Flexural behavior and design of high-strength I-shaped steel-UHPC composite beams

Abstract

High-strength materials and corresponding composite members are gradually accepted in the bridge industry to achieve the purposes of large span and lightweight. Three simply supported I-shaped steel-concrete composite beams were tested to investigate their flexural behavior when different combinations of materials were used, i.e., high-strength steel-ultra high performance concrete (HS-UHPC), high-strength steel-conventional strength concrete (HS-CC), and conventional strength steel-ultra high performance concrete (CS-UHPC). The test finds that the HS-UHPC composite beam has the highest flexural strength and reasonable plastic deformation ability as compared to the other two beams. A finite element (FE) model was also constructed and benchmarked. Subsequently, the model was utilized to conduct parametric studies, aimed at exploring in-depth the flexural behavior of steel-concrete composite beams. The results from the tests and FE analyses were employed to assess the suitability of existing design specifications (i.e., AISC 360-22, GB 50017-2017, and Eurocode 4) in estimating the flexural strength of HS-UHPC composite beams. The evaluations indicated that GB 50017-2017 can reasonably estimate the flexural strength of HS-UHPC composite beams. This research provides valuable insights into the design and construction of HS-UHPC composite beams in bridges.

Keywords

UHPC High strength steel Flexural behavior Composite beam Design

Introduction

The steel-concrete (SC) composite beams are composed of concrete decks, steel beams, and stud, they have been widely used in bridges engineering. When subjected to a positive bending moment, the concrete deck mainly bears the compressive force and the steel beam mainly sustains the tensile force, allowing for the efficient utilization of the advantages offered by both steel and concrete materials (Lin et al., 2021).

The application of high-strength steel (HS) and ultra-high-performance concrete (UHPC) in bridges has been increased, due to the high strength and energy dissipation efficiency of HS (Ban et al., 2018) and the high strength and blast of UHPC (Hung et al., 2021; Lai et al., 2023). For example, HS beams were used in the Tokyo Gate Bridge in Japan and the Nanpu Bridge in China; UHPC decks were used in the Sherbrooke Bridge in Canada and the Dongting Bridge in China. The use of HS and UHPC in bridges typically often results in a thinner concrete deck and smaller steel beams, and thus reduces self-weight and material consumption (Carlos et al., 2023; Deng et al., 2021; Jun et al., 2018; Lai et al., 2022; Zhang et al., 2020).

Numerous researchers have studied the performance of SC composite beams using HS or UHPC. Yoo et al. (2016), Liu et al. (2020), Zhu et al. (2020), Hu et al. (2020), Zhu et al. (2021), Lin et al. (2022), and Guan et al. (2024) conducted tests on conventional-strength steel-UHPC (CS-UHPC) composite beams. It was found that: (i) steel-UHPC composite beams could significantly enhance the bending performance, (ii) UHPC resulted in sufficient ductile behavior, (iii) the shear lag effect became more significant as the load increased but less significant if UHPC was used, and (v) the increasing of the width for longitudinal rib leaded to a reduction in the stiffness of the waffle deck. Ban et al. (2013), Su et al. (2018), Nguyen et al. (2021), and Du et al. (2022) conducted tests on high-strength steel-conventional strength concrete (HS-CC) composite beams. It was shown that: (i) the no-linear deformation behavior of composite beams lessened with an increase of the steel beam applied, (ii) the lower compressive strain of concrete limited the extent of the full plastic zone of high-strength steel section, and (iii) the concrete strength had limited impact on the flexural strength of this composite beam. The previous research identified several factors that influence the mechanical performance of SC beams, including the connector type, yield stress of steel, reinforcement ratio, concrete type, and cross-section dimension.

