Seismic performance of CFRP-confined circular high-strength concrete columns with high axial compression ratio

Construction and Building Materials 134 (2017) 91–103

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Seismic performance of CFRP-confined circular high-strength concrete columns with high axial compression ratio Daiyu Wang a,b, Zhenyu Wang a,b,⇑, Scott T. Smith c, Tao Yu d a

Key Lab of Structures Dynamic Behavior and Control of the Ministry of Education (Harbin Institute of Technology, Harbin, China School of Civil Engineering, Harbin Institute of Technology, Harbin, China c School of Environment, Science and Engineering, Southern Cross University, Lismore, NSW, 2480, Australia d Faculty of Engineering and Information Sciences, University of Wollongong, Wollongong, NSW, Australia b

h i g h l i g h t s
Seismic performance of high-strength concrete (HSC) columns confined with CFRP.
Confinement of potential plastic hinge regions with CFRP wraps.
Experimental investigation of half-scale concrete columns.
Energy dissipation, stiffness degradation and damping characteristics quantified.

a r t i c l e

i n f o

Article history: Received 12 July 2016 Received in revised form 28 October 2016 Accepted 21 December 2016

Keywords: FRP High-strength concrete Confinement Columns Seismic performance

a b s t r a c t This paper presents an experimental investigation on the seismic performance of circular high-strength concrete (HSC) columns confined with carbon fiber-reinforced polymer (CFRP) composites. A total of eleven 1/2 scale columns were constructed of which nine were confined with CFRP wraps at potential plastic hinge regions. All columns were tested under combined high axial compression load and cyclic lateral displacement excursions. The primary variables of the tests were the axial compression load level, concrete strength, and the extent of the CFRP wrapping at the plastic hinge region. In order to evaluate the residual seismic capacity of CFRP-confined columns, three of the confined specimens were initially loaded to induce damage. The load was then removed after which the same columns were loaded to failure. The failure modes, hysteretic responses, energy dissipation and stiffness degradation characteristics, and the equivalent viscous damping ratios of the tested columns were then presented and interpreted. The test results showed that CFRP wraps applied at potential hinge regions resulted in significantly improved ductility and energy dissipation capacities of the columns even when tested under a high axial compression ratio. The plastic deformation capacity of the CFRP-confined columns was observed to decrease with an increase of axial compression ratio though. In addition, pre-damaged CFRP-confined columns may have insufficient residual seismic capacity due to the damage and failure in the unconfined regions under high axial compression load levels. Finally, empirical models of the degradation of effective and unloading stiffness are provided based on the test results. Ó 2016 Elsevier Ltd. All rights reserved.

1. Introduction In the past two decades, high-strength concrete (HSC) has been increasingly applied in structural engineering due to superior material properties related to strength, stiffness, and durability. It is widely accepted, however, that the ductility of HSC generally decreases with an increase of compressive strength. High strength ⇑ Corresponding author at: School of Civil Engineering, Harbin Institute of Technology, Harbin, China. E-mail addresses: [email protected] (D. Wang), [email protected] (Z. Wang), [email protected] (S.T. Smith), [email protected] (T. Yu). http://dx.doi.org/10.1016/j.conbuildmat.2016.12.108 0950-0618/Ó 2016 Elsevier Ltd. All rights reserved.

concrete structural members therefore generally exhibit a lack of ductile behavior and hence brittle failure. The application of HSC structural members in potential earthquake regions, where significant ductility and energy dissipation capacity are required, poses a real challenge. In order to provide a potential solution to this challenge, previously conducted studies indicate that appropriate lateral confinement may prove viable [1,2]. Externally bonded fiber-reinforced polymer (FRP) composites applied in the hoop direction can be an effective method for improving the ductility of confined concrete columns. This has been demonstrated by considerable investigations on the axial compressive behavior [3–9] and seismic performance [10–18] of such

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columns. The majority of these existing studies have concentrated on FRP-confined normal-strength concrete (NSC) columns. At present, a number of experimental and analytical studies by comparison have also been conducted on the axial compressive behavior of FRP-confined HSC columns [19–29]. HSC is normally defined as concrete with a cylinder compressive strength exceeding 50 MPa [23]. This is also the definition adopted in the present paper. However, what is limited in the literature though is research on the seismic performance of FRP-confined HSC columns [30–37]. Ozbakkaloglu et al. [30–32] conducted a series of experimental investigations on the seismic behavior of HSC columns confined by FRP tubes. The FRP tube was used as a stay-in-place formwork as well as lateral confinement. Key parameters of axial load level, concrete strength, column cross-sectional shape, amount and type of FRP, and corner radius were investigated and evaluated. The test results showed that the both circular and square HSC filled FRP tube columns were capable of developing considerable inelastic deformation capacities under simulated seismic loading. Recently, Ozbakkaloglu and Idris [33,34] extend their research to the seismic performance of a new type of FRP-concrete-steel double-skin tubular column, which was originally proposed by Teng et al. [38]. Their test results indicated that FRP-HSC-steel double-skin tubular columns can also develop excellent inelastic deformation capacities under seismic load compared with concrete-filled FRP tubes. The axial compression ratios in all the Ozbakkaloglu et al.’s tests were less than 0.45. In addition, Zohrevand and Mirmiran’s [35–37] studies focused on FRP tube confined columns made of ultra-high performance concrete. The compressive strength of the concrete was up to about 183 MPa, however, the axial compression ratios were also relatively low (i.e. 0.03). In an RC frame building, the actual axial compression ratio can be larger when subjected to severe earthquakes, especially for those with a significant vertical acceleration component [39,40]. It is widely known that the ductility of columns decreases with an increase of axial compression ratio. Therefore, there is a need to investigate the seismic performance of FRP-confined HSC columns with a relatively high axial compression ratio. In addition, research on the residual seismic performance of damaged FRP-confined columns has so far been very limited. The seismic safety of FRP-confined columns after major earthquake damage also needs to be investigated and evaluated. In light of the identified research demands, this paper presents the results of an experimental investigation on eleven 1/2 scale circular HSC columns subjected to combined constant axial compression and reversed cyclic lateral displacements. Nine of the specimens were confined with externally bonded carbon FRP (CFRP) at potential plastic hinge regions at the column ends. The main test parameters were (i) the axial compression ratio, which was up to 0.65, (ii) concrete strength, and (iii) the height of the confined region at the plastic hinge regions. The residual seismic capacity of the test specimens was also evaluated by testing damaged FRP-confined columns to failure. The damage was applied by initially loading the same confined columns after which the load was removed. The seismic performance of the columns was investigated in detail with respect to failure modes, lateral loaddisplacement hysteretic responses, energy dissipation, stiffness degradation and equivalent viscous damping ratio. Empirical models for the degradation of effective and unloading stiffness are also provided based on the test results.

