Full-scale seismic testing of concrete building columns reinforced with both steel and CFRP bars

Accepted Manuscript Full-scale seismic testing of concrete building columns reinforced with both steel and CFRP bars Zhong-Kui Cai, Daiyu Wang, Zhenyu Wang PII: DOI: Reference:

S0263-8223(17)30514-7 http://dx.doi.org/10.1016/j.compstruct.2017.06.020 COST 8607

To appear in:

Composite Structures

Received Date: Revised Date: Accepted Date:

15 February 2017 24 April 2017 7 June 2017

Please cite this article as: Cai, Z-K., Wang, D., Wang, Z., Full-scale seismic testing of concrete building columns reinforced with both steel and CFRP bars, Composite Structures (2017), doi: http://dx.doi.org/10.1016/j.compstruct. 2017.06.020

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Full-scale seismic testing of concrete building columns reinforced with both steel and CFRP bars

Zhong-Kui Cai1,2,3, Daiyu Wang1,2,3, Zhenyu Wang1,2,3,*

1.Key Lab of Structures Dynamic Behavior and Control of the Ministry of Education, Harbin Institute of Technology, Harbin 150090, China. 2.Key Lab of Smart Prevention and Mitigation of Civil Engineering Disasters of the Ministry of Industry and Information Technology, Harbin Institute of Technology, Harbin 150090, China. 3.School of Civil Engineering, Harbin Institute of Technology, Harbin 150090, China.

* Corresponding Author. Tel.: +86 451 8628 3856; Fax: +86 451 8628 3856. E-mail

address:

[email protected]

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(Zhen-yu

WANG).

Abstract: This paper presents an experimental study of an innovative concrete building column which is longitudinally reinforced with both steel bars and fiber-reinforced polymer (FRP) composite bars. The main objectives of this study include: (1) investigating the seismic failure mechanism of such FRP-steel reinforced concrete (FSRC) columns under relatively large gravity load and (2) analyzing the effects of the FRP bars on the seismic performance of the FSRC columns. To this end, a total of six full-scale cantilever columns with variable carbon FRP (CFRP) bar and steel reinforcement ratios were tested under combined constant axial load and lateral displacement reversals. The specimens had 400mm square sections and effective heights of 1800mm. Test results showed that adding additional CFRP bars into the conventional steel reinforced concrete (SRC) columns was efficient in improving the postyield stiffness ratios and mitigating the residual displacements, while the hysteretic energy dissipation could be maintained. Failure modes of the tested FSRC specimens were characterized by crushing of FRP bars and buckling of steel bars at drift ratios larger than 2.4%. Furthermore, confining the FSRC column with external CFRP wraps was effective in delaying the crushing of internal CFRP bars and reducing the post-earthquake residual displacement.

Keywords: building columns; FRP bars; seismic performance; failure mechanism; post-yield stiffness; residual displacement.

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1. Introduction The devastating social and economic impacts of earthquakes in urban areas have resulted in an increasing awareness of the importance of post-earthquake serviceability for buildings [1]. Previous studies have revealed that the building residual displacement, namely the residual roof drift and maximum residual inter-story drift, was a crucial index of the functionality and repairability of the damaged buildings [2]. For example, in both 1985 Mexico City earthquake and 1995 Hyogo-ken Nanbu earthquake, some reinforced concrete (RC) buildings had to be demolished due to overlarge residual drifts even though they did not collapse [3]. Over the last two decades or so, there has been a wealth of investigations on how to quantify and mitigate the residual displacement of RC structures. Earlier research conducted by MacRae and Kawashima [4,5] revealed that the residual drifts of the RC structure decrease significantly with the increase of the post-yield stiffness ratio of RC columns. This conclusion was later supported by the study of Borzi et al. [6]. After that, several investigations have been conducted with the aim of improving the post-yield stiffness of RC columns so as to reduce the residual drifts. Sakai and Mahin [7] proposed to replace part of the longitudinal mild steel bars with un-bonded pre-stressing strands. The seismic performance of such a column system was investigated by finite element (FE) analysis based on fiber model approach. Jeong et al. [8] further studied the self-centering capability of partially pre-stressing RC columns with un-bonding longitudinal bars or external steel jackets. Both shake table tests and FE analysis were performed. Saiidi et al. [9] conducted quasi-static tests on a new self-centering column, in which super-elastic shape memory alloy and engineered cementitious composites were incorporated in the plastic hinge zone. Results of the above three studies [7-9] showed that all three approaches were effective in improving the post-yield stiffness and the self-centering capacities, as compared to the conventional steel reinforced concrete (SRC) columns. However, it was also noted by Sakai and Mahin [7] and Jeong et al. [8] that due to the decrease in hysteretic energy dissipation ability, the maximum displacement response of the proposed self-centering columns under earthquake was larger than those of the SRC columns. Furthermore, the self-centering columns in the test of Saiidi et al. [9] also exhibited much narrower hysteretic loops than the SRC reference specimen. Since poor hysteretic energy dissipation may lead to overlarge seismic demands for self-centering structures [10], therefore, it is important to mitigate the post-earthquake residual displacement of the column without compromising its energy dissipation ability. 3/32

More recently, fiber reinforced polymer (FRP) composite materials were employed in concrete columns to improve the post-yield stiffness ratios and the self-centering capabilities. The FRP has many advantages against steel material, including small density, high strength, corrosion resistance, non-magnetic, and in particular, the elastic stress-strain relationship. In order to improve the post-yield stiffness of concrete columns from the material level, a novel composite bar made with FRP skin over a steel core was proposed by Wu et al. [11]. The seismic performance of columns longitudinally reinforced with such composite bars was also investigated by Fahmy et al. [12] and Sun et al. [13,14]. Research results showed that these new columns had a designable post-yield stiffness and exhibited reduced residual displacement. However, the production process of such composite bar is complicated, as pointed out by Ibrahim et al. [15]. A method of improving the post-yield stiffness of the bridge column from the section level was further proposed in reference [15], i.e., reinforcing the column simultaneously with both FRP bars and conventional steel bars. Such FRP-steel reinforced concrete (FSRC) bridge columns were tested under combined lateral displacement reversals and relatively small axial load corresponding to an axial compression ratio of 0.033. The cross sections of the cantilever specimens were 200mm×200mm and the effective heights were 850mm. Basalt FRP (BFRP) bars were employed in the specimens. Subsequently, a 3D finite element for the FSRC columns was also developed by Ibrahim et al. [16] with great emphasis on the bond condition between FRP bars and the surrounding concrete. Numerical and experimental results indicated that the FSRC columns exhibited evident post-yield stiffness and improved self-centering capabilities. Moreover, the FSRC bridge columns had hysteretic energy dissipation ability that was comparable to or surpass that of the SRC counterparts. So far, to the best of the authors’ knowledge, studies on FSRC columns have been quite limited and only focused on bridge columns. This paper contributes to extending the FSRC concept to building columns so as to improve the post-earthquake serviceability of concrete buildings. One distinct difference between a bridge column and a building one was that the axial compression ratio of the later was much larger than that of the former. Previous investigations have demonstrated that for FRP bars, the strength and strain capacity in compression were relatively low compared to those in tension [17,18]. Moreover, the compressive behavior of FRP bars is far less well-studied than the tensile response [19-21]. Therefore, the effectiveness of the FRP bars in FSRC building columns remains unknown. The main objectives of this study include: (1) experimentally investigate and explain the seismic failure mechanism of FSRC columns under relatively large gravity load and (2) 4/32