However, research on composite beams using both high-strength steel and UHPC is still limited, although the potential application of HS-UHPC composite beams has attracted attention from the bridge industry. Moreover, the applicability of design methods in current specifications for designing HS-UHPC composite beams is questionable, which limits the application in practice (He et al., 2022; He, Xu, et al., 2023; He, Yang, et al., 2023; Tang et al., 2023). To tackle this, three composite beams (i.e., HS-CC, CS-UHPC, and HS-UHPC) was examined through experimental means first in this paper. A 3D FE model was benchmarked. Subsequently, the benchmarked model was used to further conduct the parametric analyses for composite beams. Finally, the research assessed and analyzed the adequacy of design specifications in predicting the flexural strength of HS-UHPC composite beams.

Experimental program

Specimens design

The specimens including an HS-CC beam, a CS-UHPC beam, and an HS-UHPC beam were tested. The specimens measured 800 mm in width and 6400 mm in length. Figure 1 shows the cross-section details of specimens. The height, top flange width, and bottom flange width of the I-shaped steel beams were 530 mm, 200 mm, and 280 mm, respectively; the height of the concrete deck was 75 mm and the width of the concrete deck was 800 mm. Details of the built-up I-shaped steel beams include the following: (1) a steel plate (thickness was 10 mm) was selected as the top flange, (2) a steel plate (thickness was 12 mm) was selected as the web, and (3) a steel plate (thickness was 16 mm) was selected as the bottom flange. Three rows of shear studs (diameter was 16 mm) with a 70 mm spacing were employed to connect the concrete deck and steel beam.

Figure 1.

Details of specimens (unit: mm).

The concrete deck had two layers of reinforcements (diameter was 8 mm) for all three specimens. The spacing of reinforcements was 100 mm and 116 mm in longitudinal and transverse, respectively. Specimen HS-CC was cured at room temperature. Specimens CS-UHPC and HS-UHPC were cured with steam at 95°C for 48 hours and then cured at room temperature until 28 days.

Materials

The measured properties are summarized in Table 1. Steel coupons were cut from each steel plate, and the stud and reinforcing bar used the same material as the specimens and were welded to the steel gasket. Three repeat tests were conducted to limit the possible discrepancy. These properties were measured per GB 228.1-2010 (Standards Press of China, 2010). Table 2 and Table 3 summarize the mixed proportions of the CC and UHPC, respectively. The tensile strength is 2900 MPa for steel fibers used in UHPC. According to the GB 50081-2019 (China Architecture and Building Press, 2019) and GB 31387-2015 (Standards Press of China, 2015), The compressive cube strength of the CC and UHPC on the day of testing was 51.5 MPa and 174.5 MPa, respectively.

Table 1.

Measured Properties of steel.

Material	Thickness (mm)	f_y (MPa)	f_u (MPa)	E_s (GPa)	δ (%)
CS	10	351	477	210	25.5
CS	12	443	610	197	29.7
CS	16	422	565	201	30.8
HS	10	617	692	210	19.9
HS	12	650	738	202	20.2
HS	16	718	780	200	21.5
Stud	-	340	438	199	-
Reinforcement	-	340	426	200	-

Table 2.

Mix Proportions of Conventional Concrete.

Cement (kg/m³)	Sand (kg/m³)	Coarse aggregate (kg/m³)	Water (kg/m³)	Superplasticizer (kg/m³)	Water-to-binder ratio
490.0	745.0	1111.0	151.9	7.4	0.31

Table 3.

Mixture Proportions of UHPC.

Cement (kg/m³)	Silica fume (kg/m³)	Quartz sand (kg/m³)	Water (kg/m³)	Superplasticizer (kg/m³)	Water-to-binder ratio	Steel fiber (%)
852.5	255.7	1000.8	221.7	21.3	0.2	2

Test setup

Figure 2 shows the specimens were loaded incrementally using a rigid transfer beam and a 5000 kN hydraulic jack. The following loading protocol was used. Firstly, preloading up to 20 kN was conducted. Then, the load was incrementally applied at a rate of 50 kN per increment with a loading speed of 0.5 kN/s. For each increment, the load was sustained for 2 minutes to observe the specimens. The steel beam yielded, displacement-controlled loading was used and the loading speed is 1 mm/min. When the specimens failed, the test was ended.

