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Article

Experimental Study on the Structural Performance of Reinforced Truss Concrete Composite Slabs during and after Fire

by
Bing Li
1,
Zhengshang Li
1,
Zhijun Chen
2,
Zhoulin Yang
1 and
Yang Zhang
1,*
1
College of Civil Engineering and Architecture, Hainan University, Haikou 570228, China
2
Haikou Qiongshan District Fire and Rescue Brigade, Haikou 570228, China
*
Author to whom correspondence should be addressed.
Buildings 2023, 13(7), 1615; https://doi.org/10.3390/buildings13071615
Submission received: 3 May 2023 / Revised: 20 June 2023 / Accepted: 21 June 2023 / Published: 26 June 2023
(This article belongs to the Section Building Structures)
Browse Figures
Versions Notes

Abstract

:
Standard fire resistance tests and post-fire tests on residual mechanical properties were carried out on two four-edge simply supported reinforced truss concrete two-way composite slabs with wide seam splicing (“integral seam”) and close seam splicing (“separated seam”), and the effects of the different splicing forms on the mechanical properties of the reinforced truss concrete composite slabs during and after exposure to high temperatures were explored. The results indicate that there are significant differences in terms of crack developments, moisture loss times on the slab surface, spalling characteristics of the bottom of the slab and cross-section temperature gradients of concrete and steel reinforcement along the thickness direction between the wide seam spliced reinforced truss concrete composite two-way slab (referred to as “S1 composite slab”) and the close seam spliced reinforced truss concrete composite two-way slab (referred to as “S2 composite slab”) in the fire test. This is mainly due to differences in the splicing forms and structural measures of the prefabricated base slabs at joints, which led to changes in the heat transfer path at the joints. As the temperature at the close seam can be released quickly, the stiffness recovery of the S2 composite slab is significantly greater than the stiffness weakening due to the thermal hysteresis of the concrete, so the deflection of the S2 composite slab recovers immediately after the fire stops. Overall, both the S1 and S2 composite slabs exhibit two-way high temperature deformation characteristics. The post-fire residual static load tests show that the static load-carrying capacity of the S1 composite slab is 12.08% higher than that of the S2 composite slab and the ultimate displacement is 8.74% lower than that of the S2 composite slab. It is appropriate to calculate the residual load capacities of the S1 and S2 composite slabs after the fire as a one-way slab.

1. Introduction

In recent years, the Ministry of Housing and Urban Rural Development in China has successfully promoted a large number of prefabricated building demonstration projects across the country, and has created a number of national leading enterprises, enabling prefabricated buildings to develop rapidly. Among them are prefabricated concrete buildings, prefabricated steel buildings and prefabricated composite buildings, which have proliferated [1]. In order to meet the national prefabrication and assembly rates for prefabricated buildings and to minimize the cost increments in the construction of prefabricated buildings, reinforced truss concrete composite slabs are often actively used as mainstream components in engineering construction. Reinforced truss concrete composite floor slabs, due to their high degree of industrial production and low weight (prefabricated base slab approximately 60 mm thick), are widely used in Europe, Japan and China in prefabricated construction. At present, some research results at room temperature have been included in the technical regulations of prefabricated construction [2,3] and relevant standard atlas [4,5], providing a solid technical guarantee for the engineering application of reinforced truss concrete composite floor slabs. Although a series of standards, regulations and codes related to prefabricated buildings have been introduced in the country and around the world, there is still a large knowledge gap in the independent standard system for fire protection of prefabricated buildings [6]. To improve the fire protection standards for prefabricated buildings, we need to rely on in-depth research to accelerate the establishment of fire resistance systems for prefabricated buildings.
Currently, fires are frequent and potentially dangerous, sometimes leading to catastrophic collapses of buildings. For cast-in-situ reinforced concrete structures, determining their fire resistance is mainly based on the fire resistance rating or fire resistance limit of individual components, such as the “Code of Design on Building Fire Protection and Prevention” [7] and the relevant Eurocode parts [8], which use the above ideas. For the fire resistance performance design of structures, practical design is far from universal. The findings of a study of reinforced truss concrete composite slabs at room temperature show that the stress concentration at the joints of the composite slabs [9,10,11], the lack of stiffness at the splicing area [12,13,14,15,16,17] and the large crack width at the time of damage [18,19,20,21] lead to a slight reduction in the bending load capacity at the joints, so the objective existence of the splice on the side of the slab becomes the weak point of the composite slab. At the same time, the influence of the number of splices cannot be ignored, and an increased number of splices can share the width of the bending crack at the splice and delay the development of the crack [22], but the increase in the number of splices will clearly deteriorate the fire resistance of the composite slab. Therefore, there is an urgent need to investigate the fire performance of spliced reinforced truss concrete composite slabs. Deng et al. [23,24,25,26] conducted fire resistance tests on composite slabs by varying boundary restraints, reinforcement types and mortar coatings set on the bottom of the slabs, and provided some suggestions for improving the fire resistance of composite slabs. Wei [27] conducted fire resistance tests on simply supported composite concrete slabs with different composite face heights and different composite face treatments. The results showed that when the composite face height ratio (the height ratio of the composite cast-in-situ layer to the precast slab) was less than 1 and the composite face was roughly treated, the composite slabs had the best fire resistance and the smallest deformation under fire. Wang et al. [28,29,30,31,32] carried out fire tests on simply supported composite slabs and obtained important data such as the temperature fields of the composite slabs, the displacements inside and outside the plane and the corner of the slab edge, which provided excellent test data for the correct understanding of the fire resistances of composite slabs. Ren [33] conducted fire resistance tests on reinforced concrete composite slabs with wide seams and found that the concrete spalling at the seams of precast slabs was the most serious and eventually developed into through cracks, resulting in the floor slabs being burnt through. Li et al. [34] studied the effects of post-fire regimes with different reinforcement rates and fire exposure times on the span deflections of two-way reinforced concrete slabs. An increase in the reinforcement ratio and a decrease in the fire time can slow down the development of the span deflection, and formulae for calculating the stiffness and deflection of two-way reinforced concrete slabs after fire were derived. Wang et al. [35,36] determined the residual ultimate bearing capacities of reinforced concrete continuous slabs after fire based on different theoretical models. Yang et al. [37,38] experimentally studied the performance of two-way reinforced concrete slabs in a full-scale integrated structure under the action of fire. The results showed that unlike two-way slabs in a single structure, the maximum out-of-plane displacement eventually occurred not far from the center of the slab and leaned towards the edge of the inner plate. Zhu et al. [39,40,41,42,43,44,45,46] conducted fire tests on two-way composite reinforced concrete slabs in different boundary modes, and the results showed that the boundary conditions play a crucial role in the fire resistance duration, the development of cracks on the slab surface and the damage patterns on the slab surface. Other research on the fire resistance limits of floor slabs is based on common floor slabs [45,46,47,48,49,50,51] and composite floor made of profiled steel sheets [52,53], but the results may not be applicable to reinforced truss concrete composite slabs.
The above-mentioned research on high temperature studies mainly investigated the influences of parameters such as boundary constraints, composite surface configurations and reinforcement trusses on the high temperature performances of composite slabs. The research is mainly focused on composite slabs without seams, and the findings to be developed and the research targets need to be expanded. It is clear that the results of a study on the fire resistance of seamless composite slabs may not be applicable to reinforced truss concrete composite two-way slabs with seams due to the existence of these seams. Fire resistance tests and theoretical studies on four-sided simply supported two-way composite concrete slabs with close and wide seams under boundary constraints are carried out to understand the influences of the type and configuration of the seams and the setting of the reinforcement truss on their fire resistance performance in order to reveal the mechanism of high temperature damage in the case of close and wide seams and to establish a fire resistance design method for reinforced truss concrete two-way composite slabs. Therefore, this paper designs and fabricates two full-scale reinforced truss concrete two-way composite slabs with two types of splicing: close seam and wide seam, and conducts an experimental study on the standard fire behavior and the residual mechanical properties of the composite slabs after fire.

