Tests on cyclic performance of FRP–concrete–steel double-skin tubular columns

ARTICLE IN PRESS Thin-Walled Structures 48 (2010) 430–439

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Tests on cyclic performance of FRP–concrete–steel double-skin tubular columns Lin-Hai Han a,, Zhong Tao b, Fei-Yu Liao a, Yi Xu b a b

Department of Civil Engineering, Tsinghua University, Key Laboratory of Structural Engineering and Vibration of China Education Ministry, Beijing 100084, PR China College of Civil Engineering, Fuzhou University, Fuzhou, Fujian Province 350108, PR China

a r t i c l e in f o

a b s t r a c t

Article history: Received 3 March 2009 Received in revised form 9 December 2009 Accepted 21 January 2010 Available online 12 February 2010

Eight FRP–concrete–steel double-skin tubular columns were tested under constant axial load and cyclically increasing flexural loading. The main parameters in the tests are axial load level and number of fibre reinforced polymer (FRP) layers. The influence of those parameters on the strength, ductility, stiffness, and energy dissipation was investigated. It was found that, in general, FRP–concrete–steel double-skin tubular columns exhibit high levels of energy dissipation prior to the rupture of the longitudinal FRP, but experience a sudden drop in the lateral load capacity after that. The ductility of the specimens can be improved to some extent due to the existence of the axial compressive load in current tests. & 2010 Elsevier Ltd. All rights reserved.

Keywords: Fibre reinforced polymer (FRP) Concrete Steel tube Double-skin Hysteretic behaviour Ductility

1. Introduction In recent decades, due to the advantages (such as high strength and resistance to corrosion, etc.) that fibre reinforced polymer (FRP) offers over traditional structural materials (like steel and concrete), there has been a growing interest in the use of FRP composites in civil engineering, both in retrofit of the existing structures and using in new constructions [1,2]. Therefore, lots of research work has been done to examine the behaviour of structures consisting of FRP composites in combination with other materials [1]. FRP–concrete–steel double-skin tubular column (DSTC) is an innovative type of composite members, which was proposed by Teng et al. [3,4]. Fig. 1 gives a schematic view of the typical cross-sections. The composite column form consists of a steel tube inside, a FRP tube/ wrapping outside and concrete in between. This type of composite column is expected to combine the advantages of all three types of materials and those of the structural form of concrete-filled doubleskin steel tubular columns. A series of stub columns and beams of the FRP–concrete–steel DSTC were conducted by Teng et al. [4], Yu et al. [5]. The test results of the stub columns had confirmed that the sandwiched concrete in the column could be effectively confined by the outer FRP tube, thus showed a ductile response, and the effectiveness of confinement was

not affected by the inner void in the double–skin column [4]. The test result of the beams showed that the composite construction demonstrated very ductile flexural behaviour [5]. However, there is still a lack of information on the composite members under cyclic loading. It indicates a need for further research in this area. Eight FRP–concrete–steel DSTCs were thus tested under constant axial load and cyclically increasing flexural loading in this paper. The main parameters in the experiment are axial load level and number of FRP layers. The influence of those parameters on the strength, ductility, stiffness, and energy dissipation was investigated.

2. Experimental programme 2.1. General Eight FRP–concrete–steel DSTCs, including four columns with square sections and four columns with circular sections, were tested. Test parameters were axial load level (n) and number of FRP layers (nf). Two kinds of cross-sections were selected as shown in Fig. 1(a) and (b), respectively, i.e., circular column with circular inner steel tube and square column with circular inner steel tube. The axial load level (n) in this paper is defined as following, i.e.:

 Corresponding author. Tel.: + 86 10 62797067; fax: + 86 10 62781488.

E-mail addresses: [email protected], [email protected] (L.-H. Han). 0263-8231/$ - see front matter & 2010 Elsevier Ltd. All rights reserved. doi:10.1016/j.tws.2010.01.007

n¼

No Nu

ð1Þ

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Nomenclature B D Ds E Ec Es Ki Ks M Mu

width of square section diameter of circular section diameter of the inner steel tube energy dissipation concrete modulus of elasticity steel modulus of elasticity initial section flexural stiffness of the composite column serviceability-level section flexural stiffness of the composite column moment moment capacity

where No is the axial load applied on a composite specimen; Nu is the axial compressive capacity of the composite column. The value of Nu was determined by using a theoretical model, which is based on the fibre element method and described in Tao and Yu [6]. Concrete strength at the time of tests was used in the calculations. The details of each column are presented in Table 1, where D and B are the outside dimensions of circular and square members, respectively; L is the overall length of a specimen; Ds and ts are the diameter and thickness of the inner steel tube, respectively; nf is the number of FRP layers; tf is the overall nominal thickness of the wrapped FRP. In Table 1, the specimens with ‘‘C’’ and ‘‘S’’ in the specimen labels refer to columns with circular and square cross-sections, as shown in Fig. 1(a) and (b), respectively.

