Flexural behavior of textile reinforced mortar-autoclaved lightweight aerated concrete composite panels

Abstract

To improve the deficiencies of prefabricated autoclaved lightweight aerated concrete (ALC) panel such as susceptibility to cracking and low load-bearing capacity, a textile-reinforced mortar-autoclaved lightweight aerated concrete (TRM-ALC) composite panel was developed in this study. One group of reference ALC panels and five groups of TRM-ALC panels were fabricated and subjected to four-point flexural tests. TRM was applied on the tensile side of the ALC panels to create TRM-ALC. The variable parameters were the plies of textile (one or two), type of textile (basalt or carbon), and whether the matrix (without textile) was applied on the compression side of panel. The results showed that a bonding only 8-mm-thick TRM layer on the surface of the ALC panel could increase the cracking load by 180%−520%. The flexural capacity of the TRM-ALC panel increased as the number of textile layers increased. Additional reinforcement of the matrix on the compressive side could further enhance the stiffness and ultimate load-bearing capacity of the TRM-ALC panel. Such panels with basalt textile failed in flexural mode, with the rupture of fabric mesh. Those with carbon textile failed in shear mode due to the ultra-high tensile strength of carbon. In addition, analytical models related to the different failure modes were presented to estimate the ultimate load-carrying capacity of the TRM-ALC panels.

Graphical abstract

Keywords

prefabricated autoclaved lightweight aerated concrete panel / textile-reinforced mortar / cracking load / flexural capacity / stiffness

1 Introduction

Pre-assembled structures are increasingly attracting attention due to the fast construction efficiency, high building quality, and reduction of labor demand [1,2]. Prefabricated autoclaved lightweight aerated concrete (ALC) panels are being widely used as floor slabs, roofs, interior partitions, and peripheral walls in assembly buildings because of advantages including ultra-light weight, heat and sound insulation, high fire ratings, and ease of installation [3]. Nonetheless, the disadvantages, including susceptibility to cracking and low load-bearing capacity, limit the structural application of these panels [4,5].

Some researchers have proposed externally bonding fiber-reinforced polymer (FRP) composites on the surface of the ALC panels to enhance their structural performance [4–7]. The results showed that the application of FRP significantly enhanced the flexural stiffness and load-carrying capacity. However, this solution has some drawbacks, including poor resistance to fire, low ductility level resulting from the brittle fracture of FRP or debonding, and high cost.

Textile-reinforced mortar (TRM) is a composite material made of fiber textile and inorganic matrix. Carbon textile, basalt textile, and glass textile were generally chosen as the reinforcing textile of the TRM [8]. The matrix of the TRM can be cement-based mortar or fine-grained concrete [8–11]. TRM material exhibited high strength and large deformation under tension and bending loads [9,12–15]. Owing to the non-flammable nature of the inorganic matrix, the fire resistance of TRM is much better than that of FRP [16,17]. Furthermore, the TRM possesses better bonding behavior with concrete [18]. Therefore, TRM can be an alternative to FRP for improving the structural behavior of concrete elements. Some novel precast sandwich panels, made of a core layer of polymeric rigid thermal insulation and two external layers of TRM, have been created [19]. However, the use of the TRM in the ALC panel is rarely reported in the literatures. Schladitz et al. [20] and Koutas and Bournas [21] reported that the stiffness and flexural load-bearing capacity of RC slabs were significantly increased after applying the TRM on the tensile side of slabs. However, some shortcomings have been observed in the application of TRM, such as the low cracking strength and large width of cracks resulting from the brittle fracture of the mortar, which limit the application of TRM [22–24]. Researches [24–26] have indicated that the incorporation of short fibers could improve the toughness of the matrix, and further increase the cracking strength of the TRM and reduce the width of cracks. To date, many investigations [21–24] have focused on the mechanical performance of the RC beams strengthened by TRM without short fibers. However, a few researches have also been reported on application of short-fiber-added TRM on the panel element. To confirm the influence of the TRM with short fibers on the anti-cracking performance and load-carrying capacity of the ALC panel, corresponding experimental study and theoretical analysis should be explored.

In this work, TRM with short polyvinyl alcohol (PVA) fibers was applied on the surface of the ALC panel to form a kind of novel lightweight panel (TRM-ALC). The TRM-ALC panel was expected to combine the advantages of both ALC panel and the TRM material, including light weight, high-strength and high cracking resistance, making it suitable to be used for structural or non-structural members in pre-assembly buildings. This study aimed at understanding the structural response of the developed TRM-ALC composite panel. The effects of the number of plies of textiles, type of textile, and additional application of the matrix on the compression side of the TRM-ALC panel, on the mechanical properties of the TRM-ALC panels were investigated. The failure mode, load–deflection curve, strength, ductility, and energy absorption capacity of the TRM-ALC panels were presented and discussed. Finally, to predict the mechanical behavior of the TRM-ALC panels, a model was established to predict the cracking load of the panels based on force equilibrium and strain compatibility. Two calculation methods were presented to respectively estimate the flexural and shear capacity of the TRM-ALC panels.

