Experimental investigation on the seismic performance of GFRP-wrapped thin-walled steel tube confined RC columns

Engineering Structures 110 (2016) 269–280

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Experimental investigation on the seismic performance of GFRP-wrapped thin-walled steel tube confined RC columns Zhongkui Cai a,b, Daiyu Wang a,b, Scott T. Smith c, Zhenyu Wang a,b,⇑ a

Key Lab of Structures Dynamic Behavior and Control of the Ministry of Education, Harbin Institute of Technology, Harbin 150090, China School of Civil Engineering, Harbin Institute of Technology, Harbin 150090, China c School of Environment, Science and Engineering, Southern Cross University, Lismore, NSW 2480, Australia b

a r t i c l e

i n f o

Article history: Received 2 February 2015 Revised 24 November 2015 Accepted 26 November 2015 Available online xxxx Keywords: Reinforced concrete Steel tube confined columns Local buckling Confinement FRP Seismic performance

a b s t r a c t In order to eliminate buckling in thin-walled steel tube confined reinforced concrete (RC) columns, the external fiber-reinforced polymer (FRP) composite wraps may be applied. This paper reports the details and results of a series of seismic tests on FRP-wrapped thin-walled steel tube confined RC columns. A total of six full-scale circular cantilever columns were tested under combined constant axial load and lateral cyclic displacement excursions. The test specimens consisted of a control RC column, a carbon FRP (CFRP) confined RC column, a glass FRP (GFRP) confined RC column, a thin-walled steel tube confined RC column, and two GFRP-wrapped thin-walled steel tube (GST) confined RC columns of varying axial load levels. The diameter and height of the RC columns were 400 mm and 1600 mm, respectively, and the outer diameter to thickness ratio of the steel tubes was 135. The seismic responses of the test specimens were compared in terms of failure mode, hysteric behavior, ductility, stiffness degradation, energy dissipation capacity and equivalent viscous damping ratio. The effects of the GFRP wraps and the steel tube on the behavior of the GST confined column were revealed respectively. Test results showed that the steel tube confined column experienced local buckling and abrupt welding seam splitting resulting from high biaxial stresses, whereas such premature failure was prevented in the GST confined RC columns. For the specimens designed and tested in this study, both components in the GST contributed much to the seismic performance of the GST confined column, while the thin-walled steel tube was more effective than the GFRP wraps in improving the ultimate drift ratio. In addition, as the axial compression ratio was increased from 0.2 to 0.45, there was almost no decrease in displacement ductility for the GST confined columns. Ó 2015 Elsevier Ltd. All rights reserved.

1. Introduction Concrete-filled steel tubes (CFSTs) are becoming increasingly popular in high-rise buildings and bridges [1–3]. They can facilitate rapid construction because of the use of steel tubes as permanent formwork. In addition, as the steel tubes are continuous and can therefore resist axial force and bending moment, internal steel reinforcement is not required. There are, however, some limitations to the technology. Firstly, steel or reinforced concrete (RC) beams that need to be connected to the columns require elaborate connection strategies [4,5]. In addition, the steel tubes are vulnerable to premature buckling failure near the beam-column junction [6–9]. In order to overcome the limitations associated with beam ⇑ Corresponding author at: Key Lab of Structures Dynamic Behavior and Control of the Ministry of Education, Harbin Institute of Technology, Harbin 150090, China. Tel./fax: +86 451 8628 3856. E-mail address: [email protected] (Z. Wang). http://dx.doi.org/10.1016/j.engstruct.2015.11.043 0141-0296/Ó 2015 Elsevier Ltd. All rights reserved.

connections and buckling of steel tubes, the steel tube confined RC column system was developed [10,11]. In this system the steel tube is stopped near the ends of the RC column segments and while it does provide confinement to the column, it does not resist the axial load directly. Therefore, internal steel reinforcement is still needed in order to provide axial and flexural strength. Due to economic reasons, a thin-walled steel tube made by welding thin steel plates is usually used to provide the lateral confinement for steel tube confined RC columns. A number of studies have been carried out to investigate the axial compressive behavior and seismic performance of the system [12–17]. The test results indicate that the steel tube confined RC columns exhibited excellent seismic behavior, as well as excellent ductility and energy dissipation capacity. However, severe outward local buckling of the steel tube in such columns has been observed in the experiments of Liu and Zhou [18] and Seangatith and Thumrongvut [19]. The reason may be that the transfer of axial load from the core concrete to the external thin-walled steel tube is still possible in these