2.2. Material properties Tension tests were carried out to determine the material properties of the inner steel tube. The measured average yield

FRP

FRP Concrete

ro

Concrete

n nf No Nu P Pue tf ts

m s f D

e

431

axial load level ( =No/Nu) number of FRP layers axial compressive load compressive capacity of the composite column lateral load experimental lateral strength nominal thickness of FRP jacket wall thickness of steel tube ductility coefficient stress curvature lateral displacement strain

strength (fy), tensile strength (fu) and modulus of elasticity (Es) are 300, 352 and 1.8  105 MPa, respectively. The concrete mix was designed for compressive cube strength (fcu) at 28 days of 30 MPa. The mix proportions were as follows:

   

Cement: 538 kg/m3 Water: 205 kg/m3 Sand: 598 kg/m3 Coarse aggregate: 1109 kg/m3.

In the concrete mix, the fine aggregate used was silica-based sand, the coarse aggregate was carbonate stone. To measure the concrete strength, six 150 mm cubes were cast and cured in conditions similar to the FRP–concrete–steel DSTC specimens. The average cube strengths (fcu) of the concrete used to fabricate specimens at 28 days and the time of tests were 33.6 and 49.6 MPa, respectively. The modulus of elasticity (Ec) of concrete at 28 days was 27,200 MPa. Bi-directional carbon fibre sheets, which had a nominal thickness of 0.17 mm per layer, were used to fabricate the FRP. The tensile properties of the cured FRP determined from tensile tests of flat coupons according to ASTM D3039 [7], are given in Table 2, where the values presented are the averages from test coupons made from both fibre directions, calculated on the basis of the nominal thickness of 0.085 mm for the fibre sheet. The test results demonstrated that the material properties in both directions were nearly the same. The tensile properties of the epoxy as provided by the supplier are also given in Table 2. 2.3. Specimen preparation

Steel tube

Steel tube Fig. 1. A schemic view of typical cross-sections.

All inner steel tubes were made of cold-formed steel tubes. Each tube was welded to a square steel base plate 16 mm thick. Then a steel mold was installed for concrete pouring. The concrete was filled in layers and was vibrated by a poker vibrator. After

Table 1 Summary of test information. No.

Specimen label

nf

tf (mm)

D(B)  L (mm)

Ds  ts (mm)

No (kN)

Nu (kN)

n

1 2 3 4 5 6 7 8

S-0.2-1 S-0-2 S-0.2-2 S-0.5-2 C-0.3-1 C-0-2 C-0.3-2 C-0.6-2

1 2 2 2 1 2 2 2

0.17 0.34 0.34 0.34 0.17 0.34 0.34 0.34

150  1500 150  1500 150  1500 150  1500 150  1500 150  1500 150  1500 150  1500

75  2.2 75  2.2 75  2.2 75  2.2 75  2.2 75  2.2 75  2.2 75  2.2

205 18 205 410 200 14 205 410

872 891 891 891 666 683 683 683

0.23 0.02 0.23 0.46 0.3 0.02 0.3 0.6

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Table 2 Material properties of CFRP and epoxy. Material Elastic modulus (GPa)

Tensile strength (MPa)

Ultimate strain Thickness (%) (mm)

CFRP

3950

1.60

Epoxy

247 2.7

42.8

1.85

0.17 (nominal) –

MTS Actuator 100t Jack

Rigid stub

provide effective confinement along the middle 180 mm of the FRP–concrete–steel DSTC specimens with circular sections, and 150 mm of the specimens with square sections, respectively. The stub was attached with a hydraulic ram. The in-plane displacements were measured at locations along the FRP–concrete–steel DSTC specimen test-length by displacement transducers. The tests were performed under rotation controlled cyclic loading until the strength of a specimen decreased to about 50% peak load, where the typical pattern of lateral loading is shown in Fig. 3. The rotation increment was determined as 0.005 rad, and the corresponding mid-span displacement of the specimen was 3.75 mm. Therefore the cyclic loading history including three cycles at the mid-span displacement was 3.75, 7.5, 11.25 mm y. No displacements at the reaction blocks were observed until the FRP–concrete–steel DSTC specimen bowed after reaching the failure load.