2 Experimental program

2.1 Specimen design

Four-point flexural tests were performed on one group of reference ALC panels and five groups of TRM-ALC panels, as listed in Tab.1. The dimensions of the reference ALC panels, were 2000 mm × 600 mm × 100 mm (length × width × height). Cost-effective basalt textile and high-strength carbon textile were chosen as potential reinforcing textiles of the TRM layer. The reference ALC panels were named as Group P-R. Further groups, named P-B1, P-B2, and P-C1, were strengthened with the TRM layer on the tensile side. The TRM layer was intended to bear the tensile stress, which could increase the load-carrying capacity of the tensile side of the panel. One and two plies of basalt textiles were used in the TRM layer in groups P-B1 and P-B2, respectively. One ply of carbon textile was used in group P-C1. To avoid the premature crushing of top ALC before yield of longitudinal tensile reinforcements, an inorganic matrix layer was applied on the compressive side of the ALC panel to enhance the strength of the top compressive region. For groups P-B2D and P-C1D, the TRM and matrix layer were applied on the tensile and compressive sides of the ALC panels, respectively. Longitudinal steel bars in both the compression and tension sides consisted of three plain bars. The thickness of the protective layer of the inside steel bars was 25 mm in the ALC panels. The textile grid only required a 2−3 mm anchoring thickness due to the anti-corrosion effect of fiber textiles in mortar. To minimize the weight and thickness of the TRM-ALC panel, the thickness of both the TRM layer on the tensile side and the matrix layer on compressive side was designed as 8 mm. The geometric details of the reference ALC panels and TRM-ALC panels are shown in Fig.1.

The construction steps of the TRM-ALC panels were as follows: 1) the surface of the prefabricated ALC panel was made wet with clean water; 2) a 2−3 mm thick matrix layer was applied by trowel on the surface of panel; 3) textile was pressed into the matrix; 4) another layer of matrix was applied to completely cover the fabric mesh. For the TRM-ALC panels with more than one layer of textile, steps (3) and (4) were repeated. Wooden formworks were positioned on either side of the panel to ensure that the thickness of the TRM layer was 8 mm. After being cured by spraying water for 28 d, the panels were loaded until failure.

2.2 Materials

2.2.1 ALC and steel bar

The ALC panels produced by Shaanxi Huda Building Materials Co., Ltd. were chosen to be the tested specimens. According to GB/T 11969-2020 [27], six cubes with dimensions of 100 mm × 100 mm × 100 mm were cut from an ALC panel to test its mechanical properties. The cubic compressive strength and splitting tensile strength of ALC were 5.1 MPa (COV = 3.5%) and 0.57 MPa (COV = 2.0%), respectively.

The diameter of the steel bars used in the ALC panels was 5 mm. The average yield strength (f_y) and ultimate strength (f_u) of the steel bars were 631 MPa (COV = 4.1%) and 742 MPa (COV = 2.5%), respectively.

2.2.2 Textile

Basalt and carbon textile are shown in Fig.2. The textiles were both in the form of bidirectional grids. The mechanical properties of the two textile grids were determined in accordance with GB/T 36-262 standards [28], as presented in Tab.2.

2.2.3 Matrix

Zhang et al. [26] investigated the effect of short PVA fiber volume fraction on the anti-cracking capacity of the TRM composite. The result showed that the TRM specimens with a 1.2% PVA fibers exhibited high cracking strength. Based on that result [26], short PVA fibers were mixed into the matrix at a volume fraction of 1.2%. The amounts of the materials used to produce the matrix of the TRM are listed in Tab.3. The length and diameter of the PVA fiber were 12 mm and 39 μm, respectively. The elastic modulus and tensile strength of the PVA fibers were 40 GPa and 1600 MPa, respectively.

The tensile performance of the matrix was obtained by referring to the tested method for fiber-reinforced concrete in Ref. [29]. The tensile stress–strain curve of the matrix is shown in Fig.3. After being cured for 28 d, the measured average cracking strength and cracking strain of the TRM matrix, respectively, were 3.5 MPa (COV = 1.2%) and 0.058% (COV = 4.5%); the measured average tensile strength and ultimate tensile strain respectively were 4.0 MPa (COV = 1.7%) and 0.33% (COV = 8.3%). The measured average cubic compressive strength of the matrix at 28-curing d was 53.8 MPa (COV = 4.3%), which was obtained from three cubes with dimensions 100 mm × 100 mm × 100 mm.

2.3 Test setup

All panels were subjected to four-point bending, as shown in Fig.4. The vertical concentrated load was exerted by a hydraulic jack with a maximum load of 1000 kN. This load was evenly distributed to two loading points by a spreader steel beam. The loading rate was 0.2 mm/min. The total distance between the specimen supports was 1800 mm. The effective length of the flexural span was 300 mm. Deflection measurements were taken at the two loading points and the mid-span point of the panels by three linear variable differential transformers (LVDTs).