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columns because of the existence of friction or other bond stresses at the bi-material interface. The steel tube is therefore subjected to compression in the axial direction and tension in the hoop direction. As the axial stress in the steel tube becomes significant, buckling is likely to occur especially at the two ends of the columns. A solution to this problem might be put forward by drawing experience from Xiao’s [20] investigation on the local buckling of CFST, although the CFST is quite different from the steel tube confined column. In his research, the end portions of the CFST were confined with steel tube segments or fiber-reinforced polymer (FRP) wraps to constrain the outward local buckling, and this novel composite column was referred to by Xiao and colleagues as a confined CFST. Subsequent experimental and numerical investigations [21–25] have also validated the effectiveness of the external FRP wraps on improving the seismic performance of CFSTs. Inspired by the concept of confined CFST, glass FRP (GFRP) was externally wrapped at the potential plastic zones of a thin-walled steel tube confined column in order to eliminate the local buckling problem. This type of column is referred to as a GFRP-wrapped thin-walled steel tube confined RC column (or GST confined RC column) in this paper. Several advantages of GST confined columns might be expected as compared with conventional steel tube confined ones. (i) On account of lateral dilation of the concrete, the confining pressure in a thin-walled steel tube confined column is approximately constant after the steel tube yields. However, for the GST confined columns, the confining pressure may continue increasing due to the elastic nature of the GFRP wraps. This feature leads to more effective confinement to the concrete core. (ii) The outward local buckling of the steel tube might be delayed or even prevented by the external confinement provided by the GFRP wraps. (iii) Both corrosion of the steel tube and the welding seam splitting due to poor welding quality might be prevented by the application of GFRP wraps. Likewise, the GST confined column can be regard as an evolution from the FRP wrapped concrete column (FWCC) and the concrete-filled FRP tube (CFFT). Axial compression tests [26,27] have showed that the compressive behavior of circular FWCCs is similar to that of CFFTs, except when the influence of long term concrete shrinkage in CFFTs is not negligible [28]. Furthermore, studies on lateral cyclic behavior of both FWCCs [29–32] and CFFTs [33–36] have demonstrated that the energy dissipation and displacement ductility are significantly improved comparing with their un-strengthened RC counterparts. Nonetheless, there are still two advantages that might be expected from the GST confined RC column over the aforesaid two column systems. Firstly, the inner thin-walled steel tube of the GST could protect the outer FRP wraps or tubes from cutting by the internal crushed concrete. In addition, the GST confined column might be more cost-effective than the FWCC or CFFT as the steel tube is cheaper than FRP. To sum up, the GST confined RC column is attractive according to the above comparisons with the conventional steel tube confined columns, FWCCs and CFFTs. There, however, exists a knowledge gap related to this type of columns especially with regards to its seismic performance. This paper reports an experimental investigation of 400 mm diameter full-scale GST confined RC columns subjected to a constant axial load and lateral cyclic loading. To the best of the authors’ knowledge, tests have not been reported on such large scale specimens. Each tube was made from 3 mm thick steel and the steel tube was not designed to resist longitudinal load directly. Such column dimensions translate to an outer diameter-to-thickness ratio of about 135. GFRP was used instead of more conventional CFRP, as GFRP has a large strain capacity which is a desirable feature for seismic applications. In addition, GFRP does not suffer from galvanic corrosion problems which may be a concern when CFRP is directly bonded onto steel [37]. The main objects of this paper are as follows: (i) examining

whether or not the external GFRP wraps could effectively suppress the local buckling of the inner steel tube of the GST confined column; (ii) evaluating the respective effects of the GFRP wraps and the steel tube on improving the seismic performance of the GST confined columns; (iii) investigating the failure mode, displacement ductility, stiffness degradation and energy dissipation of the GST confined columns under cyclic loads, and as well as the effects of the axial load ratio on these items. 2. Experimental program 2.1. Specimen design A total of six full-scale specimens were designed and prepared consisting of (i) one ordinary RC column (ORC), (ii) one CFRPconfined RC column (C-RC), (iii) one GFRP-confined RC column (G-RC), (iv) one thin-walled steel tube confined RC column (SRC), and (v) two GST confined RC columns (GS-RC-1 and GS-RC2) of which the level of axial load was varied. The test matrix is shown in Table 1 and basic details of the test specimens are illustrated in Fig. 1. All the columns were constructed with nominally identical dimensions and steel reinforcement layouts. The diameter and effective height (from lateral loading point to top of footing) of the columns were 400 mm and 1400 mm, respectively. Six 20 mm diameter hot-rolled deformed steel bars were adopted for the longitudinal reinforcement which resulted in a longitudinal steel reinforcement ratio of 1.5%. Also, 8 mm plain steel bars were used as hoop reinforcement. These bars were spaced at 200 mm centers which resulted in a small hoop reinforcement volumetric ratio of 0.3%. The concrete cover thickness was 30 mm and it was measured to the outside of the hoop steel reinforcement. The columns were constructed with a stub footing of dimensions 1500
600
600 mm in order to simulate a fixed-end boundary condition. In addition, an RC cuboid of dimensions 600
400
400 mm was cast integrally at the top of each column in order to accommodate the connection of the lateral loading actuator. Both the rectangular footing and column-top cuboid were heavily reinforced in order to eliminate any premature failure during testing. 2.2. Specimen construction Fig. 2 shows the extent of the confinement schemes for all six test specimens. Specimen ORC, which served as a plain RC control, did not contain any external confinement. Specimens C-RC and G-RC were confined by three layers of uni-directional CFRP and GFRP wraps, respectively, at the potential plastic hinge region at the base of the column. The height of the confined zone was 500 mm which was about 1.25 times the diameter of each column. Specimen S-RC was a thin-walled steel tube confined RC column, in which the steel tube was made by rolling a 3-mm-thick steel plate onto a cylindrical shell with the seam continuously welded. The position of welding seam, as shown in Fig. 2(f), is the most

Table 1 Details of test specimens. Specimen

Confinement

n

fc (MPa)

ORC C-RC G-RC S-RC GS-RC-1 GS-RC-2

No confinement CFRP wrap GFRP wrap Steel tube GST GST

0.45 0.45 0.45 0.45 0.45 0.20

27.1 27.1 27.1 24.2 24.2 24.2

Note: GST = GFRP-wrapped thin-walled steel tube; n = axial compression ratio; fc = 28-day compressive strength of concrete.

