Axial compressive behavior of GFRP composite spiral ties-reinforced ECC-encased HCFST columns

Abstract

Aiming to solve the issues of conventional concrete-encased high-strength concrete-filled steel tube (HCFST) columns, a novel GFRP composite spiral ties (GC)-reinforced engineered cementitious composite (ECC)-encased HCFST column was proposed and tested under axial compression in this study. The test variables included encasing cement-based materials, the thicknesses of ECC and steel tube, the compression strength of HSC, and the tie configuration. The test results indicated that the failure mode of the composite columns changed after replacing the outer concrete with ECC, showing that the ECC-encased HCFST columns had excellent integrity. When ECC with the same compression strength was used instead of concrete, the bearing capacity, strength index (SI), ductility (μ) and energy dissipation (E_d.) of the composite columns were increased by 14.2%, 7.7%, and 81%, respectively. Increasing the thicknesses of ECC and the steel tube significantly improved the composite columns' bearing capacity, and the latter was superior in enhancing the μ and E_d. of the columns. Compared with other tie configurations, GC had an outstanding performance in improving μ and E_d., reflecting the excellent restraint performance of GC. Furthermore, a calculation method for the axial compressive bearing capacity of GC-reinforced ECC-encased HCFST columns was proposed based on the composite constraint model and the unified theory of strength.

Keywords

ECC GFRP HCFST composite column axial compression

Introduction

Concrete-filled steel tubes (CFSTs), a type of steel-concrete composite structure, have received extensive research and applied in civil engineering (Sakino et al., 2004; Chitawadagi et al., 2009). CFSTs have the benefits of light weight and high bearing capacity since the steel tube and concrete have an excellent combination effect. However, the following issues exist in the CFST: (1) Without external support, the steel tube easily buckles outwards under the axial load (Fam et al., 2004); (2) In direct exposure to the external environment, the steel tube has poor durability and is prone to corrosion and damage, resulting in high maintenance costs (Han et al., 2014); and (3) When a fire occurs, the strength of the steel tube decreases linearly, and the restraining effect on the concrete is reduced (Yang et al., 2008).

To address the abovementioned issues regarding CFSTs, a new type of composite column (concrete-encased CFST column) was proposed (Han et al., 2014). The outer concrete layer enhanced the durability and fire resistance of the concrete-encased CFST column compared to the CFST (Cai et al., 2018). Various properties of concrete-encased CFST columns, such as compression (Han et al., 2014), bending (An et al., 2014), tension (Han et al., 2016), and seismic performance (Ji et al., 2014; Liao et al., 2014), have been investigated. Kang and Qian (Kang et al., 2006) examined the behavior of concrete-encased CFST columns under axial compression. The test results show that the outer concrete had severe peeling damage and that the core CFST retained axial load-bearing capacity even if the outer concrete failed. Han et al. (Han et al., 2014) conducted a numerical investigation on the axial compression performance of concrete-encased CFST columns. The research results show that both the composite column and outer concrete reached the peak load simultaneously. Despite reaching the peak load, the CFST remained in the elastic-plastic stage, indicating poor combined performance of the outer concrete and inner CFST.

Due to the low creep and high elastic modulus, high-strength concrete (HSC) is introduced into the CFST to improve the stiffness of the composite column. The concrete-encased HSC-filled steel tube (HCFST) column was further studied by researchers. Kang et al. (Kang et al., 2010) conducted research on the axial compression performance of concrete-encased HCFST columns. The failure mode of the composite column was that the outer concrete compressive crashed first, and then the CFST failed. Niu et al. (Niu et al., 2022) studied the axial compressive performance of concrete-encased HCFST columns, and test results showed that they have serious spalling of concrete and poor ductility. Ke et al. (Ke et al., 2021) experimentally studied the seismic performance of concrete-encased HCFST columns and found that the failure characteristics were concrete crushing and longitudinal reinforcement yielding. The ductility and energy dissipation capacity could be significantly enhanced by increasing the volumetric tie ratio in the range of 1.33%–1.63%.

However, the following issues in concrete-encased CFST/HCFST columns still need to be noted: (1) Compared to the inner CFST/HCFST, the outer concrete layer has lower compressive ductility, and its peak load is ahead of that of the inner CFST/HCFST. This leads to poor combined effect between the inner and outer layers (Cai et al., 2018; Han et al., 2014). (2) The steel tie is located at the outermost position of the reinforcement cage, which is more likely to corrode in harsh environments, resulting in weakened lateral restraint performance and reduced service life of the composite column (Zhang et al., 2021).

To address the issues of low compressive ductility of the outer concrete and poor codeformation ability between the outer and inner parts, the engineered cementitious composite (ECC) with superior ductility and strain-hardening behaviour (Li et al., 1992; Li et al., 1991) can be adopted to replace the ordinary concrete in the concrete-encased HCFST column. The ultimate tensile strain of ECC is approximately 3%–7%, approximately 300-700 times that of ordinary concrete, which can better codeform with the internal HCFST. ECC effectively distributes stress during loading, significantly reducing crack formation and enhancing crack resistance (Li et al., 1992). The fiber-reinforced components within ECC’s microstructure improve toughness and ductility while markedly enhancing corrosion resistance. This enables ECC to robustly protect internal steel tubes in corrosive environments, thereby strengthening the overall reliability of the structure (Cai et al., 2018). In addition, the corrosion issue can be solved by using glass fibre-reinforced polymer (GFRP) ties that have great durability instead of steel ties. For the rectangular FRP tie, lateral loads are resisted by bearing against the concrete while also experiencing longitudinal stresses along the fibre direction. The lower load-bearing capacity of FRP perpendicular to the fibres creates significant radial stress, leading to stress concentration at the bend that weakens the structure and causes premature failure. Additionally, breaking at bends is common due to this stress concentration, compounded by uneven lateral restraint and the presence of an arched “non-effective restraint area” (Mander, 1988). For the circular spiral tie, uniform lateral restraint can be achieved, whereas it is unsuitable for the square column (Zhang et al., 2022). The GFRP composite spiral ties (GC) proposed by the authors in this paper (Zhang et al., 2022), consisting of outer GFRP rectangular ties and an inner GFRP spiral tie, can provide composite restraint to the core area, dramatically improve the lateral restraint effect and be suitable for the square section.