Figure 2.

Test setup.

Figure 1(b) also shows the arrangement of gauges. 12 strain gauges were arranged at the loading points and the midspan. Five displacement gauges (i.e., LVDTs) were employed to collect the deflection of the specimens. The supports were arranged with a spacing of 1000 mm from the midspan. The LVDT was also arranged on each side of the ends of concrete deck and steel beam.

Analysis of the test results

Failure modes

The failure mode of the three specimens were shown in Figure 3. During the initial loading stage, no obvious phenomena were observed for all three specimens. Upon the yielding of the bottom flange, the specimens entered the plastic stage, and the deformation of the specimen gradually enlarged until reaching the peak load. For specimens HS-UHPC and HS-CC, the UHPC deck and concrete deck was crushed at a load of 2455 kN and 1789 kN, respectively; for specimen CS-UHPC, the bottom flange fractured at a load of 1255 kN.

Figure 3.

Failure model of specimens.

The comparison between HS-UHPC and HS-CC indicates that using UHPC improves both flexural strength and ductility. For specimen HS-UHPC, the addition of steel fibers along with the high compressive strength of UHPC results in a stronger compressive region and a limited failure surface, effectively leading the steel beam to fully enter the plastic stage. For specimen HS-CC, however, the concrete deck with a large area in the midspan was crushed, but the bottom flange had just yielded, indicating that the elastic-plastic properties of the steel were not fully utilized. Therefore, using UHPC and high-strength steel is a better match for high-strength composite beams. Specimen CS-UHPC failed because of the fracture of the bottom flange. This was possibly caused by local defects of steel materials (which will be discussed later in Section 3.3), leading to the stress concentration phenomenon after yielding.

Moment - midspan displacement curves

The moment - midspan displacement curves was shown in Figure 4. Three stages (elastic, elastoplastic, descending) was divided for each curve. During the elastic stage (less than 0.8 M_u), all three specimens exhibited an almost linear moment - displacement response. Thereafter, the bottom beam flange yielded and the stiffness of the composite beam started to degrade. The specimens entered the elastoplastic stage (0.8 M_u to 1.0 M_u), and the slope of the moment - displacement curves decreased accordingly. Upon reaching the peak moment (1.0 M_u), the specimens experienced failure, and the corresponding curve transitioned into the descending stage. Specimen HS-UHPC showed a much better deformation ability at the peak load.

Figure 4.

Moment - midspan displacement curve.

Deflections of the SC composite beams were measured using the five displacement gauges. Figure 5 shows the different load levels of the measured deflections, which matched with the half-sine wave, especially at the elastic stage (i.e., less than 0.8 M_u).

Figure 5.

Deflections along the span at different load levels.

Longitudinal strain

The longitudinal strain of the concert deck and steel beam at the midspan (up to 1.0 M_u) was shown in Figure 6. At the elastic stage, the longitudinal strain was linearly distributed along the beam height, indicating that the composite beam complies with the assumption of plane section. After the load exceeds 0.8 M_u, the bottom flange of specimens HS-UHPC and CS-UHPC was no strain data given, as shown in Figure 6(a) and (c). This happened because the strain gauges exceeded their range after the load exceeded 0.8 M_u. With the load increased, the neutral axis moved upward from the steel beam web. When approaching the ultimate load, the neutral axis moved faster but still remained in the steel beam. This indicates that both the concrete deck and UHPC deck experienced compressive forces throughout the entire loading procedure.

Figure 6.

Distributions of the longitudinal strain at the midspan.