2. Experimental Overview

2.1. Mechanical Properties of Materials

The detailed concrete mixes are shown in Table 1. The cement in the study is Hainan “Huaxin” PC32.5 cement. The coarse aggregate is graded crushed stone, the maximum particle size is 30 mm and the apparent density is 2600 kg/m3. The fine aggregate is medium sand, the fineness modulus is 3.0 and the apparent density is 2700 kg/m3. The setting time represents the time from the mixing of concrete with water to the loss of plasticity of cement and directly affects the transportation and pouring of concrete mixes. The stability refers to the uniformity of the volume change of the concrete during the hardening process, and the stability index determines whether the concrete components will produce expansion cracks. The initial expansion cracks of the concrete are not allowed in the process of use.
The reinforcements are “Zhongzheng” HRB400 steel bars with nominal diameters of 10 mm and 8 mm, where 8 mm bars are used for the webs of reinforcing trusses and the remaining 10 mm bars are used for other parts. The reinforcement ratio is 0.56%. The number of trusses for each precast slab was 5. The strength index of the reinforcement is shown in Table 2.

2.2. Specimen Design

The truss-reinforced concrete two-way composite slab specimens with the integral joint type and isolation joint type were named S1 and S2 composite slabs, respectively. The slab dimensions and reinforcement configurations are shown in Figure 1. Both slabs have dimensions of 4200 mm × 3600 mm and a thickness of 130 mm. The prefabricated baseplate of the S1 composite slab was manufactured by joining two prefabricated boards with the dimensions of 1950 mm × 3600 mm × 60 mm at a spacing of 300 mm, and reinforcement bars were arranged in the spacing as per the requirements of the specification. Finally, a 70 mm thick concrete layer was cast in place above the baseplate to yield the integral joint-type composite slab. Similarly, the prefabricated baseplate of the S2 composite slab was also prepared by connecting two prefabricated boards of 2100 mm × 3600 mm × 60 mm, and a 70 mm thick concrete layer was cast in place above the baseplate to afford the isolation joint-type composite slab. Both prefabricated boards and the cast-in-situ layer were made of C30 concrete. The reinforcement bars were all HRB400 ribbed steel bars. The representative value of compressive strength of the 150 mm concrete cubes for making prefabricated boards was measured to be 33 MPa. The measured representative value of compressive strength of the cast-in-situ layer was 37 MPa. The measured yield strength and tensile strength of the reinforcement bars were 441 MPa and 664 MPa, respectively. The truss extended from the prefabricated baseplate protection layer to the cast-in-situ protection layer. Each prefabricated board comprises 5 reinforced trusses, and each truss was made by welding 1 upper chord, 2 lower chords and wave-shaped bent abdominal reinforcement bars together. The diameter of the abdominal reinforcement bars was 8 mm, and the diameter of both the upper and lower chord reinforcement bars was 10 mm. Double-layer two-way reinforcement bars of 10 mm in diameter were arranged at a spacing of 200 mm on the top and bottom of S1 and S2 composite slabs. Bent-up reinforcement bars of 10 mm in diameter were spaced at a distance of 200 mm at the joints of the S1 composite slab, with a continuous bending angle of 60° and a projected length of 320 mm. The bending height was the distance from the bottom reinforcement bars to the top bars. With regard to the S2 composite slab, reinforcement bars of 10 mm in diameter were placed symmetrically to the joints, 200 mm away from each other, with a total of 800 mm in length. Positions 1 and 2 in Figure 1c,d indicate the details of the wide seam splicing and close seam splicing, respectively.