Therefore, the GC-reinforced ECC-encased HCFST column was proposed in this study. The axial compression performance of the proposed column was investigated. The effects of the encasing cement-based material, the thicknesses of ECC and steel tube, the compression strength of HSC, and the tie configuration on the failure mode, bearing capacity, ductility, and energy dissipation of the composite column were studied. In addition, based on the composite constraint model and unified theory of the CFST, a calculation method for the axial compressive bearing capacity of GC-reinforced ECC-encased HCFST columns was proposed.

Experimental investigation

Specimen details

A total of 10 square composite columns were designed for the axial compression test, the details of which are displayed in Figure 1. All composite columns had a 300 mm × 300 mm cross-section and a 900 mm height (except for column E^#-GC-6-80 with a cross-section of 340 mm × 340 mm). The column height-to-side length ratio was 3 (2.65 for column E^#-GC-6-80), and the stability factor of all composite columns was one per GB/T 50152-2012 (2012); therefore, the composite column could be recognized as a short column without considering the second-order effect. Eight longitudinal steel bars with a 16 mm diameter were positioned outside the tube, and the longitudinal reinforcement ratio α_l was 1.79% (α_l = A_l/A_c, where A_l and A_c are the total area of the longitudinal reinforcement and the cross-sectional area of the composite column, respectively). The concrete cover thickness was 25 mm, which was the minimum distance from the outer edge of the tie to the composite column surface.

Figure 1.

Details of the composite column (unit: mm).

Table 1 displays the test matrix for all composite columns. Experimental parameters included the encasing cement-based material (ECC or concrete), the thickness of ECC (60 mm or 80 mm), the thickness of the steel tube (6 mm or 10 mm), the compression strength grade of the HSC (C60 or C80), and the tie configuration. Five tie configurations were used, namely, GFRP composite spiral tie (GC), steel rectangular tie (SR), GFRP rectangular tie (GR), GFRP spiral tie (GS), and no tie (O), as shown in Figure 2. The GC was made up of multiple GFRP rectangular ties and one GFRP spiral tie at a spacing of 100 mm (see Figure 1). The rectangular tie was lapped at corners with a lap length of 125 mm. The size of the GFRP rectangular tie was 250 mm × 250 mm, and the diameter of all ties was 12 mm. The tie spacing within the range of 150 mm was reduced to 50 mm and CFRP sheets of 150 mm height were attached at the column ends to avoid premature failure of the column ends.

Table 1.

Test Matrix.

Label	Encasing material	ECC thickness/mm	Tube thickness/mm	Strength grade	Tie configuration
N-SR-6-80	NC	N/A	6	C80	SR
N-GC-6-80	NC	N/A	6	C80	GC
E-GC-6-80	ECC	60	6	C80	GC
E^#-GC-6-80	ECC	80	6	C80	GC
E-GC-10-80	ECC	60	10	C80	GC
E-GC-6-60	ECC	60	6	C60	GC
E-O-6-60	ECC	60	6	C60	O
E-SR-6-60	ECC	60	6	C60	SR
E-GR-6-60	ECC	60	6	C60	GR
E-GS-6-60	ECC	60	6	C60	GS

Note: The column label is divided into four parts: encasing cement-based material (N and E), tie configuration (GC, SR, GR, GS, and O), steel tube thickness (6 mm or 10 mm), and compression strength grade of the HSC (C60 or C80). In addition, E and E^# represent 60-mm-thick and 80-mm-thick ECC layers, respectively.

Figure 2.

Tie configurations.

Materials

The raw materials of ECC included PO.42.5 ordinary Portland cement, fly ash (FA), silica fume (SF, SiO₂≥93%), quartz sand (particle size of 74 μm), water, superplasticizer (water reduction rate of 35%), thickener (hydroxypropyl methylcellulose), and PE fibres, as listed in Table 2. The tensile strength, elastic modulus, and elongation at break of the PE fibres were 3000 MPa, 110 GPa, and 3%, respectively. To test the compressive and tensile properties of ECC, cubic and dog bone tensile specimens were tested per JC/T 2461-2018 (2018). The compressive strength, tensile strength, and ultimate tensile strain of ECC were 45.1 MPa, 4.6 MPa, and approximately 6.5% (650 times higher than that of conventional concrete, see Figure 3(a)), respectively. As shown in Figure 3(b), dense microcracks appeared in the gage section of the dog bone, which had a gage length of 80 mm, indicating that ECC had excellent crack control ability and ductility.

Table 2.

Mix Proportion of ECC (kg/m³).

Binder			Quartz sand	Water	Superplasticizer	Thickener	PE fibre (vol = 2%)
Cement	FA	SF
315	750	974	240	360	1.8	0.48	19.4

Figure 3.

Tensile behaviour of ECC.

Table 3 shows the mix ratios for HSC (C60 and C80) and concrete (C40). Concrete cubes of 150 mm × 150 mm × 150 mm in size were tested for compressive strength per GB/T 50081-2019 (2019). The cubic compressive strengths of C40, C60, and C80 grade concrete were 44.8 MPa, 63.9 MPa, and 84.9 MPa, respectively. The steel tube was made of Q345b steel, and the properties of the steel were tested according to GB/T 228.1-2010 (2010). The experiments were conducted using a computer-controlled CMT5305 hydraulic universal testing machine at a loading rate of 5 MPa/s. Three samples of 6 mm and 10 mm thick Q345b steel were tested respectively, and the average value of the three samples was taken as the material properties, as shown in Table 4. The testing methods for the HRB400 steel bar were the same as those for Q345b, with the measured yield strength and ultimate strength being 438.5 MPa and 569.7 MPa, respectively.

Table 3.

Mix Ratios of HSC and concrete (kg/m³).

Concrete	Water	Cement	Sand	Stone	SF	SP
C40	195	433	555	1184	N/A	N/A
C60	157	462	588	1230	14.1	2.2
C80	155	532	596	1208	15.4	3.1

Table 4.