The maximum strains measured of the UHPC deck and concrete deck for specimens HS-UHPC and HS-CC were 4400 με and 2400 με, which exceeded the peak compressive strain of the UHPC and CC, respectively. The maximum recorded strains at the HS bottom flange exceeded 5000 με, which surpassed the yield strain (i.e., 3200 με) significantly. For specimen CS-UHPC, after the bending moment reached 0.8 M_u, the strain of the web near the bottom plate experienced a rapid increase. After the bottom flange yielded, the fracture section produced stress concentration, owing to the initial defects of the material. The strain of the UHPC deck was 1500 με, which is below the nominal ultimate compressive strain for the UHPC. Consequently, the failure mode of specimen CS-UHPC manifested as a fracture in the bottom flange. That was why specimen CS-UHPC showed relatively brittle behavior.

Load - slip curves

The slip was determined by subtracting the average longitudinal displacement measurements of the steel beam and concrete deck at both ends. The slip was small before reaching the ultimate moment (M_u) as shown in Figure 7. When M_u was reached, the slip would increase rapidly until the specimen failed; however, the slip was still negligible, where the maximum slips for specimens HS-CC, CS-UHPC, and HS-UHPC were 0.40, 0.13, and 0.30 mm, respectively. This confirmed the designed shear studs were sufficient in ensuring full shear connection. Specimen HS-CC has a largest slip, due to the natural bond strength of shear studs embedded in UHPC is greater than that in CC, and the elastic modulus of UHPC is slightly higher than CC.

Figure 7.

Moment - slip curves.

Finite element (FE) analysis

Model details and benchmarking

Figure 8(a) shows a typical FE model using ABAQUS (2016). The S4R shell element was employed for the steel beam. The C3D8R solid element was employed for the concrete deck. The beam element (B31) was employed to model shear studs. The zero-length connector elements (CONN3D2) were used to attached to the studs and the top flange. The truss element (T3D2) was employed to model rebars. The mesh size was set as 40 mm based on the mesh sensitivity analysis. The explicit dynamic procedure was adopted.

Figure 8.

Overview of the typical FE model.

A bilinear stress-strain curve was used to simulate the steel beam and stud with a linear hardening branch, the post-yield stiffness was 0.01E_s (Wang et al., 2022). The concrete damage plasticity (CDP) was used to employ the behavior of concrete deck. For CC in compression, the stress-strain curve proposed by Tao et al. (2013) was used. The uniaxial of concrete in tension was modeled by CEB-FIP (2010). For UHPC in compression and tension, the stress-strain curves proposed by Yang et al. (2021) were used.

The penalty method (friction coefficient equal to 0.5) was used to model the tangential direction, while hard contact was used to model the normal direction for the contact interaction between concrete slab and steel beam (Korkmaz et al., 2022; Tang et al., 2023; Tong et al., 2022). The shear behavior was modeled using shear force-relative slip behavior model (Ollgaard et al., 1971). The rebars were embedded in the concrete. As shown in Figure 8, the loading surface was coupled to the reference points RP-1 and RP-2, following which the vertical load was imposed at these reference points. The support surface was coupled to the reference points RP-3 and RP-4.

Verification of the FE model

Figure 9 compares the failure modes for three specimens. In this figure, the black area indicates the position of concrete crushing. As shown, the FE model successfully simulated the concrete deck crushing process for specimens HS-UHPC and HS-CC. Besides, by removing the mesh from the bottom flange, the fracture location was successfully identified, showing the great agreement of the captured fracture in the steel beam bottom flange with tests. Figure 10 compares moment - deflection curves, and the curves extrapolated from the FE analysis closely matched the measured curves.

Figure 9.

Comparison of measured and predicted failure mode.

Figure 10.

Comparisons of the moment - deflection curves between FE analysis and experimental tests.

Parametric studies

For a more in-depth exploration of the impact of the height of the steel beam (h_s), yield stress of steel beam (f_y), compressive cylinder strength of concrete deck (f’_c), and thickness of bottom flange (t_fb) on the SC composite beams, a total of 30 FE parametric analyses were conducted in this section. Table 4 summarizes the details and main results of these analyses. In this table, μ is the displacement ductility factor (μ = u_m / u_y); u_y and u_m are the deflection at the yield and the peak load, respectively.

Table 4.

