Behaviour of FRP-confined compound concrete–filled circular thin steel tubes under axial compression

Abstract

This article presents test results of a recent study on the axial compressive behaviour of fibre-reinforced polymer–confined compound concrete–filled thin steel tubes. The usage of compound concrete, which is a mixture of fresh concrete and large pieces of recycled concrete lumps, can recycle waste concrete in a simple but effective way. Totally, three series of tests were conducted, with the parameters including the relative strength between fresh concrete and recycled concrete lumps, the volumetric percentage (i.e. mix ratio) of recycled concrete lumps, the diameter-to-thickness ratio of the steel tubes, and the thickness of the fibre-reinforced polymer jackets being investigated. The stress–strain curves of the steel tube and compound concrete core were derived and the effects of different parameters were then examined and discussed. An existing stress–strain curve model of fibre-reinforced polymer–confined normal concrete-filled steel tubes was also found performing well in predicting the behaviour of fibre-reinforced polymer–confined compound concrete-filled steel tubes.

Keywords

confinement fibre-reinforced polymer recycled concrete lumps recycling stress–strain behaviour

Introduction

Due to the shortage of natural aggregates, many researchers have investigated the recycling of waste concrete (Khalaf and DeVenny, 2004) through adoption of recycled coarse aggregates (RCAs; He et al., 2015; Xiao et al., 2012; Zhao et al., 2015). It is well understood that there is a thin layer of mortar adhered on the surface of RCAs, leading to a large water absorption, a lower specific gravity, a poor workability, and a weakened interfacial transition zone (ITZ; Poon et al., 2004; Zhao et al., 2015). As a result, the performance of recycled aggregate concrete (RAC) generally decreases with the increase in the mix ratio of RCA (i.e. the ratio between the weight of RCA and the total weight of coarse aggregates).

Using large pieces of recycled concrete lumps (RCLs) to fabricate RAC has been proposed by Wu et al. (2008), in whose study, the waste concrete was crushed into large pieces instead of small-sized aggregates. The RCLs were then dropped into casting mould and mixed with fresh concrete (FC) to form a mixture of FC and RCLs. Note that this mixture of FC and RCLs has subsequently been referred to as compound concrete by Teng et al. (2016). Associated with RCLs, the cost of recycling concrete can be reduced and the recycling rate can be increased as both aggregates and mortar in the waste concrete can be recycled. This recycling method is especially suitable to be used together with an outer confining tube (i.e. steel tube, fibre-reinforced polymer (FRP) tube), where there is no internal steel cage which may lead to a risk of congestion of RCLs.

Extensive research has been conducted by Wu’s group (Wu et al., 2013, 2014, 2015, 2016, 2018a, 2018b, 2018c, 2019; Wu and Jin, 2019; Zhao et al., 2016) in the previous decade, demonstrating the effectiveness of this recycling method. The Wu’s group research is mainly focused on compound concrete–filled steel tubes. Based on their research, the following suggestions were given (Wu et al., 2016): (1) the maximum volumetric ratio of RCLs can be 30%–35% and (2) the maximum size of RCLs should not exceed half of the minimum dimension of the specimen cross section. Teng et al. (2016) investigated the behaviour of FRP-confined compound concrete under axial compression, and they concluded that the negative effect of RCLs can be largely offset by providing sufficient lateral confinement. Recently, Yu et al. (2019) further developed the idea of compound concrete, where the FC is replaced by a calcium sulfoaluminate (CSA)-based cementitious material (or high-water material) which can easily fill the voids among RCLs without mechanical vibration.

Many studies have been conducted on FRP-confined normal concrete and RAC and FRP-confined normal concrete–filled steel tubes (e.g. Hu et al., 2011; Teng et al., 2013; Wang et al., 2019; Xiao et al., 2005; Yu et al., 2014). However, behaviour of FRP-confined compound concrete–filled steel tubes under compressive loadings has never been studied. In lights of research demands, experimental programme on compound concrete–filled FRP-confined thin steel tubes under axial compression was conducted in the present study to look at the axial compressive behaviour of FRP-confined compound concrete–filled circular thin steel tubes. The parameters investigated include the relative strength between FC and RCLs, the volumetric percentage (i.e. mix ratio) of RCLs, the diameter-to-thickness ratio of steel tubes, and the thickness of the FRP jacket. The test results are carefully interpreted and the performance of an existing stress–strain model for FRP-confined normal concrete filled steel tube is assessed.