Specimen

3. Experimental results and specimen behaviour

Support bar

The failure modes of all tested specimens are generally close. From the observations during the whole failure process, it was found that rupture of the longitudinal FRP jacket on one side of

0.03 0.02

that, the specimens were placed upright to air-dry. During curing, a very small amount of longitudinal shrinkage occurred at the top of the column. A high-strength epoxy was used to fill this longitudinal gap so that the concrete surface was flush with the inner steel tube at the top. Finally, the specimens were wrapped by FRP using a hand lay-up method. The finishing end of a sheet overlapped the starting end by 150 mm. The radius of chamfered angle (ro) for the square section, as shown in Fig. 1(b), is 25 mm in present study. The wrapped specimens were left to cure in the laboratory environment at room temperature for about 3 months before testing. 2.4. Cyclic test apparatus

Rotation amplitude

Fig. 2. Arrangement of test set-up.

0.01 0

-0.01 -0.02

-0.03

0

3

6 Number of Cycles

9

12

Fig. 3. Typical pattern of lateral loading.

All FRP–concrete–steel DSTC specimens were tested under combined constant axial load (No) and cyclically increasing flexural load (P). The in-plane curvatures were measured at locations along the FRP–concrete–steel DSTC specimen by two displacement transducers. Eight strain gauges mounted on the FRP sheet surface were used for each specimen to measure strains at the mid-span. Fig. 2 gives a schematic view of the test set-up. The ends of the FRP–concrete–steel DSTC specimen were attached to cylindrical bearings and were free to rotate in-plane, and thus simulating pin–pin end conditions. The axial load (No) was applied and maintained constant by a hydraulic ram. A hydraulic pump was used to control the axial load (No). The flexural loading was applied by imposing cyclically lateral loading in the middle of the FRP–concrete–steel DSTC specimen. The specimen was confined in the middle part by a rigid stub made of high-strength steel as shown in Fig. 2. The rigid stub was designed and was made of two separate halves of a box with a concentric hole that can exactly fit the FRP–concrete–steel DSTC specimen. The two halves were pushed against the specimen and connected together using high-strength bolts. The stub may

Longitudinal FRP rupture Circumferential FRP rupture Crushed concrte

Local inward indent

Fig. 4. Typical failure modes of the tested specimens.

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80

80

Longitudinal FRP rupture

0 -40 -80 -50

0 -40 Longitudinal FRP rupture

Longitudinal FRP rupture -25

0

25

-80 -50

50

25

50

Longitudinal FRP rupture 40

P (kN)

P (kN)

0

80

Logitudinal FRP rupture

40 0 -40 -80 -50

-25

(mm) S-0-2 (n =0.02, nf =2)

(mm) n S-0.2-1 ( =0.23, nf=1) 80

Longitudinal FRP rupture

40

P (kN)

P (kN)

40

433

0 -40 Longitudinal FRP rupture

Longitudinal FRP rupture -25

0

25

50

(mm) S-0.2-2 (n =0.23, nf =2)

-80 -50

-25

0

25

50

(mm) S-0.5-2 (n =0.46, nf =2)

Fig. 5. Cyclic load (P) versus lateral displacement (D) curves for square specimens.

the stub occurred at the peak load, accompanied with a sudden drop in the lateral load capacity. After that, the rupture of the longitudinal FRP fibres developed rapidly toward the neutral axis of the specimen section with increasing mid-span displacement. Then rupture of the circumferential FRP fibres along the longitudinal direction was also observed due to the volume dilation of the concrete caused by the reverse loading, shown in Fig. 4(a). Obvious flexure cracks of the concrete could be observed at the location where the FRP jacket ruptured. Fig. 4(b) shows the crushed concrete after the FRP sheet was stripped. It was observed that the rupture process of circumferential FRP fibres seems more sluggish than those of longitudinal FRP fibres. After the sudden drop of the specimen strength, the lateral load tended to keep stable at a lower level. With the lose of the functionality of the FRP jacket at this stage, only the concrete and the inner steel tube could sustain the lateral load. During the whole test, no debonding between the concrete and FRP jacket was observed. After the test, local buckling was observed for the inner steel tube. Fig. 4(c) shows the inner steel tube after the knock off of its outer concrete. It was found that the inner tube shown an inward indent near the mid-span due to the support from the around concrete. The test curves of lateral load (P) versus mid-span displacement (D) for square and circular specimens are shown in Figs. 5 and 6, respectively, and the points for the rupture of the longitudinal FRP jackets are also marked on these figures. It can be seen from Figs. 5 and 6 that, the rupture of the longitudinal FRP fibres led to a sudden drop in strength due to the fact that these fibres were located outside the cross-section and thus could undertook a big proportion of the bending moment. After that, the P–D curve shows a relatively stable