3 Test results and discussion

3.1 Failure mode

The final crack pattern of the specimens is shown in Fig.5.

1) Reference group P-R

For group P-R, the first crack occurred on the tensile side of mid-span zone of the panels. After cracking, the tensile stress sustained by the ALC was transferred to the longitudinal tensile steel bars. With an increase in the load, the number of cracks increased, and the existing cracks extended upwardly. Subsequently, the compressive depth of the panel gradually decreased, and the strain at the top of compressive region gradually increased. Finally, the ALC near the loading point was crushed. The reference panels (P-R) failed in a typical flexural mode.

2) Groups P-B1, P-B2, and P-B2D

For groups P-B1, P-B2, and P-B2D, cracks first appeared on the TRM layer of the flexural span. As the loading increased, more cracks formed on the TRM layer, and one of which extended upward and developed into a main crack. Subsequently, the textiles in the tensile side were ruptured, the TRM layer broke into two portions along the main crack, and the load gradually decreased. Groups P-B1, P-B2, and P-B2D failed due to the rupture of textiles, with a major flexural crack but without top surface crushing. Compared with the reference ALC panels, the cracks on bottom surface of the TRM-ALC panels were finer, denser and more numerous.

3) Groups P-C1 and P-C1D

For groups P-C1 and P-C1D, some fine cracks first appeared on the bottom TRM layer. As the applied load increased, the cracks on the shear span region extended obliquely toward the nearby loading point. Then, the main shear crack penetrated the shear-span cross section and suddenly widened. Finally, the TRM-ALC panels with carbon textile failed in shear mode. This was due to the enhanced flexural capacity of the panels making the specimens vulnerable to shear failure.

Compared with the reference ALC panels, the average crack spacing of the TRM-ALC was significantly smaller. The number of cracks in the pure flexural region of TRM-ALC panels was 1.7−2.8 times of that in the ALC panels. The bridging effect of short PVA fibers and textiles grids could transfer the extra tensile stress and restrain the width of cracks. Adding a second layer of basalt textile on the tensile side of the TRM-ALC panels had almost no effect on the crack pattern. Shear cracks opened on the shear-span region of the TRM-ALC panel due to the fact that flexural capacity of the panel could be significantly enhanced by substituting basalt textile with carbon textile.

3.2 Load–deflection response

The load–deflection analysis curves for all panels are presented in Fig.6.

As shown in Fig.6, the load−deformation analysis curve of the specimens can be divided into three different phases: 1) un-cracked phase; 2) cracking up to peak phase; 3) post-peak phase. During the first phase, the curve gradients for the TRM-ALC panels were higher than those of the ALC panels. This indicated that bonding a TRM thin layer can significantly increase the initial stiffness of the ALC panel. During the second stage, after cracking, the load of the reference ALC panels showed an abrupt drop. This may be due to the brittleness of ALC, suggesting that the ALC in the tension side was out of work when cracked. However, due to the bridging stress of textile and PVA fibers, the load of the TRM-ALC panels steadily increased after cracking, unlike the reference ALC panels. After cracking, the tensile steel bars and textiles were activated to carry the tensile stress. Therefore, the load increase rate of the TRM-ALC panels was faster than that of the ALC panels. At the post-peak stage, the strength attenuation of the TRM-ALC panels was relatively slower than that of the reference ALC panels. This can be attributed to the local rupture of the single basalt/carbon bundle.

3.3 Strength

Tab.4 lists cracking load (P_cr), peak load (P_m), yield deflection (Δ_y), ultimate deflection (Δ_u), and ductility factor (μ). The yield deflection was obtained by the farthest point method in Ref. [30]. The ultimate deflection was defined as the deflection corresponding to 85% of the peak load at the post-peak stage.

The cracking load of the TRM-ALC panels was 180%–520% above that of the ALC panels. The cracking strength of the matrix was about seven times that of the ALC. The high cracking strength of the matrix contributed to improvement of the cracking load of the panels. As the number of the textile layers increased, the cracking load of the TRM-ALC panels increased.

Compared with that of group P-R, the peak load of group P-B1, P-B2, and P-B2D increased 94.4%, 144.5%, and 209.4%, respectively. The TRM layer on the tensile side provided an additional resistance to tensile force; thus, the flexural capacity of the TRM-ALC panels was higher than that of the ALC panel. As the number of the textile layers increased, the flexural capacity of the TRM-ALC panels increased; however, the effective utilization of the textiles decreased. This is due to the dominant factor affecting peak load of the TRM-ALC being the breaking of the bottom-most textile grid. The flexural capacity of the double-face strengthened TRM-ALC panels was higher than that of the single-face strengthened TRM-ALC panels. This is due to the high compressive strength of the matrix causing increase of the load-bearing capacity offered by the compression region of the panel.