2. Experimental program 2.1. Test specimens and material properties A total of 11 I-shaped circular HSC columns were prepared and tested under combined constant axial compression load and lateral

reverse displacement. Each specimen was comprised of nominally identical 180 mm diameter circular cross section (D), 1260 mm height (H), and two integrally cast stubs with dimension of 300  370  780 mm at the column ends. In addition, the shear span ratio (H/2D) of each specimen was 3.5. Six deformed steel bars of 12 mm diameter were evenly arranged in a circle to provide the longitudinal reinforcement which resulted in a longitudinal steel reinforcement ratio of 2.7%. All longitudinal deformed steel bars were effectively anchored into the two end stubs. Circular hoop steel of 4 mm diameter undeformed steel bars was utilized to provide transverse reinforcement. Finally, the concrete cover within the length of each column was 8 mm. To aid in the evaluation of the confinement provided by the FRP, the steel transverse reinforcement was designed according to an older Chinese seismic design code (GBJ11-89) [41] (referred to as substandard columns). In this case, the spacing of the hoop steel at the two potential plastic hinge regions of each column was 60 mm, leading to a volumetric ratio of about 0.6%. The spacing of the hoop steel increased to 120 mm outside the potential plastic hinge regions. The height of the potential plastic hinge regions was assumed to be 220 mm (i.e. about 1.2D). Tensile tests on three hoop steel bars and three longitudinal steel bars were conducted according to ASTM E8/E8M [42] and the average yield stresses were found to be 354 MPa and 402 MPa, respectively. The dimension and reinforcement details of the specimens are illustrated in Fig. 1. The details of material properties are provided in Table 1. All specimens were cast vertically using two different grades of HSC. The two grades of concrete were mixed at the laboratory and they were comprised of water-to-cement ratios of 0.32 and 0.25, respectively. At column testing, the average compressive standard cylinder strengths of 54.8 MPa and 71.2 MPa, respectively, were measured. Nine column specimens were retrofitted by wrapping four layers of unidirectional carbon fiber sheet laterally in the potential plastic hinge zones. The carbon fiber sheet was of 0.167 mm nominal thickness and the FRP wrap was formed in a wet layup manner. The average tensile strength, elastic modulus and ultimate tensile strain of the FRP in accordance with ASTM D3039 [43] were 3430 MPa, 230 GPa and 1.8%, respectively. These properties are also summarized in Table 1. The height of the wrapped region was designed to be 200 mm or 320 mm (i.e. 1.1D or 1.8D), based on existing studies on the plastic hinge length of RC columns and FRP-confined RC columns [44–47]. For example, Sheikh and Khoury [44] suggested a simple model that specified the plastic hinge length to be 1.0H for columns under high axial loads. In addition, Tirasit and Kawashima’s [45] tests showed that the length of the damage zone was between 0.5D and 1.5D for columns under different levels of load. Similar results were also obtained upon the application of other models [46,47]. Consequently, the wrapping height was conservatively chosen to be 1.2D (i.e. 200 mm) and 1.8D (i.e. 320 mm) for the relative lower and higher axial compression ratios, respectively. The specimen details are summarized in Table 2. In the table, the specimen naming convention starts with a ‘‘C” or ‘‘R” to represent control specimens or CFRP-confined specimens. The following number refers to the height of the confined region with 0 denoting an unconfined column, and 2 and 3 denoting confined heights of 200 mm and 320 mm, respectively. Next, the letter S and the following number 5 or 7 represents concrete strengths of 54.8 MPa and 71.2 MPa, respectively. The third letter A and the following number represent the axial compression ratio (n) (i.e. the intensity ratio of the gross section capacity), where 45, 55, and 65 represent the axial compression ratio of 0.45, 0.55 and 0.65, respectively. For the last three specimens in the table, the final letter D refers to predamage. For example, specimen R3S1A55D refers to an initially damaged FRP-confined column with wrapped length of 320 mm

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(a) Dimension and reinforcement details

(b) Steel cage

Fig. 1. Specimen dimensions and reinforcement details.