descending trend, resulting from the sluggish rupture of the circumferential FRP fibres, the development of the concrete cracking and the gradual yielding of the inner steel tube. The maximum lateral loads (Pue) obtained in the tests are presented in Table 3. Specimens S-0.2-2 (square section, axial load level of 0.23, 2 layers of FRP) and C-0.2-2 (circular section, axial load level of 0.3, 2 layers of FRP) are selected to demonstrate the typical response of the measured moment (M) versus curvature (f) graph, shown as in Fig. 7(a) and (b), respectively. In general, there is an initial elastic response, followed by inelastic behaviour with gradually decreasing stiffness, until the plastic bending moment capacity is reached asymptotically. A careful examination of the test results reveals that, in general, the M f relations enter the inelastic stage at about 20% of the moment capacity (Mu). Consequently, the initial cross-section flexural stiffness (Ki) is defined as the secant stiffness corresponding to a bending moment of 0.2Mu. The M f relations can also be used to determine the serviceability-level flexural stiffness (Ks) of the FRP–concrete–steel DSTC cross-section, which is defined as the secant stiffness corresponding to the serviceability-level bending moment of 0.6Mu [8]. The points corresponding to 0.2Mu, 0.6Mu and Mu in both positive and negative directions were labelled in Fig. 7, respectively. Table 3 gives the initial cross-section flexural stiffness (Ki) and serviceabilitylevel cross-section flexural stiffness (Ks) of the tested specimens. It can be found that Ki and Ks increase with the increasing number of FRP layers (nf) for the specimens with the same axial load level (n). Ki increased by about 30% both for square and circular specimens when nf increased from 1 to 2, whilst Ks increased by about 40% and 34%, respectively, for square and circular sections. This is attributed to the improved flexural stiffness by the increased longitudinal FRP, as well as the somewhat enhancing confinement effect resulting from the

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60

60

Longitudinal FRP rupture

30

P (kN)

P (kN)

30 0 -30

Longitudinal FRP rupture

0 -30

Longitudinal FRP rupture

Longitudinal FRP rupture -60 -50

-25

0

25

-60 -50

50

-25

(mm) C-0.3-1 (n =0.3, nf =1) 60

25

50

60

Longitudinal FRP rupture

Longitudinal FRP rupture

30

30

P (kN)

P (kN)

0

(mm) C-0-2 (n =0.02, nf =2)

0 -30

0 -30

Longitudinal FRP rupture -60 -50

-25

0

25

50

-60 -50

(mm) C-0.3-2 ( n =0.3, nf =2)

Longitudinal FRP rupture -25

0

25

50

(mm)

C-0.6-2 (n =0. 6, nf =2)

Fig. 6. Cyclic load (P) versus lateral displacement (D) curves for circular specimens.

Table 3 Measured results. No.

Specimen label

Dy (mm)

Du (mm)

Pue (kN)

Mue (kN m)

Ki (kN m2)

Ks (kN m2)

Ductility coefficient m

1 2 3 4 5 6 7 8

S-0.2-1 S-0-2 S-0.2-2 S-0.5-2 C-0.3-1 C-0-2 C-0.3-2 C-0.6-2

9.05 10.7 8.4 7.9 8.1 9.95 8.5 8.05

23.6 21.75 20.55 22.4 22.35 20.05 21.5 25.3

55.8 50.9 62.6 66.3 43 42.4 51.7 55

24.6 19.8 27 32.1 19.65 16.07 22.94 29.45

1107 601 1442 1640.5 984.7 549.5 1280.3 1422.2

688.5 438.8 973 1175 717.9 399.6 958.8 1016.7

2.61 2.03 2.45 2.84 2.76 2.02 2.53 3.14

30

Mu

0.6Mu

15

M (kN.m)

M (kN.m)

15

0.2Mu 0

-0.2Mu -15 -30 -0.06

0.2Mu -0.2Mu -0.6Mu

4. Analysis of test results and discussions

-Mu

-Mu 0 0.03 (1/m) S-0.2-2 ( n =0.23,nf =2)

0

0.6Mu

-15

-0.6Mu

-0.03

additional circumferential FRP. Meanwhile, Ki and Ks also increase with the increase of the axial load level (n) for the specimens with the same FRP layers (nf), due to the fact that, within the current testing limitations, the increased compressive stress would enlarge compressive area and restrain the development of the cracks.