The average peak load of group P-C1 increased 216.4% above that of group P-R. The ultra-high tensile strength of carbon textile improved the flexural capacity of the panel, leading to that the flexural capacity exceeding the shear capacity. Therefore, group P-C1 failed in shear mode. The peak load of group P-C1D increased 17.8% in comparison with group P-C1. The double-face strengthening was also favorable in improving the shear capacity of the panels.

The gain in load-bearing capacity of group P-C1 was significantly higher than that of group P-B2 even though the textile reinforcement ratio of P-C1 was smaller than that of P-B2. This is due to the elastic modulus and tensile strength of carbon textile being far higher than those of basalt textile. Comparing the increase of load-bearing capacity of group P-B2 with that of group P-B1, the peak load of the TRM-ALC panels with three layers of basalt textile was still smaller than that of the TRM-ALC with one layer carbon textile. For the same cost, the TRM-ALC panel with carbon textile possesses higher load-carrying capacity than that of the same panel with basalt textile.

3.4 Deformation and ductility

The yield deflection and ultimate deflection of the TRM-ALC panels were smaller than those of the reference ALC panels. The yield point and ultimate point of the ALC panels were dominated by the yield of the longitudinal tensile reinforcements and the crushing of the top ALC, respectively. For the TRM-ALC panels, the increase of the section stiffness resulted in the reduction of the yield deflection. The ultimate points of the TRM-ALC panels were controlled by the fracture of the bottom textile. The deformability of the textile was much lower than that of the steel bars. Therefore, bonding the TRM layer on the ALC panels resulted in a decreased deformation ability. As the amounts of the textile layer increased, the ultimate mid-span deflection and ductility factor of the TRM-ALC panels increased.

The ductility factor of group P-C1 was lower than that of group P-B2 due to the shear failure mode. Compared with group P-B2, group P-B2D exhibited a lower ultimate deflection while a higher ductility factor. This may be attributed to the matrix on the compressive side increasing the stiffness of the panels, resulting in the decrease of the yield deflection.

3.5 Energy absorption capacity

Fig.7 depicts the comparison of the average absorbed energy value of each group. The area under a load–deflection curve represents the absorbed energy of the panel. Compared with group P-R, the absorbed energy values of group P-B1, P-B2, and P-B2D showed increases of 54.3%, 132.3%, and 160.0%, respectively. Applying only 8-mm thick TRM layer on the ALC panels could significantly increase the energy absorption capacity of the panels. With the increase in the number of textile layers, the absorbed energy value of the TRM-ALC panels increased. Compared with group P-R, the absorbed energy values of group P-C1 and P-C1D saw their absorbed energy values increase by 194.1% and 227.9%, respectively. The carbon-TRM strengthening scheme achieved a greater improvement of the energy absorption capacity, compared to the basalt-TRM strengthening scheme.

4 Estimating bearing capacity

To predict the mechanical performance of the TRM-ALC composite panel, the following assumptions are applied: 1) the strain distribution in the panel section is linear; 2) the stress–strain relationship models of the ALC, steel bar and fiber textile are shown in Fig.8; 3) the tensile strengths of the ALC and matrix are only considered before cracking. These assumptions have also been made in Refs. [7,31].

4.1 Cracking load

When the tensile strain of the matrix at the tensile edge reaches the cracking strain (ε_cr = 0.00058), the TRM-ALC panels will crack. In the cracking state, the strain and stress distributions on the cross-section of the TRM-ALC panel are shown in Fig.9. The force equilibrium can be expressed as Eq. (1).

(1)C_{\mathrm{c}}+C_{\mathrm{s}}^{\prime }=T_{\mathrm{c}\mathrm{t}}+T_{\mathrm{s}}+T_{\mathrm{t}}+T_{\mathrm{m}},

where C_{\mathrm{c}}, C_{\mathrm{s}}^{\prime }, T_{\mathrm{c}\mathrm{t}}, T_{\mathrm{s}}, T_{\mathrm{t}}, and T_{\mathrm{m}} are the forces afforded by compressive ALC, compressive steel bars, tensile ALC, tensile steel bars, textile grid, and matrix, respectively.

According to the strain compatibility, when the tensile strain of the matrix at the bottom tensile edge ({\epsilon }_{\mathrm{c}\mathrm{r}}) is given, then the tensile strain of ALC in the tensile side ({\epsilon }_{\mathrm{c}\mathrm{t}}), the tensile strain of the tensile steel bar ({\epsilon }_{\mathrm{s}\mathrm{t}}) and textile ({\epsilon }_{\mathrm{f}\mathrm{t}}), and the compressive strain of ALC at the compressive edge ({\epsilon }_{\mathrm{c}}) and compressive steel bar ({\epsilon }_{\mathrm{s}}^{\prime }), can all be determined:

(2)\{\begin{array}{rl}& {\epsilon }_{\mathrm{c}}=\frac{x_{\mathrm{a}}}{h+t-x_{\mathrm{a}}}{\epsilon }_{\mathrm{c}\mathrm{r}},\\ & {\epsilon }_{\mathrm{s}}^{\prime }=\frac{x_{\mathrm{a}}-a_{\mathrm{s}}^{\prime }}{h+t-x_{\mathrm{a}}}{\epsilon }_{\mathrm{c}\mathrm{r}},\\ & {\epsilon }_{\mathrm{c}\mathrm{t}}=\frac{h-x_{\mathrm{a}}}{h+t-x_{\mathrm{a}}}{\epsilon }_{\mathrm{c}\mathrm{r}},\\ & {\epsilon }_{\mathrm{s}\mathrm{t}}=\frac{h-x_{\mathrm{a}}-a_{\mathrm{s}}}{h+t-x_{\mathrm{a}}}{\epsilon }_{\mathrm{c}\mathrm{r}},\\ & {\epsilon }_{\mathrm{f}\mathrm{t}}=\frac{h-x_{\mathrm{a}}+t/2}{h+t-x_{\mathrm{a}}}{\epsilon }_{\mathrm{c}\mathrm{r}},\end{array}

where x_a is the compressive depth of the panel at the cracking state; h is the thickness of the ALC panel; t is the thickness of the TRM layer; a_{\mathrm{s}} is the distance from the centroid of the tensile steel bars to the tensile edge of the ALC; a_{\mathrm{s}}^{\prime } is the distance from the centroid of the compressive steel bars to the compressive edge.

Substituting Eq. (2) into Eqs. (1) and (3) can be obtained.

(3)\begin{array}{rl}& \frac{1}{2}{E}_{\mathrm{c}}{\epsilon }_{\mathrm{c}}b{x}_{\mathrm{a}}+A_{\mathrm{s}}^{\prime }{E}_{\mathrm{s}}{\epsilon }_{\mathrm{c}}^{\prime }\\ =& {\int }_0^{h-x_{\mathrm{a}}}{\sigma }_{\mathrm{c}\mathrm{t}}{\epsilon }_{\mathrm{c}\mathrm{t}}b\mathrm{d}y+A_{\mathrm{s}}{E}_{\mathrm{s}}{\epsilon }_{\text{st}}+n{A}_{\mathrm{f}}b{E}_{\mathrm{f}}{\epsilon }_{\mathrm{f}\mathrm{t}}+tb{f}_{\mathrm{m}\mathrm{c}\mathrm{r}},\end{array}

where E_{\mathrm{c}}, E_{\mathrm{s}}, and E_{\mathrm{f}} are the elastic modulus of the ALC, steel bar, and textile, respectively; b is the width of the panel; n is the number of textile grid layers; A_{\mathrm{f}} is the section area of textile within 1 m spacing; f_mcr is the cracking strength of the TRM matrix.

The compressive depth at the cracking state x_a can be calculated using Eq. (3). Then, the cracking moment (M_cr) of the TRM-ALC panels can be obtained using Eq. (4).

(4)\begin{array}{rl}{M}_{\mathrm{c}\mathrm{r}}=& {\int }_0^{h-x_{\mathrm{a}}}{\sigma }_{\mathrm{c}\mathrm{t}}{\epsilon }_{\mathrm{c}\mathrm{t}}b\mathrm{d}y\frac{h}{2}+A_{\mathrm{s}}{E}_{\mathrm{s}}{\epsilon }_{\mathrm{s}\mathrm{t}}(h-a_{\mathrm{s}}-\frac{x_{\mathrm{a}}}{2})\\ & +n{A}_{\mathrm{f}}b{E}_{\mathrm{f}}{\epsilon }_{\mathrm{f}\mathrm{t}}(h-\frac{x_{\mathrm{a}}}{2}+\frac{t}{2})+tb{f}_{\mathrm{m}\mathrm{c}\mathrm{r}}(h-\frac{x_{\mathrm{a}}}{2}+\frac{t}{2}).\end{array}

Therefore, the cracking load (P_cr) can be derived by Eq. (5).

(5)P_{\text{cr}}=2M_{\text{cr}}/l_{\mathrm{s}},

where l_s is the length of the shear span region.

4.2 Flexural capacity

The flexural failure of the TRM-ALC panels was dominated by the fracture of the textile. According to the flexural failure mode of the TRM-ALC panels (Fig.5), the compressive force can be assumed to be concentrated at or near the top of the panels. The peak moment of the TRM-ALC panels (M_u) can be calculated using the model associated with this flexural failure mode, from Ref. [32]. This analytical model is shown in Fig.10. When the tensile strain of the textile reaches the ultimate strain (ε_tu), the TRM-ALC panels will reach peak load state. The peak moment of the TRM-ALC panels (M_u) can be calculated by Eq. (6). The strain of the steel bar and textile at peak state are listed in Eq. (7).