Experimental programme

RCLs

The RCLs used in this study were produced from waste concrete of two different strength grades. The first batch of RCLs, referred to as RCLs-1, was produced by crushing some reinforced concrete beams (Figure 1(a)) used for teaching purpose in the structural lab of Wenzhou University. The age of the beams was about 7 years at the time of crushing. Before crushing, six concrete cores, with diameter × height = 75 mm × 75 mm, were taken out from the three beams and tested under axial compression. The average compressive strength and the standard deviation of the concrete cores were 34.2 and 3.83 MPa, respectively. The cube compressive strength of the waste concrete is taken to be the same as core compressive strength (JGJ/T 384, 2016). The second batch of RCLs, referred to as RCLs-2, was produced by crushing some 150 mm concrete cubes cast in advance. The average cube compressive strength was 48.8 MPa and the concrete age was about 5 months at the time of crushing.

Figure 1.

Recycled concrete lumps (RCLs): (a) waste concrete beams, (b) photograph of RCLs, and (c) size distributions of RCLs.

Both batches of waste concrete were crushed manually, and only RCLs with sizes between 30 and 80 mm were adopted in the present study (Figure 1(b)). It should be clarified that the waste concrete was produced by small-size aggregate (i.e. <20 mm); as a result, even RCLs of 30 mm would contain more than one aggregate from waste concrete. The size distributions of the RCLs are shown in Figure 1(c). The densities of RCLs-1 and RCLs-2 at dry condition are 1900 and 2100 kg/m³, respectively. The water absorption of the RCLs was tested by immersing the RCLs into water for 24 h and then dried in oven at 105°C for another 24 h. The water absorption was decided as the ratio between the weight loss and the dry weight of RCLs, which were 9.33% and 6.25% for RCLs-1 and RCLs-2, respectively.

Test matrix

The tests were conducted in three batches, referred to as Batches A, B, and C, as summarized in Table 1. Three different types of steel tubes (i.e. A, B, and C) were adopted in three batches, respectively. The diameter and height of all specimens were 255 and 510 mm, respectively, leading to a height-to-diameter ratio of 2. In each batch, except for specimens with bare steel tube, two different FRP jackets were adopted by applying one or two layers of carbon FRP (CFRP) jacket around the steel tube via the wet layer-up process. For each tube, two different mix ratios of RCLs (i.e. 0% and 30%) were adopted. The mix ratio is defined as the ratio of the volume of RCLs and the total volume of compound concrete. For all cases, two nominally identical specimens were prepared to check the repetition of the tests and the reliability of the test results. As a result, there were 12 specimens in each batch and 36 specimens in total. Of them, 12 are confined by bare steel tubes and the other 24 are confined by FRP-confined steel tubes. The differences among Batches A, B, and C lie in the steel tube adopted and the relative strength between FC and RCLs. The outer diameter ( $D_{o}$ ) and thickness ( $t_{s}$ ) of the steel tubes are given in Table 2. The diameter-to-thickness ratios ( $D_{o} / t$ ) of the steel tubes were 100, 185, and 113 for Batches A to C, respectively. In Batches A and B, RCLs-1 were adopted, together with FC with a higher strength grade (i.e. ∼55 MPa). In Batch C, RCLs-2 were adopted together with FC with a lower strength grade (i.e. ∼35 MPa). In this study, RCLs and FC of different relative strengths were used together to investigate the effect of strength difference between RCLs and FC, as in real project the strength of RCLs may not be the same as that of FC.

Table 1.

Specimen details.