2.3. Heating System

The fire test was performed using a horizontal fire testing furnace from the Fire Laboratory at Huaqiao University in Xiamen, Fujian Province. A 1.5 m high steel beam–column composite reaction frame surrounding the furnace was provided, and the distance between the outer wall of the furnace and the frame was 40 mm. A 1.0 m thick aluminum silicate fireproof cotton layer was placed between the outer and inner walls of the furnace. The net dimensions of the furnace chamber were 4.2 m × 6.0 m × 1.7 m, and a flue with dimensions of 0.5 m × 0.5 m × 4.2 m was installed along the length direction at the bottom of the furnace for suctioning gas out of the furnace by maintaining a pressure balance inside the furnace. There were 14 gas burners around the furnace, and the burners were turned on and off by computer programs. The temperature inside the furnace was measured in real time through an inserted ceramic tube thermocouple. The structure of the fire testing furnace is illustrated in Figure 2. The ISO-834 standard temperature rise curve was used to characterize the temperature in the fire testing furnace. An initial furnace temperature of 28 °C was measured prior to the start of the fire test using the instrumentation, and the furnace was left to cool naturally.

2.4. Boundary Conditions and Loading Plan

As required by GB/T 50152-2012 “Standard for Test Method of Concrete Structures” [54] and GB/T 9978.1-2008 “Fire-resistance tests—Elements of building construction” [55], the four-edge simply supported conditions of the composite slabs were realized by steel rollers and balls. Both the rollers and balls had a diameter of 80 mm, and the roller length was 500 mm. The testing furnace was modified based on the test requirements, and provided with rigid supports. Specifically, two brick walls were constructed along the short side of the furnace, and one brick column was built along the long side. One end of the I-beam was placed on the brick column, and the other end was placed on the reaction frame. Before the fire test, the furnace body, brick walls, brick columns and I-beams were wrapped with 2 to 3 layers of fire-resistant cotton, which fulfilled the requirement for rigidity to support the boundary during the test. During the test, it was found that vertical displacements were all within 4 mm, and the boundary support conditions met the requirements of the test. The modifications to the interior of the furnace are depicted in Figure 3. The net dimensions of the modified furnace plane were 3240 mm × 4100 mm. In the fire test, a constant-load standard heating–natural cooling mode was adopted, a uniform load of 2.0 kN/m2 was applied to the cast iron blocks. The testing arrangement is illustrated in Figure 4.

2.5. Experiments and Measurement Point Distributions

The testing arrangement of the composite slabs during the test is shown in Figure 4 and detailed in Figure 5. On the S1 composite slab, there were horizontal displacement measurement points V1-V4 mainly distributed at the center of each plate side, vertical displacement measurement points W1-W12 mainly distributed at the joints, 1/4 and 1/2 of the plate surface, measurement points Y1-Y3 at the plate edge corner, concrete temperature measurement points T1-T7 and reinforcement measurement points T2-S, T5-S, T6-S and T7-S near 1/4 of the long side of the board surface, at the joints and at the joint center. In addition, there were also loading points S1–S4 for the residual bearing capacity test after fire. The composite slab was divided into four regions, labelled 1#–4#, each of which was further subdivided into four sub-parts.
The temperature data were collected using self-made K-type thermocouples. The cross-section distribution of the thermocouples is shown in Figure 6. The temperatures of the concrete and reinforcement bars along the plate thickness direction were measured using thermocouple strings, and each string was made up of 5 thermocouples.

