the strengthened specimens, caused initially by the unconfined concrete deterioration between consecutive CFRP strips
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and then by the crushing of confined concrete. It should be noted that specimens reinforced with high axial stiffness composite materials manifested greater secant stiffness than other specimens, due to the confinement effect of concrete, which reduces the cyclic damage.
The cumulative dissipation energy due to the cyclic loading is reported as a function of the drift ratio for the seven specimens in Fig. 7b. The results of cumulative energy are also reported in Table 4. The
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cumulative dissipation energy is evaluated only for the cyclic part of specimen S7 load-drift ratio relationship. A slight difference in the cumulative dissipation energy is observed for the control specimens, S1 and S5, and the strengthened that failed due to shear before the flexural yielding, S2 and
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S3. Thus in specimens failed in shear the energy dissipation was not affected by the CFRP external confinement. For the members that exhibited a “quasi-ductile” behaviour, specimens reinforced with high
AC C
axial stiffness material dissipated more energy with respect to other specimens of same concrete class. Table 4. Cumulative energy at different drift ratios. 5. Strain analysis
The effective strains on the CFRP strips due to combined shear action and concrete lateral dilatation is investigated. The CFRP strains were measured by means of strain gauges located at mid-span and midheight of each side of the first four strips starting from the base column. The layout of strain gauges on CFRP strips is shown in Fig. 8. The ratio between CFRP experimental strain recorded at the positive peak drift of each cycle, εf, and ultimate composite material strain reported in Table 2, εfu, is presented for different positive imposed drift ratios ∆+ up to conventional failure along the column shear span, LS (i.e.
ACCEPTED MANUSCRIPT 6% LS, 24% LS, 42% LS and 61% LS). The effective strains recorded for specimen S7 are plotted up to the last cycle achieved before the monotonic loading. The strain profiles for negative imposed drift ratios are similar to those obtained for positive drift ratios and are not herein reported for brevity. Carbon strips work simultaneously for shear action, when crossed by a concrete diagonal crack, and for confinement,
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contrasting the concrete lateral dilatation due to a non-uniform distribution of compressive stress given by the combined action of axial load and bending moment. Thus in the positive load action direction, column side C is subjected to compression whereas side A is in compression for small imposed drift ratios otherwise is in tension. Strips in sides B and D work mainly for the shear action.
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The carbon fibres effective strains recorded on the four sides of the same generic CFRP strip resulted quite uniformly distributed for almost all specimens. As expected, the most significant strains were
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recorded on sides B and D with respect to other sides, probably due to the formation of new diagonal crack, which causes stress/strain concentration.
The strains recorded on specimens reinforced with two plies of composite material are slightly lower with
AC C
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respect to those recorded for one ply reinforcement. This is particularly evident for specimen S7.
AC C
EP
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Fig. 8. CFRP strain distribution along the column height.
ACCEPTED MANUSCRIPT The strain profiles along the member shear length for a generic imposed drift ratio follow a similar trend for all specimens: higher strains are achieved from the second (i.e. 24% LS) and third (i.e. 42% LS) strips starting from the base column if compared with strain developed in first and forth strip. This is reasonable, as large diagonal cracks developed in the concrete region between second and third strips can
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be observed for the “as built” specimens. Maximum CFRP strains recorded are generally lower or equal to 40% of the ultimate composite material strain εfu, except for specimens S3 and S6 that achieved maximum strains up to 75% εfu. However, it should be noted that S3 and S6 are strengthened with type B composite material, characterized by a lower εfu than other CFRP types.
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International codes provisions for shear capacity of RC columns strengthened with FRP follow the truss analogy method and just add the FRP contribution, Vf, to the original capacity of the member. The FRP
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contribution to shear capacity is commonly calculated as an equivalent transverse reinforcement, as reported in Eq. (1). =
,
(1)
where εf,eff is an effective fibres strain, which is usually lower than the ultimate strain εfu. It is defined as the strain to be multiplied for the FRP elastic modulus and the available FRP cross sectional area to
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provide the shear contribution carried by the FRP. Thus, it is not related to the experimental strain recorded during tests, but it is derived by the knowledge of members’ capacity before and after the strengthening. Current codes limit the effective fibres strain, εf,eff, due to debonding considerations [32,
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32, 35, 36] or for preserving the concrete integrity inside fibres and the aggregate interlock mechanism (i.e. εf,eff = min{0.4%; 75%εfu} [34]). However, debonding problems are rare in case of fully wrapped
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configurations, where the FRP tensile rupture usually governs the failure mode. Only the fib bulletin 14 [35] provides a specific formulation for evaluating the effective fibres strain in case of fully wrapped configurations and fracture controlled failure modes [14], even though it is developed using experimental tests on beams.