Specimen	Steel tube	f_cu of FC (MPa)	f_cu of RCLs (MPa)	CFRP layers	Mix ratio (%)
SA-C0-R0-1,2	A	56.3	34.2	0	0
SA-C0-R30-1,2				0	30
SA-C1-R0-1,2				1	0
SA-C1-R30-1,2				1	30
SA-C2-R0-1,2				2	0
SA-C2-R30-1,2				2	30
SB-C0-R0-1,2	B	54.3	34.2	0	0
SB-C0-R30-1,2				0	30
SB-C1-R0-1,2				1	0
SB-C1-R30-1,2				1	30
SB-C2-R0-1,2				2	0
SB-C2-R30-1,2				2	30
SC-C0-R0-1,2	C	35.4	48.8	0	0
SC-C0-R30-1,2				0	30
SC-C1-R0-1,2				1	0
SC-C1-R30-1,2				1	30
SC-C2-R0-1,2				2	0
SC-C2-R30-1,2				2	30

FC: fresh concrete; RCL: recycled concrete lump; CFRP: carbon fibre–reinforced polymer.

Table 2.

Properties of steel tubes.

Tube	Outer diameter $D_{o}$ (mm)	Thickness $t_{s}$ (mm)	$D_{o} / t_{s}$	Elastic modulus $E_{s}$ (GPa)	Yield stress $f_{y}$ (MPa)	Ultimate strength $f_{u}$ (MPa)	Strain at hardening $ε_{sh}$ (%)	Strain at ultimate strength $ε_{su}$ (%)
A	255	2.54	100	196.7	272.7	375.1	2.6	12.0
B	254	1.37	185	200.0	302.5	371.0	2.5	16.5
C	255	2.25	113	209.9	284.0	371.2	3.0	15.0

Each specimen is given a name according to the following rule: (1) two letters (i.e. SA, SB, SC) indicating the type of steel tube adopted; (2) the letter ‘C’ followed by a digital number (i.e. 0, 1, or 2) indicating the number of CFRP layers; (3) the letter ‘R’ and a number (i.e. 0 or 30) indicating the volumetric percentage (i.e. mix ratio) of RCLs; and (4) a digital number (i.e. 1 or 2) to differentiate between two nominally identical specimens. For example, specimen ‘SB-C1-R30-2’ refers to the second specimen adopting steel tube ‘B’ and confined by a one-layer CFRP jacket, which is filled with compound concrete incorporating 30% of RCLs.

Material properties

As discussed in section ‘RCLs’, two different RCLs (RCLs-1 and RCLs-2) were adopted, with their original compressive strength being 34.2 MPa (75 mm × 75 mm core) and 48.8 MPa (150 mm cube strength), respectively. For Batches A and B, FC of the same mix proportion was used, and the 28-day cube compressive strength was 56.3 and 54.3 MPa, respectively. For Batch C, the 28-day cube compressive strength of the FC was 35.4 MPa. The compressive strength of RCLs and FC concrete is also given in Table 1.

All the steel tubes were manufactured by welding thin mild steel sheet along its longitudinal direction, as seamless steel tubes available in the market generally possess a much smaller diameter-to-thickness ratio (i.e. $D_{o} / t_{s} < 40$ ). Three steel coupons were cut from each type of steel sheet and tested according to BS EN ISO 6892-1 (2009). The mechanical properties of the steel sheet based on coupon tests were summarized in Table 2, including the elastic modulus ( $E_{s}$ ), yield stress ( $f_{y}$ ), ultimate tensile strength ( $f_{u}$ ), strain at the initiation of hardening ( $ε_{sh}$ ), and strain ( $ε_{su}$ ) at ultimate strength. The strain corresponding to strain hardening is generally above 2.5%.

All the FRP tubes were formed by wrapping unidirectional CFRP sheet around the hoop direction of the steel tubes. The nominal thickness of the CFRP sheet is 0.15 mm. Five FRP coupons were prepared and tested following ASTM D7565 (2010). The average elastic modulus ( $E_{frp}$ ) and tensile strength ( $f_{u, frp}$ ) of the FRP sheet were 257 GPa and 4050 MPa, respectively.