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8@100mm 15mm gap

FRP-wrapped steel tube or steel tube

A

A

6 20mm Long. bars

15mm gap 8@200mm

400mm A-A

Fig. 1. Specimen dimensions and steel reinforcement details.

3-layers CFRP

(a) ORC

3-layers GFRP

(b) C-RC

(c) G-RC

500mm

Internal steel tube

700mm

Pull

Thin-walled steel tube

500mm

500mm

No external confinement

Push

1-layer GFRP Welding seam 3-layers GFRP

150mm FRP overlap zone

(d) S-RC

(e) GS-RC-1, GS-RC-2

(f) Position of welding seam and FRP overlap zone

Fig. 2. Confinement details of test specimens.

unfavorable position in terms of the stress state. The outer diameter to thickness ratio of the steel tube was 135. An additional two steel tubular RC columns that were identical to specimen S-RC were also prepared. These additional specimens, namely specimens GS-RC-1 and GS-RC-2, were confined with uni-directional GFRP wraps. Three layers of wrap were applied to a 500 mm region

which extended from the column footing, while one layer of wrap was applied to the remaining 700 mm length of column. All the FRP wraps used in the entire experimental program were applied in a wet lay-up manner with a 150 mm overlap, and the overlap zone is shown in Fig. 2(f). It is worth noting that the steel tubes utilized in specimens S-RC, GS-RC-1 and GS-RC-2 were applied over

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Table 2 Material properties of steel and fiber sheets. Material Longitudinal bar Hoop bar Steel tube Carbon fiber sheet Glass fiber sheet

d or t (mm)

fy (MPa)

d = 20 d=8 t = 3.0 t = 0.167 t = 0.359

fu (MPa)

348 417 315  

383 616 462 3400 3240

ey or eu (le)

E (MPa) 5

2.0
10 2.0
105 2.0
105 2.45
105 7.24
104

ey = 1740 ey = 2085 ey = 1575 eu = 13,878 eu = 44,751

Note: d = diameter; t = thickness of the steel tube or single ply fiber sheet; fy = yield strength; fu = ultimate strength; E = elastic modulus; ey = yield strain of steel; eu = ultimate tensile strain of fibers.

Rollers Axial loading apparatus

Load cell Hydraulic jack Steel hinge

Specimen Reaction wall Ground anchor

(a) Schematic

(b) Photograph

Fig. 3. Test setup.

the full height of the columns but terminated 15 mm from both ends of the columns. The purpose of these gaps was to avoid direct transfer of axial load to the thin-walled steel tube. Therefore, by comparing the test results of specimen GS-RC-1 with specimen S-RC, the confinement effect provided by the external GFRP wrap in the GST confined RC columns can be investigated. Similarly, comparing the test results of specimen GS-RC-1 with specimen G-RC can reveal the contribution of the inner thin-walled steel tube to the seismic performance of the column system. 2.3. Materials The six test column specimens were casted in two batches of three due to space limitations of the laboratory. Three cylinders of 150 mm diameter of 300 mm height were cast along with each batch of columns and they were tested in accordance with ACI 31808 code [38]. The average 28-day cylinder compressive strength of concrete for the first and second batches was 27.1 MPa and 24.2 MPa, respectively. These strengths are indicated in Table 1. The difference in strength is minimal and hence it is not considered to impact noticeably on the column test results. Table 2 is a summary of the mechanical properties of the steel and glass/carbon fiber sheets. The properties of the steel materials (i.e. reinforcing bars and steel plate) were obtained from tension tests in accordance with ASTM A370 [39], while the properties of the fiber sheets (before resin impregnation) are provided by the manufacturer. 2.4. Testing and instrumentation The column specimens were tested under constant axial load and cyclic lateral displacement using the test setup shown in Fig. 3. As provided in Table 1, two different axial compression ratios of 0.2 and 0.45 were applied. The axial compression ratio

for specimens ORC, C-RC, G-RC, S-RC and GS-RC-1 was maintained at 0.45, and it was more than twice the ratio of 0.2 adopted for specimen GS-RC-2. The axial compression ratio, n, was calculated from Eq. (1):