Table 1 Mechanical properties of steel reinforcement and CFRP wrap. Material

Diameter/Thickness (mm)

Yield strength (MPa)

Yield/Rupture strain (%)

Elastic modulus (GPa)

Ultimate strength (MPa)

Steel

12 4 0.167

354 402 –

0.18 0.20 1.80

198 201 230

516 808 3430

CFRP

Table 2 Specimen details. Specimen

D (mm)

H (mm)

Longitudinal steel

Hoop steel

fc (MPa)

n

FPR sheet layers

Wrapped height (mm)

C0S5A45 R2S5A45 R3S5A55 R3S5A65 C0S7A45 R2S7A45 R3S7A55 R3S7A65 R3S5A55D R3S5A65D R3S7A65D

180

1260

6/12

/4@60 (/4@120)

54.8 54.8 54.8 54.8 71.2 71.2 71.2 71.2 54.8 54.8 71.2

0.45 0.45 0.55 0.65 0.45 0.45 0.55 0.65 0.55 0.65 0.65

0 4 4 4 0 4 4 4 4 4 4

0 200 320 320 0 200 320 320 320 320 320

Note: D = diameter of cross section; H = height of specimens; fc = average compressive standard cylinder strengths; n = axial compression ratio.

at the column ends, constructed with 54.8 MPa concrete compressive strength, and subjected to an axial compression ratio of 0.55. 2.2. Test setup and instrumentation All the columns were tested under combined constant compression and incrementally increased lateral reverse displacement. The tests were conducted in a purpose built testing frame, shown

schematically in Fig. 2, of which second-order effects (i.e. P–D effect) could be incorporated. The constant axial load was applied through a hydraulic jack of 2000 kN load capacity to the top of the specimens before applying the lateral displacement. The applied axial load was determined based on the considered axial compression ratio which varied from 0.45 to 0.65, as shown in Table 2. The axial compression ratios, n, were calculated according to the Chinese design code of GB50011-2010 [48] as per:

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Fig. 2. Test setup.

n ¼ N=Afc

ð1Þ

where N is the applied axial load, fc is the cylinder compressive strength and A is the gross area of the column section. The reverse lateral load was applied by an electro-hydraulic actuator with a 600 kN load capacity. The lateral loading scheme was the same for both control and FRP-confined columns, as shown in Fig. 3. The first and second displacement levels were applied under a load control mode and they were reversed only one time per each level. The goal of the first two load levels was 30% and 60% of the theoretical peak lateral load (Pc) of the specimens, respectively. Based on finite element analysis using the Open System for Earthquake Engineering Simulation (OpenSees) software, the average value of Pc for the control and FRP-confined specimens

Fig. 3. Cyclic lateral loading scheme.

was found to be approximately 65 kN. Consequently, the target lateral load of the first and second cycles was set to 20 kN and 40 kN, respectively. After completing these two cycles, the subsequent displacement levels were applied under displacement control mode and reversed twice for each level. The displacement controlled levels increased with an increment of yield displacement (Dy) of the control specimens. The Dy refers to the lateral displacement causing tensile yield of the longitudinal steel bars. Yield was determined by monitoring the strains in the instrumented bars (as shown in Fig. 4) during the first two load cycles. It was found that Dy (i.e. the increment of the displacement level) of the control specimens was about 5 mm. For the pre-damaged specimens (i.e. R3S5A55D, R3S5A65D and R3S7A65D), the specimens were loaded to approximately Pc in three to four levels (i.e. 1/550 lateral drift ratio, 0.3Pc, 0.6Pc and Pc) under a load control mode. Then, the specimens were unloaded to their initial stage and retested under the same cyclic loading scheme as the other specimens. Finally, the tests were stopped when the lateral load reduced to 80% of the peak load or when the lateral displacement reached the limitation of the actuator. Pre-loading was utilized in order to simulate frequent or moderate seismic actions. The current Chinese seismic design code (GB50011-2010) [48] requires that structures should not experience severe damage under frequent and moderate earthquakes. For RC frames, the specified upper limit of the elastic and inelastic interstory drift ratio is 1/550 and 1/50, respectively. In this study, the drift ratio of the pre-damaged specimens corresponded to the elastic drift ratio limit of 1/550 at the first predamage level. The drift ratio was about 1/100 at the final predamage level, which was between the elastic and inelastic drift ratio limit.

Fig. 4. Instrumentation layout.

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All test specimens were instrumented with load cells, linear variable displacement transducers (LVDTs) and electrical strain gauges in order to measure the axial and lateral loads, the lateral displacements, and strain responses, respectively, as shown in Fig. 4. The lateral displacement response was measured by a total of four LVDTs positioned along the length of each column, as also shown in Fig. 2. Four electrical strain gauges of 2 mm gauge length were placed on the longitudinal steel with a further eight of the same strain gauges installed on the hoop steel at the two ends of the columns. A total of 8 strain gauges (i.e. 4 in the lateral and 4 in the axial directions) of 50 mm gauge length were adhered to the outer surface of the CFRP wraps or concrete at the two plastic hinge regions of the columns. In Fig. 4, the symbols of HC and LC refer to strain gauges on the CFRP wrap in the horizontal and longitudinal directions, respectively, while HS and LS refer to strain gauges on the hoop and longitudinal bars, respectively. 3. Experimental results and discussions 3.1. General test observations and failure modes Fig. 5 shows the typical failure modes of the test specimens. The control specimens exhibited a similar failure process and they failed in a flexural manner. Initially, horizontal flexural cracks first formed on the tension side of a plastic hinge region from the second cycle (i.e. 0.6Pc level). Then, more flexural cracks occurred and the initial cracks widened and extended with an increase of