30

Mu

0.06

-30 -0.08

-0.04

0 0.04 (1/m) C-0.3-2 (n =0.3,nf =2)

Fig. 7. Typical moment (M) versus curvature (f) relations.

0.08

4.1. Effects of axial load level and number of FRP layers Fig. 8 shows the influence of axial load level (n) on the lateral load (P) versus lateral displacement (D) envelope curves of the FRP–concrete–steel DSTC specimens. It can be found from this

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80 S-0.2-1 (nf=1)

P (kN)

40

S-0.2-2 (nf=2)

0 -40 -80 -50

-25

0 25 (mm) Square specimen (n=0.23)

50

60 C-0.3-1 (nf=1)

30 P (kN)

figure that, the influences of the axial load level (n) on both square and circular members are generally similar. In general, the elastic stiffness of the specimen increases with the increase in axial load level (n). The P versus D envelope curves exhibit a sudden drop after the peak load (Pue) due to the rupture of the FRP jackets. The residual strength for a specimen with an axial load level n of 0.02 is lower than that for the specimen with higher axial load level (n). This is owing to the fact that the existence of the axial compressive load may be effective in restraining the concrete cracking and in suppressing the rupture of the FRP jacket to some extent. Fig. 9 shows the influence of number of FRP layers (nf) on the lateral load (P) versus lateral deflection (D) envelope curves. It was found that, in general, the number of the FRP layers has little influence on the elastic stiffness, whilst the elastic-plastic stiffness of the specimen increases with the increasing of nf. For square specimens, the residual strength of the specimen S-0.2-1 with nf =1 is generally close to that of the specimen S-0.2-2 with nf = 2, shown in Fig. 9(a). However, for circular sections, the residual strength of the specimen C-0.3-2 with nf =2 is higher than that of the specimen C-0.3-1 with nf = 1, as shown in Fig. 9(b). This is attributed to the fact that more fibres concentrated near the neutral axis for a circular section than for a square section. These fibres kept intact during the unloading stage and made an obvious contribution to the residual strength for circular members. Fig. 10 shows the relationship between ultimate strength (Pue) versus axial load level (n). It can be concluded that, in general, the

435

C-0.3-2 (nf=2)

0 -30

80

P (kN)

40

-60 -50

n=0.02 n=0.23 n=0.46

0 25 (mm) Circular specimen (n=0.3)

50

Fig. 9. Influence of number of FRP layers (nf) on P  D envelope curves.

0

-40

-80 -50

-25

0 25 (mm) Square specimen (nf = 2)

50

60

30

P (kN)

-25

n=0.02 n=0.3 n=0.6

4.2. Axial shortening and longitudinal strain

0

-30

-60 -50

-25

0 25 (mm) Circular specimen (nf =2)

ultimate lateral strength (Pue) increases with the increase in axial load level (n). Once again, the reason is that the axial compression load can restrain the crack development of the concrete and postpone the rupture of the FRP jackets. Fig. 11 shows the relationship between ultimate strength (Pue) versus number of FRP layers (nf). It can be found that, in general, the ultimate lateral strength (Pue) increases with the increasing number of FRP layers (nf). The number of FRP layers (nf), however, has a more significant effect on the ultimate strength (Pue) for the circular members than that for the square members, shown in Fig. 11. Compared with columns wrapped with one layer of FRP, the strength enhancement are 20.2% and 12.2% for the circular and square columns, respectively, strengthened with two layers of FRP.

50

Fig. 8. Influence of axial load level (n) on P  D envelope curves.