3. Fire Test Results and Analysis

3.1. Main Test Results and Causes

The red lines in Figure 7 indicate the bottom joint position on the S1 and S2 composite slabs. Two joints at a spacing of 300 mm were observed on the bottom of the S1 composite slab, and natural seams between the prefabricated baseplates were evident on the bottom of the S2 composite slab.
Before turning on the fire, a total load of 2.0 kN/m2 was applied to the S1 composite slab at room temperature through four steps of loading, with 0.5 kN/m2 applied each time. After loading, no obvious changes were observed on the specimen surface, and the maximum displacement of the slab span was 3 mm. To enable the S1 composite slab to bear the load stably, the burner was turned on 30 min after loading, and the initial temperature of the furnace was set to 28 °C. After 15 min of exposure to fire, fine cracks occurred in the 3#-2, 4#-3 and 1#-4 areas on the plate surface, and water vapor came out of the cracks. Meanwhile, water stains were observed on the side of the plate. Twenty minutes after heating, water gradually gathered in the above-mentioned areas, forming a puddle. After being burned for thirty minutes, the board began to break, making discontinuous “bang” sounds, and two approximately parallel cracks propagated along the joints of the S1 composite slab. After 35 min of heating, water bubbles kept coming out of the cracks at the joints, and a large amount of water accumulated at the cracks in the plate. Five minutes later, the board surface clearly concaved downward, and the corners warped. The moisture on the plate surface flowed towards the plate center due to gravity, and water on the other areas of the plate surface almost completely evaporated. The maximum depth of the accumulated water at the center reached 3 cm. After fifty minutes of heating, a large amount of white water vapor was generated on the plate surface, and the cracks spread from the joints to the surrounding areas, becoming increasingly larger. Meanwhile, a large number of approximately parallel cracks occurred in the trusses. The temperature in the furnace reached 900 °C. With the increase in the exposure time to fire, a large amount of water accumulated at the center of the composite slab and evaporated, and the number and widths of cracks on the plate surface increased. At the 180th min, the average temperature rise (160 °C) at the center of the unexposed surface of the plate exceeded 140 °C. The composite slab was considered no longer fire-proof based on GB/T 9978.1-2008 “Fire-resistance tests—Elements of building construction” [55]. The fire test was stopped and the furnace was allowed to cool naturally. There was still some water at the center, and the maximum warping displacement of the plate edge increased to 7 cm.
Figure 8a is the original image of the S1 composite slab surface after fire, and the main cracks on the slab surface are marked in Figure 8c. The red line in Figure 8c indicates the position of the joints after fire. It was found that the two initial cracks almost overlapped over the joints, and the cracks mainly spread along the length direction of the plate, parallel to the joints. Most cracks occurred on the trusses and at the joints. A similar crack propagation pattern was observed in a previous study, which found that cracks mainly occurred in the area around the joints of an integral joint-type composite slab with a seam width of 450 mm. However, the present study obtained a different propagation mode of the four corner cracks. In the present study, numerous radially distributed diagonal cracks developed at the corners during the test. The reason for this is that the difference in warping deformation between the long and short sides of the composite slab caused a torque effect at the corners, putting the concrete on the diagonal lines under tension and making the concrete in the upper parts of the corners crack radially. Most cracks on the plate side were connected with the cracks on the plate surface. The cast-in-situ layer was not separated from the prefabricated layer, but small cracks near the corners extended to the slab surface. At the same time, shear stress was generated at the interface between the prefabricated and cast-in-situ layers due to the temperature gradient effect. With the rising temperature, the horizontal shear stress gradually increased, and the reinforced truss bore a large portion of it, so the prefabricated plates and the cast-in-situ layer worked together. Physically, tensile stress concentrated in the cast-in-situ concrete on the reinforced truss. In addition, with good thermal conductivity, the upper reinforcement bars of the truss expanded and deformed greatly, resulting in tensile stress to the concrete warped by the bars. As a result, the tensile stress of the concrete around the bars on the upper surface of the composite slab increased, making it easy for cracks to form along the bars. Furthermore, on the cross-section of the joints of the S1 composite slab, there was a weak bonding surface characterized by low bonding strength between the strip cast-in-situ concrete plate and the prefabricated plate. Longitudinal cracks first occurred on this surface during the fire because the large tensile stress generated by temperature gradients was mainly borne by the cast-in-situ layer.
The S2 composite slab adopted the same load pattern as the S1 composite slab. After loading, no obvious changes were observed on the surface of the S2 composite slab, and the maximum displacement in the mid-span of the slab was 5 mm. The fire test commenced 30 min after loading, and the initial temperature inside the furnace was set to 28 °C. After 8 min of exposure to fire, cracks and water stains were first generated in the 4#-3 area, followed by the 1#-4 and 2#-1 areas successively. The average temperature inside the furnace was 550 °C. Five minutes later, the 3#-1 area cracked and water stains were produced. After heating the slab for 15 min, water stains were generated on the slab sides and at the slab center, and there was also a small amount of water vapor at the slab center. At the 20th min, “bang” sounds came from the slab bottom and cracks rapidly spread along the joints at the center. A large amount of water vapor was generated, which gathered and quickly formed a small puddle. In the process, the moisture was intensely gasified. Thirty minutes after heating, more and more water stains were produced on the slab surface. After 50 min of heating, a large amount of liquid water gathered at the joints on the slab surface, and the long cracks along the joints greatly widened, from which a large amount of water vapor came out. After 60 min of exposure to fire, the plate surface was evidently concaved, and the four corners of the slab were significantly warped. The mid-span displacement of the slab increased to 80 mm, and the average furnace temperature was 950 °C. At the 105th min, the slab made “bang” sounds. It was inferred that the concrete at the joints on the plate bottom cracked again based on the damage on the slab bottom. The moisture on the slab surface evaporated completely. Subsequently, the four corners of the slab warped. At the 165th min, the average temperature rise (142 °C) of the unexposed surface of the slab exceeded 140 °C, indicating the slab was no longer fire-resistant. The fire test was stopped and the furnace was allowed to cool down naturally.
Figure 8a shows the original image of the S2 composite slab surface after fire, and Figure 8c marks the major cracks on the slab surface. The red line in Figure 8d indicates the position of the joint after the fire. The major cracks occurred almost at the joints on the slab surface, and most cracks spread along the joints. One difference between the S1 and S2 composite slabs is that the latter had fewer but wider cracks. The propagation mode of cracks on the S2 slab sides was similar to that of cracks on the S1 slab sides, but the horizontal cracks on the bonding surface of the S2 slab extended farther than those of the S1 slab. Therefore, the bonding surface between the prefabricated and cast-in-situ layers of the S2 composite slab bore a greater shear force. However, no interface slip was observed. The reason for this is that the joints in the prefabricated baseplates of the S2 composite slab enabled the effective release of the temperature expansion stress in the middle. Moreover, the joints also reduced the integral stiffness of the S2 composite slab and the torque effect at the corners. Hence, there were few diagonal cracks at the corners.
The two fire tests were stopped when the heat insulation properties of the S1 and S2 composite slabs were destroyed. However, there was a large difference in the time when the joints on the unexposed surface of the composite slabs began to crack during the tests. The prefabricated plates of the S2 composite slab were connected only by cement paste, and the cast-in-situ layer bore a large portion of the tensile stress generated by the temperature gradients, so the tensile strength of the S2 slab was far lower than the bonding strength between the strip cast-in-situ concrete plate and the prefabricated plate at the joints of the S1 slab. Therefore, the S2 slab cracked on the interface at the joints first. In addition, the moisture contents in the S1 and S2 composite slabs were measured to be 3.70% and 3.71% before the test, respectively, which were essentially the same. The water on the surface of the S1 composite slab was completely evaporated by the residual heat 180 min after the fire test, but the water on the surface of the S2 composite slab disappeared before the end of the fire test. The huge difference in the time taken for water to evaporate between the two slabs was attributed to the wider cracks at the joints of the S2 slab and more intense vaporization activity in the S2 slab. Moreover, the water evaporation and migration paths varied in slabs of different joint types [56], and the joints provided additional channels for water to overflow from the slab. When the S2 slab was heated, the moisture in it evaporated and moved towards the unexposed surface and joints, absorbing heat from the slab and thus sharply reducing the temperature inside the slab. Furthermore, cement paste infiltrated into the S2 slab joints when pouring the cast-in-situ concrete, so it had lower thermal conductivity and lower temperature at the joints.
Figure 9a presents the spalling concrete on the bottom of composite slab S1 after the fire was stopped. The most severe spalling occurred in the cast-in-situ concrete strip, where wide cracks were connected, and the steel bars at the bottom of some composite slabs were exposed due to the spalling of protective concrete layers, mainly concentrated in regions No. 3# and No. 4# on the surface of the composite slabs, with a maximum depth of around 35 mm and a maximum length of around 140 mm. Figure 9b shows the spalling concrete on the bottom of composite slab S2 after the fire was stopped. The spalling mainly occurred at local joints, causing protective concrete layers to fall off. The cracks on the concrete layer were about 1600 mm long and 800 mm wide at maximum, making the steel bars on the bottom of corresponding sections completely exposed.
The spalling characteristics of composite slabs S1 and S2 were directly related to their joint structures on the bottom. Along the joint cast-in-situ belt of composite slab S1, all the stressed steel bars in the vertical joint direction of the two prefabricated bottom plates were connected by lap joints, and the force was mainly transferred through the steel bars. Thus, there were essentially no obvious cracks in the concrete. However, the force was transferred through the concrete in the joint direction, and concrete materials at the joint deteriorated sharply due to the fire, which led to a significant decline in the tensile strength. Consequently, cracks and damage were very likely to arise at the concrete joint, even accompanied by local spalling. In contrast, joints in composite slab S2 were visible, and the cement slurry that infiltrated into the composite layer in the process of filling joints with concrete was far from being sufficient. Therefore, concrete at the joints was subjected to fire on both sides and suffered from spalling.