lateral displacement. Vertical cracks and concrete cover spalling were observed to occur when the lateral force reached the peak strength of the columns and the hoop steel began to yield simultaneously. Diagonal shear cracks were, however, not noticed. The columns finally failed due to severe spalling and crushing of the concrete and buckling of the longitudinal bars in a plastic hinge region. In addition, the amount of damage of the higher concrete strength specimen (i.e. C0S7A45) was more extensive than that of the specimen with relative low concrete strength (i.e. C0S5A45), as shown in Fig. 5(a and b). The region concrete cover which suffered from severe spalling or crushing at the column ends was positioned between 15 mm to 240 mm away from the column-footing interface for specimen C0S5A45, and between 10 mm to 350 mm for specimen C0S7A45. The damage of FRP-confined columns was effectively reduced compared to the two control specimens. In the initial stage of the tests, no apparent damage was noticed except audible epoxy cracking. With an increase in lateral displacement to about 35 mm corresponding to a drift ratio of about 3%, a few horizontal cracks on the CFRP wraps were formed due to excessive flexural deformation of the columns, as shown in Fig. 5(c–h). Similar observations were also reported by Ozbakkaloglu et al. [30–32] where that localized changes in FRP color occurred within the plastic hinge region of HSC filled FRP tube columns at 3% drift ratio. However, no rupture of the CFRP wraps was observed after the tests in this study. This is difference with Ozbakkaloglu et al.’s test results where fiber rupture was the failure mode for almost all the FRP tube columns

Horizontal cracks

Interface crack

(a) C0S5A45

(b) C0S7A45

(c) R2S5A45

(d) R3S5A55

(e) R3S5A65

Interface crack (g) R3S7A55

(h) R3S7A65

(i) R3S5A55D

(j) R3S5A65D (k) R3S7A65D

Fig. 5. Failure modes.

(f) R2S7A45

removed (l) CFRP

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[30–32]. The interface of the column end and the footing cracked when the lateral displacement reached about 40 mm. This was also the final failure mode of the FRP-confined specimens. The unconfined regions of the columns were not observed to exhibit significant damage although several vertical cracks were found at the end of the tests. For specimen R3S7A65 which had a high concrete strength (i.e. 71.2 MPa) and was subjected to a high axial compression ratio (i.e. 0.65), the damage of the unconfined region was found to be more severe than that of the other confined specimens (Fig. 5h). In order to better observe the damage of the confined regions, CFRP wraps were removed after the tests. Removal of the CFRP wraps showed that the concrete cover was severely spalled and the longitudinal bars had buckled in most cases, as shown in Fig. 5(l). The severely damaged regions were located 50 mm to 300 mm away from the column-footing interface and this corresponded to 0.3 to 1.7 times the cross section diameter (D). Ozbakkaloglu et al.’s [30–42] study observed similar results where the most extensive damage occurred at the ends of HSC filled FRP tube columns at about 0.5D to 1.3D from the columnfooting interface. The plastic hinge length based on the variation of FRP tube hoop strains along column height was 1.0D to 2.0D. Previous studies also have confirmed that the degree of lateral confinement can affect the extent of the damage zone of FRP-confined columns. The calculated plastic hinge length with reference to Gu et al.’s model [49] was about 110 mm (i.e. 0.61D) and 165 mm (i.e. 0.91D) for the two concrete strength columns of this study, respectively. While the prediction was slightly smaller compared with the observed severe damage zones in this study, the variation trends of the plastic hinge length increased with an increase of concrete strength. This was consistent with test observations. The pre-damaged FRP-confined specimens generally exhibited a similar failure process as the undamaged specimens, except for specimen R3S7A65D which had a high concrete strength (i.e. 71.2 MPa) and was subject to a high axial compression ratio (i.e. 0.65), as shown in Fig. 5(i–k). For specimen R3S7A65D, no significant damage was noticed in the FRP-confined regions, but the unconfined region was severely damaged (see Fig. 5(k)). The specimen fractured outside of the confined area and the longitudinal bars were seriously buckled. Such failure in the unconfined region was sudden and brittle. It is demonstrated that pre-damaged FRPconfined columns with high concrete strength and significant axial compression ratio demanded a larger wrapped area in the plastic hinge region. Otherwise, the columns will exhibit insufficient residual seismic capacity. A65D exhibited brittle failure in the outside of the CFRPconfined region. 3.2. Lateral load-displacement hysteretic curves The experimental lateral load-displacement hysteretic curves of all the specimens are shown in Fig. 6. It can be observed that the unconfined control specimens exhibited narrow hysteretic loops and rapid decreases in lateral resistance after the peak point at small lateral displacement levels. In addition, the ductility and energy dissipation capacity of the columns decreased with an increase of the concrete strength. For example, specimen C0S7A45 exhibited rapid post-peak strength degradation, a less number of hysteretic loops, and inferior hysteretic behavior than that of specimen C0S5A45, as shown in Fig. 6(a) and (b). The hysteretic behavior of CFRP-confined columns was superior to that of the control specimens, as shown in Fig. 6(c–h). Compared to the control specimens, the confined columns exhibited much wider hysteretic loops and more gradual post-peak strength degradation behavior. The CFRP wraps located in the plastic hinge regions substantially enhanced the ductility and energy dissipation capacity of the columns. It can be found that the seismic perfor-