analyze the effects of the FRP bars on the seismic performance of the FSRC building columns, in terms of post-yield stiffness, residual displacement, deformation capacity and energy dissipation ability. To this end, four FSRC columns and two SRC reference columns were tested under combined constant axial load (corresponding to an axial compression ratio of 0.3) and lateral displacement reversals. Carbon FRP (CFRP) bars were incorporated in this test. In addition, testing of the compressive strength of CFRP bars was also conducted and presented in this paper, considering that studies on FRP bar mechanical properties under compression are limited and that the standard test method has not been established as reported in ACI 440.3R code [22]. 2. Mechanical properties of CFRP bars with embedded sensors 2.1 Reasons for using CFRP bars with embedded optical fiber sensors Commercially produced CFRP bars with embedded optical fiber Bragg grating (OFBG) sensors [23] were used in this study, as shown in Fig. 1(a). During the fabrication of CFRP bars, the optical fiber is placed at the center of the carbon fiber strand. Multiple Bragg grating sensors can be constructed in this optical fiber at the points where strain responses are to be monitored. Such CFRP bars were employed in this experimental investigation for the following reasons. (i) When measuring the strain distribution of FRP bars, the bond between bars and concrete might be adversely affected by closely attached strain gauges [24], while the embedded OFBG sensors will not bring disturbance to the bond condition. (ii) The diameter of the optical fiber is so small that it will not influence the mechanical properties of the CFRP bar [25]. (iii) As the OFBG sensors are encapsulated within FRP bars, they thus have better corrosion resistance than conventional strain gauges. In addition, response signals of the CFRP bars are transmitted through the optical fiber and a jumper wire to the data acquisition terminal while additional lead wires are not needed. 2.2 Mechanical properties of CFRP bars in tension The nominal diameter of the CFRP bars used in this study is 10mm. The surface texture of the CFRP bar is smooth due to technological limitations on the fabrication of CFRP bars with OFBGs. Tensile properties of the CFRP bars were tested in accordance with ASTM D7205 code [26]. A total of five CFRP bar samples were tested using an MTS servocontrolled hydraulic system, as shown in Fig 1 (b). Tensile tests were conducted at a constant strain rate of 0.003 min-1. The strain rate was selected so as to produce failure of the sample within 1 to 10 minutes. Fig. 1(c) shows the CFRP bar sample after failure. As indicated in 5/32

Table 1, the average tensile strength and modulus of the CFRP bar were 1592.4 MPa and 134GPa, respectively, resulting in an ultimate tensile strain of 1.19%.

(a) Before test

(b) MTS test setup

(c) After failure

Fig. 1. CFRP bar tensile tests Table 1. Summary of material properties Material

d or t (mm)

f y (MPa)

f u (MPa)

E (GPa)

 y or  fu

CFRP bar

d = 10

—

1592.4

134

 fu  0.0119

Carbon fiber sheet

t = 0.168

—

4340.0

241

 fu  0.0180

Steel hoop

d=8

354.9

550.3

200

 y  0.0018

Longitudinal steel bar

d = 14

493.2

679.5

200

 y  0.0025

Longitudinal steel bar

d = 18

441.2

612.2

200

 y  0.0022

Notes: d = diameter of bars; t = thickness of single ply carbon fiber sheet; fy = yield stress; fu = tensile strength; E = elastic modulus in tension;  y = yield strain of steel bars; and  fu = ultimate tensile strain of FRP materials.

2.3 Compressive strength of CFRP bars So far, few studies have been conducted to evaluate FRP bar mechanical properties under compression. Although test methods for a wide range of material properties of FRP bars have been developed by both the ACI Committee 440 [22] and the ASTM International, the compressive properties are not covered. A testing approach for FRP bar compressive strength was specified in Chinese code GB/T1448 [27] by referring ASTM D695 code for rigid plastics [28]. However, as reported by Deitz et al. [17], ASTM D695 code is not

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precisely applicable for FRP bars. In this study, three test methods were used to measure the compressive strength of CFRP bars and the test results were compared. (i) Firstly, the test method specified in Chinese code GB/T1448 [27] was adopted as method I. As shown in Fig. 2(a), the length of tested CFRP bar sample was 20mm, resulting in a length to diameter ratio of 2. Five samples were loaded directly using a hydraulic compression machine. Failure mode was characterized by longitudinal splitting and premature crushing at either end of the bar sample. The average compressive strength of tested CFRP bar samples was 400.4MPa. (ii) The second test method was modified from method I. In order to prevent the premature crushing at two ends as observed in tests using method I, both ends of the CFRP bar sample were confined by steel rings, as shown in Fig. 2(b). The unbraced length between two rings was 20mm, resulting in a length to diameter ratio of 2, which was identical to that of method I. Split failure was observed in this test method. The average compressive strength of five tested samples was 570.4MPa. (iii) In the third test method, the confinement effects of surrounding concrete on FRP bar compressive behavior were taken into consideration. As shown in Fig. 2(c), four CFRP bars with lengths of 300mm were longitudinally cast into a CFRP-wrapped concrete cylinder with dimensions of 150×300mm. These four CFRP bars spaced uniformly around the perimeter of a circle with a diameter of 100mm, as illustrated in Fig. 2(c). The CFRP wrap, which was applied in a wet lay-up manner, consisted of three layers of uni-directional carbon fiber sheets and the thickness of single ply sheet was 0.168mm. Mechanical properties of the carbon fiber sheet were tested in accordance with ASTM D3039 [29] and the results were shown in Table 1. Five such cylinders were prepared, together with another five ones with the same CFRP wraps but no longitudinal CFRP bars. These ten cylinders were axially loaded per ASTM C39 code [30]. For cylinders with CFRP bars, failure initiated by the rupture of external CFRP wraps and followed by the crushing of CFRP bars. The average axial load capacities of cylinders with and without longitudinal CFRP bars were 1960.9kN and 1709.6kN, respectively. Therefore, the contribution of each CFRP bar in resisting compression can be expressed as