Summary of parametric Analysis for SC Composite Beams.

No.	L (mm)	b_c (mm)	h_s (mm)	f_y (MPa)	f’_c (MPa)	t_fb (mm)	M_u,FE (kN · m)	μ
HS-UHPC-1	6000	800	300	550	133	16	1397.8	3.25
HS-UHPC-2	6000	800	430	550	133	16	2081.2	4.03
HS-UHPC-3	6000	800	530	550	133	16	2440.4	4.29
HS-UHPC-4	6000	800	650	550	133	16	3377.7	3.17
HS-UHPC-5	6000	800	750	550	133	16	4081.2	3.54
CS-UHPC-1	6000	800	300	345	133	16	937.3	5.34
CS-UHPC-2	6000	800	430	345	133	16	1368.5	5.84
CS-UHPC-3	6000	800	530	345	133	16	1740.0	7.46
CS-UHPC-4	6000	800	650	345	133	16	2240.3	7.43
CS-UHPC-5	6000	800	750	345	133	16	2696.7	7.62
HS-CC-1	6000	800	300	550	50	16	1155.0	1.31
HS-CC-2	6000	800	430	550	50	16	1679.7	1.28
HS-CC-3	6000	800	530	550	50	16	2102.0	1.25
HS-CC-4	6000	800	650	550	50	16	2743.1	1.10
HS-CC-5	6000	800	750	550	50	16	3221.6	1.17
CS-UHPC-6	6000	800	530	235	133	16	1242.0	8.97
CS-UHPC-1	6000	800	530	345	133	16	1740.0	7.46
HS-UHCP-6	6000	800	530	460	133	16	2220.5	5.57
HS-UHCP-3	6000	800	530	550	133	16	2440.4	4.29
HS-UHCP-7	6000	800	530	690	133	16	3124.0	3.06
HS-CC-1	6000	800	530	550	50	16	1996.9	1.23
HS-CC-6	6000	800	530	550	60	16	2061.8	1.29
HS-HC-1	6000	800	530	550	80	16	2300.8	2.13
HS-HC-2	6000	800	530	550	100	16	2387.2	2.91
HS-UHPC-3	6000	800	530	550	133	16	2440.4	4.29
HS-UHPC-8	6000	800	530	550	133	12	2274.2	3.94
HS-UHPC-9	6000	800	530	550	133	14	2400.5	3.99
HS-UHCP-3	6000	800	530	550	133	16	2440.4	4.29
HS-UHPC-10	6000	800	530	550	133	18	2764.8	4.28
HS-UHPC-11	6000	800	530	550	133	20	2915.1	4.38

Impact of the height of steel beam (h_s)

Three groups of composite beams (i.e., HS-CC, CS-UHPC, and HS-UHPC) were analyzed. Each group consisted of five components with different steel beam web heights, i.e., 300, 430, 530, 650, and 750 mm. t_fb was set as 16 mm, for HS-CC group f_y was set as 550 MPa and f’_c was set as 50 MPa, for CS-UHPC group f_y was set as 345 MPa and f’_c was set as 133 MPa, for HS-UHPC group f_y was set as 550 MPa and f’_c was set as 133 MPa. As shown in Figure 11, for all three groups of specimens, M_u,FE increased and u_m decreased with increasing h_s. For example, when h_s increased from 300 to 750 mm, M_u,FE increased by up to 180%, and u_m decreased by up to 101%. Under the same h_s, the u_m of HS-UHPC specimens is about three times higher than HS-CC specimens. The material type had an appreciable influence on u_m. The CS-UHPC specimens had the best ductility. The HS-UHPC specimens exhibit acceptable ductility and higher flexural strength as compared to CS-UHPC specimens.

Figure 11.

Effect of the height of steel beam.