n¼

N f cA

ð1Þ

where N is the axial load, fc is the cylinder compressive strength and A is the cross section area of the RC column. Constant axial load and cyclic lateral displacement were applied by a vertically mounted 3000 kN hydraulic jack and a horizontally positioned 1000 kN MTS actuator, respectively. As illustrated in Fig. 3, a purpose designed and built loading apparatus was utilized in order to enable a constant axial load to be maintained and moved with the upper part of the column during testing. Therefore, second-order effects (i.e. P-delta effect) could be incorporated and hence simulated during testing. The axial load was applied to the column first and it was maintained by a pressure-relief valve which was part of a hydraulic system. Displacement control was used for the lateral loading. Owing to the considerable difference in ductility between the ordinary RC column specimen (i.e. specimen ORC) and the strengthened RC columns, two different displacement control loading methods were adopted. The displacement levels for specimen ORC were incremented by 4 mm with two cycles per drift level until failure. For the other five specimens, the increment of the imposed displacement level was 8 mm, with two cycles per level until 48 mm and one cycle for the following displacement levels until failure. A limited amount of lateral displacement was programmed into the actuator controller in order to protect the experimental set-up. Finally, tests were stopped when the lateral resistance of the columns reduced to 70% of their tested peak load capacity, or the limitation of the actuator displacement was reached. Fig. 4 illustrates the instrumentation details. The lateral force was measured by a load cell embedded within the actuator, while

Z. Cai et al. / Engineering Structures 110 (2016) 269–280

Push

Pull

LVDT1 LVDT2 North

South

LVDT3

A 50

200

350

LVDT4 A

LVDT5

HN200

HS200 VS200

Horizontally attached strain gauges

VN200 Vertically attached strain gauges VM200 A-A Fig. 4. Layout and labeling of LVDTs and strain gauges.

a total of five linear voltage displacement transducers (LVDTs) were mounted along the length of the column to measure deformation of the specimen. In addition, fifteen electrical strain gauges of 10 mm gauge length were adhered onto the surface of each test specimen. Five gauges were adhered at positions 50 mm, 200 mm and 350 mm above the top of the footing for each column. Of each set of five gauges, three of these gauges were oriented vertically while the remaining two gauges were oriented in the hoop direction. The axial load applied to each specimen was measured by a load cell as shown in Fig. 3. Finally, the readings from the LVDTs, strain gauges and load cells were collected automatically using a computer controlled high speed data acquisition system. In addition, the plastic hinge behavior is of great importance to the performance of flexural members, and yet it was not measured in the current study for the following reason: The plastic hinge length is typically measured indirectly by strain gauges on longitudinal bars [40] or rotation measurements [41], while such instrumentation might have an adverse impact on the performance of tested specimens. An innovative approach of determining the plastic hinge length of CFFTs was proposed by Ozbakkaloglu and Saatcioglu [42,43]. In their studies, the hoop-strain profile of FRP tubes was recorded and used to establish the plastic hinge region behavior. Evidently, this method is more convenient and efficient than the conventional approaches, and it will be adopted to study the propagation and the length of plastic hinge of the GST confined column in the further research. 3. Test results and discussions 3.1. Failure modes The final failure modes of all specimens are shown in Fig. 5 while the following is a summary of the main test observations:

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(1) Specimen ORC suffered flexural failure at a small drift ratio of 2.0%. The drift ratio is defined as the ratio of lateral displacement as per LVDT1 minus LVDT5 (see Fig. 4 for LVDT labeling) to the effective height of the columns of 1400 mm. Flexural cracks initially formed in the plastic region at a displacement level of 4 mm, and started to propagate diagonally at a displacement level of 16 mm. Subsequently, the flexural-shear cracks widened and lengthened with the increment of lateral displacement. At the end of the test, extensive spalling of crushed concrete and slight buckling of the longitudinal bars were observed, as shown in Fig. 5(a). (2) For specimens C-RC and G-RC, there was almost no evident damage during the loading stage. Tests for both specimens were stopped at a drift ratio of 5.1% when the lateral force had reduced to about 70% of the peak load. There was no observed indication of severe damage for these specimens after removal of the FRP wraps, as is evident in Figs. 5 (b) and (c). (3) For specimen S-RC, local buckling occurred at the bottom of the steel tube when the specimen reached a displacement of 40 mm for the first time. Subsequently, the welding seam of the steel tube in the buckled region abruptly split. The reason for the splitting may be that the welding seam was at the location (as shown in Fig. 2) with the most disadvantaged stress state. This may have been exacerbated on account of poor weld quality. The split, which initiated at the base of the column, propagated gradually with increased lateral displacement. Finally, crushed concrete and buckled longitudinal internal reinforcing steel bars were observed post-test after removal of the steel tube, as shown in Fig. 5 (d). (4) The development of damage in specimens GS-RC-1 and GSRC-2 was similar. Initially, there were no visual signs of damage in both specimens at drift ratios smaller than 5.1%. At drift ratios of about 5.7%, a gap between the column and the footing appeared. For specimen GS-RC-1, testing was stopped at a drift ratio of 7.4% when the lateral force in the push direction reduced to about 70% of its peak load. For specimen GS-RC-2, testing was stopped when the stroke limit of the MTS actuator was reached, and the positive and negative lateral strength had already dropped to 77.0% and 82.6% of the average peak strength, respectively. Unlike specimen S-RC, neither local buckling nor welding seam splitting was observed for both GST confined specimens throughout the tests. After removal of the FRP and steel tubes post-test, there was no apparent damage to the original RC columns as evident in Fig. 5(e) and (f). In addition, there was no de-bonding between the external GFRP wraps and the internal steel tube. 3.2. Strain response In the earlier studies on the steel tube confined columns [15,16], it was expected that the axial loads did not transfer from the core concrete to the external steel tube and that the effect of steel tube was to provide lateral confinement only. However, outward local buckling of the steel tube was observed in this study and another two experimental investigations [18,19]. In order to identify the reason for the premature local buckling and welding seam splitting of the steel tube confined column tested in this study, the strain responses of the steel tube in specimen S-RC are investigated. The strain gauge results are shown in Fig. 6 in accordance with the following notation. HN200 and VN200 are the strain gauges that were horizontally and vertically attached, respectively, at 200 mm above the top of the footing on the north-