Preparation of specimens

The specimens were prepared by the following steps: (1) fabrication of steel tubes using automatic arc welding. (2) For specimens with one or two layers of CFRP, the CFRP jackets were formed by wrapping continuous unidirectional CFRP sheet around the hoop direction of the steel tubes via the wet layer-up process. The finish end of the CFRP overlapped the start end by one-fourth of the circumference of the steel tube (i.e. 200 mm), while the middle of the overlapping zone coincided with the welding seam of the steel tube. In addition, one additional layer of CFRP strip was applied near the 30 mm end of the tubes to prevent unwanted end failure there. (3) The steel tubes with or without FRP tube were protected by a plastic film and fixed to a flat base plate before casting of concrete. (4) For specimens with RCLs, the RCLs were measured and saturated in water for 24 h before casting. The surface of RCLs was wiped by a cloth before casting, so that the RCLs were used in a saturated surface dried condition. During casting, a thin layer of FC was first placed at the bottom of the tubes. Then, FC and RCLs were placed into the tubes simultaneously and a needle vibrator was used to consolidate the compound concrete. At the top end of the specimens, the surface of the concrete core was slightly lower than that of steel tube, which would be levelled by a thin layer of high-strength gypsum plaster before testing. (5) All the specimens were cured in room temperature for 28 days and then tested.

Test set-up and loading scheme

For each specimen, four pair of foil type strain gauges was attached at the mid-height section of the columns, which were uniformly distributed along the hoop direction. The first pair of strain gauges lies at the middle of the overlapping zone (i.e. at the same position with welding seam of steel tube) and the other three pairs all lie outside the overlapping zone (Figure 2(a)). A compressometer was attached to the middle part of the columns, with four linear variable displacement transducers (LVDTs) installed on it to measure the axial shortening of the column in the 250-mm mid-height region. Another four LVDTs were used to measure the total shortening of the columns. The positions of LVDTs coincide with that of strain gauges in hoop direction (Figure 2(b)).

Figure 2.

(a) Arrangement of strain gauges and (b) test set-up.

All the tests were conducted on a loading machine with capacity of 5000 kN. After the pre-loading, load-controlled loading scheme was adopted first, followed by displacement-controlled loading scheme. A data log was used to collect the readings of strain gauges and LVDTs.

Test results and analysis

General behaviour and failure modes

During the test, three specimens (i.e. SA-C0-R30-1, SB-C0-R30-1, and SC-C0-R30-1) failed by mistake and were excluded from the following discussions. In this section, compression loads, stresses, and strains were taken as positive and tensile components were taken as negative.

For specimens without FRP jacket, the failure mode is elephant-foot buckling near one end of the specimen (Figure 3(a)). For specimens with FRP jacket (except specimen SA-C2-R30-1), the failure was generally caused by rupture of FRP in the middle zone, and the buckling of steel tube was generally not observed. Specimen SA-C2-R30-1 failed by debonding of FRP jacket in the overlapping zone. The photos of typical specimens after test are shown in Figure 3.

Figure 3.

Specimens after tests: (a) SA-C0-R30-1, (b) SB-C1-R30-2, and (c) SC-C1-R30-2.

Key test results are summarized in Table 3, including the peak load (P_u), ultimate hoop strain of FRP jacket (ε_{h, u}), peak axial stress (f_cc), and corresponding axial strain (ε_cc) of compound concrete core. The peak load (P_u) was the maximum load recorded by the loading machine. The ultimate hoop strain (ε_{h, u}) of FRP jacket was taken as the maximum average strain recorded by three hoop strain gauges outside the overlapping zone (i.e. SG6-8 in Figure 2(a)). For some specimens (e.g. SA-C1-R0-1, SA-C1-R0-2, SA-C2-R0-1), one or more hoop strain gauges were damaged well before the rupture of FRP jacket. In this case, the ultimate hoop strain shown in Table 3 may be substantially lower than the actual rupture strain of FRP jacket, and the peak load was taken as the maximum load before the damage of hoop strain gauges. The peak axial stress (f_cc) of compound concrete core was obtained by the method discussed later in section ‘Stress–strain behaviour of the concrete core’. The axial strains of concrete core are assumed to be the same as the axial strains of the outer steel tube which were based on the four mid-height LVDTs.