Mechanical Properties of Q345b Steel.

t_e (mm)	ε _y	σ_y (MPa)	σ_u (MPa)	E₀ (×10⁵ MPa)	σ_y/σ_u	δ (%)
5.86	0.0023	368.5	508.1	2.11	0.725	27.3
10.08	0.0022	370.7	516.6	2.13	0.718	27.9

Note: t_e is the measured thickness; ε_y is the yield strain; σ_y is the nominal yield strength; σ_u is the tensile strength; E₀ is the elastic modulus; σ_y/σ_u is yield-to-strength ratio, and δ is ultimate elongation.

The tensile properties of 12-mm-diameter GFRP ties at the straight portion were determined per GB/T 30022-2013 (2013), as listed in Table 5. The tensile strength of the GFRP rectangular ties at the bend portion (

f_{b r}

) can be significantly reduced due to the stress concentration compared with the straight portion . The measured tensile properties of the bend portion of the GFRP rectangular ties based on ACI 440.3R-04 (2013) and CSA S806-12 (2012) (see Table 5). Similarly, the tensile strength (

f_{b s}

) of the spiral tie, experiencing strength loss, was calculated by ACI 440.1R-06 (2004), as expressed in equation (1) (refer to Table 5 for results).

f_{b s} = (0.05 \frac{r_{b}}{d_{b}} + 0.3) f_{u} \leq f_{u}

(1)

where

\frac{r_{b}}{d_{b}}

is the ratio of the bend radius to the bar diameter.

f_{u}

is the GFRP ties at the straight portion’s ultimate tensile strength.

Table 5.

Tensile Properties of the GFRP Ties.

E (GPa)	$f_{u}$ (MPa)	$ε_{u}$ (%)	$f_{b r}$ (MPa)	$η_{1}$	$f_{b s}$ (MPa)	$η_{2}$
45.1	1093.7	2.4	516.8	0.47	843.1	0.77

Note: E and $ε_{u}$ are the GFRP ties’ tensile modulus of elasticity and ultimate strain at the straight portion. $η_{1}$ is the strength reduction factor for the rectangular tie, defined as the ratio of $f_{b r}$ to $f_{u}$ . $η_{2}$ is the strength reduction factor for spiral ties, defined as the ratio of $f_{b s}$ to $f_{u}$ .

Test setup and loading protocol

All composite columns were loaded by a multifunctional structural test system with a load capacity of 12000 kN (see Figure 4(a)). Before formal loading, a preload with less than 10% of the predicted peak load was used to check the testing device. The load was first controlled by force with speed of 1 kN/s before 80% of the predicted peak load. To better track the postpeak behaviour of the composite column, the displacement-controlled loading method was used at a speed of 0.1 mm/min until the load decreased to 80% of the peak load.

Figure 4.

Test setup and measurement.

The axial load was automatically recorded by the force sensor at an accuracy of 1 kN. Four displacement sensors with a range of 50 mm were arranged surrounding the middle of the composite column, and the average displacement was taken as the target axial deformation of the composite column. The strain variation of the column surface was monitored by pasting strain gauges vertically and horizontally at the middle height (see Figure 4(b)). The strains of the steel tube along the axial and transverse directions were measured by pasting vertical and horizontal strain gauges at the middle height. Strain gauges were also attached to rectangular ties, spiral ties, and longitudinal bars to assess the reinforcement strain.

Test results

Test observations and failure mechanisms

In general, the behaviour of composite columns was mainly affected by the encasing cement-based materials. Compared with concrete-encased HCFST columns, ECC-encased HCFST columns showed more ductile failure and had better integrity after failure. The failure mechanisms were also dependent on the thicknesses of ECC and steel tube, HSC compression strength, and tie configuration. The detailed behaviour and failure mechanisms of the composite columns are detailed in the following sections.

A vertical concrete microcrack appeared at the middle height of column N-SR-6-80 as the load increased to 4990 kN (78.0% of the peak load). The initial crack developed along the vertical direction and more cracks appeared as the load increased. When the outer concrete reached the ultimate compressive strength, the concrete was crushed and peeled off in a large area, as shown in Figure 5(a). Then, the HCFST mainly carried the load, and the steel tube eventually failed due to local buckling that was observed by removing the loose outer concrete after the test. Buckling of the ties and longitudinal bars can be found in Figure 5(b). The failure mode of specimen N-GC-6-80 was similar to that of N-SR-6-80, as shown in Figure 5(c). However, the degree of concrete peeling of N-GC-6-80 was lower, indicating that the GFRP composite spiral ties could mitigate concrete peeling. The GFRP rectangular and spiral ties were found to be broken after removing the outer loose concrete (see Figure 5(d)).

Figure 5.

Failure mode for N-SR-6-80, N-GC-6-80, and E-GC-6-80.

The GC-reinforced ECC-encased HCFST columns had a different failure mode from the concrete-encased HCFST columns. Compared to concrete-encased HCFST columns, GC-reinforced ECC-encased HCFST columns typically demonstrated superior ductility, with failure modes characterized by the gradual formation of multiple microcracks leading to a main crack, rather than sudden cracking. In contrast, concrete-encased HCFST columns were more prone to brittle failure. ECC exhibited significantly better crack resistance than ordinary concrete, resulting in slower crack propagation. Therefore, under load, GC-reinforced ECC-encased HCFST columns were able to more effectively control crack width, while the cracks in concrete-encased HCFST columns tended to expand rapidly.