Fig. 12 shows the typical response of the measured axial shortening (d) of the tested FRP–concrete–steel DSTC specimens through progressive loading cycles (D/Dy). It can be seen that, the axial shortening (d) keeps increasing with the increasing of the loading cycles, though the axial load (No) is kept constant. This phenomenon is expected due to the continuous damage of the concrete rigidity under the reverse lateral loading. For a FRP– concrete–steel DSTC with square section, d increased in a smooth way during the elastic cycles and then started to increase significantly. For a FRP–concrete–steel DSTC specimen with circular section, the shortening (d) was noticed to be more gradual. It can also be found that the total axial deformation for

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80

nf =2

n=0.23

60

Pue (kN)

Pue (kN)

n=0.46 n=0.02

nf =1 40 20 0

0

0.15

0.3 0.45 n Square specimen (nf =2)

0

0.6

2 nf Square specimen (n=0.23)

3

60

60

nf =2

n=0.6 45

45

n=0.3

Pue (kN)

Pue (kN)

1

n=0.02 30

nf =1 30 15

15

0 0

0

0.15

0.3 0.45 n Circular specimen (n f=2)

0.6

Fig. 10. Relationship between lateral strength (Pue) and axial load level (n).

4.3. Rigidity degradation Column flexural stiffness (Rk) is used to assess the rigidity degradation of the FRP–concrete–steel DSTC specimens with

1

2 nf Circular specimen (n=0.3)

3

Fig. 11. Relationship between lateral strength (Pue) and number of FRP layers (nf).

increasing lateral deflections. Rk can be given as Rk ¼

the circular column is larger than that of its square courterpart, which demonstrates the fact that the former has higher residual strength, and can resist the axial load much longer than the latter. Specimen C-0.3-2 is selected to demonstrate the typical response of the measured lateral load (P) versus the measured fibre longitudinal strain (e) for the circular and the square members, shown in Fig. 13. The presented strains in Fig. 13 are defined as tension corresponding to positive values and compression to negative values, respectively. It should be noted that no strain reading was available after the rupture of FRP at peak load. Therefore, no softening branch was shown in Fig. 13. It can be found from this figure that, generally, the development of the longitudinal strains (e) is not symmetrical. The tensile strain is higher than the compressive strain at a same load level. This is similar to the feature of axially compressive FRP-confined RC column described earlier by Teng et al. (2002) [1]. It was also found that the measured rupture strains of the tensile fibres for all specimens are below 0.01, and lower than the ultimate tensile strain (ef = 0.016) measured from material tests.

0

EI ðEIÞD ¼ 0

ð2Þ

in which, EI is the measured flexural stiffness of the specimens corresponding to different lateral deflection (D). The deflection of a beam–column loaded with a concentrated lateral load at mid-span, including the second-order effect, can be given by the following [9]:   PL3 3ðtan uuÞ ð3Þ D¼ 3 48EI u where rffiffiffiffiffiffiffiffiffiffiffi 1 No L2 u¼ ; EI 2 in which, EI can be obtained by substituting the measured peak load (Pp) at different loading cycles (D/Dy) and the corresponding deflection (Dp) into Eq. (3) and repeating iterations for it. It should be noted that Eq. (3) was derived based on a structural mechanics model [9]. FRP–concrete–steel DSTC specimens with circular sections were selected to illustrate the rigidity degradation. Fig. 14(a) shows the effect of axial load level (n) on the rigidity degradation. Generally, there is obvious rigidity degradation for the composite columns subjected to cyclic loading. From this comparison, it can be found that the rigidity degradation for members with higher axial load level (n) was less significant compared with specimens with lower axial load level, especially after the FRP

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1.2

4

2

C-0-2(n=0.02) C-0.3-2(n=0.3) C-0.6-2(n=0.6)

0.8

Rk

y

/

437

0

0.4 -2

FRP rupture 0

-4 0

1

2

3

4

0

1

2

3

5

4

y

d (mm)

Axial load level ( n ), nf = 2

Square specimen (S-0.5-2, n=0.46, nf =2) 6

1.2

3

0.8

C-0.3-1(nf=1)

/

y

Rk

C-0.3-2(nf=2)

0

0.4

-3

0

0

-6 0

1

2

3

4

1

2

5

3

4

y

d (mm)

Number of FRP layers (nf ), n=0.3

Circular specimen (C-0.6-2, n=0.6, nf =2)

Fig. 14. Effects of axial load level (n) and number of FRP layers (nf) on the rigidity degradation.

Fig. 12. Axial shortening (d).