3.2. Comparison of the Temperature Rise Curves in the Fire Furnace

Comparisons between the average temperatures in the fire furnace, the temperatures at several main measuring points and the ISO-834 standard temperature curve are shown in Figure 10. As can be seen from the figure, the temperature control in the fire furnace was in excellent agreement with the ISO-834 standard temperature curve, implying that this experiment was performed in strict accordance with the ISO-834 international standard temperature rise curve.

3.3. Temperature Curves of the Concrete

In order to analyze and compare the temperature differences at the joints of composite slabs S1 and S2, the temperature gradients of concrete sections in regions T4, T5 and T6 on composite slab S1 and in regions T3, T7 and T6 on composite slab S2 were analyzed. Figure 11 and Figure 12 show the temperature and time curves of these concrete sections on composite slabs S1 and S2. It can be seen from the figures that the distribution of the temperature field measured using the thermocouple series at each temperature measuring point of S1 was essentially consistent, and at the same section height, the concrete temperature at the joint given by the measuring point T5 of S1 was slightly higher than that at the measuring points T4 and T6 along the pouring belt after the joint because spalling occurred on the side of the concrete subjected to fire at the measuring point T5, which directly increased the fire exposure area and resulted in a sharp rise in temperature. Due to the thermal inertia of concrete, the temperature at each measuring point of the concrete continued to rise for a period of time after the fire was stopped and then dropped rapidly. However, the temperature of the concrete at the joint at measuring point T7 of the S2 slab was much lower than that at the measuring points T1 and T6, mainly because the overflow cement slurry filled part of the joint when the joint of the prefabricated plate was constructed using concrete on the upper cast-in-situ layer, resulting in a significant decrease in the thermal conductivity of the joint as the cement paste had a thermal conductivity of 0.35~0.45 W/(m·k) and, thus, a lower temperature. The temperature at the joint of the concrete measuring point dropped rapidly after the fire was stopped, except for a slow decline after a slight continuous rise in the temperature of the cast-in-situ layer of concrete. The temperature of the cast-in-situ layer given at measuring point T7 was essentially consistent with that at other measuring points. Because of the presence of dense cracks, the water migration and evaporation route in the slab, respectively, moved towards the side that was not subjected to the fire and the joint. As a result, the water in the slab completely evaporated after the fire was stopped, and more heat energy was absorbed. For this reason, the temperature of the concrete and steel bars in the S2 slab was lower than those in the S1 slab.