1960.9  1709.6

4  62.8kN . Thus, it can be computed that the average compressive

strength of tested CFRP bars by this method was 800.4MPa. The following is a discussion of the above test results. Comparison of the results of the

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method I and II showed that confining the ends of the CFRP bar samples increased the measured compressive strength by 42%. Furthermore, the tested strength using method III was approximately twice of that for method I. This improvement was contributed from the effective confinement provided by surrounding concrete. Therefore, it can be concluded that the tested compressive strength of an FRP bar was sensitive to the confinement of the bar sample. This result may be attributed to the fact that the confinement is effective in delaying the lateral dilation of the FRP bar and the outward buckling of the fibers. Therefore, it is suggested that when investigating the compressive properties of FRP bars, the confinement effects of surrounding concrete should be considered.

(a) Test method I

(b) Test method II

(c) Test method III

Fig. 2. CFRP bar compressive tests 3. Seismic experimental program 3.1 Specimen design In the quasi-static tests, a total of six full-scale cantilever column specimens with identical geometric dimensions were designed and prepared. As shown in Fig. 3(a), each specimen was composed by a footing of 1500×700×700mm, a column with a square section of 400×400mm and a loading stub of 700×400×400mm. The effective height of the specimen (from the top of the footing to the center of the loading stub) was 1800mm, leading to a shear span ratio of 1800/400=4.5. This 1800-mm-high cantilever column represented the lower half 8/32

of a first-story frame column with a story height of 3.6m (between the foundation and the point of inflection). The specimen was fixed to the laboratory floor by two 100mm threaded steel rods which passing through the ducts inside the footing.

(a) Geometric dimensions

(b) Reinforcement details.

(c) CFRP bar anchors

Fig. 3. Specimen design Longitudinal reinforcement of the six specimens was illustrated in Fig. 3(b) and further summarized in Table 2. Designations of tested specimens are S1F0, S1F1, S1F2, S2F0, S2F2 and S2F2-C. Here, “S1” and “S2” represent that the specimen was longitudinally reinforced with twelve deformed steel bars with diameters of 14mm and 18mm , respectively; “F0”, “F1” and “F2” mean that zero, eight and twelve 10mm CFRP bars were incorporated in the specimen, respectively. In addition, the above compressive test results have revealed that the confinement has a significant impact on FRP bar compressive behavior. Accordingly, the sixth specimen S2F2-C was designed to study the influence of confinement on the utilization

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efficiency of FRP bars for an FSRC column. The longitudinal reinforcement of specimen S2F2-C was identical to that of specimen S2F2 but the former was confined by additional CFRP wraps. The steel reinforcement ratio  s , FRP reinforcement ratio  FRP and the ratio

 FRP  s for each specimen are shown in Table 2 as well. Here, ρs and ρFRP are defined as the ratios of steel bar area As and FRP bar area AFRP to the column cross-sectional area Ag, respectively, i.e., s  As Ag 100% and FRP  AFRP Ag 100% . Typically, the steel reinforcement ratio  s of a building column ranges from 1.0% to 2.0%. Therefore, the designed  s (1.15% and 1.91%) for the specimens of this study are representative. Table 2. Reinforcement details of tested specimens Specimen

Steel bars

FRP bars

 s (%)

 FRP (%)

 FRP s (%)

Confinement

S1F0

12 14mm

0

1.15

0

0

hoops

S1F1

12 14mm

8 10mm

1.15

0.39

34%

hoops

S1F2

12 14mm

12 10mm

1.15

0.59

51%

hoops

S2F0

12 18mm

0

1.91

0

0

hoops

S2F2

12 18mm

12 10mm

1.91

0.59

31%

hoops

S2F2-C

12 18mm

12 10mm

1.91

0.59

31%

hoops + CFRP wraps

Notes:  s  steel reinforcement ratio;  FRP  CFRP reinforcement ratio.

The design of the anchorage of longitudinal CFRP bars was presented as follows. As specified in Chinese code GB50608 [31], the development length of the CFRP bar la can be calculated from Eqs.(1) to (3):

f d _ FRP   f  fu _ FRP la 

f d _ FRP 8 ft

d FRP

f t  0.62 f c

(1) (2)

(3)

where, f d _ FRP and f u _ FRP are the design and ultimate strengths of FRP bars in tension, respectively;  f  1 1.4  70% is the strength reduction factor of FRP bars; d FRP is the nominal diameter of the FRP bar; f t and f c are the tensile and compressive strength of unconfined concrete, respectively. In this study, the designed f c was 25MPa. Accordingly, 10/32

la =461mm and la is smaller than the height of the specimen footing, i.e., 700mm. Therefore, the development length of the CFRP bar was adequate based on Chinese code GB50608. To further prevent the premature anchorage failure of CFRP bars, steel anchors were fabricated and applied at bar ends. As shown in in Fig. 3(c), the anchor consisted of a steel tube with inner screw threads and six shear studs. The length, outer and inner diameters of the anchor was 300mm, 28mm and 18mm, respectively. The tube of the anchor was filled with expansive cement grout, as recommended by ASTM D7205 [26]. All the six specimens were transversely confined by 8mm steel hoops with a space of 100mm along the columns. The amount of transverse steel reinforcement was determined per Chinese seismic design code for buildings [32]. The concrete cover thickness was 25mm and it was measured to the outside of the hoop. Thus, hoop volumetric ratios  v for all the specimens were 1.26%. In addition, specimen S2F2-C was externally confined by three layers of unidirectional CFRP wraps at the potential plastic hinge region at the base of the column. The height of the confined zone was 500mm which was 1.25 times the side length of the column. As stated earlier, this specimen was designed to investigate the effects of confinement on the seismic performance of FSRC columns. The mechanical properties of FRP and steel materials were tested before the construction of specimens. Tensile and compressive strengths of CFRP bars have been introduced in the second part of this paper. The CFRP wraps adopted in specimen S2F2-C were same as those used in tests of CFRP bar compressive strength. The tensile properties of a single ply carbon fiber sheet are shown in Table 1. Steel bars with diameters of 8mm, 14mm, and 18mm were used in the specimens. Their yield stresses and tensile strengths were tested in accordance with ASTM A370 [33] and are summarized in Table 1. All the six specimens were cast from one batch of ready-mixed concrete. Standard cylinders with diameters of 150mm and heights of 300mm were cast along with the columns. They were tested in accordance with ASTM C39 code [30] to measure the compressive strength of unconfined concrete f c . The tested f c at 28 days is 24.7MPa. 3.2 Design of the CFRP confinement of S2F2-C The external CFRP wraps of specimen S2F2-C was so designed that internal concrete and CFRP bars within the potential plastic hinge zone can be effectively confined. The design of the confinement scheme was described as follows. Firstly, the confinement level was determined in accordance with ACI 440.2R code [34]. In order to assure a non-softening branch in the stress-strain relationship of CFRP-confined 11/32