mance was all significantly improved for the specimens with CFRP wrapping despite the wrapping height. The seismic performance of HSC columns can be effectively improved when the plastic hinge region is sufficiently confined, even when subjected to a high axial compression ratio. Comparison between specimens under the same axial compression but different concrete strength (i.e. C2S5A45 and C2S7A45, C3S5A65 and C3S7A65) revealed that the peak lateral strength and energy dissipation capacity of the confined columns was improved with an increase of the concrete strength when subjected to a relative lower axial compression ratio (i.e. 0.45). However, the influence of concrete strength on the hysteretic behavior was reduced with an increase of axial compression ratio (i.e. 0.55 and 0.65). It was also found from the comparison between columns with the same concrete strength of 54.8 MPa but subjected to different axial compression ratios (i.e. C2S5A45, C3S5A55 and C3S5A65), the peak load and energy dissipation capacity generally increased with an increase of axial compression ratio. However, for the confined specimens of 71.2 MPa compressive strength but different axial compression ratios, a marginal difference in hysteretic behavior was noticed. It is evident that the influence of axial compression ratio became less obvious with an increase of concrete strength. The pre-damaged loading history for CFRP-confined columns is shown in Fig. 6(i). The confined columns were loaded approximately to the peak load of the corresponding unconfined control specimen. Then, the specimens were unloaded to their initial position and retested under the same cyclic loading scheme as the other specimens. The hysteretic behaviors of the pre-damaged CFRP-confined columns are shown in Fig. 6(j–l). It can be seen that the hysteretic behaviors of pre-damaged confined specimens were still much better than the control specimens, but less than the corresponding undamaged CFRP-confined specimens. The initial stiffness, peak lateral strength, and energy dissipation capacity of the columns was generally reduced due to the preload damage. This was especially the case for specimen R3S7A65D which had a high concrete strength and was subjected to a high axial compression ratio. This is consistent with the previous test observations that the specimen R3S7. 3.3. Envelope curves and performance indexes The envelope curves of all the specimens obtained from the experimental hysteretic curves are presented in Fig. 7. It can be observed that the CFRP-confined columns generally had higher peak lateral load, considerably better ductility and much more gradual post-peak strength degradation, compared to the corresponding control specimens (i.e. R2S5A45 to C0S5A45 and R2S7A45 to C0S7A45). Upon comparison between the CFRPconfined columns under different axial compression ratios, it was found that the peak lateral load slightly increased but the ductility slightly decreased with an increase of axial compression, as shown Fig. 7(a–b). It was also noted from the comparison between predamaged and undamaged CFRP-confined columns that the initial stiffness and peak load were slightly reduced due to the predamaged load (refer Fig. 7(c–e)). To investigate and evaluate quantitatively the improvement in the seismic performance contributed by the CFRP wraps, the performance indexes of the average displacement and the corresponding load in push and pull direction at the yield (i.e. Dy and Py), peak (i.e. Dc and Pc), and ultimate (i.e. Du and Pu) points were obtained and they are summarized in Table 3. The performance indexes were calculated from the average envelope curves. For uniformity, the yield point of each specimen was determined by the energy method proposed by Mahin and Bertero [50,51] and illustrated in Fig. 8. The ultimate point refers to the position where the load was reduced to 85% of the peak load.

D. Wang et al. / Construction and Building Materials 134 (2017) 91–103

97

Fig. 6. Experimental lateral load-displacement curves.

It can be seen from Table 3 that both the yield displacement (Dy) and load (Py) of the column were slightly increased after confinement with CFRP wraps at the plastic hinge regions, especially for the specimens with a high concrete strength. For example, the yield displacement and corresponding lateral load of specimen R2S5A45 were enhanced by 17.7% and 7.6%, respectively, com-

pared with the corresponding control specimen C0S5A45. The yield displacement and corresponding lateral load applied to specimen R2S7A45 was 38.9% and 20.6%, respectively, higher than that of specimen C0S7A45. It is also found that the enhancement in both the yield displacement and load considerably increased with an increase of axial compression ratio. For example, from the compar-

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Fig. 7. Envelope curves.

Table 3 Summary of test results. Specimen

Dy & Incr. (mm) (%)

dy (%)

Py & Incr. (kN) (%)

Dc & Incr. (mm) (%)

dc (%)

Pc & Incr. (kN) (%)

Du & Incr. (mm) (%)

du (%)

l & Incr. Du/Dy

(%)

C0S5A45 R2S5A45 R3S5A55 R3S5A65 C0S7A45 R2S7A45 R3S7A55 R3S7A65 R3S5A55D R3S5A65D R3S7A65D

6.2 7.3 8.7 12.8 5.4 7.5 8.1 8.8 15.9 11.0 11.4

0.49 0.58 0.69 1.02 0.43 0.60 0.64 0.70 1.26 0.87 0.90

50.1 53.9 66.6 69.9 57.8 69.7 77.9 83.8 67.0 69.0 84.0

13.9 21.2 25.0 40.0 11.1 18.3 25.0 25.0 30.7 18.1 22.9

1.10 1.68 1.98 3.17 0.88 1.45 1.98 1.98 2.44 1.44 1.82

61.3 72.0 86.3 87.2 71.8 88.3 100.4 106.9 78.2 80.8 97.7

25.1 74.7 71.9 77.5 17.4 85.4 79.5 74.0 74.6 68.5 58.0

1.99 5.93 5.71 6.15 1.38 6.78 6.31 5.87 5.92 5.44 4.60

4.05 10.23 8.26 6.05 3.22 11.39 9.81 8.41 4.69 6.23 5.09

– 152.6 104.0 49.4 – 253.7 204.7 161.2 15.8 53.8 58.1

– 17.7 40.3 106.5 – 38.9 50.0 63.0 156.5 77.4 111.1

– 7.6 32.9 39.5 – 20.6 34.8 45.0 33.7 39.5 45.3

– 52.5 79.9 187.8 – 64.9 125.2 125.2 120.9 30.2 106.3

– 17.5 40.8 42.3 – 23.0 39.8 48.9 27.6 31.8 36.1

– 197.6 186.5 208.8 – 390.8 356.9 325.3 197.2 172.9 233.3

Note: Dy = yield displacement; dy = yield drift ratio; Py = yield load corresponding to Dy; Dc = peak displacement; dc = peak drift ratio; Pc = peak load corresponding to Dc; Dy and du = ultimate displacement and drift ratio; l = Du/Dy ductility factor.