3.4. Temperature Curves of the Steel Bars

The temperature fields obtained at individual steel bar measuring points on composite slabs S1 and S2 were essentially the same. Therefore, the temperature at points T5-S on the S1 slab and T2-S on the S2 slab was measured, as specifically shown in Figure 5. The temperature–time curves of the distributed truss reinforcement and stressed reinforcement in the thickness direction are shown in Figure 13. In terms of S1, T5-S-1 was the measuring point on the bottom of the stressed reinforcement and truss reinforcement, T5-S-5 was the temperature measuring point on the top of the stressed reinforcement and truss reinforcement and T5-S-2 to T5-S-4 in the middle were the measuring points of the truss reinforcement. The temperature of the steel bars at T2-S-1 and T2-S-5 was essentially consistent with that of the concrete at the same positions. The temperature differences of the steel bars at the middle three measuring points were small, and the temperature was higher than that of the concrete in the same positions. In addition, the curves in the horizontal section were shorter than those of the concrete in the horizontal section. The above results were attributed to the good thermal conductivity of the truss reinforcement. The temperature of the steel bar at each measuring point in the S2 slab was lower than that in the same position in the S1 slab. Furthermore, the temperature in the S2 slab suddenly decreased and then increased before the curve was about to enter the horizontal section, which was even more obvious in the region closer to the bottom of the prefabricated plate. This was related to the low temperature at the joint and the migration of water to the joint. On the whole, the temperature of the concrete in the S1 and S2 slabs also showed similar distribution patterns.

3.5. In-Plane Displacements

Figure 14 shows the in-plane displacement changes of the S1 and S2 slabs in the heating and cooling phases, where V2 and V4 represent the horizontal displacements along the long edge, while V1 and V3 represent the displacements along the short edge. The specific scheme is shown in Figure 5. Due to the differences in joint splicing modes, in-plane displacement variations differed greatly between the S1 and S2 slabs. The horizontal deformations of the S1 slab at high temperatures were similar to those of the ordinary reinforced concrete floor due to wide joint splicing, with expansion deformations in both directions. The sum of the displacements along the long edge was roughly equal to that of those along the short edge, with a maximum displacement sum of about 21 mm. There was a minor increase in expansion deformation before it began to decline. The expansion thermal strain in the long edge direction on the S2 slab was partially released when subjected to fire due to tight joint splicing, resulting in declines in the horizontal displacements. However, the sum of the displacements along the long edge was similarly roughly equal to that along the short edge, with a maximum displacement of about 14 mm, indicating that the in-plane stiffness of the S2 slab was weakened, but its horizontal deformation characteristics remained consistent and the overall deformation still conformed to the deformation law of two-way slabs.

3.6. Out-of-Plane Displacements

The out-of-plane displacements of the S1 and S2 slabs are shown in Figure 15. After the fire was stopped, the deflection of the S1 slab continued to increase over a period of time and then decreased rapidly, essentially consistent with the change pattern in the temperature field mentioned earlier, which was directly related to temperature conduction lag. During the experiment, the maximum deflection of the S1 slab was 100 mm, and the displacement variations recorded at W2, W4 and W7 were essentially consistent. The integrity of the S1 slab was maintained well despite considerable lateral cracking along the short edge and local spalling at the joints. The maximum deflection of the S2 slab was 107 mm, and the displacement variations recorded at W3, W6, W11 and W12 were almost the same. Therefore, the overall deformations of the S1 and S2 slabs conformed with the high temperature deformation characteristics of two-way slabs. Due to visible splicing, the overall stiffnesses were greatly weakened. Therefore, the deformation rate of the S2 slab was greater than that of S1 after fire exposure, with a more severe overall deformation, as shown in Figure 16. Though the temperature of the concrete and steel bars in the S2 slab was lower than those in the S1 slab, the overall stiffness of the S2 slab was more significantly weakened, with a weaker anti-deformation capacity than the S1 slab because of tight joint splicing. The surface deflection of the S2 slab started to recover immediately after the fire was stopped, which was significantly different from that of the S1 slab, because the stiffness recovery caused by the rapid temperature release at the tight joint was much greater than the stiffness weakening caused by the thermal conductivity lag of the concrete. Upon the discontinuation of test data collection, the overall residual deformation of the S2 slab was slightly less severe than that of the S1 slab, showing better deformation recovery ability, mainly because the S2 slab had a lower internal temperature, less fire damage and greater residual stiffness after the fire. Figure 17 shows the variations in the deflections and average furnace temperatures of the S1 and S2 slabs. In the heating phase, the deflection deformation of the S1 slab increased linearly before the furnace temperature reached 570 °C; the deflection deformation of the S2 slab exhibited a linear increase before the furnace temperature reached 390 °C, mainly because the initial stiffness of the S2 slab was small, and high temperature deformation was more sensitive to the furnace temperature. The horizontal section began to emerge in the curves of the S1 and S2 slabs after the fire was stopped, and their deflections rapidly began to recover when the furnace temperature decreased to around 600 °C. The overall residual deformation of the S1 slab was slightly greater than that of the S2 slab after the furnace temperature was reduced to about 100 °C, indicating the weaker residual stiffness but better deformation recovery capacity of the S2 slab.

3.7. Rotation Angles

The rotation angles along the edges of the S1 and S2 slabs are shown in Figure 18, where Y1 and Y3 represent the rotation angles about the long edge, and Y2 represents that about the short edge. Figure 18a shows that the rotation angle about the long edge in the S1 slab was significantly greater than that about the short edge, mainly because the long edge cracked in parallel to the short edge, the in-plane stiffness damage in the long edge direction was severe and the surface force was mainly transferred in the short edge direction.