(6)\begin{array}{rl}{M}_{\mathrm{u}}=& A_{\mathrm{s}}{\sigma }_{\text{s,p}}^{\prime }(a_{\mathrm{s}}^{\prime }/2)+A_{\mathrm{s}}{\sigma }_{\text{st,p}}(h-a_{\mathrm{s}}^{\prime }/2-a_{\mathrm{s}})\\ & +n{A}_{\mathrm{f}}b\alpha f_{\mathrm{f}}(h-a_{\mathrm{s}}^{\prime }/2+t/2),\end{array}

(7)\{\begin{array}{rl}& {\epsilon }_{\text{s,p}}^{\prime }=\frac{a_{\mathrm{s}}^{\prime }/2}{h-a_{\mathrm{s}}^{\prime }/2+t/2}{\epsilon }_{\text{tu}}\\ & {\epsilon }_{\text{st,p}}=\frac{h-a_{\mathrm{s}}^{\prime }/2_{\mathrm{s}}-a_{\mathrm{s}}}{h-a_{\mathrm{s}}^{\prime }/2+t/2}{\epsilon }_{\text{tu}}\\ & {\epsilon }_{\text{ft,p}}={\epsilon }_{\text{tu}}\end{array},

where {\sigma }_{\text{s,p}}^{\prime }, and {\sigma }_{\text{s,p}} are the stress of the upper and lower steel bars at the peak load state, respectively. f_f is the tensile strength of grid. α is the equivalent reduction coefficient of the tensile strength of textile. α can be calculated by the method in Ref. [26]. As the number of textile layers increased, the effective utilization rate of textile decreased. \alpha = 0.005e^{\left(2.2\sqrt{\beta /\nu }\right)}+0.53, v is the ratio of the cross-sectional area of the textile grids to the TRM layer, β is the PVA short-fiber volume fraction.

The adopted stress–strain constitutive model of the steel bar is presented in Eq. (8).

(8)\{\begin{array}{ll}{\sigma }_{\mathrm{s}}^{\mathrm{\prime }}=E_{\mathrm{s}}{\epsilon }_{\mathrm{s}},& 0\le {\epsilon }_{\mathrm{s}}\le {\epsilon }_{\mathrm{y}},\\ {\sigma }_{\mathrm{s}}=f_{\mathrm{y}},& {\epsilon }_{\mathrm{y}}\le {\epsilon }_{\mathrm{s}},\end{array}

where f_{\mathrm{y}} and {\epsilon }_{\mathrm{y}} are the yield strength and yield strain of the steel bars, respectively. {\epsilon }_{\mathrm{y}}=0.0022.

The flexural capacity (P_f) can be derived by Eq. (9).

(9)P_{\mathrm{f}}=2M_{\mathrm{u}}/l_{\mathrm{s}}.

4.3 Shear capacity

An analytical model for calculating the shear capacity of RC members without stirrups was proposed in Ref. [33]. In this model (Fig.11), the shear capacity of RC members without stirrups was dominated by materials in two locations: a) the concrete in shear-compression zone not penetrated by shear cracks; and b) the concrete in shear-tension zone above the neutral axis and penetrated by shear cracks.

The shear capacity of the TRM-ALC composite panels (V_u) can obtain by summing the shear contributions of the concrete in shear−compression zone and the concrete in shear−tension zone, as expressed in Eq. (10).

(10)V_{\mathrm{u}}=V_{\text{sc}}+V_{\text{st}} =0.5f_{\mathrm{c}}^{\prime }bc{\left(\frac{x}{c}\right)}^2+ f_{\mathrm{t}}bc\left(1-\frac{x}{c}\right),

where V_sc and V_st are the shear contributions from the shear−compression zone and shear−tension zone, respectively; d_v is the effective depth of the RC member; x is the depth of the shear−compression zone; c is the depth of the concrete compression zone; c can be attained based on the strain compatibility and linear-elastic bending theory; c=d_{\mathrm{v}}\left[\sqrt{(n^{\prime }\rho {)}^2+2n^{\prime }\rho }-n^{\prime }\rho \right]; n' is the ratio of elastic modulus between the longitudinal reinforcements and concrete, n^{\prime }= E_{\mathrm{s}}/E_{\mathrm{c}}; \rho is the ratio of longitudinal tensile reinforcements. For the single-face strengthened panel, d_v = h + t/2. For the double-face strengthened panel, d_v = h + 1.5t. The matrix on the compression zone can contribute a part of shear capacity. This can be considered in the shear contribution of the shear−compression zone. For the double-face strengthened TRM-ALC panel, V_{\text{sc}}=0.5f_{\mathrm{m}}^{\prime }bt{\left({\displaystyle \frac{x}{c}}\right)}^2+0.5f_{\mathrm{c}}^{\prime }b\left(c-t\right) {\left({\displaystyle \frac{x}{c}}\right)}^2. f_{\mathrm{m}}^{\prime } is the compressive strength of the matrix. f_{\mathrm{c}}^{\prime } is the compressive strength of ALC. For TRM-ALC panels, \rho ={\displaystyle \frac{A_{\mathrm{s}}}{b{h}_0}}+{\mu }^{\prime }{\displaystyle \frac{nb{A}_{\mathrm{f}}}{b{d}_{\mathrm{v}}}}. {\mu }^{\prime }=E_{\mathrm{f}}/E_{\mathrm{s}}.