concrete, the effective confinement pressure f l provided by the CFRP wraps should satisfy the following requirement:

fl 

2 E f nt f  fe 2b

 0.08 f c

(4)

where E f and t f are the tensile modulus and thickness of single ply carbon fiber sheet, respectively; n refers to the number of layers of fiber sheet used to form the CFRP wrap;  fe is the effective strain of the fiber sheet and equals to 55% of the fiber ultimate strain  fu ; b = the side length of the column square section; and f c = cylinder compressive strength of unconfined concrete at 28 days. Substituting the following values, namely, E f =241GPa ,

t f =0.168mm ,  fu =0.0180 , b  400mm and f c=24.7MPa , into Eq.(4) gives the minimum layer of carbon fiber sheet n  2 . In order to assure an effective confinement, three layers of carbon fiber sheet were adopted for specimen S2F2-C. Accordingly, the actual confinement ratio CR , defined as f l f c , can be computed as: CR =0.17 . Secondly, the confinement height was determined based on the existing plastic hinge length models proposed for SRC columns and FRP-confined SRC columns. Paulay and Priestley [35] estimated the equivalent plastic hinge length of SRC columns using the following equation:

Lp 0  0.08L  0.022db f y

(5)

where L is the shear span of the column; while d b and f y are the diameter and yield strength of the longitudinal bar, respectively. Substituting

L  1800mm ,

d b  18mm

and

f y =441.2MPa into Eq.(5) gives Lp 0  319mm . After the SRC column is confined by FRP wraps, the plastic hinge length Lp _ FRP can be estimated from the model proposed by Jiang et al. [24]. For confinement ratio CR = f l f c lies within 0.1 to 0.5, the Lp _ FRP can be calculated by Eq.(5), (6) and (7):

 2r  Lpc   c   b 

0.72

Lp _ FRP  Lp 0  Lpc

(6)

  0.51  2.30CR  2.28CR2   L

(7)

where rc is the corner radius and was taken as 0.15b for specimen S2F2-C. Substituting 12/32

Lp 0  319mm , rc  0.15b , CR =0.17 and L  1800mm into Eq.(7) gives Lp _ FRP  459mm . Consequently, the height of the confined zone was conservatively taken as 500mm. To sum up, for specimen S2F2-C, three layers of unidirectional CFRP wraps were applied to a 500mm region which extended from the column-footing intersection. 3.3 Test setup and loading protocols The quasi-static test setup is shown in Fig. 4. Each column was tested under combined constant axial load and lateral cyclic displacement excursions. The axial compression was imposed through a purpose designed loading apparatus, as shown in Fig. 4. The axial load applied by the 3000kN hydraulic jack was able to move with the loading stub and remain in the gravity direction, thus taking the P-delta effects into consideration. Furthermore, a loading cell was connected to the control system of the hydraulic jack in order to keep the axial load constant. The axial compression ratio N for each tested specimen was 0.3, where

N was computed using the following equation: N 

N f cAg

(8)

Here, N is the axial load, f c is the concrete compressive strength at 28 days and Ag is the column cross-sectional area.

(a) Schematic

(b) Photograph

Fig. 4. Test setup The lateral displacement was applied by a horizontal 1000kN MTS actuator. In order to assure that the specimen can only deform within the plane parallel to the horizontal loading 13/32

direction (i.e., the x-y plane indicated in Fig. 4(a)), two roller supports were placed at each side of the loading stub, as shown in Fig. 4(b). The out-of-plane movements of the specimen were thus constrained. It is worthy to note that these roller supports will not introduce a disturbance to the lateral resistance of the specimen. In addition, displacement control was used for the lateral loading. The loading protocol was shown in Fig. 5. For the sake of clarity, displacements in push and pull direction were defined as positive and negative, respectively. Displacement levels were first incremented by 4mm until 68mm and were further incremented by 8mm until failure. Two repetitive cycles were performed for each of the displacement levels. Finally, tests were stopped when the lateral resistance of the specimens had dropped under 85% of their tested peak loads.

Fig. 5. Lateral loading protocols 3.4 Instrumentations Instrumentation details were illustrated in Fig. 6. For each tested specimen, the imposed lateral displacement was measured by a linear voltage displacement transducer (LVDT), which was labeled as LVDT1 in Fig. 6. This LVDT was mounted at the point of lateral loading and at the height of 1800mm from the top of the footing. Another LVDT, labeled as LVDT2, was mounted at the top of the footing to monitor the possible displacement of the footing. Thus, the actual displacement of the specimen equals to LVDT1 minus LVDT2. In addition, the corresponding lateral resistance was measured by the inbuilt loading cell of the MTS actuator. In this study, electrical strain gauges and OFBG sensors were used to measure the strain response of steel bars, CFRP bars and wraps. (1) Firstly, for each specimen, one strain gauge was attached at each of the four steel bars at the column corners. These four bars were labeled as No.1 to No.4 in Fig. 6. All the strain gauges located at the height of 150mm from the top of 14/32

the footing. (2) Secondly, for the four FSRC columns, OFBG sensors were embedded in the CFRP bars which located at the farthest positions from the neutral axis of the column section. These CFRP bars were labeled as No.1 to No.6 in Fig. 6. Three OFBG sensors were constructed in each of the CFRP bars. One of the three OFBG sensors located at the columnfooting intersection, while the other two were 150mm above and below the middle one, respectively. (3) Thirdly, for specimen S2F2-C which was confined by CFRP wraps, two strain gauges were adhered onto the surface of the wraps at the position of 150mm above the top of the footing. The cross-sectional positions of the gauges were illustrated in Fig. 6. Finally, the readings from the strain gauges and the OFBG sensors were collected automatically using a computer controlled data acquisition system.