4. Residual Properties of the Composite Slabs after Fire

Figure 19a demonstrates the test method for determining the residual bearing capacities of the S1 and S2 slabs after fire. Four-point loading was adopted for the two slabs. The specific loading locations S1-S4 are shown in Figure 5. Figure 19b shows the test results of the residual bearing capacities of the S1 and S2 slabs. The residual ultimate bearing capacities of the two slabs were 538 kN and 480 kN, respectively. The load–time curves show that the residual bearing capacity of the S1 slab was always greater than that of the S2 slab. Although the stiffness damage of the S1 slab was obviously greater than that of the S2 slab after the fire, the residual bearing capacity of the S1 slab was greatly improved by the recovery of the steel strength at the joint after fire, resulting in a greater residual bearing capacity. During the experiment, the S1 and S2 slabs cracked on the bottom. The damage patterns on the bottoms of the S1 and S2 slabs are shown in Figure 20. In addition to spalling concrete and bare steel bars at the joints, the stressed steel bars on the bottom of the precast bottom plate were exposed due to external force. In addition, the S1 slab was damaged more severely than the S2 slab, mainly reflected in cracking and spalling concrete on the bottom.
Figure 21 presents the deflection–time relationships of the composite boards in the residual bearing capacity test after the fire, and the arrangement of the deflection measurement points in the test is the same as that during the fire, as shown in Figure 5. The horizontal steps represent the load duration. In the experiment, the deflections of the wide seam under static load are essentially the same in the direction perpendicular to the vertical joint, so the wide seam structure after the fire can ensure the overall force of the composite boards. The maximum ultimate displacement at the mid-span was 94 mm. The difference between the displacement at the measurement point at the quarter short side and that at the quarter long side gradually increased, indicating that the failure process gradually changed from bidirectional forcing to unidirectional forcing mode, and the force transmission path mainly depended on the truss reinforcement and force reinforcement. The residual bearing capacities of the S1 composite boards after the fire should be calculated according to the unidirectional board. The truss reinforcement had a significant influence on the force performance of the composite boards during and after the fire. The ultimate displacement of the S2 composite boards was 103 mm, slightly larger than that of the S1 composite boards, showing better deformation ability. The overall displacement changes of the S2 composite boards were similar to those of the S1 composite boards, and its residual bearing capacities after fire should also be calculated according to the unidirectional board.

5. Conclusions

Through a standard fire resistance test and residual mechanical properties test on two reinforced truss concrete composite two-way plates with different splicing methods, the influences of joint forms on the mechanical properties of composite slabs during and after high-temperature exposure were explored. The main conclusions obtained are as follows.
1. In the standard fire test, both the S1 and S2 composite slabs were damaged due to their thermal insulation performances, and the fire resistance limit of the S1 composite slab was increased by 9.1% compared with that of the S2 slab. The crack development mode of the S1 composite slab surface cracks was parallel through long cracks along the steel truss and joints, while the S2 composite slab had fewer long cracks and main cracks with wider crack widths. The spalling locations of the concrete on the fire surface of the two laminates were concentrated at the joints. Overall, the S1 composite slab was more damaged than the S2 composite slab.
2. In the standard fire test, the temperatures of the concrete and reinforcement in the S2 composite slab section were lower than those in the S1 composite slab. This is due to the existence of close seams, which caused the migration and evaporation of moisture from the slab to the backfire surface and the joints, respectively, and, in turn, resulted in the evaporation of moisture from the slab before the fire stopped and the absorption of more heat. The relatively low temperatures at the joints of the S2 composite slab were mainly due to the fact that the joints were partially filled with cement slurry during the construction of the upper cast-in-situ layer of concrete, resulting in a significant decrease in the thermal conductivity of the joints.
3. After the fire was stopped, the mid-span deflection of the S1 composite slab was 6.54% lower than that of the S2 composite slab. In terms of the linear phase of the deflection or deformation, the S1 composite slab was 46.15% more resistant to high temperature deformations than the S2 composite slab. The residual deformation of the S1 composite slab was 14.5% higher than that of the S2 composite slab after the fire was stopped, and the deflection of the S2 composite slab started to recover immediately after the fire was stopped, as the recovery of the stiffness due to the rapid release of temperature along the close seam was significantly greater than the weakening of the stiffness due to the thermal hysteresis of the concrete. Overall, both the S1 and S2 composite slabs exhibited high temperature deflection characteristics of two-way slabs.
4. The stiffness damage of the S1 composite slab during the fire was significantly greater than that of the S2 composite slab, but the recovery of the reinforcement strength at the joints after the fire significantly increased the residual load-carrying capacity of the S1 composite slab. The results show that the static load-carrying capacity of the S1 composite slab was 12.08% higher than that of the S2 composite slab, and the ultimate displacement is 8.74% lower compared to that of the S2 composite slab. The residual load capacity of the S1 and S2 composite slabs after fire should be calculated as a one-way slab.

Author Contributions

Project administration, Supervision, Validation and Writing—review & editing, B.L.; Formal analysis, Z.L.; Investigation, Z.L. and Z.C.; Resources, Z.C.; Data curation, Y.Z. and Z.Y.; Writing—original draft, Z.L.; Visualization, Z.Y.; Funding acquisition, Y.Z. All authors have read and agreed to the published version of the manuscript.

Funding

This research was funded by the Hainan Province Key R&D Plan Project (Grant no.: ZDYF2022SHFZ338), Lateral Science Foundation of Hainan University (Grant no.: HD-KYH-2020039), and NSFC (Grant no.: 51768017).