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(a) ORC

(b) C-RC

(c) G-RC

(d) S-RC

(e) GS-RC-1

(f) GS-RC-2 Fig. 5. Failure modes of test specimens.

ern surface of the specimen. As stated previously in this paper, the cross-sectional positions of these strain gauges can be found in Fig. 4. As illustrated by VN200 in Fig. 6, the steel tube was generally compressed in the axial direction during testing, although the steel tube did not carry axial load directly. This indicated that the axial load could be transferred from the core concrete to the external thin-walled steel tube because of friction or bond stresses. At the time local bucking of the steel tube occurred (i.e., the specimen reached a displacement of 40 mm for the first time), the compressive strain of VN200 and tensile strain of HN200 were 2050 le and 1480 le, respectively. These correspond to points A and B as identified in Fig. 6. Accordingly, the Mises equivalent stress rs at the location of strain gauge VN200 can be calculated according to Eq. (2) as the steel tube was approximately in a biaxial stress state: Fig. 6. Strain responses of specimens S-RC.

rs ¼

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r21  r1 r2 þ r22

ð2Þ

Z. Cai et al. / Engineering Structures 110 (2016) 269–280

where ri ¼ Es ei (i = 1, 2), and Es is the elastic modulus of steel. Here, e1 ¼ 1480 le and e2 ¼ 2050 le. According to the calculation, rs is equal to 614 MPa, which is much larger than the yield strength of the steel tube in S-RC (i.e. yield strength = 315 MPa). Therefore, the thin-walled steel tube had already yielded according to the Mises yield criterion. Furthermore, according to Eq. (2), with an increase of longitudinal compressive stress (negative), the steel tube would yield at a smaller hoop tensile stress (positive), thus leading to less effective confinement. Therefore, once part of the steel tube yielded, outward local buckling was likely to occur as observed in specimen S-RC. In addition, as shown in Fig. 6, the maximum responses per loading cycle for both VN200 and HN200 decreased at the late stage of testing. This may be attributed to weld seam splitting, which may then lead to de-bonding between the steel tube and concrete core. Based on the conclusion that the steel tube is subjected to a biaxial stress state and that the axial stress is significant at large drift ratios, the weld seam may be vulnerable to splitting especially when the welding quality is suspect. Obviously, such weld failure would lead to an apparent decrease in the confinement effect for the concrete core, resulting in a less ductile response for the column. Therefore, it was necessary to add external confinement onto the thin-walled steel tube confined RC column to reduce the risk of premature buckling and welding failure. 3.3. Lateral load versus drift response The lateral load-drift ratio hysteretic responses of all test specimens are shown in Fig. 7. Characteristics of the hysteretic curves of tested specimens are summarized as follows. Specimen ORC exhibited very poor hysteretic behavior as observed by narrow hysteretic loops arising from rapidly decreasing lateral resistance at small drift. This response is expected because specimen ORC is a conventional RC column with an inadequate hoop reinforcement volumetric ratio of only 0.3%. Compared to specimen ORC, specimens C-RC, G-RC and S-RC exhibited much wider hysteretic loops with more gradual decrease in lateral resistance. All three of these specimens failed at a drift ratio more than twice that of specimen ORC. However, specimen S-RC exhibited an apparent decline in positive lateral resistance from a drift ratio of 3% (as indicated in Fig. 7(d)) to the end of testing. This may be attributed to the local buckling. Furthermore, this premature failure of specimen S-RC led to an extremely asymmetrical decrease in lateral resistance. At a drift ratio of 5.1%, the positive lateral resistance declined to 65% of its positive peak load, whereas the negative lateral resistance was still larger than 90% of its negative peak load. Compared to behavior of FRP-confined RC columns C-RC and GRC, as well as the steel tube-confined column S-RC, specimen GSRC-1 exhibited superior hysteretic behavior. This latter specimen showed a more gradual decrease in lateral resistance with failure at a much larger drift ratio. The hysteretic curves of specimens GS-RC-1 and S-RC are very similar at a drift ratio less than about 3%, as shown in Fig. 7(d). Although, upon the drift ratio exceeding 3.4% at which local buckling and weld splitting occurred in specimen S-RC, specimen GS-RC-1 maintained a much smoother decrease in lateral resistance than specimen S-RC. Accordingly, it can be concluded that the external GFRP wraps play a key role in maintaining the steady hysteretic behavior of a GST confined RC column at a comparatively large drift ratio. The influence of compression ratio can be determined from inspection of the hysteretic behaviors arising from specimens GS-RC-1 and GS-RC-2. These two specimens were subjected to compression ratios of 0.45 and 0.2, respectively, of which the former is considerably larger than the latter. As shown in Figs. 7 (e) and (f), the decrease in lateral resistance of specimen GS-RC-2