50

P Pue

P (kN)

25

0.85Pue Py

0

-25

-50 -8000

-4000

0

4000

O

8000

y

max

u

Fig. 15. Envelope of cyclic lateral load (P) versus lateral deflection (D) response. Fig. 13. Typical lateral load (P) versus FRP longitudinal strain (e) relations.

It can be expressed as [10]

rupture occurred. This phenomenon can also be explained by the fact that higher axial load level can restrain the development of concrete cracking and suppress the further rupture process of the FRP jackets to some extent. It was found that nf has moderate influence on the rigidity degradation for the composite specimens within the scope of the current tests. Fig. 14(b) shows the effect of number of FRP layers (nf) on the rigidity degradation.

4.4. Ductility and dissipated energy For convenience of analysis, ductility coefficient (m) is used to quantify the ductility of the FRP–concrete–steel DSTC specimens.

m¼

Du Dy

ð4Þ

where Dy is the yielding displacement, and Du is the displacement when the axial load falls to 85% of the ultimate strength (Pue), shown in Fig. 15. The ductility coefficients (m) so determined are presented in Table 3. It can be seen from Table 3 that, the ductility coefficients (m) for the specimens in current tests vary from 2.02 to 3.14. The average ductility coefficient (m) for the circular members is 2.61 whilst that for the square members is 2.48. As can be seen, m for a FRP–concrete–steel DSTC specimen is generally lower than that of a concrete-filled steel tubular (CFST) column [10,11], which is usually expected to exceed 5.5 for circular section and 4 for

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4

4

3

3

n=0.6

n=0.46 n=0.3

n=0.23

2

2

n=0.02

n=0.02

1

1 0

0.2

0.4

0

0.6

0.2

0.4

0.6

0.8

n (2) Circular specimens ( n f =2)

n (1) Square specimens (n f =2)

Axial load level (n ) 4

4

3

3

nf=1

nf=1

nf=2

nf=2 2

2

1

1 0

1

2

3

0

1

2

3

nf nf (2) Circular specimens (n=0.3) (1) Square specimens (n=0.23) Number of FRP layers ( nf ) Fig. 16. Effects of axial load level (n) and number of FRP layers (nf) on the ductility coefficient (m).

square section, respectively [10,11]. This is conceivable since the FRP jacket locates at the periphery of the cross-section and bears a relatively high proportion of the lateral load. Therefore, once the longitudinal FRP jacket ruptured, the lateral load would mainly be undertaken by the concrete and inner steel tube, which results in the significant drop in the lateral load capacity. It is thus suggested that, if this type of composite columns are to be utilised in an earthquake-prone region, it is necessary to limit the load-bearing proportion of the FRP jackets for the purpose of improving the earthquake-resistant ability of the composite columns. Fig. 16(a) shows the relationship between the ductility coefficient (m) and the axial load level (n). It can be found that, m increases with the increase in axial load level (n). As can be seen, this trend for the FRP–concrete–steel DSTC specimens is much different from that occurred for the normal RC, CFST [10,11], or concrete filled double skin steel (CFDST) [12] column, which all exhibit a reduced ductility behaviour under a higher axial load level (n) generally. This is mainly attributed to the fact that a higher axial load level (n) would improve the compressive stress and thus postpone the rupture of the FRP jacket to some extent. Fig. 16(b) shows the relationship between the ductility coefficient (m) and the number of FRP layers (nf). A careful examination of the test results revealed an interesting phenomenon that the number of FRP layer (nf) has a moderate influence on the ductility in the current tests. In general, the specimen with two layers of FRP shown a slightly lower ductility than that of the specimen wrapped with only one layer of FRP. As

expected, longitudinal FRP will bear higher lateral load with the increase of FRP layers, thus leading to a more significant drop in load when the rupture of the longitudinal FRP fibres occurs. However, since bi-directional fibre sheets were utilised, it is expected that the increasing confinement from the FRP jacket may partially counteract the effect of longitudinal FRP fibre rupture. Energy dissipation ability is one of the most important indices to evaluate the seismic performance of a member or structure. The dissipated energy in each cycle is calculated from the lateral load (P) versus lateral displacement (D) curve as the area bound by the hysteretic hoop of that cycle. Fig. 17 shows the relationship of cumulative energy dissipation (E) versus relative lateral displacement (D/Dy). It can be found that, generally, cumulative energy dissipation (E) under the same incremental lateral displacement (D/Dy) increases with the increase of the axial load level (n), as shown in Fig. 17(a). But the number of FRP layers (nf) has a moderate influence on the E versus D/Dy curves, as shown in Fig. 17(b).