The most recent literature review in European seismic design provisions [38] reports another formulation, valid for both shear and confinement applications, for evaluating the effective strain of fully wrapped members, based on geometrical considerations and accounting for debonding of external layers. In Fig. 8, the ratios εf/εfu at peak load evaluated along the sides B and D of specimens S2 and S3, which achieved a brittle failure even after the strengthening, are reported together with some analytical previsions of fibres
ACCEPTED MANUSCRIPT effective strains. The experimental strains are reported for both the positive (i.e. full markers) and negative (i.e. empty markers) load action directions. It should be noted that the ultimate strains εfu are different for composite materials used for specimens S2 and S3, see Table 1. The following analytical formulations are investigated for comparison with experimental outcomes: the
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ACI 440 (2008) [34], which limits the effective strain to 0.004; the Eurocode 8 - Part 3 (2004) (i.e. EC83) formulation [32]; the fib bulletin 14 (2001) formulation [33] and the formulation suggested in Pantazopoulou et al. (2016) [38]. Other approaches based on debonding [33] usually give effective strain lower than that of ACI 440 (2008) which is the most conservative of these formulations; for this reason,
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they are not herein considered. In Fig. 8 is also reported the ratio εf,eff,exp/εfu, where εf,eff,exp is the experimentally derived effective strain and represents the strain that is needed to increase the shear
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capacity of the unstrengthened members, evaluated according to Eurocode 8-3 formulation [32], up to the experimental maximum loads achieved from specimens S2 and S3.
The comparison between the derived ratio εf,eff,exp/εfu and the analytical formulations results shows that the strain limitation to 0.004 gives effective strains compatible with the experimental recorded strains but conservative with respect to the derived ratio εf,eff,exp/εfu. Very similar results are obtained by using the
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approach based on debonding considerations suggested by EC8-3 (2004). Conversely, the formulation suggested in Pantazopoulou et al. (2016) overestimates the CFRP contribution so as to the approach indicated in fib bulletin 14 (2001), even if it is more close to the derived ratio εf,eff,exp/εfu. These results
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confirm the need of experimental research on RC columns governed by shear and strengthened with FRP,
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in order to improve the accuracy in the evaluation of the fibres effective strain.
(a) (b) Fig. 8. Effective strains at peak load for specimens S2 (a) and S3 (b).
ACCEPTED MANUSCRIPT Conclusions L’Aquila earthquake (2009) confirmed the high occurrence of columns brittle failure due to shear, especially in case of reduced shear length (i.e. band-type windows, semi-basement, etc.). Indeed, among 284 heavily damaged RC buildings that were demolished after the earthquake, 56 (i.e. 20%) achieved
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columns shear failure. The use of external reinforcement systems, such as CFRP, is herein experimentally investigated as a reliable method for increasing the capacity of short RC members and prevent the brittle failure caused by shear for developing plastic deformations.
•
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Based on the experimental evidences, the following conclusions are reported:
For specimens characterized by poor quality concrete, CFRP configurations with axial rigidity,
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Ef ρf, lower or equal to 0.31GPa increased from 56% to 67% the original shear capacity of the short RC columns. However, the external reinforcement was in both cases not enough to avoid the brittle failure of the member. •
An external reinforcement ratio of 0.34 was able to prevent the brittle failure of the poor quality concrete members, allowing the development of plastic deformations up to drift ratios of +7.5%
•
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(-7.2%) and ductility 4.0.
Specimens characterized by medium quality concrete and strengthened with two plies of composite material, fully developed plastic deformations, with ultimate drift ratios equal or
•
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greater than 8.5%.
High composite material axial stiffness and high external reinforcement ratios increased the
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peak-to-peak secant stiffness of short columns and allowed a greater energy dissipation with respect to other configurations.
•
The CFRP strain profiles along the column shear length underlined the non-uniform distribution of strains among the different strips, due to the formation of concrete diagonal cracks in the region between 24%LS and 42%LS.
•
The peak strains recorded were lower than 40% of ultimate CFRP strain for all specimens, except for specimens S3 and S6, where maximum strain of 75%εfu were achieved.
ACCEPTED MANUSCRIPT •
A big dispersion of analytical formulations results provided by international codes and literature reviews for evaluating the FRP effective strains, necessary to predict the FRP shear contribution, has been observed.
The experimental results presented in this work and the recent earthquakes (i.e. Central Italy 2016) have
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proven the effectiveness of FRP techniques for increasing the shear capacity of existing RC members and preventing their brittle failures. However, the comparison between the experimentally derived effective strains and those provided by available analytical formulations confirmed the need of more experimental research for calibrating specific formulations for fully wrapped RC columns, in order to optimize the
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design process of the strengthening interventions, especially in terms of cost-benefit ratio. Acknowledgement
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