Figure 5(e)–(g) shows the failure mode of column E-GC-6-80 as an example. The initial crack developed vertically in the middle height of column E-GC-6-80 at 87.6% of the peak load (5890 kN), which was later than that in column N-SR-6-80 (78.0% of the peak load), indicating that the composite column could control cracks better by replacing concrete with ECC. In column E-GC-6-80, at the initiation of the initial crack, the width was measured at 0.02 mm. This indicates that ECC can effectively control the early development of cracks. The multiple microcracks presented an oblique development pattern, along with the sound of PE fibres being pulled out of the matrix, as shown in Figure 5(e) and (f). When ECC reached their ultimate compressive strength, the composite column also attained its ultimate load-bearing capacity, after which the HCFST primarily bore the remaining load. When the column E-GC-6-80 reached its peak load, the width of the main crack was 0.2 mm, which was still at a relatively low level. After the peak load, the width of the main crack developed rapidly, and by the end of the test (when the load had dropped to 80% of the peak load), the width of the main crack had reached 5 mm. ECC could absorb and disperse the applied load energy, reducing localized stress concentration. This characteristic allowed cracks to develop rapidly after reaching peak load, but their growth rate was significantly slower than that of concrete. The application of ECC significantly enhanced the overall structural integrity of the composite column, with no observable delamination, further demonstrating ECC’s superior performance in crack resistance and corrosion protection. The bridging effect of PE fibers at the main crack substantially improved the crack resistance of ECC and effectively facilitated stress transfer across the crack surfaces (see Figure 5(g)). This bridging effect not only effectively mitigated the propagation of cracks but also significantly enhanced the durability of the composite column, allowing it to exhibit higher toughness and strength under extreme loading conditions. Therefore, the application of ECC played a crucial role in enhancing the crack resistance and corrosion protection of the composite column. Except for E-O-6-60, the failure mode of the SR-, GR- or GS-reinforced ECC-encased HCFST composite columns was similar to that of E-GC-6-80, as shown in Figure 6. Column E-O-6-60 without ties had poor integrity, indicating that the composite column necessitates tie reinforcement to improve structural integrity.

Figure 6.

Failure mode and crack distributions.

Load and axial displacement responses

The load-axial displacement curves of all composite columns can be divided into four stages: elastic stage, elastic-plastic stage, descending stage, and softening stage, as shown in Figure 7. The axial displacement increased linearly with the load in the elastic stage. The outer concrete of the concrete-encased HCFST columns began to crack, and the curves exhibited nonlinearity in the elastic-plastic stage. The plastic deformation of the internal HCFST caused the initial nonlinearity of the curves for ECC-encased HCFST composite columns, due to the lower elastic modulus of ECC than concrete (Cai et al., 2018). When the outer cement-based material reached the compressive strength, the composite column reached the peak load, after which the load rapidly decreased due to the severe damage of the outer cement-based material, and the curves entered the descending stage. The internal HCFST primarily provided the bearing capacity of the composite column during the softening stage, and the descending speed of the curve slowed down.

Figure 7.

Load-axial displacement curves of all composite columns.

The test results of all composite columns are shown in Table 6. The ratio of the peak load to the superposition of the bearing capacities of each component is defined as the strength index (SI) of the composite column, which characterizes the combined effect of each component. The calculation of SI is defined by the following formula.

S I = \frac{N_{u}}{A_{E} f_{E} + A_{s} f_{y s} + A_{l} f_{y l} + A_{c} f_{c}}

(2)

where

N_{u}

is the peak load.

A_{E}

A_{s}

A_{l}

, and

A_{c}

are the cross-sectional areas of the outer cement-based material, steel tube, longitudinal bar, and inner HSC, respectively.

f_{E}

is the prism compressive strength of ECC, calculated from the cubic compressive strength per GB 50010-2010 (2010).

f_{c}

is the cylinder compressive strength of HSC, calculated from the cubic compressive strength according to GB 50010-2010 (2010).

f_{y s}

and

f_{y l}

are the yield strengths of the steel tube and longitudinal steel bar, respectively.

Table 6.

Test Results.

Column label	N_u/kN	δ_u/mm	δ_u85/mm	SI	μ	E_d/kJ
N-SR-6-80	6398	5.78	6.88	1.05	1.19	29.3
N-GC-6-80	5882	5.25	6.14	1.05	1.17	23.1
E-GC-6-80	6722	7.97	10.0	1.20	1.26	41.9
E^#-GC-6-80	7700	7.00	9.66	1.16	1.38	46.9
E-GC-10-80	7530	9.17	15.4	1.22	1.68	86.4
E-GC-6-60	6471	7.78	12.2	1.22	1.57	50.7
E-O-6-60	4366	7.31	8.19	0.96	1.12	19.8
E-SR-6-60	6272	7.29	8.68	1.19	1.19	28.9
E-GR-6-60	5767	7.20	8.78	1.10	1.22	30.8
E-GS-6-60	6299	7.68	9.68	1.20	1.26	35.9

The ductility of each composite column was quantitatively assessed using the ductility coefficient μ (Cai et al., 2018). The ratio of axial displacement $δ_{u 85}$ to $δ_{u}$ , as shown in equation (3), was defined as μ.

μ = \frac{δ_{85}}{δ_{u}}

(3)

where

δ_{u 85}

and

δ_{u}

are the axial displacements when the load drops to 85% of the peak load

N_{0.85 u}

and reaches the peak load

N_{u}

, respectively.

The energy dissipation E_d, is obtained by integrating the load-axial displacement curve over the axial displacement interval from 0 to $δ_{u 85}$ , as shown in equation (4).

E_{d} = \int_{0}^{δ_{85}} f (δ) d δ

(4)

Strain analysis

The relationship between the axial load and strains of ties, steel tubes, and ECC is shown in Figure 8. To investigate the strain law of the different tie configurations, the tie strain responses of columns E-GC-6-60, E-GR-6-60, and E-GS-6-60 are compared in Figure 8(a). During the initial loading stage, the tie strain increased linearly and slowly and was maintained at a low level. This phenomenon is explained as follows. At the initial stage, the transverse deformation of ECC was minor, which did not squeeze the ties outwardly. As the load increased further, the increased transverse deformation of ECC squeezed the tie, causing a rapid increase in the tie strain due to the passive restraint characteristic of the GFRP tie. At the peak load, the strains of the E-GC-6-60’s GFRP rectangular and spiral ties were 2469 με and 589 με, respectively, with corresponding stresses of 111.4 MPa and 26.6 MPa. The strain of the former was much greater than that of the latter, indicating that the GFRP rectangular ties played a more significant role in restraining the core area at the peak load. The strains of the E-GC-6-60’s rectangular and spiral ties were less than those of E-GR-6-60 and E-GS-6-60 at an equal load level before the peak load, indicating that the rectangular and spiral ties in the GFRP composite spiral ties were jointly stressed, with a lower strain level than the single tie configuration.