Fig. 6. Layout of LVDTs, strain gauges and OFBG sensors. 4. Seismic test results and discussions 4.1 Observed behavior and failure modes The main test observations and the final failure modes of all specimens were first described. After that, an analysis of the failure modes of FSRC columns was presented. For the sake of clarity, the six specimens were divided into three categories: two conventional SRC specimens, three FSRC specimens without CFRP confinement and one FSRC specimen confined by both rectangular hoops and external CFRP wraps. Failure modes for all the six specimens were illustrated in Fig. 7. In the following discussion, the drift ratio  is defined 15/32

as the ratio of lateral displacement to the effective height of the specimen of 1800mm. (1) SRC column specimens. The development of damage in S1F0 and S2F0 were similar and both specimens suffered typical flexural failure. Flexural cracks initially formed at a displacement level of 8mm and increased quickly with the increment of lateral displacement levels. These horizontal cracks started to propagate diagonally towards the web of the column at the drift ratio of 1.11%, which was attributed to the combined effects of flexural and shear stresses. Extensive crushing and spalling of the concrete cover of specimen S1F0 and S2F0 were observed at drift ratios of 1.78% and 2.44%, respectively. Slight buckling of the longitudinal steel bars was observed after removal the crushed concrete, as shown in Fig. 7(a) and (b). (2) FSRC column specimens without CFRP wraps. For specimen S1F1, the propagation of concrete cracks was similar to those of the SRC column specimens. The concrete cover was severely crushed at the drift ratio of 2.0%. A loud crack sound was heard during the second cycle of the drift level of 2.44%. This sound was attributed to the crushing of No.2 CFRP bar (see Fig. 6 for CFRP bar labeling), which was observed on the compressive side of the column end. Moreover, the display device of the OFBG sensor system also showed that No.2 CFRP bar was in a compressive state when it failed. In addition, evidently buckled steel bars were also observed next to the crushed CFRP bar, as shown in Fig. 7(c). Specimen S1F2 was reinforced with four more CFRP bars than S1F1. Spalling of crushed concrete occurred during the drift levels of 2.44% and 2.67%. When the specimen was pulled to   2.44% for the first time, the crushing of No.4 CFRP bar (see Fig. 6 for labeling) was found at the compressive side of the column end. Besides, buckling of the longitudinal steel bar next to No.4 CFRP bar was also observed, as shown in Fig. 7(d). As for specimen S2F2, extensive spalling of crushed concrete was observed at the drift level of 2.89%, followed by a loud crack sound which was attributed to the crushing of No. 4 CFRP bar. After that, No. 2 CFRP bar crushed as well when the drift level increased to 3.78%. Both CFRP bars failed on the compressive side of the column end. The spalling of concrete, crushing of the CFRP bar, and buckling of the steel bar of specimen S2F2 were shown in Fig. 7(e). In addition, it is worth to note that damage occurred at larger displacement levels for the FSRC columns, as compared to the SRC counterparts.

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(a) S1F0

(b) S2F0

(c) S1F1

(d) S1F2

(e) S2F2

(f) S2F2-C

Fig. 7. Failure modes of tested specimens. (3) FSRC column with external CFRP confinement. For specimen S2F2-C, there were no visual signs of damage until the specimen was pushed to 16mm, at which flexural cracks started to form at the unconfined region. Part of the CFRP wraps started to bulge at the drift ratio of 3.11%, which was attributed to the lateral dilation of the internal concrete. Three crack sounds were heard when the specimen was loaded to drift ratios of 4.67%, 5.11% and 5.56%, respectively. The display device of the

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OFBG sensor system showed that these three sounds resulted from the fracture of No. 2, No. 4 and No.5 CFRP bars (see Fig. 6 for labeling), respectively. Moreover, all the three CFRP bars were in compressive states when they failed. After removal of the wraps, only marginally damage was found in the plastic hinge region. For the sake of better observation, the concrete cover was further removed, and the crushed CFRP bars and slightly buckled steel bars were observed, as shown in Fig. 7(f). (4) Analysis of the failure modes for FSRC column specimens From the above test observations, it was found that the failure modes for FSRC building column specimens were characterized by crushing of FRP bars and buckling of steel bars. This result can be explained as follows. When the FSRC column was loaded to a large drift ratio, spalling of concrete cover and lateral dilation of concrete core would take place. After that, outward local buckling was very likely to occur for both CFRP bars and steel bars which located on the compressive side of the column end, as observed in this study. Subsequently, the CFRP bars would crush due to the elastic nature. As for specimen S2F2-C, both the lateral dilation of the concrete and the outward buckling of longitudinal bars were effectively suppressed by the externally CFRP wraps. Therefore, the crushing of CFRP bars was evidently delayed as well, which can be found by comparing the observations of specimen S2F2 to S2F2-C. In addition, it can be found that for all FSRC specimens, the crushing of CFRP bars took place at drift ratios greater than 2.4%, which was larger than the allowable inter-story drift ratio (2.0%) specified by Chinese seismic design code for concrete buildings [32]. It is worth to note that the above-mentioned failure mode was more likely to occur for building columns than bridge columns. Generally, the axial load level for a building column was much larger than that of a bridge column [15]. Consequently, for an FSRC building column, the compressive stress level of the CFRP bar is higher and the dilation of concrete is more severe, thus making the CFRP bars more vulnerable to the outward buckling and crushing. 4.2 Strain response Strain response of CFRP bars and steel bars were analyzed here aiming at a better understanding of the cyclic behavior of FSRC columns. Due to limited space, only parts of the strain results were presented and discussed. Fig. 8(a) shows the strain response of specimen S1F2, which was representative of the results of tested FSRC column specimens. Strain history for No.1 and No.4 CFRP bars and