Table 3.

Key test results.

Specimen	Peak load $P_{u}$ (kN)	Ultimate hoop strain $ε_{h, u}$ (%)	Unconfined concrete strength $f'_{co}$ (MPa)	Peak concrete stress $f_{cc}$ (MPa)			Strain at peak stress $ε_{cc}$ (%)
Specimen	Peak load $P_{u}$ (kN)	Ultimate hoop strain $ε_{h, u}$ (%)	Unconfined concrete strength $f'_{co}$ (MPa)	Test $f_{cc, \exp}$	Prediction $f_{cc, pred}$	$\frac{f_{cc, \exp}}{f_{cc, pred}}$	Test $ε_{cc, \exp}$	Prediction $ε_{cc, pred}$	$\frac{ε_{cc, \exp}}{ε_{cc, pred}}$
SA-C0-R0-1	2995	–	47.9	53.10	57.43	1.08	0.68	0.80	1.18
SA-C0-R0-2	2889	–	47.9	53.56	57.43	1.07	0.38	0.80	2.11
SA-C0-R30-1	–	–	42.2	–	–	–	–	–	–
SA-C0-R30-2	2723	–	42.2	53.93	51.93	0.96	0.60	0.86	1.43
SA-C1-R0-1	3406	0.96	47.9	62.67	64.83	1.03	1.10	1.13	1.03
SA-C1-R0-2	3484	1.03	47.9	64.14	65.09	1.01	1.21	1.17	0.97
SA-C1-R30-1	3121	0.77	42.2	57.86	57.70	1.00	0.84	0.97	1.15
SA-C1-R30-2	3398	1.10	42.2	62.15	60.10	0.97	1.34	1.30	0.97
SA-C2-R0-1	4030	1.17	47.9	72.94	75.49	1.03	1.59	1.55	0.97
SA-C2-R0-2	4453	1.71	47.9	81.43	82.81	1.02	2.13	2.26	1.06
SA-C2-R30-1	3516	0.97	42.2	63.16	66.62	1.05	1.37	1.35	0.99
SA-C2-R30-2	4275	1.55	42.2	77.59	75.05	0.97	1.97	2.14	1.09
SB-C0-R0-1	2688	–	46.2	52.08	50.35	0.97	0.53	0.51	0.96
SB-C0-R0-2	2813	–	46.2	51.64	50.35	0.98	0.43	0.51	1.19
SB-C0-R30-1	–	–	41.0	–	–	–	–	–	–
SB-C0-R30-2	2613	–	41.0	48.42	50.42	1.04	0.49	0.56	1.14
SB-C1-R0-1	3153	1.13	46.2	61.01	58.96	0.97	1.24	1.17	0.94
SB-C1-R0-2	3133	1.19	46.2	60.55	59.25	0.98	1.29	1.23	0.95
SB-C1-R30-1	2894	1.07	41.0	56.33	54.64	0.97	1.03	1.17	1.14
SB-C1-R30-2	3033	1.46	41.0	58.93	56.96	0.97	1.29	1.54	1.19
SB-C2-R0-1	3913	1.24	46.2	73.87	71.67	0.97	1.64	1.56	0.95
SB-C2-R0-2	3948	1.50	46.2	74.93	75.41	1.01	1.88	1.91	1.02
SB-C2-R30-1	3482	1.34	41.0	65.92	68.60	1.04	1.53	1.80	1.18
SB-C2-R30-2	3639	1.21	41.0	68.88	66.48	0.97	1.48	1.62	1.09
SC-C0-R0-1	1959	–	30.1	36.06	39.78	1.10	0.78	1.09	1.40
SC-C0-R0-2	2017	–	30.1	37.15	39.78	1.07	0.80	1.09	1.36
SC-C0-R30-1	——	–	33.5	–	–	–	–	–	–
SC-C0-R30-2	2123	–	33.5	43.29	40.50	0.94	1.03	1.02	0.99
SC-C1-R0-1	2821	1.42	30.1	49.37	49.40	1.00	1.67	1.77	1.06
SC-C1-R0-2	3002	1.68	30.1	53.64	51.19	0.95	1.79	2.08	1.16
SC-C1-R30-1	3011	1.70	33.5	52.82	54.80	1.04	1.94	2.04	1.05
SC-C1-R30-2	2928	1.67	33.5	51.63	54.73	1.06	1.88	2.00	1.06
SC-C2-R0-1	3625	1.52	30.1	64.50	61.06	0.95	2.62	2.40	0.92
SC-C2-R0-2	3250	1.39	30.1	55.57	59.14	1.06	2.32	2.17	0.94
SC-C2-R30-1	3269	1.17	33.5	63.89	59.78	0.94	1.68	1.75	1.04
SC-C2-R30-2	3805	1.45	33.5	67.43	63.89	0.95	2.30	2.19	0.95
Average						1.00			1.11
SD						0.05			0.22