275

was slightly slower than specimen GS-RC-1. However, the lateral load-capacity as well as the area enclosed within the hysteretic curves of specimen GS-RC-1 was larger than that of specimen GS-RC-2. This demonstrates that with an increase of the axial compression ratio in a certain range for GST confined RC columns, the increase in lateral load-carrying capacity and energy dissipation capacity is apparent while the decrease in ductility is not as obvious. Although the specimens tested in this study do not fall into the category of the short column since the shear span ratio is 3.5, the shear effects on the lateral load-drift ratio hysteretic responses are still not negligible, especially considering that the axial load ratio is 0.45 and the hoop reinforcement volumetric ratio is 0.3%. Therefore, pinching effect can be found more or less in the hysteretic curve for each tested specimen. Furthermore, as shown in Fig. 7, the pinching effects of specimens GS-RC-1 and S-RC are less obvious as opposed to specimens C-RC and G-RC. Hence, it might be concluded that the steel tube is, to some extent, more effective in resisting shear effects than the FRP wraps with fibers oriented only in the hoop direction. However, such a conclusion requires further experimental and analytical studies for validation. 3.4. Envelop curves In order to compare the ductility and plastic deformation capacity of the tested specimens, the displacement ductility ratio l and the ultimate drift ratio du are defined, respectively, in Eqs. (3) and (4):

l ¼ Du =Dy

ð3Þ

du ¼ Du =L
100%

ð4Þ

where Dy and Du are the yield displacement (as defined by the energy method) and ultimate displacement, respectively. Both of these parameters are defined in Fig. 8. Note that the lateral loads for all specimens have dropped under 85% of the peak strength at the ending of tests, thus the definitions of l and du are consistent in this paper. In addition, L is the effective height of the specimen (i.e., 1400 mm). The envelop curves of all specimens obtained from the hysteretic curves are shown in Fig. 9. Based on the envelop curves, Dy, Du, l, du as well as the peak loads Pp are calculated and summarized in Table 3. Meanwhile, the normalized peak loads Pp/(fcA) are also calculated in order to remove the effect of variation in concrete strength. Additionally, considering the influences of Pdelta effect on the load carrying capacity, the bending moment capacity Mmax at the fixed end is calculated from the following equation and also provided in Table 3:

Mmax ¼ maxðMP þ MN Þ ¼ maxðP  L þ N  DÞ

ð5Þ

where MP and MN are the bending moments caused by the lateral load P and the constant axial load N, respectively; L is the effective height of the column and D is the lateral displacement. Note that for the above variables, the average values of push and pull results are compared in this study. As shown in Table 3, the increases in the normalized peak lateral loads of the strengthened columns C-RC, G-RC, S-RC and GSRC-1 varied from 8% to 18%, as compared to the control specimen ORC, and similarly, the increases in bending moment capacities Mmax ranged from 10% to 19%. Thus, the improvement in the load carrying capacity by the strengthening schemes studied in this paper was not remarkable. However, the displacement ductility ratio and ultimate drift ratio were significantly improved for these strengthened specimens, and this was especially the case for the two GST confined RC columns. Nonetheless, larger ultimate drift could lead to more severe P-delta effect, especially when the axial

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(a) Hysteretic curve of ORC vs. GS-RC-1

(b) Hysteretic curve of C-RC vs. GS-RC-1

(c) Hysteretic curve of G-RC vs. GS-RC-1

(d) Hysteretic curve of S-RC vs. GS-RC-1

(e) Hysteretic curve of GS-RC-1

(f) Hysteretic curve of GS-RC-2 vs. GS-RC-1

Fig. 7. Load–displacement hysteretic curves.

load level is high. In order to quantify the influence of the secondorder effects, a parameter kpd was defined in Eq. (6):

    MN ND ¼ max ; D 6 Du kpd ¼ max MP þ MN PLþND

ð6Þ

Evidently, the ‘‘P-delta effect index” kpd indicates the maximum proportion of MN in the total bending moment. The kpd for all specimens were calculated and summarized in Table 3. As it is shown, the P-delta effect accounts for more than 20% of the bending moment for the strengthened specimen, including specimen GSRC-2 of which the axial load ratio is 0.2. Based on the test results,