Impact of the yield stress of steel (f_y)

The study considered five different yield stress of steel beam values, i.e., 235, 345, 460, 550, and 690 MPa h_s was set as 530 mm, f’_c was set as 133 MPa, and t_fb was set as 16 mm. As shown in Figure 12, M_u,FE increased but u_m decreased with increasing f_y. When f_y increased from 235 to 690 MPa, M_u,FE increased by up to 151%, and u_m decreased by up to 42%. The reason is that the HS has higher strength but lower ductility.

Figure 12.

Effect of the yield stress of steel beam.

Impact of compressive strength of concrete (f’_c)

The study considered five different compressive strength values of concrete deck, i.e., 50, 60, 80, 100, and 133 MPa h_s was set as 530 mm, f_y was set as 550 MPa, and t_fb was set as 16 mm. As shown in Figure 13, both M_u,FE and u_m increased with increasing compressive strength of concrete. When compressive strength of concrete increased from 50 to 133 MPa, M_u,FE increased by up to 260%, and u_m increased by up to 185%. The reason is that higher-strength concrete has higher peak strains, resulting in the increased plasticity of the high-strength steel. Hence, the UHPC has a better match with high-strength steel as compared to CS.

Figure 13.

Effect of the compressive strength of the concrete.

Impact of thickness of steel beam bottom flange (t_fb)

The study considered five different thickness values of the steel beam bottom flange, i.e., 12, 14, 16, 18, and 20 mm h_s was set as 530 mm, f_y was set as 550 MPa, and f’_c was set as 133 MPa. Figure 14 shows the impact of the thickness of the steel beam bottom flange. As shown, M_u,FE increased but u_m decreased with increasing t_fb. When t_fb increased from 12 to 20 mm, M_u,FE increased by up to 28%, and u_m decreased by up to 15%. The reason is that as t_fb increases, the flexural strength also increases.

Figure 14.

Effect of the thickness of steel beam bottom flange.

Evaluation of design specifications

In this section, the applicability of several design specifications was evaluated, including GB 50017-2017 (China Architecture and Building Press, 2017), Eurocode 4 (CEN, 2004), and AISC 360-22 (AISC, 2022).

Current design specifications

GB 50017-2017

GB 50017-2017 (China Architecture and Building Press, 2017) uses two methods to calculate the flexural strength of composite beam (M_n), depending on the placement of the plastic neutral axis Figure 15. When the plastic neutral axis located within the steel beam, M_n is estimated using equation (1):

M_{n} = b_{e} h_{c} f_{c} y_{1} + A_{s} f_{y} y_{2}

(1)

where b_e is the effective deck width, h_c is the deck thickness, A_s is the compression area on the steel beam cross-section.

Figure 15.

Plastic stress distributions for the composite beam.

When the plastic neutral axis located within the concrete deck, M_n is estimated using equation (2):

M_{n} = b_{e} x f_{c} y

(2)

where x is the height of the compression zone of the concrete flange, y is the distance between the centroid of the tensile force in the steel beam and the centroid of the compressive force in the concrete deck.

Eurocode 4

Eurocode 4 (CEN, 2004) uses a similar approach as GB 50017-2017 (China Architecture and Building Press, 2017), while the concrete stress is reduced to 0.85 f’_c. When the plastic neutral axis located within the steel beam, M_n is estimated using equation (3):

M_{n} = 0.85 b_{e} h_{c} f_{c} y_{1} + A_{s} f_{y} y_{2}

(3)

When the plastic neutral axis located within the concrete deck, M_n is estimated using equation (4):

M_{n} = 0.85 b_{e} x f_{c} y

(4)

AISC 360-22

AISC 360-22 (AISC, 2022) calculates M_n using equation (5a), where C is the compressive strength of the concrete deck (the smaller of Equations (5b) and (5c)).