Figure 8.

Strain variations.

Figure 8(b) shows the steel tube strain of composite column E-GC-6-60. The initial ratio of the transverse strain to vertical strain of the steel tube was 0.33, which agreed well with the elastic Poisson’s ratio of the steel (ν = 0.3). This is explained as follows. Due to the lower Poisson’s ratio of the HSC (ν = 0.2-0.28 (Committee, 1984)) than the steel tube, the HSC did not squeeze the steel tube outwards at the initial loading. The ratio of transverse strain to vertical strain of the steel tube increased to 0.4 at a load of 3780 kN due to the squeeze between the core HSC and steel tube. The transverse and vertical strains rapidly developed as the steel tube yielded when the load reached 6080 kN.

Figure 8(c) illustrates the transverse strain of ECC at the middle height of composite columns E-GC-6-60, E-O-6-60, E-SR-6-60, E-GR-6-60, and E-GS-6-60. The transverse strain of ECC could represent the degree of lateral expansion of the composite column and reflect the different tie configurations’ restraint effect. The transverse strain of all representative composite columns increased linearly during the initial loading stage. The transverse strain of E-O-6-60 was greater than those of the other composite columns at an equal load level, as no ties were adopted in the encasing material. E-GC-6-60 had the smallest transverse strain, revealing that the restraint provided by the GC was superior to the single tie configuration. The transverse strain of the composite column grew faster with the development of ECC microcracks after the peak load.

Parametric discission

Effect of the encasing cement-based material

Figure 9(a) depicts the effect of the encasing cement-based material on the composite column. The initial stiffness is the ratio of the load to the axial displacement of the composite column in the elastic stage. The initial stiffness of column E-GC-6-80 (975.6 kN/mm) was 32.0% less than that of N-GC-6-80 (1433.7 kN/mm) because ECC had a relatively lower elastic modulus than ordinary concrete. This indicated that the initial stiffness was weakened by using ECC instead of concrete. The δ_u of column E-GC-6-80 (7.97 mm) was significantly larger than that of column N-GC-6-80 (5.25 mm), as ECC with a higher peak compressive strain could better coordinate deformation with the inner HCFST at the peak load. In the failure stage, column E-GC-6-80 had a slower load descending speed than column N-GC-6-80 after the peak load, indicating a higher ductility.

Figure 9.

Load versus displacement curves.

The peak loads of columns N-GC-6-80 and E-GC-6-80 were 5882 kN and 6722 kN, respectively, and the bearing capacity increased by 14% after replacing the concrete with ECC with the same compressive strength. This phenomenon could be attributed to the significantly higher ultimate compressive strain of ECC compared to that of ordinary concrete (Cai et al., 2020), which enabled it to withstand greater axial deformation. As a result, the collaborative load-bearing capacity between ECC and the internal HCFST was higher than that between concrete and HCFST, leading to a better combined effect under axial compression conditions. The improvement of the combined effect also could be seen from the comparison between the SI of columns N-GC-6-80 and E-GC-6-80, which were 1.05 and 1.20, respectively. The μ values of columns N-GC-6-80 and E-GC-6-80 were 1.17 and 1.26, respectively, showing that the utilization of the high ductility ECC increased the ductility. The E_d. was significantly improved by using ECC outside the steel tube, as proven by the E_d. of E-GC-6-80 being 1.81 times that of N-GC-6-80. This was due to the development of ECC microcracks and the energy dissipated by PE fibres pulling out of the matrix.

Effect of thicknesses of the ECC and steel tube

The effect of the thicknesses of the ECC and steel tube on the composite column is shown in Figure 9(b). The initial stiffnesses of columns E^#-GC-6-80 with an ECC thickness of 80 mm (1193.7 kN/mm) and E-GC-10-80 with a tube thickness of 10 mm (1167.8 kN/mm) were 22.3% and 19.7% greater than those of column E-GC-6-80 with 60-mm-thick ECC and 6-mm-thick tube (975.6 kN/mm) in the elastic stage, respectively. The peak loads of columns E^#-GC-6-80 (7700 kN) and E-GC-10-80 (7530 kN) were 20.4% and 17.7% higher than that of E-GC-6-80 (6722 kN), respectively. Increasing the thicknesses of the ECC and steel tube could significantly increase the composite column’s initial stiffness and peak load.

The μ values of columns E^#-GC-6-80 and E-GC-10-80 were 1.38 and 1.68, respectively, which were 9.5% and 33.3% higher than that of column E-GC-6-80, showing that the μ of composite column was positively correlated with the thicknesses of the ECC and steel tube. In addition, the tube thickness provided a more prominent contribution to ductility, as the internal HCFST carried most of the bearing capacity after the peak load. Compared with column E-GC-6-80, the E_d. of columns E^#-GC-6-80 and E-GC-10-80 were increased by 11.9% and 106.2%, respectively. E_d. did not increase noticeably when the ECC thickness was increased, but it did increase significantly by increasing the thickness of the steel tube.

Effect of the compression strength of HSC

Figure 9(c) depicts the effect of the compression strength of HSC. The concrete elastic modulus increased as the compression strength rose (Nematzadeh et al., 2012); therefore, column E-GC-6-80 showed a higher initial stiffness than column E-GC-6-60. Compared to column E-GC-6-60, the load capacity of column E-GC-6-80 did not show a significant increase. This is primarily due to the fact that as compressive strength increased, HSC became more prone to brittle failure under compression, and sudden failure could occur under ultimate loads. Therefore, although the theoretical increase in compressive strength should lead to an increase in load capacity, the actual results often fell short of expectations. The strength indexes of columns E-GC-6-80 and E-GC-6-60 were similar, 1.20 and 1.22, respectively, which demonstrated that the combined effect of the composite column was not significantly affected by the compression strength of HSC. The μ values of columns E-GC-6-80 and E-GC-6-60 were 1.26 and 1.57, respectively. μ decreased by 19.7% when the compression strength grade of HSC increased from C60 to C80, as the brittleness of the concrete increased as the compression strength rose. A similar variation law could be found in E_d, i.e., the E_d. of column E-GC-6-60 (50.7 kJ) was 21.0% higher than that of column E-GC-6-80 (86.4 kJ).