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No.1 steel bar of S1F2 were presented (see Fig. 6 for labeling of longitudinal bars). As depicted in this figure, the maximum strain response of No.1 and No.4 CFRP bars increased with the displacement levels. Finally, No. 4 CFRP bar crushed at the displacement of -44mm (   2.44% ), while No.1 CFRP bar survived the whole cyclic loading procedure. The maximum compressive strain of No.1 and No. 4 CFRP bars were 4315με and 3785με, respectively. The strain response of No.1 CFRP bar was further compared with that of No.1 steel bar, as shown in Fig.8(a) as well. Note that the effective depths of these two bars were the same. Hence, based on the plane section assumption, their strain response would be identical. However, the tensile strain in No.1 steel bar was generally larger than that of No.1 CFRP bar and the difference between them increased with the lateral displacement. At the displacement level of 24mm (   1.33% ), the tensile strain of No.1 steel bar was 2200με greater than that of No.1 CFRP bar. This result implied that slippage took place between the CFRP bars and the surrounding concrete for specimen S1F2. Such slip also made it impossible to make full use of the tensile strength of the CFRP bars. For example, the maximum tensile strain of No.1 and No.4 CFRP bars were 3413με and 2684με, respectively, less than 30% of the ultimate tensile strain 11900με. Actually, the slip between CFRP bars and concrete was also found more or less in the other FSRC specimens. This slip might be triggered by the following reasons: (1) the smooth surface texture of CFRP bars resulted in imperfect bond condition; (2) the placing quality of concrete in the specimen footing might be defective due to the congestion of the rebar cage; and (3) the develop length of the CFRP bars might be insufficient and the anchors might not be effective enough. Measures will be taken in the future experimental research to prevent or mitigate the FRP bar slippage in FSRC columns. Nevertheless, experimental and numerical studies of Ibrahim et al. [15,16] have revealed the displacement ductility of FSRC columns could be improved by the controlled slippage of FRP bars. Fig. 8(b) depicts the CFRP bar strain response of specimen S2F2-C, from which the effects of external CFRP wraps can be investigated. As shown in this figure, No.2 CFRP bar crushed when S2F2-C was pushed to +84mm (   4.67% ) for the second time at the displacement level of 84mm, corresponding to a compressive strain of 3827με. As for No. 4 CFRP bar, fracture occurred at -92mm (   5.11% ) of the second cycle at the displacement level of 92mm. The compressive strain at failure was 4602με. By assuming that the compressive modulus of the FRP bar equal to the tensile modulus [17,18,20], the ultimate

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compressive

strain

of

the

CFRP

bars

in

this

study

can

be

given

as

800.4MPa 134GPa  5973 (using test method III of section 2.3). Therefore, the compressive strains developed in No.2 and No.4 CFRP bars were 64% and 77% of the ultimate compressive strain, respectively. Furthermore, the maximum tensile strain of No.2 and No.4 CFRP bars were 5535με and 6351με, respectively, corresponding to about 50% of the ultimate tensile strain, i.e., 11900με. It was found that both the compressive and tensile strain of CFRP bars in specimen S2F2-C were larger than those of the other three FSRC specimens. Therefore, externally confining the FSRC column with CFRP wraps can effectively delay the compressive buckling and crushing of the FRP bar, thus improving the FRP bar utilization efficiency.

(a) Specimen S1F2

(b) Specimen S2F2-C

Fig. 8. Strain response. 4.3 Hysteretic curves The lateral resistance versus displacement and drift ratio curves of the tested specimens were illustrated in Fig. 9. Major observations for each specimen were also indicated in the hysteretic curves at the corresponding displacements. Characteristics of hysteretic response of tested specimens were analyzed as follows. Hysteretic curves of specimens S1F0, S1F1 and S1F2 were shown in Fig. 9(a)~(c), respectively. These three specimens were reinforced with the same amount longitudinal steel bars (  s  1.15% ). For SRC specimen S1F0, apparent degradation in lateral resistance occurred at a drift ratio of 1.8%, almost immediately after the positive and negative peak loads were achieved. This degradation was initiated by the spalling of the concrete cover on

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(a) S1F0

(b) S1F1

(c) S1F2

(d) S2F0

(e) S2F2

(f) S2F2-C

Fig. 9. Hysteretic curves. both sides of the column end. In contrast, for FSRC specimen S1F1, the lateral strength beyond the peak points (corresponding to the peak loads) degraded so marginally that two strength “plateaus” had developed in both push and pull loading directions, as indicated in Fig. 9(b). The positive plateau was terminated by the crushing of No.2 CFRP bar at a drift ratio of 2.44%, while the negative one stopped at the 2.00% drift ratio corresponding to the 21/32

spalling of concrete cover. At the end of the strength plateaus, the positive and negative lateral resistance dropped by only 4% and 3%, respectively. Likewise, lateral strength plateaus were found in hysteretic curves of specimen S1F2, as shown in Fig. 9(c). Both the positive and negative lateral resistance fell by only 3% at the end of the plateaus. The existence of strength plateaus demonstrated that the cyclic behavior of an FSRC column was more ductile than its SRC counterpart before the crushing of FRP bars. Another characteristic found in hysteretic curves of FSRC specimens S1F1 and S1F2 was the slight decline of lateral resistance resulted from FRP bar failure. The crushing of No.2 CFRP bar in S1F1 and No.4 CFRP bar in S1F2 resulted in lateral load declines of about 12.3kN and 9.4kN, respectively. This result indicated that a portion of compressive force was carried by the CFRP bars. Hysteretic curves of specimens S2F0 and S2F2 were depicted in Fig. 9(d) and (e), respectively. As compared to the hysteretic curves of the aforesaid three specimens, S2F0 and S2F2 exhibited wider hysteretic loops, which can be contributed to the significant increase of longitudinal steel reinforcement ratio  s . In addition, for specimen S2F2, the declines in lateral strength resulted from the crushing of No.2 and No.4 CFRP bars were 15.0kN and 7.5kN, respectively. Nevertheless, FSRC specimen S2F2 maintained its lateral resistance to a larger displacement level than the SRC specimen S2F0. Hysteretic curves of specimen S2F2-C were illustrated in Fig. 9(f). As compared to the other specimens, S2F2-C exhibited superior hysteretic behavior. This latter specimen showed an evident lateral resistance plateau which extending to a drift ratio of 4.7%, while all the aforementioned five specimens failed before drift ratios of 3.00% or 3.78%. Lateral resistance degradation of S2F2-C was characterized by a series of declines caused by the crushing of CFRP bars. Finally, the test was stopped at a drift ratio of 6.00% when both the positive and negative lateral resistance have dropped under 85% of the peak loads. 4.4 Envelope curves Generally, the positive and negative branches of an envelope curves are asymmetrical. To facilitate the comparison of the envelope curves, the average envelope curve (i.e., the average curve of positive and negative branches) of each specimen was calculated, as shown in Fig. 10(a) and (b). The yield points, peak point and ultimate points for each specimen were also indicated in the figures. Here, the yield point was defined by the energy method [36] (as illustrated in Fig. 10(a)); and the ultimate point corresponded to the point at which lateral resistance decreased to 85% of the peak load. The drift ratios and lateral loads at yield points 22/32

(i.e.,  y and Vy) and peak points (i.e.,  p and Vp), as well as the drift ratios at ultimate points

 u were further summarized in Table 3. Note that all these variables were calculated according to the average envelope curves. In order to consider the influences of P-delta effect on the load carrying capacity, the maximum bending moment M max at the fixed end was calculated from Eq. (9) and also provided in Table 3:

M max  max  MV  M N   max V  L  N  D 

(9)

where, M V and M N are the bending moments resulted from the lateral load V and the constant axial load N, respectively; L is the effective height of the column and D is the lateral displacement. In addition, for the purpose of comparing the ductility of the specimens, displacement ductility ratio d and load degradation ratio [15] k3 were defined in the following Eq.(10) and (11), respectively:

d 

k3 

u  100% y Vu  V p



u

pL

(10)

(11)

where, Vu is the ultimate load, i.e., Vu  0.85Vp . Actually, the load degradation ratio k3 represents the slope of the line connecting the peak point and the ultimate point, as illustrate in Fig. 10(b). The computation results for d and k3 were summarized in Table 3 as well.