M_{n} = C (d_{1} + d_{2}) + f_{y} A_{s} (d_{3} - d_{2})

(5a)

C = A_{s} f_{y}

(5b)

C = 0.85 f_{c}^{,} A_{c}

(5c)

Evaluations

Figure 16 presents a comparison of the flexural strength derived from various design specifications with the flexural strength determined through testing and FE analyses. All three specifications assume full plasticity through the cross-section (i.e., full yielding in the steel and full crushing in the concrete), but they are generally conservative as shown in Figure 16. The main reason is the strain hardening of the steel beam is ignored. Table 5 lists the statistics of M_n/M_u ratio. All three specifications have better applicability to members using HS beams. This is because the strain hardening (increase of strength beyond the yield stress) of these members are lower than that using CS beams (about 20% and 10%, respectively) when the beams reach the peak moment. The specifications have better applicability to members using CC decks. This is because these members experience crushing of concrete in the compressive region before significant strain hardening in the tensile region of the steel beam was developed. Eurocode 4 (CEN, 2004) and AISC 360-22 (AISC, 2022) are more conservative than GB 50017-2017 (China Architecture and Building Press, 2017). This is because the maximum stress developed in the concrete is reduced from f’_c (in GB 50017-2017) to 0.85 f’_c (in Eurocode 4 and AISC 360-22).

Figure 16.

Comparison of calculated strength with experimental or FE results.

Table 5.

Summary of the M_n/M_u for Different Design Specifications.

Specimens	GB 50017-2017		Eurocode 4		AISC 360-22
Specimens	Mean	Standard deviation	Mean	Standard deviation	Mean	Standard deviation
HS-UHPC	0.94	0.10	0.82	0.07	0.84	0.09
CS-UHPC	0.85	0.12	0.81	0.11	0.82	0.11
HS-CC	0.98	0.08	0.83	0.07	0.92	0.05
All	0.89	0.10	0.83	0.09	0.85	0.10

Conclusions

This paper examines the flexural behavior of high-strength SC composite beams through experimental and numerical investigation. The study assesses and discusses the suitability of existing design equations in accurately predicting the flexural strength of high-strength SC composite beams. The following conclusions can be gated:

(i) Three test specimens had flexural failure at the midspan. The HS-UHPC specimen had the highest flexural strength and satisfactory ductility. The conventional concrete deck does not match with an HS steel beam, i.e., this combination for composite beam reduces both the ductility and strength.

(ii) The FE models developed were able to exactly forecast the flexural behavior of SC composite beams. Parametric studies showed that the flexural strength increased with increasing height of the steel beam, yield stress of steel, compressive strength of concrete, and thickness of bottom flange; the ductility reduced with increasing height of the steel beam, yield stress of steel, or decreasing compressive strength of concrete.

(iii) Due to the strain-hardening behavior of steel beams was not taken into account, existing specifications, i.e., GB 50017-2017, Eurocode 4, and AISC 360-22, are generally conservative. GB 50017-2017 gives a relatively reasonable estimation of flexural capacity for HS-UHPC members, and is recommended for designing high-strength I-shaped steel-UHPC composite beams.

Footnotes

Declaration of conflicting interests

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

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: The research was supported by the National Natural Science Foundation of China (Award No.: 51978170, 52108122).

ORCID iD

Xiaoqiang Yang

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Flexural behavior and design of high-strength I-shaped steel-UHPC composite beams

Abstract

Keywords

Introduction

Experimental program

Specimens design

Materials

Test setup

Analysis of the test results

Failure modes

Moment - midspan displacement curves

Longitudinal strain

Load - slip curves

Finite element (FE) analysis

Model details and benchmarking

Verification of the FE model

Parametric studies

Impact of the height of steel beam (hs)

Impact of the yield stress of steel (fy)

Impact of compressive strength of concrete (f’c)

Impact of thickness of steel beam bottom flange (tfb)

Evaluation of design specifications

Current design specifications

GB 50017-2017

Eurocode 4

AISC 360-22

Evaluations

Conclusions

Footnotes

Declaration of conflicting interests

Funding

ORCID iD

References

Impact of the height of steel beam (h_s)

Impact of the yield stress of steel (f_y)

Impact of compressive strength of concrete (f’_c)

Impact of thickness of steel beam bottom flange (t_fb)