Effect of the tie configuration

Figure 9(d) depicts the effect of the tie configuration on the composite column. The initial stiffness of columns with various tie configurations was approximately the same in the elastic stage. The peak loads of columns E-GC-6-60, E-O-6-60, E-SR-6-60, E-GR-6-60, and E-GS-6-60 were 6471 kN, 4366 kN, 6272 kN, 5767 kN, and 6299 kN, respectively. Column E-GC-6-60 with the tie configuration GC had the highest peak load, which is explained by the composite constraint to the core area provided by the GC leading to the considerable increase in the axial compressive strength of ECC. For rectangular ties, the peak load of column E-SR-6-60 with steel rectangular ties was 8.8% higher than that of column E-GR-6-60 with GFRP rectangular ties. This may be due to premature rupture at the bends of the GFRP rectangular ties. For GFRP ties, the peak load of column E-GS-6-60 with GFRP spiral ties was 9.2% higher than that of column E-GR-6-60 with GFRP rectangular ties. The GFRP rectangular ties showed a tendency to rupture earlier at the bends and had poor lateral restraint performance, resulting in a lower peak load of column E-GR-6-60.

The highest SI (1.22) was observed in column E-GC-6-60, whereas the lowest SI (0.96) was found in column E-O-6-60. For the composite columns with the configuration of single ties, the SI ranged from 1.10 to 1.20. This highlighted the significance of the composite constraint provided by the GC. The μ values for the columns with tie configurations GC, O, SR, GR and GS (E-GC-6-60, E-O-6-60, E-SR-6-60, E-GR-6-60, and E-GS-6-60) were 1.57, 1.12, 1.19, 1.22, and 1.26, respectively. The μ of column E-GC-6-60 with GC was significantly higher than that of columns with a single tie constraint. Column E-O-6-60 had the lowest μ and failed with poor integrity due to the lack of ties. The E_d. values of columns E-GC-6-60, E-O-6-60, E-SR-6-60, E-GR-6-60, and E-GS-6-60 were 50.7 kJ, 19.8 kJ, 28.9 kJ, 30.8 kJ and 35.9 kJ, respectively. Compared with other tie configurations, GC could effectively improve the E_d. of composite columns. In columns with the single tie configuration, the E_d. of column E-GS-6-60 was the highest, as the best restraint provided by the GFRP spiral tie among SR, GR, and GS enhanced the energy dissipation capacity. For column E-O-6-60 without ties, the lowest E_d. was observed.

Load-carrying capacity analysis

Calculation methods in current design codes

At present, there are three main methods for calculating the axial compression bearing capacity of the concrete-encased CFST column in the current design codes (see Table 7): (1) the simple superposition method in American code ACI 318-14 (2014), as expressed in equations (2), and (5)–(7) the superposition method considering the restraint contribution of the steel tube in Japanese code AIJ (1997; 2001), as expressed in equations (8)-(10), and (3) the superposition method considering the stability effect of the column and the restraint effect of the steel tube in Chinese code T/CECS 188-2019 (2019), as expressed in equation (11).

Table 7.

Calculation Methods in the Current Design Codes.

Code	Calculation formula
American code ACI 318-14	$N_{1} = N_{R C} + N_{C F S T}$	(5)
	$N_{R C, 1} = 0.85 f_{c y l, 150, 1} (A_{R C} - A_{y l}) + f_{y l} A_{y l}$	(6)
	$N_{C F S T, 1} = 0.85 f_{c y l, 150, 2} A_{c o r e} + f_{y s} A_{y s}$	(7)
Japanese code AIJ	$N_{2} = N_{R C} + N_{C F S T}$	(8)
	$N_{R C, 2} = 0.85 f_{c y l, 100, 1} (A_{R C} - A_{y l}) + f_{y l} A_{y l}$	(9)
	$N_{C F S T, 2} = 0.85 f_{c y l, 100, 2} A_{c o r e} + (1 + η) f_{y s} A_{y s}$	(10)
Chinese code T/CECS 188-2019	$N_{3} = 0.9 φ [f_{c o, 1} A_{R C} + f_{y l} A_{y l} + f_{c o, 2} A_{c o r e} (1 + 1.8 θ)]$	(11)

Note: $N_{R C}$ and $N_{C F S T}$ are the bearing capacities of the reinforced concrete and CFST, respectively. $f_{c y l, 150, 1}$ and $f_{c y l, 150, 2}$ are the cylinder (d = 150 mm and h = 300 mm) compression strengths of the external and core concrete, respectively. $f_{c y l, 100, 1}$ and $f_{c y l, 100, 2}$ are the cylinder (d = 100 mm and h = 200 mm) compression strengths of the external and core concrete, respectively. According to GB 50010-2010 (2010), $f_{c y l} = 0.76 f_{c u}$ , and $f_{c u}$ is the cubic compression strength. $A_{R C}$ and $A_{c o r e}$ are the cross-sectional areas of the external and core concrete, respectively. $A_{y l}$ and $A_{y s}$ are the cross-sectional areas of the longitudinal bar and steel tube, respectively. $f_{y l}$ and $f_{y s}$ are the yield strengths of the longitudinal bar and steel tube, respectively. η is the constraint coefficient, which is 0.27 in Japanese code AIJ. $φ$ is the compression stability coefficient, and $φ$ is one in the case of the short column. $f_{c o, 1}$ and $f_{c o, 2}$ are the prism compressive strengths of the external and core concrete, respectively, which were converted from the cubic compressive strengths per GB 50010-2010 (2010). θ is the hoop coefficient of the CFST, θ = $f_{y s} A_{y s} / (f_{c o, 2} A_{c o r e})$ .