Fig. 8. Schematic diagram of energy method.

ison between the specimen of R3S5A65 and R2S5A45, the yield displacement and load increased 75.3% and 29.7%, respectively, with an increase in the axial compression from 0.45 to 0.65. The enhancement was up to 106.5% and 39.5% when compared with the control specimen under an axial compression ratio of 0.45 (i.e. C0S5A45). Table 3 reveals that the presence of the CFRP wraps resulted in significant enhancement in peak displacement (Dc) but slight improvement in peak load (Pc). In addition, both the peak displacement and load increased with an increase in the axial compression ratio. Compared with the control specimens, the peak load of the confined columns under the same axial compression ratio improved about 20%, while the corresponding displacement varied from about 50% to 65%. As the axial compression ratio increased from 0.45 to 0.65, the enhancement in the peak load increased from about 40% to 50%, while the improvement in peak displacement was from 80% to 190%.

D. Wang et al. / Construction and Building Materials 134 (2017) 91–103

It is also evident from Table 3 that the most significant improvement due to the wrapped CFRP was the ultimate lateral displacement (Du), especially for specimens made from high concrete strength. The improvement in Du after FRP confinement increased by about 200% and 400% for specimens made from 54.8 MPa and 71.2 MPa concrete strength, respectively. It was also noted from the comparison between FRP-confined specimens with different axial compression ratio and confined height in the plastic hinge region that the axial compression ratio and wrapped height of CFRP had a slightly influence on the ultimate displacement capacity, even for specimens under a high axial compression ratio of 0.65. This suggests that the ductility of HSC columns can be effectively improved if the plastic hinge region of the columns is well confined by CFRP wraps with sufficient length (i.e. larger than 1.1D in the present study) even under high axial compression ratio. Moreover, it was observed that FRP-confined columns with relative higher concrete compressive strength developed larger lateral drift ratio than corresponding columns with lower concrete strength under the same axial compression ratio. Similar results were also observed by Ozbakkaloglu et al. [30–32] for the seismic performance of HSC filled FRP tube columns. The displacement ductility factor (l) directly reflects the plastic deformation capacity of structures or members. A larger ductility factor means larger plastic deformation capacity and energy dissipation capacity. The results provided in Table 3 show that the ductility factors of the CFRP-confined columns with the concrete strength of 54.8 MPa and 71.2 MPa were significantly increased by up to about 150% and 250% compared to the control specimens, respectively. In addition, the ductility factors of confined specimens decreased with an increase of axial compression ratio. For example, the improvement in ductility factor was 152.6%, 104.0% and 49.4% for specimens R2S5A45, R3S5A55 and R3S5A65, respectively, compared with C0S5A45. The improvement was 253.7%, 204.7% and 161.2%, respectively, for specimens of R2S7A45, R3S7A55, and R3S7A65 as compared with C0S7A45. This demonstrates that although the CFRP-confined HSC columns have excellent ultimate displacement capacity under a high axial compression ratio, the plastic deformation capacity of the columns reduces with an increase of axial compression ratio. Nonetheless, the plastic deformation capacity of CFRP-confined specimens under high axial compression ratio is still superior to unconfined columns. For the three pre-damaged CFRP-confined columns (i.e. R3S5A55D, R3S5A65D and R3S7A65D), the yield displacement was much larger than the corresponding undamaged CFRPconfined columns (i.e. R3S5A55, R3S5A65 and R3S7A65) due to the reduction of initial stiffness. The yield load was, however, approximately the same. In addition, the pre-damaged load history had no obvious influence on the peak displacement, but it slightly reduced the peak load and ultimate displacement capacity of the columns. The ductility factor of the columns was also found

(a) Effective stiffness

99

reduced after pre-damage, especially for the specimen with the relative lower concrete strength and axial compression ratio (i.e. R3S5A55D). This demonstrates that the plastic deformation of CFRP-confined columns reduces after pre-damaged. 3.4. Stiffness degradation One of the most important parameters to evaluate the hysteretic behavior of RC members and structures is the stiffness degradation ratio. In the present study, both the effective stiffness (Ke) and the unloading stiffness (Ku) of the specimens were obtained and investigated based on the experimental hysteretic curves. The effective stiffness Ke and the unloading stiffness Ku at each displacement level are defined as the secant stiffness and the tangent stiffness at the unloading point, respectively [17,52]. The experimental values of Ke and Ku at different displacement levels for all specimens are plotted in Fig. 9. Both Ke and Ku were calculated in the positive and negative direction of each hysteretic loop considering the slight asymmetry of the experimental lateral load-displacement loops. In addition, to eliminate the influence of concrete strength and for ease of comparison, both Ke and Ku were normalized with respect to the theoretical lateral stiffness (KD) of the corresponding control column, which is defined as follows:

K D ¼ 12EI=H3

ð2Þ

where E is the elastic modulus of concrete; I is the moment of inertia of column cross-section; H is the corresponding storey height which is equal to the clear cantilever height of the specimens. It is evident from Fig. 9(a) that the effective stiffness of CFRPconfined columns decreased with an increase of the lateral drift ratio in the initial stage, while the rate of degradation become much more gradual when the lateral drift ratio increased to about 1.5–2.0%. Table 3 shows that the drift ratio of 1.5–2.0% is approximately equal to the drift ratio at the peak lateral load of these columns. This is similar to the observation reported by Ozbakkaloglu et al. [30–32]. For the control specimens, the degradation of effective stiffness is approximately the same as CFRP-confined columns before the drift ratio reached 1.0%. This is about equal to the peak displacement level of the control specimens. After the peak load point, the control specimens generally exhibited a more serious degradation in the effective stiffness than the CFRP-confined columns. This is consistent with the previous observation that control columns exhibited rapid post-peak strength degradation behavior. It is also evident from Fig. 9(a) that there is little difference in the normalized effective stiffness degradation for different specimens, which demonstrates that the axial compression ratio, concrete strength, pre-damaged load history, and height of CFRP-confined area at column ends have little effect on the effective stiffness degradation of CFRP-confined columns. Based on the experimental results of all the CFRP-confined specimens presented in Fig. 9, the

(a) Unloading stiffness

Fig. 9. Degradation of stiffness.

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D. Wang et al. / Construction and Building Materials 134 (2017) 91–103

degradation of effective stiffness was found to reduce exponentially. The best-fit exponential expression is:

K e ¼ K D ð1:45e1:08d þ 0:10Þ

ð3Þ

where d (%) is the lateral drift ratio of the columns. It should be noted that there was effective stiffness degradation before yielding of the columns based on the test results, although the initial stiffness of the columns is generally assumed equal to the yield stiffness (Ky) in the lateral load-displacement hysteretic model. In this case, there is no degradation of effective stiffness and it is equal to the yield stiffness before the yield point of the columns is reached. Consequently, the proposed empirical equation provided in Eq. (3) can be used to predict the post-yield effective stiffness degradation when developing a lateral load-displacement hysteretic model of CFRP-confined HSC columns. Fig. 9(b) shows that the unloading stiffness exhibits similar degradation behavior as the effective stiffness. The unloading stiffness of all the columns reduced rapidly before a drift ratio of around 1.5%. After that, the degradation rate became increasingly gradual with an increase of the lateral drift ratio. It is also evident that the variation of axial compression ratio, concrete strength, pre-damaged load history and wrapped CFRP hardly influences the normalized unloading stiffness degradation of all the CFRPconfined HSC columns. It was found that there was little difference between the unloading stiffness of the two repeated loops at the same displacement level, suggesting that the effect of repeated loading cycles on the unloading stiffness was minor. Comparison between the control and CFRP-confined specimens shows that the unloading stiffness degradation of the control columns was similar with that of CFRP-confined columns under the same displacement level, except the control specimens failed at a small drift ratio. Based on the experimental results of all the CFRP-confined specimens presented in Fig. 9(b), it was found that the degradation of unloading stiffness can be best represented by a power equation as follows:

K u ¼ 0:659dð0:326Þ K D

ð4Þ

It should be noted that the value of the unloading stiffness based on Eq. (4) will be infinite if the drift ratio is equal to 0. This is, however, not true in reality. It is can be found from Fig. 9(b) that the largest experimental value of unloading stiffness under very small drift ratio is about 1.5 times the theoretical lateral stiffness (KD) of the corresponding control columns. Consequently, the upper limit of Eq. (4) is suggested to be equal to 1.5 KD. It should also be noted that the best-fit trendline expression was obtained on limited studies and further studies need to be conducted to verify. 3.5. Energy dissipation The energy dissipation capacity is an important index to evaluate the seismic performance of RC members and structures. In earthquake events, RC structures with insufficient energy dissipation capacity are vulnerable and even collapse due to cumulative energy under small displacements rather than a single large deformation. In the present study, the total cumulative energy dissipated by all the preceding hysteretic loops (Esum) against the lateral drift ratio was plotted in Fig. 10. It is obvious that the CFRP-confined columns had considerably larger Esum than the control specimens, especially for the high concrete strength columns. For example, the final Esum of specimen R2S5A45 was about 6 times higher than that of the corresponding control specimen C0S5A45, while for the columns with a higher concrete strength the final Esum of the confined column R2S7A45 was about up to 45 times higher than that of control specimen

C0S7A45. In addition, the cumulative energy dissipation capacity of confined columns of concrete strength 54.8 MPa was found to slightly increase as the axial compression ratio increased from 0.45 to 0.55. The energy dissipation was, however, approximately the same for confined columns with an axial compression ratio of 0.55 and 0.65. Moreover, the variation of axial compression ratio had little influence on the energy dissipation capacity of confined columns with a high concrete strength of 71.2 MPa. It is also evident that the dissipated cumulative energy of the pre-damaged confined columns was slightly lower than that of the corresponding confined columns under the same drift ratio, especially for the high strength columns. For example, the final Esum of predamaged specimen R3S7A65D reduced to about half compared with the corresponding specimen R3S7A65. 3.6. Equivalent viscous damping ratio The equivalent viscous damping (EVD) ratio (neq) is another generally utilized index to represent the hysteretic energy dissipation capacity of RC members. It is also a crucial factor in the displacement-based design (DBD) method. The EVD ratio can be calculated from the lateral load-displacement hysteretic curves of RC members [17,53]. In the present study, the definition proposed by Clough and Penzien [52] is adopted to calculate the EVD ratio as follows:

neq ¼

1 ED 4p E S

ð5Þ

where ED is the area of one complete idealized load-displacement hysteretic loop representing the dissipated energy; Es = VmDm is the recoverable elastic strain energy stored in an equivalent linear elastic system under static conditions with the same effective stiffness; Vm and Dm are the average peak force and displacement at the peak force of the hysteretic loop. Fig. 11 shows the EVD ratio versus the lateral drift ratio for all the specimens. It can be generally observed that the EVD ratio increased with an increase of lateral drift ratio. In addition, the EVD ratio of the control and pre-damaged specimens was found to be slightly larger than the corresponding confined specimens under the same drift ratio. This is due to the fact the average peak force (Vm) of each hysteretic loop of the control and damaged columns is generally smaller than the corresponding confined columns under the same displacement level. The smaller Vm will result in the smaller recoverable elastic strain energy (Es) and further result in the larger EVD ratio. However, the final EVD ratio of confined columns is obviously larger than that of the control and damaged columns. As higher EVD ratios generally result in better energy dissipation capacity, this result is consistent with the observation of cumulative energy dissipation. Furthermore, it was found that the variation of axial compression ratio, concrete strength, and height of CFRP-confined areas also had limited influence on the EVD ratio of confined columns. 3.7. Strain responses To evaluate the confinement efficiency of CFRP wraps and steel hoops, the measured largest lateral strains on both the CFRP wraps and the internal hoop steel reinforcement at each loading cycle were plotted in Fig. 12. In this figure, the lateral tension strain was defined as positive. It is evident from Fig. 12 that the lateral strains on CFRP wraps and steel hoops were both increased with an increase of the lateral displacement levels (i.e. the increase of loading cycles) and also generally increased with an increase of axial compression ratio. The measured largest lateral strain on steel hoops for all the specimens is generally larger than the yield

D. Wang et al. / Construction and Building Materials 134 (2017) 91–103

101

Fig. 10. Cumulative energy dissipation.

Fig. 11. Equivalent viscous damping ratio.

Fig. 12. Strain responses.

strain of the steel (i.e. 0.2%). It is noted that the hoops have yielded for both control and confined columns at the last stage of the tests. Fig. 12 also shows that the maximum lateral strains on the CFRP wraps varied from 0.2% to about 1.0%, which was about 0.11 to 0.56 times of rupture strain of flat CFRP coupons. It is thus suggested that the applied CFRP wraps had surplus deformation capacity to sustain the dilation of the confined plastic hinge region of the

HSC columns. This was also consistent with the experimental observations that no rupture failure of CFRP occurred during the tests. Although the final lateral strain on CFRP wraps was much smaller than the rupture strain, it was still generally larger than the maximum lateral strain value on the steel hoops. The confinement of CFRP wraps is still more effective than the steel hoops. Furthermore, the lateral strain on the CFRP wraps of confined columns

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with a higher concrete strength (i.e. 71.2 MPa) was generally larger than that confined columns with a lower concrete strength (i.e. 54.8 MPa) under similar displacement levels. It is demonstrated that the CFRP wraps provided larger and more effective confinement when applied to high concrete strength columns. This is consistent with the previous experimental observations that the confined high strength columns exhibited larger improvement in ductility and energy dissipation capacity when compared with the corresponding control specimens. 4. Conclusions This paper has presented an experimental investigation on the seismic performance of CFRP-confined circular HSC columns. The main test variables included the axial compression ratio, the concrete strength, the height of the CFRP-confinement, and the predamaged load. Based on interpretations and discussions of the test results, the following conclusions can be drawn: 1. Externally bonded CFRP wraps applied in the potential plastic hinge regions only can effectively and significantly improve the seismic performance of circular HSC columns even under a high axial compression ratio. Compared to unconfined control columns, the ultimate lateral displacement, ductility factor and energy dissipation capacity of corresponding retrofitted columns were noticeably improved, especially for columns with relative higher concrete strength. 2. The peak lateral load and corresponding displacement increased with an increase of the axial compression ratio. The ultimate lateral deformation capacity only slightly decreased though with an increase of axial compression ratio. FRPconfined HSC columns still exhibited excellent deformation capacity even under a high axial compression ratio of 0.65. However, the plastic deformation capacity (i.e. ductility factor) reduced with an increase in the axial compression ratio. 3. The efficiency of CFRP wraps in improving the ductility and energy dissipation capacity of HSC columns increased with an increase in concrete strength. Compared to unconfined control columns, the ultimate displacement capacity of the corresponding confined columns with 54.8 MPa and 71.2 MPa concrete strength increased 2.0 and 4.0 times, respectively. In addition, the total cumulative energy dissipation capacity respectively increased by about 6.0 and 45.0 times. Confined columns made from high concrete strength exhibited a larger enhancement in ductility and energy dissipation capacity. 4. The seismic performance of the HSC columns improved when the height of the confined area at the plastic hinge region was larger than 1.1 times the cross-section diameter. Pre-damaged confined columns, however, need larger confined length in the plastic hinge region to ensure sufficient residual seismic capacity and avoid potential serious damage and failure in the unconfined regions under high axial compression ratio. The lateral load capacity, ductility and energy dissipation capacity of predamaged confined columns were slightly lower than that of the corresponding confined columns. 5. The axial compression ratio, concrete strength, pre-damaged load history and height of CFRP-confined area at plastic hinge region had limited influence on the degradation of effective and unloading stiffness and the equivalent viscous damping ratio of CFRP-confined HSC columns. Acknowledgements This research was supported by the National Natural Science Foundation of China (Grant No. 51408153, No. 51478143, and

No. 51278150), the National Key Basic Research Program of China (973 Program, Grant No. 2012CB026200) and the China Postdoctoral Science Foundation (Grant No. 2014M551252 and No. 2015T80354).

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