With the increase of the shear span-to-depth ratio (λ), the shear failure mode of the RC members changes from shear−compression mode to diagonal-tension mode. Therefore, as the shear span-to-depth ratio increases, the height of shear−compression zone gradually decreases, while the height of shear−tension zone gradually increases. The relationship between the coefficient of shear−compression zone height ({\displaystyle \frac{x}{c}}) and the shear span-to-depth ratio (λ) is {\displaystyle \frac{x}{c}}={\displaystyle \frac{1}{\lambda }} [33]. λ = l_s/d_v.

Tab.5 displays the predicted results of the cracking load and ultimate load-bearing capacity for all TRM-ALC panels. It can be clearly seen that the calculated loads are in good agreement with experimental data. The models presented in this study are reliable for predicting the cracking loads and peak loads of the TRM-ALC panels.

5 Conclusions

In this study, the structural response of the novel TRM-ALC composite panels was experimentally investigated. The main conclusions are summarized below.

1) The TRM-ALC panel showed significant improvements in the anti-cracking behavior and load-carrying capacity over those of the ALC panel.

2) As the number of the textile layers increased, the cracking strength and flexural strength of the TRM-ALC panels increased. The effective utilization of the textiles decreased with the increase of the textile layers.

3) The TRM-ALC panels with basalt textile failed in flexural mode, characterized by the rupture of the textile grid. The TRM-ALC panels with carbon textile failed in shear mode, due to the ultra-high tensile strength of the carbon textile and a low shear-span ratio. At the same cost, the ultimate load-carrying capacity of the TRM-ALC panels with carbon textile is higher than that of the TRM-ALC panels with basalt textile; however, the deformation capacity of the TRM-ALC panels with carbon textile is smaller. Carbon textile is better in terms of combined effect.

4) The additional reinforcement of the matrix on the compressive side is favorable in increasing the stiffness and load-carrying capacity of the TRM-ALC panels.

5) Based on the force equilibrium and strain compatibility, a calculation method was proposed to predict the cracking load of the TRM-ALC panels.

6) The flexural and shear capacity of the TRM-ALC panels was calculated according to the analytical models associated with the flexural failure mode. characterized by the rupture of the textile grid and the shear mode of the RC member without stirrups from the literatures, respectively.

References

[1] Xu C Y, Nehdi M L, Wang K L, Guo Y. Experimental study on seismic behavior of novel AAC prefabricated panel walls. Journal of Building Engineering, 2021, 44: 103390

[2] Ma F D, Deng M K, Yang Y. Experimental study on internal precast beam–column ultra-high-performance concrete connection and shear capacity of its joint. Journal of Building Engineering, 2021, 44: 103204

[3] Yardim Y, Waleed A M T, Jaafar M S, Laseima S. AAC-concrete light weight precast composite floor slab. Construction and Building Materials, 2013, 40: 405–410

[4] Wang B, Wang P, Chen Y S, Zhou J, Kong X, Wu H, Fan H, Jin F. Blast responses of CFRP strengthened autoclaved aerated cellular concrete panels. Construction and Building Materials, 2017, 157: 226–236

[5] Salgado I D P, Silva F D A. Flexural behavior of sandwich panels combining curauá fibre-reinforced composite layers and autoclaved aerated concrete core. Construction and Building Materials, 2021, 286: 122890

[6] Mousa M A, Uddin N. Experimental and analytical study of carbon fibre-reinforced polymer (FRP)/autoclaved aerated concrete (AAC) sandwich panels. Engineering Structures, 2009, 31(10): 2337–2344

[7] Tuncer E, Binici B, Canbay E. Behavior and design of FRP bonded autoclaved aerated concrete beams. Construction and Building Materials, 2021, 282: 122712

[8] Kouris L A S, Triantafillou T C. State-of-the-art on strengthening of masonry structures with textile reinforced mortar (TRM). Construction and Building Materials, 2018, 188: 1221–1233

[9] Yin S P, Xu S L, Li H. Improved mechanical properties of textile reinforced concrete thin plate. Journal of Wuhan University of Technology. Materials Science Edition, 2013, 28(1): 92–98

[10] Carozzi F G, Bellini A, D’Antino T, de Felice G, Focacci F, HojdysŁ, Laghi L, Lanoye E, Micelli F, Panizza M, Poggi C. Experimental investigation of tensile and bond properties of Carbon-FRCM composites for strengthening masonry elements. Composites. Part B, Engineering, 2017, 128: 100–119