The axial load–strain curves of all specimens are plotted in Figure 4. In this figure, the axial load is the total axial load undertaken by the specimen. The axial strain and hoop strain were obtained in the same manner as Table 3. The load–strain responses of all specimens can be divided into two stages. In the first stage, the axial load increased with the increase in axial/hoop strain but at a decreasing rate, which is similar with normal concrete. For specimens without FRP jacket (i.e. concrete-filled steel tubes), the second stage of load–strain curve is either decreasing (i.e. Batches A and B) or almost flat (i.e. Batch C). For specimens with FRP jacket, the axial load–strain curves have a second ascending portion, and the slope of the second ascending portion increases with the increase in FRP jacket thickness (i.e. number of CFRP layers).

Figure 4.

Axial load–strain curves: effect of CFRP layers: (a) Batch A: R = 0%, (b) Batch A: R = 30%, (c) Batch B: R = 0%, (d) Batch B: R = 30%, (e) Batch C: R = 0%, and (f) Batch C: R = 30%.

The axial load–strain curves of some specimens are replotted in Figure 5 to investigate the effect of mix ratio (R), where specimens with the same steel tube and FRP jacket but different mix ratios are plotted together. The first stage of the load–strain curve of specimens with and without RCLs is generally similar with each other except small difference at the turning point between two stages. For Batches A and B, the second stage of the load–strain curve of specimens with RCLs is slightly lower than that of specimens without RCLs, as the strength of RCLs in Batches A and B are smaller than FC. But the difference become smaller with the increase in deformation, which means the negative effect of RCLs can be partially offset by the FRP confinement. For Batch C, where the strength of RCLs is higher than that of FC, the second stage of the load–strain curve of specimens with RCLs is slightly higher than those without RCLs, showing the positive effect of higher strength RCLs. The same observation has been reported by Teng et al. (2016) for compound concrete confined by glass fibre–reinforced polymer (GFRP) tubes.

Figure 5.

Axial load–strain curves: effect of mix ratio: (a) Batch A: C1, (b) Batch A: C2, (c) Batch B: C2, and (d) Batch C: C2.

Stress–strain behaviour of steel tubes

To obtain the stress–strain curve of concrete core, the axial load undertaken by the steel tube needs to be decided first. In an FRP-confined concrete-filled steel tube, the steel tube is under plane stress state (Hu et al., 2011). The axial and hoop stresses of the steel tube at any strain state can be decided by an incremental manner using classic J₂ flow theory (Chen and Saleeb, 1994; Hu et al., 2011; Teng et al., 2013). More information can be found from Teng et al. (2013).

During the prediction, the hoop strain of steel tube was taken as the average reading of three hoop strain gauges outside the overlapping zone. The axial strain of steel tube was taken as the average reading of four mid-height LVDTs. The yield stress and elastic modulus listed in Table 2 were adopted. Poisson’s ratio of steel tubes was taken as 0.3. The strain hardening effect of steel tube was not considered, as the maximum strain of steel tube from the test was generally smaller than the strain at the initiation of hardening ( $ε_{sh}$ ) of steel tube (Table 2).