The applicability of the code methods to ECC-encased HCFST columns was checked by test results in this study. Since ECC and concrete are cement-based materials, ECC strength could be used in the code calculation when involving concrete strength. The calculated and tested bearing capacities of composite columns were compared in Figure 10 and Table 8. The code ACI 318-14 (2014) provided the most conservative prediction, with an average of 21.3% lower than the test value. This was because neither the restraint of ties nor the restraint of the steel tube to the core HSC was considered by the code ACI 318-14 (2014). Although the code AIJ (1997; 2001) and the code T/CECS 188-2019 (2019) considered the restraint effect of the steel tube, the transverse restraint effect of the ties was not considered. The predicted results of codes AIJ (1997, 2001) and T/CECS 188-2019 (2019) were still lower than the test value, with average errors of 14.6% and 20.0%, respectively. The applicability of the abovementioned code methods was relatively poor for the ECC-encased HCFST column.

Figure 10.

Comparison of calculated and tested bearing capacities of composite columns.

Table 8.

Comparison of the Predicted Bearing Capacity and Tested Bearing Capacity.

Column label	$N_{u}$ (kN)	ACI 318-14		AIJ		T/CECS 188-2019
Column label	$N_{u}$ (kN)	$N_{u}^{P 1}$ (kN)	$D_{u}^{P 1}$ (%)	$N_{u}^{P 2}$ (kN)	$D_{u}^{P 2}$ (%)	$N_{u}^{P 3}$ (kN)	$D_{u}^{P 3}$ (%)
E-GC-6-80	6722	5075	24.5	5525	17.8	4994	25.7
E^#-GC-6-80	7700	5844	24.1	6329	17.8	5667	26.4
E-GC-10-80	7530	5700	24.3	6468	14.1	6129	18.6
E-GC-6-60	6471	4776	26.2	5215	19.4	4866	24.8
E-O-6-60	4366	4108	5.9	4549	4.2	4217	3.4
E-SR-6-60	6272	4773	23.9	5212	16.9	4861	22.5
E-GR-6-60	5767	4775	17.2	5213	9.6	4861	15.7
E-GS-6-60	6299	4775	24.2	5215	17.2	4862	22.8
Average value	N/A	N/A	21.3	N/A	14.6	N/A	20.0

Note: $N_{u}$ is the tested bearing capacity of the composite columns. $N_{u}^{P 1}$ , $N_{u}^{P 2}$ , and $N_{u}^{P 3}$ are the predicted results of codes ACI 318-14, AIJ, and T/CECS 188-2019, respectively. $D_{u}^{P 1}$ , $D_{u}^{P 2}$ , and $D_{u}^{P 3}$ are the predicted deviations of codes ACI 318-14, AIJ, and T/CECS 188-2019, respectively.

Proposed calculation method of the bearing capacity of the composite column

The axial compressive strength of concrete will increase when it is restrained laterally, but the current code ignores the lateral restraint effect of ties in calculating the axial compressive bearing capacity of the composite column. However, the lateral restraint performance of the GC was strong and cannot be ignored based on the experimental observation discussed in Section 4.4. Therefore, the existing bearing capacity calculation formulas cannot accurately evaluate the axial compressive bearing capacity of GC-reinforced ECC-encased HCFST columns. This study proposed a calculation method of the bearing capacity of composite columns with GC by combining the composite constraint model and the unified theory of strength of CFSTs.

Two interfaces exist in GC-reinforced ECC-encased HCFST columns, i.e., the external ECC-steel tube interface and steel tube-core HSC interface. P1 and P2 represent the contact stresses at the two interfaces, as shown in Figure 11. Han et al. (Han et al., 2014) and Cai et al. (Cai et al., 2020) numerically analysed concrete-encased CFST columns and discovered that the interface contact stress between the steel tube and external concrete was virtually negligible at peak load, indicating that there was no interaction between the steel tube and the outer structure when the composite column was under the peak load. As a result, the bearing capacity of the composite column could be calculated by adding the bearing capacities of the inner HCFST and the outer GC-reinforced ECC in this study.

Figure 11.

Contact stresses at the composite column interface.

A schematic diagram of the axial load-strain relationship of the outer ECC and inner HCFST is shown in Figure 12, where $ε_{e c}$ and $ε_{H C F S T}$ are the peak strains of ECC and HCFST, respectively. For ECC used in this study, $ε_{e c}$ is between 0.004-0.0042 (Zhou et al., 2015), and $ε_{H C F S T}$ can reach 0.0046 according to equation (12) (Han et al., 2014). It is obvious that the inner and outer parts did not reach the axial limit states at the same time. The composite column reached the peak load when the outer ECC reached the limit state, whereas the axial pressure of the inner HCFST continued to increase. Therefore, a bearing capacity reduction factor α of the HCFST should be introduced, as illustrated in equation (13).

ε_{H C F S T} = 1300 + 12.5 f_{c y l} + (600 + 33.5 f_{c y l}) θ^{0.2}

(12)

N_{u} = N_{R E} + α N_{H C F S T}

(13)

where

N_{R E}

and

N_{H C F S T}

are the bearing capacities of the GC-reinforced ECC and the HCFST, respectively.

Figure 12.

Axial load-strain curves of ECC and HCFST under axial compression.

The bearing capacity of the external GC-reinforced ECC was calculated using the composite constraint model of GFRP ties (Zhang et al., 2022). According to the different restraint conditions, the external ECC section was divided into three restraint areas: double restraint area, single restraint area and no restraint area, as shown in Figure 13. Then, the following formula can be used to calculate the bearing capacity of the external GC-reinforced ECC.

N_{R E} = f_{e 1} (A_{1} - A_{y l 1}) + f_{e 2} (A_{2} - A_{y l 2}) + f_{e o} A_{3} + f_{y l} A_{y l}

(14)

where f_e1, f_e2, and f_eo are the peak stresses of ECC in the dual, single, and no restraint areas, respectively.

A_{1}

A_{2}

, and

A_{3}

are the ECC cross-sectional areas of the dual, single, and no restraint areas, respectively.

A_{y l 1}

and

A_{y l 2}

are the cross-sectional areas of the longitudinal bars in the dual and single restraint areas, respectively.