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was increased by 91%, 135% and 191%, respectively, by these three different confining materials. Apparently, the improvement of ultimate drift ratio from GST is less than the summation of respective improvement from the steel tube and GFRP wraps. In other words, the confinement effect of GST is not equal to the simple summation of the two components. According to the above comparisons, for the specimens designed and tested in this study, the internal steel tube in the GST was more effective than the external GFRP wraps in improving the ultimate drift ratio, even though its out diameter-tothickness ratio was as large as 135. The reason for the above result might be that the elastic-to-plastic nature of the steel was more effective to increase the deformation capacity of a column than the linear elastic nature of the GFRP. Additionally, as compared to specimen C-RC, the ultimate drift ratio of specimen GS-RC-1 increased by 34%, while the displacement ductility ratio of specimen GS-RC-1 was a little smaller than that of specimen C-RC. This may be attributed to the fact that the yield displacement of specimen GS-RC-1 was 37% larger than that of specimen C-RC. As the axial compression ratio increased from 0.2 to 0.45, the peak load of specimen GS-RC-1 increased by 33% above specimen GS-RC-2. However, similar displacement ductility ratios were recorded for these two specimens. This indicates that with an increase of axial compression ratio within a certain range for GST confined RC columns, the enhancement in lateral load-carrying capacity is obvious while the decrease in ductility is not. This characteristic may prove desirable for the structural design of tall buildings and heavy industry factories in which the axial compression ratio of columns are relatively high.

Fig. 8. Definition of displacement ductility ratio.

3.5. Stiffness degradation

Fig. 9. Load–displacement envelop curves.

therefore, the P-delta effect should be taken into account in the cyclic test or analysis of such strengthened RC columns, even though the shear span ratio of is only 3.5. The effects of the external GFRP wraps and the internal thinwalled steel tube on the ductility of the GST confined column could be revealed by comparing the results of specimen GS-RC-1 to S-RC and G-RC, respectively. As compared to S-RC, the ductility ratio l and ultimate drift ratio du of GS-RC-1 increased by 27% and 24% respectively. This improvement is contributed from the additional confinement of external GFRP wraps. From the comparison with GRC, l and du of GS-RC-1 increased by 26% and 52% respectively, attributed to the internal thin-walled steel tube. In addition, the effects of the GFRP wraps and the steel tube in the GST can also be evaluated by comparing specimens G-RC, S-RC and GS-RC-1 to specimen ORC. It was found that the ultimate drift ratio of ORC

It can be seen from the envelop curves in Fig. 9 that the initial stiffness of all specimens was generally similar. This is desired since the main purpose of strengthening is to improve ductility instead of stiffness. However, as a result of concrete cracking, yielding of the longitudinal bars, and anchorage-slip between the concrete and steel reinforcement, degradation in stiffness occurred in the specimens under reversed cyclic load. In this paper, the stiffness degradation is quantified by consideration of effective stiffness. Based on the hysteretic curves, the effective stiffness Ke at a certain displacement magnitude is defined as follows:

Ke ¼

K þe þ K e 2

ð7Þ

 where K þ e and K e are the secant stiffness at positive and negative peak load in any hysteretic loop, respectively. For the displacement magnitude at which two cycles were conducted, only the effective stiffness corresponding to the first cycle was calculated. The effective stiffness at the first displacement magnitude of cyclic loading is noted as Ke0. Therefore, Ke/Ke0 is a measure of the effective stiffness degradation. Inspection of Fig. 10 reveals that the effective stiffness of specimen ORC degraded rapidly throughout the testing. As for the

Table 3 Main test results of specimens. Specimen

Pp (kN)

Pp/(fcA)

Mmax (kN m)

kPd

Dy (mm)

Du (mm)

l

du (%)

Ed (kN m)

ORC C-RC G-RC S-RC GS-RC-1 GS-RC-2

208.5 225.5 244.7 212.6 218.0 164.2

6.1e2 6.6e2 7.2e2 7.0e2 7.2e2 5.4e2

316.5 358.9 376.1 348.0 368.7 255.3

14% 24% 21% 25% 28% 21%

8.82 6.54 7.42 9.23 8.96 10.22

25.90 56.28 49.42 60.90 75.32 87.50

2.94 8.61 6.66 6.60 8.41 8.56

1.85 4.02 3.53 4.35 5.38 6.25

41.7 153.9 163.0 126.5 308.1 277.7

Note: Pp = peak lateral load; Pp/(fcA) = normalized peak load; Mmax = the bending moment capacity; kPd = P-delta effect index; Dy = yield displacement; Du = ultimate displacement; l = displacement ductility ratio; du = ultimate drift ratio; Ed = cumulative energy dissipation.