Data Availability Statement

The data will be required for subsequent finite element parametric analysis, which will be withheld for the time being.

Conflicts of Interest

The authors declare no conflict of interest.

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Figure 1. Detailed drawings of the composite slabs and reinforcing bars.
Figure 1. Detailed drawings of the composite slabs and reinforcing bars.
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Figure 2. Construction of the fire test furnace.
Figure 2. Construction of the fire test furnace.
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Figure 3. Layout of the renovated fire furnace.
Figure 3. Layout of the renovated fire furnace.
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Figure 4. Loading diagram of the uniformly distributed load.
Figure 4. Loading diagram of the uniformly distributed load.
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Figure 5. Distribution of the measuring points on the composite slabs.
Figure 5. Distribution of the measuring points on the composite slabs.
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Figure 6. Distribution of the thermocouples along the thickness direction of the composite slab.
Figure 6. Distribution of the thermocouples along the thickness direction of the composite slab.
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Figure 7. Plate bottoms of the composite slabs in the fire experiment.
Figure 7. Plate bottoms of the composite slabs in the fire experiment.
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Figure 8. Cracking patterns on the surface of the composite slabs after fire stopped.
Figure 8. Cracking patterns on the surface of the composite slabs after fire stopped.
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Figure 9. Plate bottom spalling map after fire.
Figure 9. Plate bottom spalling map after fire.
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Figure 10. Comparisons between the heating situations of the fire furnace and the standard heating curve.
Figure 10. Comparisons between the heating situations of the fire furnace and the standard heating curve.
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Figure 11. Section temperature distributions of S1 composite slab along the direction of concrete thickness.
Figure 11. Section temperature distributions of S1 composite slab along the direction of concrete thickness.
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Figure 12. Section temperature distributions of S2 composite slab along the direction of concrete thickness.
Figure 12. Section temperature distributions of S2 composite slab along the direction of concrete thickness.
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Figure 13. The cross-sectional steel bar temperature distributions along the thickness direction of the composite slab.
Figure 13. The cross-sectional steel bar temperature distributions along the thickness direction of the composite slab.
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Figure 14. Relationships between in-plane displacement and time for composite slabs.
Figure 14. Relationships between in-plane displacement and time for composite slabs.
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Figure 15. Out-of-plane displacement and time relationships of the composite slabs.
Figure 15. Out-of-plane displacement and time relationships of the composite slabs.
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Figure 16. Out-of-plane displacement and time relationships at the joint of the composite board.
Figure 16. Out-of-plane displacement and time relationships at the joint of the composite board.
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Figure 17. Relationships between the out-of-plane displacement of the composite slab and the average furnace temperature.
Figure 17. Relationships between the out-of-plane displacement of the composite slab and the average furnace temperature.
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Figure 18. Relationships between the inclination angle of the edge of the composite slab and time for composite slabs.
Figure 18. Relationships between the inclination angle of the edge of the composite slab and time for composite slabs.
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Figure 19. Residual bearing capacity testing on the composite slabs.
Figure 19. Residual bearing capacity testing on the composite slabs.
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Figure 20. Failure patterns of the bottom surfaces of the composite slabs after the residual bearing capacity testing.
Figure 20. Failure patterns of the bottom surfaces of the composite slabs after the residual bearing capacity testing.
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Figure 21. Relationships between the out-of-plane displacement and time from the residual bearing capacity test on the composite slabs.
Figure 21. Relationships between the out-of-plane displacement and time from the residual bearing capacity test on the composite slabs.
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Table
Table 1. Detailed concrete properties.
Table 1. Detailed concrete properties.
Water:Cement:Sand:Crushed StoneCoagulation Time (min)Compressive Strength (MPa)Flexural Strength (MPa)Specific Surface Product (m2/kg)
154:341:798:1078early
condensate		end
condensate		3 d28 d3 d28 d350
14027213.535.12.54.8
Table
Table 2. Detailed steel properties.
Table 2. Detailed steel properties.
Rebar TypeDiameter
(mm)		Yield Strength
(MPa)		Average
(MPa)		Ultimate Strength (MPa)Average
(MPa)		Elastic Modulus
(MPa)
HRB400104274255955951.96 × 105
426598
422591
84134135885871.95 × 105
408585
419589
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MDPI and ACS Style

Li, B.; Li, Z.; Chen, Z.; Yang, Z.; Zhang, Y. Experimental Study on the Structural Performance of Reinforced Truss Concrete Composite Slabs during and after Fire. Buildings 2023, 13, 1615. https://doi.org/10.3390/buildings13071615

AMA Style

Li B, Li Z, Chen Z, Yang Z, Zhang Y. Experimental Study on the Structural Performance of Reinforced Truss Concrete Composite Slabs during and after Fire. Buildings. 2023; 13(7):1615. https://doi.org/10.3390/buildings13071615

Chicago/Turabian Style

Li, Bing, Zhengshang Li, Zhijun Chen, Zhoulin Yang, and Yang Zhang. 2023. "Experimental Study on the Structural Performance of Reinforced Truss Concrete Composite Slabs during and after Fire" Buildings 13, no. 7: 1615. https://doi.org/10.3390/buildings13071615

APA Style

Li, B., Li, Z., Chen, Z., Yang, Z., & Zhang, Y. (2023). Experimental Study on the Structural Performance of Reinforced Truss Concrete Composite Slabs during and after Fire. Buildings, 13(7), 1615. https://doi.org/10.3390/buildings13071615

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