The stress–strain curves of steel tubes predicted using above-mentioned method are shown in Figure 6 for typical specimens. Figure 6(a) shows the axial stress versus axial strain curves as well as the hoop stress versus hoop strain curves, while Figure 6(b) shows the axial stress versus hoop stress relationship. The yield point of steel tube is indicated by small circles in Figure 6(a). The yield surface of steel tube is also indicated in Figure 6(b). From Figure 6, the following can be observed: (1) before yielding of steel tube, the axial stress generally increases with the increase in axial strain. In hoop direction, the steel tube is subjected to compression for most specimens. The compression hoop stress is mainly due to the confinement of the outer FRP jacket, which prevents the hoop deformation of the steel tube due to Poisson effect. However, the hoop stress level in this stage is generally very small. As a result of the hoop compression stress, the maximum axial stress of steel tube may be slightly larger than the yield stress obtained from coupon test. (2) After yielding of steel tube, the axial stress in steel tube generally decreases with the increase in axial deformation and finally tends to keep constant at certain stress level (e.g. ∼100 MPa for specimen SC-C0-R0-1). The final axial stress level increases with the increase in FRP jacket thickness, which means the outer FRP jacket increased the contribution of steel tube in axial direction. In hoop direction, the hoop stress generally transfers from compression to tension and the tensile stress in hoop direction finally also maintains at certain stress level. The hoop stress level generally decreases with the increase in FRP jacket thickness, which means the outer FRP jacket reduces the contribution of steel tube in hoop direction.

Figure 6.

Stress–strain responses of steel tubes: (a) stress–strain curves and (b) axial-hoop stress curve.

Stress–strain behaviour of the concrete core

Once the axial stress in the steel tube is decided following section ‘Stress–strain behaviour of steel tubes’, the axial load on concrete core can be obtained by subtracting the axial load undertaken by steel tube from the total axial load. The axial stress of concrete core can thus be decided as the ratio of axial load on concrete and concrete core area. The stress–strain curves of core concrete of all specimens with FRP confinement obtained by this procedure are shown in Figure 7. The peak axial stress (f_cc) and corresponding axial strain (ε_cc) of concrete core of all specimens are also summarized in Table 3. For specimens without FRP jacket, the second stage of stress–strain is either descending (Batches A and B) or almost flat (Batch C). As a result, the peak axial stress and corresponding strain are different from the ultimate axial stress and ultimate axial strain. For specimens with FRP jacket, the second stage of the stress–strain curves is generally ascending, and the peak axial stress and strain are also the ultimate axial stress and strain.

Figure 7.

Stress–strain curves of the core concrete: (a) Batch A: C1-R0, (b) Batch A: C1-R30, (c) Batch A: C2-R0, (d) Batch A: C2-R30, (e) Batch B: C1-R0, (f) Batch B: C1-R30, (g) Batch B: C2-R0, (h) Batch B: C2-R30, (i) Batch C: C1-R0, (j) Batch C: C1-R30, (k) Batch C: C2-R0, and (l) Batch C: R2-R30.

Performance of existing stress–strain model

Teng et al. (2013) proposed a stress–strain model for concrete core of FRP-confined steel tubes, which can be deemed as a refined version of Jiang and Teng (2007) analysis-oriented model for FRP-confined concrete to account for the additional confinement provided by steel tube. The model consists of three parts: (1) an active confinement model base model, (2) a lateral-to-axial strain relationship, and (3) an equation to calculate the confinement pressure. The details of the model can be found from Teng et al. (2013). It should be mentioned that Teng et al. (2013) proposed two versions of stress–strain models, and ‘Model I’, which adopts the same lateral-to-axial strain relationship with Jiang and Teng (2007) analysis-oriented stress–strain model, was adopted in the present study.

During the application of the model, the unconfined concrete strength $f'_{co}$ is taken as 0.85 times the cube strength (ACI 2008). For specimens with RCLs, the cube compressive strength of compound concrete ( $f_{cu, com}$ ) is calculated by the following equation proposed by Wu et al. (2015)

f_{cu, com} = (1 - R) f_{cu, FC} + R f_{cu, RCLs}

(1)

where $f_{cu, FC}$ and $f_{cu, RCLs}$ are the cube compressive strength of FC and RCLs, respectively, and R is the mix ratio of RCLs. The unconfined concrete strength $f'_{co}$ of each specimen decided by this method is summarized in Table 3.