(a) Specimens S1F0, S1F1 and S1F2

(b) Specimens S2F0, S2F2 and S2F2-C

Fig. 10. Average envelope curves.

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Table 3. Main test results of specimens y

Vy

p

Vp

u

M max

(%)

(kN)

(%)

(kN)

(%)

 kN  m

S1F0

0.63

138.4

1.63

165.2

2.60

336.8

4.1

12.20

1.49

1.42

12.2

S1F1

0.63

150.7

1.39

181.5

2.66

368.1

4.2

13.29

2.25

1.19

16.9

S1F2

0.77

152.5

1.54

182.6

2.89

377.6

3.8

11.00

2.17

1.13

19.7

S2F0

0.73

165.5

1.76

195.0

3.30

393.7

4.5

12.60

1.59

1.06

12.6

S2F2

0.83

184.9

1.72

221.8

3.78

442.6

4.6

12.38

2.30

0.90

18.6

S2F2-C

0.83

184.5

2.24

220.3

5.11

499.7

6.2

12.35

1.41

0.64

11.4

Specimen

d

k1

 kN

mm

k2

 kN

mm

k

k3

 kN

mm

(%)

Firstly, the load carrying abilities of the tested specimens were compared. As shown in Fig. 10 and Table 3, adding additional CFRP bars into SRC columns increased the yield load

V y , peak load V p and maximum bending moment M max . For example, the V y , V p and M max of specimen S1F2 were 10%, 11% and 12% larger than those of the reference specimen S1F0, respectively. It could be expected that the increase in load carrying capacities would be more significant if perfect bond condition was assured for the FRP bars. In addition, the maximum bending moment M max of specimen S2F2-C was 13% larger than that of S2F2, although they generally had the same lateral peak load V p . Secondly, the tested specimens were compared in terms of plastic deformation capacities and displacement ductility. As shown in Table 3, the ultimate drift ratios  u for FSRC column specimens were larger than those of their SRC counterparts, indicating the improvement in plastic deformation capacities. In addition, the displacement ductility ratio d of specimens S1F1 and S2F2 were larger than those of the reference specimens S1F0 and S2F0, respectively. Meanwhile, a decrease was also noticed as comparing the d of specimen S1F2 to that of S1F0, which might be attributed to a typical scatter of test values. Furthermore, the load degradation ratios k3 for FSRC column specimens were smaller than those of their SRC counterparts, indicating that the load degradation was more gradual for FSRC columns. This was because of that at large drift ratios, the FRP bars remained effective in resisting tension and compression while the steel bars had yield. In addition, both the ultimate drift ratio  u and displacement ductility ratio d of specimen S2F2 were increased by 32% by the additional CFRP confinement. Thirdly, the post-yield stiffness ratio  k was calculated from the following equation [15] 24/32

and analyzed:

k 

V  Vy y k2  100%  p   100% k1 Vy p y

(12)

where k1 is the secant stiffness at the yield point, and k 2 is the slope of the line connecting the yield point and the peak point, as illustrate in Fig. 10(b). The calculated results of k1 , k 2 and  k for all tested specimens were also provided in Table 3. As shown in this table, the  k of specimens S1F1 and S1F2 were 41% and 63% larger than that of the counterpart specimen S1F0, respectively. Meanwhile, as compared to the SRC specimen S2F0, the  k of FSRC specimen S2F2 increased by 46%. Therefore, adding additional CFRP bars into an SRC column was very effective in improving the post-yield stiffness ratio  k . On the contrary, as comparing the result of S2F0 to that of S1F0, it was found that  k increased only marginally with the longitudinal steel reinforcement ratio  s . Moreover, the  k of specimen S2F2 was even smaller than that of S1F2 due to the increase of  s . The tested results in this study indicated that increasing the steel reinforcement ratio of an SRC column can hardly contribute to improving the post-yield stiffness ratio. In addition, as compared to specimen S2F2, the  k of specimen S2F2-C decreased by 38%. The decrease of  k for S2F2-C can be attributed to the fact that the peak drift ratio  p of S2F2-C was 30% larger than that of S2F2. A similar result was reported in the experimental investigation of Ibrahim [15] as well. It was thus indicated that wrapping the columns with CFRP jacket may have an adverse impact on the post-yield stiffness ratio. 4.5 Residual drift ratios For each hysteresis loop, the residual displacements were defined as the positive and negative displacements ( Dr and Dr ) at which the lateral loads were equal to zero, and the residual drift ratio  r was defined as follows:

Dr  Dr 1 r    100% 2 L

(13)

where Dr and Dr were the positive and negative residual displacements, respectively. For the drift level at which two cycles were performed, only the  r corresponding to the first cycle was calculated.

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The residual drift ratio  r versus loading drift ratio  results for the tested specimens were illustrated in Fig. 11. As shown in Fig. 11(a) and (b), residual drift ratios of tested specimens increased with the loading drift evidently. Furthermore, for the three specimens in Fig. 11(a), the residual drift ratio  r at a certain drift level  decreased with the increase of the CFRP reinforcement ratio  FRP . For example, as compared to specimen S1F0, the  r of S1F1 and S1F2 at  =2% decreased by 15% and 23%, respectively. In addition, as compared to specimen S2F0 shown in Fig. 11(b), the  r of S2F2 at  =2% and  =3.6% decreased by 12% and 14%, respectively. This indicated that the self-centering capability of an SRC column can be improved by adding additional FRP bars. However, by comparing the  r of specimen S2F0 to those of S1F0, it was found that the increase in longitudinal steel reinforcement ratio  s hardly contributed to the decrease of  r for an SRC column. For example, at  =2% , the  r of S1F0 and S2F0 were 0.79% and 0.78%, respectively. Note that as specified in Chinese seismic design code for concrete buildings [32], the maximum allowable inter-story drift ratio is 2.0%.