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Fig. 10. Degradation of effective stiffness.

strengthened specimens, the apparent degradation in effective stiffness is observed at small drift ratio levels. For a drift ratio around 1.5%, however, the effective stiffness degradation of all strengthened columns became less rapid. It should be noted that the peak loads of these confined columns were achieved around such a drift. Furthermore, the stiffness degradation of specimen GS-RC-1 was less rapid than specimens C-RC, G-RC and S-RC. In addition, although the axial compression ratio of specimen GSRC-1 was more than twice that of GS-RC-2, almost the same stiffness degradation was exhibited in the complete drift range. This indicates that the increase of axial compression ratio within a certain range has little influence on the stiffness degradation of GST confined RC columns. 3.6. Energy dissipation Insufficient energy dissipation may lead to failure of structures in earthquake disasters. Therefore, cumulative energy dissipation is an important indicator of seismic performance of a structural member. Fig. 11 shows the hysteretic energy dissipation versus drift ratio results for the tested specimens. Cumulative energy dissipation at the end of tests of all specimens Ed is also summarized in Table 3. The results were calculated by summation of the area enclosed within the hysteretic curves. However, for the displacement magnitude at which two cycles were conducted, only the area corresponding to the first cycle was included. As shown in Fig. 11 and Table 3, the energy dissipation capacity of the ordinary RC column was significantly improved by all the confinement interventions adopted in this study. In addition, the

Fig. 12. Equivalent viscous damping (EVD) ratio.

energy dissipation of specimens C-RC, G-RC, S-RC and GS-RC-1 were very similar before a drift ratio of 5.1%, at which the former three specimens failed. After that, the cumulative energy dissipation of specimen GS-RC-1 continued increasing. As a result, with the same axial compression ratio, the final cumulative energy dissipation of specimen GS-RC-1 was almost or even more than twice as large as that of specimens C-RC, G-RC and S-RC. Therefore, both the internal steel tube and the external GFRP wraps in the GST specimens were of great importance in terms of improving the energy dissipation. As illustrated in Fig. 11, the cumulative energy dissipation of specimen GS-RC-1 was about 20–30% larger than that of specimen GS-RC-2 at the same drift ratio throughout the tests. This indicates that the energy dissipation capacity of the GST confined RC columns increased with the axial compression ratio. Therefore, on account of their superior energy dissipative characteristics, it may be appropriate to include such columns in high-rise buildings or heavy industrial factories that are located in high seismicity zones. 3.7. Equivalent viscous damping ratio The combined effect of elastic and hysteretic damping, which can be estimated by the equivalent viscous damping (EVD) ratio, is also a crucial parameter in the application of a displacement based design method [44]. The EVD versus drift ratio of each specimen is presented in Fig. 12. From this Figure it is evident that the EVD ratios of all specimens generally increased with increasing drift ratio. Under the same axial compression ratio, the EVD ratios of specimens C-RC, G-RC, S-RC and GS-RC-1 were generally larger than that of the un-confined specimen ORC. Additionally, both the initial and the final EVD ratio of specimen GS-RC-1 was higher than that of any other specimen. It is worth noting that the EVD ratio of specimen GS-RC-1 was generally higher than specimen GS-RC-2 at the same drift ratio levels. This indicates that the EVD ratio of a GST confined RC column increases with axial compression ratio within a certain range. On the contrary though, an increase in the axial compression ratio tends to decrease the EVD ratio in an ordinary RC column [44]. Accordingly, analysis of the EVD capacity of such a GST confined RC column must take this characteristic into consideration. 4. Conclusions

Fig. 11. Cumulative energy dissipation.

This paper presented an experimental investigation on the GST confined RC column, in which the external GFRP wraps were

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expected to prevent the local buckling of the inner steel tube and to increase the lateral confining pressure after the steel tube yields. Cyclic tests were conducted on full-scale circular columns to validate the design concept and to evaluate the respective effects of the GFRP and the steel tube on seismic performance of the GST confined column. Based on the test results and discussions presented herein, the following conclusions can be drawn: (1) When the thin-walled steel tube confined RC column was loaded both axial load and lateral displacement, the steel tube in the plastic region was subjected to a biaxial stress state and the axial stress might be significant at large drift ratios. This result was inconsistent with earlier analysis which expected that the axial loads did not transfer from the core concrete to the external steel tube. (2) The tested steel tube confined RC columns suffered from outward local buckling and welding seam splitting which resulted from high biaxial stresses. By contrast, such premature failure modes were prevented by additional confinement from the external GFRP wraps in the GST confined column. (3) By comparing the results of specimen GS-RC-1 to S-RC and G-RC, it was found that for the specimens designed and tested in this study, the thin-walled steel tube in the GST was more effective than the GFRP wraps in improving the ultimate drift ratio, even though its out diameter-tothickness ratio was as large as 135. (4) With an increase in axial compression ratio from 0.2 to 0.45, the decrease in ductility of the GST confined RC column was negligible. This characteristic may prove the feasibility of this kind of columns for the tall building structures and heavy industry factories in which the axial compression ratio of columns are relatively high. The focus of the current study was to experimentally evaluating the respective effects of the GFRP wraps and the steel tube on improving the seismic performance of the GST confined columns. Further researches on the GST confined RC columns, GFRP confined columns and steel tube confined columns with the same amount of later confinement are in preparation.

Acknowledgements This research was supported by the National Natural Science Foundation of China (Grant Nos. 51278150, 51478143), the National Key Basic Research Program of China (973 Program, Grant No. 2012CB026200) and the Fundamental Research Funds for the Central Universities (Grant No. HIT. NSRIF. 2015097).

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