5. Conclusions The following observations and conclusions can be drawn based on the current research reported in the paper. (1) Generally, FRP–concrete–steel DSTC specimens exhibit high levels of energy dissipation prior to the rupture of the longitudinal FRP, whilst they show a sudden drop in the lateral load capacity after that.

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the contribution of the FRP with respect to the overall strength will decrease accordingly. However the increasing dimensional size of the inner steel tube will result in a higher sectional slenderness ratio, which is likely to cause a serious local buckling for it under the reverse lateral loading. Therefore, further work should be conducted to develop a numerical model and carry out a parameter study in order to obtain a reasonable parameter for each component of the FRP–concrete–steel DSTC.

8

C-0-2 (n=0.02) C-0.3-2 (n=0.3) C-0.6-2 (n=0.6)

E (kN.m)

6

4

2

Acknowledgements

0

0

1

2

3

4

y

8

C-0.3-1(nf=1) C-0.3-2(nf=2)

E (kN.m)

6

2

0

1

2

The study of this paper is financially supported by the Research Foundation of the Ministry of Railways and Tsinghua University (RFMOR & THU) (no. J2008G011), the National Basic Research Program of China (973 Program) (grant no. 2009CB623200) and the Program for Changjiang Scholars and Innovative Research Team in University (IRT00736). The financial support is highly appreciated.

References

4

0

439

3

4

y Fig. 17. Effects of axial load level (n) and number of FRP layers (nf) on energy dissipation ability (E).

(2) The seismic performance of the FRP–concrete–steel DSTC can be improved to some extent due to the existence of the axial compressive load within the scope of the current tests. (3) The increase of the number of FRP layers can improve the ultimate strength of the FRP–concrete–steel DSTCs, but has moderate influence on the ductility and energy dissipation. It should be noted that, based on the experimental research carried out in this paper, it seems that higher ductility performance may be required for the composite columns in some cases to reduce the impact of the sudden rupture of FRP. To achieve this goal, a larger steel tube can be used and, therefore,

[1] Teng JG, Chen JF, Smith ST, Lam L. FRP-strengthened RC structures. John Wiley & Sons, Ltd; 2002. [2] Tao Z, Han LH, Wang LL. Compressive and flexural behaviour of CFRP repaired concrete-filled steel tubes after exposure to fire. Journal of Constructional Steel Research 2007;63(8):1116–26. [3] Teng JG, Yu T, Wong YL, Dong SL. Hybrid FRP–concrete–steel tubular columns: concept and behavior. Construction and Building Materials 2007;21(4):846–54. [4] Teng JG, Yu T, Wong YL. Hybrid FRP–concrete–steel double-skin tubular columns: stub column tests. In: The second international conference on steel & composite structures, Seoul, Korea, 2004. p. 1390–400. [5] Yu T, Wong YL, Teng JG, Dong SL, Lam ESS. Flexural behavior of hybrid FRP– concrete–steel double-skin tubular members. Journal of Composites for Construction ASCE 2006;10(5):443–52. [6] Tao Z, Yu Q. New types of composite columns—experiments, theory and methodology. Beijing: China Science Press; 2006. [in Chinese]. [7] ASTM D3039/D3039M. Standard test method for tensile properties of polymer matrix composite material. In: Annual book of ASTM standards. West Conshohocken, PA: American Society for Testing and Materials; 2000. [8] Varma AH, Ricles JM, Sause R, Lu LW. Seismic behavior and modeling of highstrength composite concrete-filled steel tube (CFT) beam-columns. Journal of Constructional Steel Research 2002;58(5–8):725–58. [9] Elremaily A, Azizinamini A. Behavior and strength of circular concrete-filled tube columns. Journal of Constructional Steel Research 2002;58(12): 1567–91. [10] Han LH. Concrete filled steel tubular columns—theory and practice (second version). Beijing: Science Press; 2007. [in Chinese]. [11] Han LH, Yang YF, Tao Z. Concrete-filled thin walled steel RHS beam-columns subjected to cyclic loading. Thin-walled Structures 2003;41(9):801–33. [12] Han LH, Huang H, Tao Z, Zhao XL. Concrete-filled double skin steel tubular (CFDST) beam-columns subjected to cyclic bending. Engineering Structures 2003;28(12):1698–714.