(a) Specimens S1F0, S1F1 and S1F2

(b) Specimens S2F0, S2F2 and S2F2-C

Fig. 11. Residual drift ratios In addition, as shown in Fig. 11(b), the external CFRP wraps were also effective in decreasing the residual drift ratios of an FSRC column specimen. As compared to specimen S2F2, the  r of S2F2-C at  =2% and  =3.6% decreased by 21% and 15%, respectively. The reason might be that the damage at the plastic hinge region of specimen S2F2-C was much slighter than that of S2F2 as observed in the tests. Furthermore, it was found by comparing  r of S2F2-C to S2F0 that, the combined usage of internal CFRP bars and external CFRP wraps was an efficient way to improve the self-centering capability of a

26/32

column. For example, the  r of S2F2-C at  =2% and  =3.6% were 31% and 26%, respectively, smaller than those of SRC specimen S2F0. 4.6 Energy dissipation Insufficient energy dissipation of a structure may lead to overlarge drift demand under earthquake excitations. Previous experimental studies showed that improving the selfcentering ability of an RC column generally resulted in degradation of the hysteretic energy dissipation ability [7-9,37]. However, a different conclusion can be drawn from the test results of this study. The hysteretic loops of S1F0 and S1F2 at the displacement level of 52mm (  =2.89% ) were compared in Fig. 12(a) as representative results. As clearly shown in this figure, the residual displacements of FSRC specimen S1F2 were evidently smaller than those of SRC specimen S1F0, while the areas enclosed within the two hysteretic loops were generally the same. The dissipated hysteretic energy during this individual cycle was 10.5kN  m and 11.1kN  m for S1F2 and S1F0, respectively. This result can be attributed to

the improvement in load carrying capacity and post-yield stiffness ratio of FSRC specimen S1F2, as compared to SRC specimen S1F0.

(a) Representative results

(b) Cumulative energy dissipation Ed

Fig. 12. Hysteretic energy dissipation ability To further investigate the energy dissipation ability of the FSRC column, the cumulative hysteretic energy dissipation Ed versus loading drift ratio curves for six specimens were depicted in Fig. 12(b). The Ed was calculated by summation of the area enclosed within the hysteretic loops. However, for the drift level at which two cycles were performed, only the area corresponding to the first cycle was included. As shown in Fig. 12(b), the Ed for SRC specimen S1F0, FSRC specimen S1F1 and S1F2 at a certain drift ratio were generally identical. Likewise, the Ed for SRC specimen S2F0 were approximately equal to that of

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FSRC specimen S2F2 until the end of the tests. This indicated that when adding additional CFRP bars in a column, the hysteretic energy dissipation ability can be maintained although the residual displacements were effectively decreased. In addition, specimen S2F2-C exhibited similar hysteretic energy dissipation ability to S2F2 until a drift ratio of 3.8%, at which the latter specimen failed. After that, the cumulative energy dissipation of S2F2-C continued increasing. As a result, the final Ed of S2F2-C was almost twice as large as that of specimen S2F2, demonstrating the prominent effects of external CFRP wraps in improving the energy dissipation ability. 5. Conclusions Quasi-static tests on FRP-steel reinforced concrete (FSRC) building columns were conducted in this study. Failure modes of the specimens were explained. Effects of FRP bars on seismic performance of FSRC columns were analyzed. Based on the test results and discussions presented herein, the following conclusions can be drawn: (1) The tested compressive strength of the CFRP bar was sensitive to the confinement of the bar sample. When axial compression was directly applied to the CFRP bars, the tested strength would be very conservative. By taking confining effects of surrounding concrete on FRP bars into consideration, the tested result could increase by 100%. The three test methods involved in this study might contribute to the development of a standard test method of FRP bar compressive strength. (2) Under combined lateral displacement and an axial load corresponding to a compression ratio of 0.3, the failure modes for FSRC columns were characterized by crushing of FRP bars and buckling of steel bars as a consequence of spalling of concrete cover and lateral dilation of the concrete core. Furthermore, such FRP bar crushing occurred at drift ratios larger than the maximum allowable inter-story drift ratio (2.0%) specified by Chinese seismic design code for concrete buildings. (3) When additional CFRP bars were added into SRC columns with steel reinforcement ratios  s ranging from 1.15% to 1.91%, the post-yield stiffness ratios increased by 41% to 63% and the residual drift ratios decreased by 12% to 23%, depending on the combination with the  s and CFRP reinforcement ratios. Meanwhile, the hysteretic energy dissipation capacities were maintained. In contrast, increasing  s of an SRC column could hardly contribute to the improvement of post-yield stiffness ratio and self-centering capability. 28/32

(4) As compared to SRC specimens, FSRC columns exhibited smoother lateral resistance degradation and larger ultimate drift ratios. This was because that at large drifts, the contribution from FRP bars continued increasing due to the elastic nature of FRP; and by contrast, the stress of steel bars could increase marginally after the bars yielded. (5) Confining the plastic hinge region of the FSRC column with CFRP wraps was effective in delaying the FRP bar crushing, preventing significant damage, increasing the ultimate drift and energy dissipation and decreasing the residual drift ratios. The CFRP wrapped FSRC column suffered only slightly damage up to a drift ratio of 4.7% and sustained the smallest residual drift among all tested specimens, indicating tremendous practical implications. The focus of the current research is to experimentally investigate the contributions of FRP bars to the seismic performance of FSRC building columns. Further studies are needed to examine the influence of the mechanical properties of FRP bars and the axial load levels. In addition, measures that can delay or even prevent the crushing of the longitudinal FRP bars should also be developed and examined. Acknowledgments This study is financially supported by the National Natural Science Foundation of China (Grant No. 51478143，No. 51408153 and No. 51278150) and the China Postdoctoral Science Foundation (Grant No. 2014M551252 and No. 2015T80354). References [1] Liossatou E, Fardis MN. Residual displacements of RC structures as SDOF systems. Earthq Eng Struct D 2015;44(5):713-34. [2] Christopoulos C, Pampanin S, Priestley M. Performance-based seismic response of frame structures including residual deformations. Part I: Single-degree of freedom systems. J Earthq Eng 2003;7(1):97-118. [3] Bojorquez E, Ruiz-Garcia J. Residual drift demands in moment-resisting steel frames subjected to narrow-band earthquake ground motions. Earthq Eng Struct D 2013;42(11):1583-98. [4] MacRae GA, Kawashima K. Post-earthquake residual displacements of bilinear oscillators. Earthq Eng Struct D 1997;26(7):701-16. [5] Kawashima K, MacRae G, Hoshikuma J, Nagaya K. Residual displacement response 29/32

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