[11] Zhang Y X, Zhang S J, Deng M K. Four-point bending tests of ECC: Mechanical response and toughness evaluation. Case Studies in Construction Materials, 2022, 17: e01573

[12] Portal N W, Thrane L N, Lundgren K. Flexural behaviour of textile reinforced concrete composites: Experimental and numerical evaluation. Materials and Structures, 2017, 50(1): 1–14

[13] Kariou F A, Triantafyllou S P, Bournas D A, Koutas L N. Out-of-plane response of masonry walls strengthened using textile-mortar system. Construction and Building Materials, 2018, 165: 769–781

[14] Kariou F A, Triantafyllou S P, Bournas D A. TRM strengthening of masonry arches: An experimental investigation on the effect of strengthening layout and textile fibre material. Composites. Part B, Engineering, 2019, 173: 106765

[15] Meriggi P, Caggegi C, Gabor A, de Felice G. Shear−compression tests on stone masonry walls strengthened with basalt textile reinforced mortar (TRM). Construction and Building Materials, 2022, 316: 125804

[16] Gao W Y, Hu K X, Dai J G, Dong K, Yu K Q, Fang L J. Repair of fire-damaged RC slabs with basalt fabric-reinforced concrete. Construction and Building Materials, 2018, 185: 79–92

[17] Ding Z, Xu M R, Dai J G, Dong B Q, Zhang M J, Hong S X, Xing F. Strengthening concrete using phosphate cement-based fibre-reinforced inorganic composites for improved fire resistance. Construction and Building Materials, 2019, 212: 755–764

[18] Yang X, Gao W Y, Dai J G, Lu Z D. Shear strengthening of RC beams with FRP grid-reinforced ECC matrix. Composite Structures, 2020, 241: 112120

[19] Portal N W, Flansbjer M, Zandi K, Wlasak L, Malaga K. Bending behaviour of novel textile reinforced concrete-foamed concrete (TRC-FC) sandwich elements. Composite Structures, 2017, 177: 104–118

[20] Schladitz F, Frenzel M, Ehlig D, Curbach M. Bending load capacity of reinforced concrete slabs strengthened with textile reinforced concrete. Engineering Structures, 2012, 40: 317–326

[21] Koutas L N, Bournas D A. Flexural Strengthening of two-way RC slabs with textile-reinforced mortar: Experimental investigation and design equations. Journal of Composites for Construction, 2017, 21(1): 04016065

[22] Zhang H Y, Liu H Y, Kodu V, Li M Y, Zhou Y. Flexural behavior of concrete slabs strengthened with textile reinforced geopolymer mortar. Composite Structures, 2022, 284: 115220

[23] Yin S P, Feng L L, Zhao J L. Study on the tensile and flexural mechanical properties of TRE in freezing−thawing environments. Construction and Building Materials, 2021, 268: 121150

[24] Dong Z F, Deng M K, Dai J, Song S. Flexural strengthening of RC slabs using textile reinforced mortar improved with short PVA fibres. Construction and Building Materials, 2021, 304: 124613

[25] Deng M K, Dong Z F, Zhang C. Experimental investigation on tensile behavior of carbon textile reinforced mortar (TRM) added with short polyvinyl alcohol (PVA) fibres. Construction and Building Materials, 2020, 235: 117801

[26] Zhang W, Deng M K, Han Y G, Li R, Yang S. Uniaxial tensile performance of high ductile fibre-reinforced concrete with built-in basalt textile grids. Construction and Building Materials, 2022, 315: 125716

[27] GB/T11969-2020. Test Methods of Autoclaved Aerated Concrete. Beijing: China Building Materials Federation, Standardization Administration of the People’s Republic of China, 2020 (in Chinese)

[28] GB/T36262. Fibre Reinforced Polymer Composite Grids for Civil Engineering. Beijing: China Building Materials Federation, Standardization Administration of the People’s Republic of China, 2018 (in Chinese)

[29] DBJ61/T112-2016. Technical Specification for Application of High Ductile Concrete. Xi’an: Shaanxi Provincial Building Standard Design Office, 2016 (in Chinese)

[30] FengPQiang HYeL P. Discussion and definition on yield points of materials, members and structures. Engineering Mechanics, 2017, 3: 41−51 (in Chinese)

[31] He W D, Wang X, Ding L N, Wu Z. Efficiency of different BFRP-based strengthening techniques in improving flexural behavior of RC slabs. Construction and Building Materials, 2021, 308: 125002

[32] Harajli M, ElKhatib H, San-Jose J T. Static and cyclic out-of-plane response of masonry walls strengthened using textile-mortar system. Journal of Materials in Civil Engineering, 2010, 22(11): 1171–1180

[33] ZhengK QLiu ZMengS PQinS Q. Novel calculation method for shear capacity of RC beams without web reinforcement. Journal of Southeast University (Natural Science Edition), 2017, 47 (2): 362–368 (in Chinese)

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