A_{l y}

is the cross-sectional area of the total longitudinal bars,

A_{l y}

A_{l y 1}

A_{l y 2}

Figure 13.

Constrained partitioning of ECC sections.

According to the Mander model (Mander, 1988), the peak stresses of ECC in the dual restraint area and single restraint area are calculated by equations (15) and (16), respectively.

f_{e 1} = f_{e o} (2.254 \sqrt{1 + \frac{7.94 f_{l}^{'}}{f_{e o}}} - 2 \frac{f_{l}^{'}}{f_{e o}} - 1.254)

(15)

f_{e 2} = f_{e o} (2.254 \sqrt{1 + \frac{7.94 f_{l r}^{'}}{f_{e o}}} - 2 \frac{f_{l r}^{'}}{f_{e o}} - 1.254)

(16)

where

f_{l}^{'}

and

f_{l r}^{'}

are the lateral restraint stresses of GC and GR, respectively, and the calculation formulas are as follows.

f_{l}^{'} = f_{l s}^{'} + f_{l r}^{'}

(17)

f_{l s}^{'} = k_{e 1} \frac{2 f_{b s} A_{f}}{S d_{s}}

(18)

f_{l r}^{'} = (\frac{A_{s x}}{s d_{c}} + \frac{A_{s y}}{s b_{c}}) \frac{k_{e 2} f_{b r}}{2}

(19)

where

f_{l s}^{'}

is the lateral restraint stress of GS.

k_{e 1}

and

k_{e 2}

are the restraint effectiveness coefficient for the GS and GR, which were 0.85 and 0.57, respectively (Zhang et al., 2022). As per CSA S806-12 CSA S806-12.

f_{b s}

is the GS’ bend strength, which should be less than 0.004 times the GFRP ties’ elastic modulus E (see Table 5).

A_{f}

is the GS’ cross-sectional area. S is the center to center spacing of the tie in the vertical direction.

d_{s}

is the core concrete’s diameter within the spiral.

f_{b r}

is the tensile strength of GR bent part (see Table 5) and should be less than 0.004 E.

A_{s x}

and

A_{s y}

are the total tie areas in the x and y directions, respectively, which are two perpendicular directions along the rectangular tie centerline.

b_{c}

and

d_{c}

are the core dimensions of the tie centerline in the x and y directions, respectively,

b_{c}

≥

d_{c}

In the unified theory of strength, the steel tube and the HSC were considered the same material (Zeng et al., 2022). The bearing capacity HCFST was calculated as follows:

N_{H C F S T} = A_{H C F S T} (1.212 + B θ + C θ^{2}) f_{H S C}

(20)

B = \frac{0.716 f_{y s}}{213} + 0.974

(21)

C = - \frac{0.104 f_{c y l, 150, 2}}{14.4} + 0.031

(22)

where

A_{H C F S T}

is the total cross-sectional area of the HCFST,

A_{H C F S T} = A_{H S C} + A_{y s}

. B and C are the hoop effect’s influence coefficients, as expressed in equations (21) and (22).

The tested bearing capacity of composite columns and the calculated HCFST and external reinforced ECC were used to determine the reduction factor α, as listed in Table 9. The reduction factor α for the GC-reinforced ECC-encased HCFST columns ranged from 0.945 to 0.963, and a conservative value of 0.9 was recommended with a predicted deviation of 7.8%, as expressed in equation (23).

N_{u} = N_{R E} + 0.9 N_{H C F S T}

(23)

Table 9.

Calculation Results of the Reduction Factor α.

Column label	$N_{R E}$ (kN)	$N_{H C F S T}$ (kN)	$N_{u}^{E}$ (kN)	$α$
E-GC-6-80	2873	3638	6722	0.945
E^#-GC-6-80	3921	3638	7700	0.963
E-GC-10-80	2873	4471	7530	0.960
E-GC-6-60	2873	3451	6471	0.959

Conclusions

This paper proposed a new GC-reinforced ECC-encased HCFST column. The axial compression test was conducted to discuss the effects of the cement-based material, the thicknesses of the ECC and steel tube, the compression strength of HSC, and the tie configuration. Based on the composite constraint model and the unified theory of strength, a calculation method for the bearing capacity of GC-reinforced ECC-encased HCFST columns was proposed. The following key conclusions are drawn.

(1) GC-reinforced ECC-encased HCFST columns exhibited a ductile failure mode, characterized by the gradual formation of microcracks that ultimately led to the development of a main crack. In contrast, concrete-encased HCFST columns displayed brittle failure. This difference arose from the PE fibres contained in ECC, which effectively dispersed stress, thereby limiting crack propagation.

(2) Compared to specimen N-GC-6-80, specimen E-GC-6-80 exhibited increases in bearing capacity, ductility, and energy dissipation of 14.2%, 7.7%, and 81%, respectively, with the most significant improvement observed in energy dissipation. This enhancement was primarily due to ECC’s capacity to absorb substantial energy through the formation and propagation of microcracks during loading.

(3) Increasing the thickness of the steel tube significantly enhanced the ductility and energy dissipation capacity. This phenomenon arose from the emphasis on evaluating ductility and energy dissipation based on mechanical performance after the peak load, during which the internal HCFST bore the main load, so the influence of the steel tube wall thickness was more obvious.

(4) Compared to other tie configurations, GC demonstrated superior performance in enhancing both ductility and energy dissipation capacity, highlighting the excellent composite confinement capability of this new tie configuration.

(5) Current calculation methods underestimated the bearing capacity of GC-reinforced ECC-encased HCFST columns. An improved method was proposed, based on the composite constraint model and unified strength theory.

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: This work is supported by National Program on Key R&D Project of China (2022YFE0210500) and Open Research Project of China-Pakistan Belt and Road Joint Laboratory on Smart Disaster Prevention of Major Infrastructures (2024CPBRJL-02), Open Project Funded by the Key Laboratory of Concrete and Pre-stressed Concrete Structures of Ministry of Education at Southeast University (CPCSME2023-10), and sub project of the Key Project of the National Development and Reform Commission of China (202203001).

ORCID iDs

Ye Liu

Zhang Pu

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