The stress–strain curves of concrete core of typical specimens predicted by Teng et al. (2013) model are plotted in Figure 7 together with test results. During the application of the model, the ultimate hoop strain ( $ε_{h, u}$ ) from the tests (Table 3) was directly adopted for each specimen. In Figure 7, for two nominally identical specimens, only the prediction of the specimen with the larger ultimate hoop strain is plotted.

From Figure 7, it can be seen that Teng et al. (2013) model can predict the stress–strain curve satisfactorily for specimens with or without RCLs. The peak stress ( $f_{cc}$ ) and corresponding strain ( $ε_{cc}$ ) predicted by Teng et al. (2013) model are summarized in Table 3 for all specimens, as well as compared with test results in Figure 8(a) and (b), respectively. The average ratio between the prediction and test result for peak stress ( $f_{cc}$ ) is 1.0 and the standard deviation is 0.05. The average ratio between the prediction and test results for strain at peak stress ( $ε_{cc}$ ) is 1.11 and the standard deviation is 0.22. It seems that the ultimate axial strain was slightly overestimated for most specimens. It can also be seen from Figure 8 that the performance of the model for specimens without RCLs (i.e. R = 0%) and specimens with RCLs (R = 30%) are comparable.

Figure 8.

Performance of Teng et al. (2013) model: (a) peak stress f_cc and (b) corresponding strain ε_cc.

To further assess the performance of Teng et al. (2013) model, the lateral-to-axial strain relationship predicted by the model is compared with test results in Figure 9. In this figure, the lateral strain is normalized by peak strain of unconfined concrete strength ( $ε_{co}$ ) and the axial strain was normalized by $ε_{co} (1 + 8 σ_{l} / f'_{co})$ . In Figure 9, large scatter is observed among the test curves for all three batches. The predictions generally follow the global trend of the test curves, but provide slightly larger axial strains at most cases.

Figure 9.

Performance of Teng et al. (2013) model: lateral-to-axial strain relationships: (a) Batch A, (b) Batch B, and (c) Batch C.

Conclusion

In this study, 12 compound concrete–filled steel tubular columns and 24 FRP-confined compound concrete–filled steel tubular columns were tested under axial compression in three batches. The test parameters included steel tube diameter-to-thickness ratio, thickness of FRP jacket, mix ratio of RCLs for compound concrete, and relative strength between FC and RCLs. The load–deformation behaviour of the columns was carefully measured and analysed. The stress–strain responses of steel tubes and concrete core were also derived. Based on the present study, the following conclusions can be made:

Compound concrete–filled steel tubes generally fail by buckling of steel tube, while FRP-confined columns generally fail by rupture of FRP jacket, and buckling of steel tube was not observed.

Bilinear load–deformation curve is observed in all specimens. The second stage of the bilinear curve of specimens without FRP jacket is either descending or horizontal due to insufficient confinement, while that of specimens with FRP jacket is generally ascending and the slope of ascending increases with the increase in FRP jacket thickness.

The steel tube is subjected to plane stress state. After the yielding of steel tube, FRP jacket can increase the contribution of steel tube in axial direction but reduce its contribution in hoop direction.

When relatively low-strength RCLs were mixed with relatively high-strength FC, the mechanical strength of specimens with compound concrete is slightly lower than that of specimens with pure FC; when relatively high-strength RCLs were mixed with relatively low-strength FC, the mechanical strength of specimens with compound concrete is improved compared with specimens with FC.

Teng et al. (2013) model for concrete filled in FRP-confined steel tubes can satisfactorily predict the stress–strain behaviour of compound concrete filled in FRP-confined steel tubes, although the ultimate axial strain is slightly overestimated.

Footnotes

Declaration of Conflicting Interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: Natural Science Foundation of China (Project Nos 51608392, 51678454, and 51578423).

ORCID iD

Junliang Zhao

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