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Strength enhancement of high strength steel beams by engineered cementitious composites encasement
Engineering Structures 207 (2020) 110288
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Strength enhancement of high strength steel beams by engineered cementitious composites encasement
T
Md. Imran Kabira, C.K. Leea, , Mohammad M. Ranaa, Y.X. Zhangb ⁎
a b
School of Engineering and Information Technology, The University of New South Wales, Canberra, ACT 2600, Australia School of Computing, Engineering and Mathematics, Western Sydney University, NSW 2751 Australia
ARTICLE INFO
ABSTRACT
Keywords: High strength steel Engineered cementitious composite encasement Flexural resistance Bond-slip behaviours Finite element modelling
This study proposes a method of using Polyvinyl Alcohol Engineered Cementitious Composites (PVA-ECC) encasement to provide continuous restraints along the compression flange of High Strength Steel (HSS) section so that it will reach its sectional plastic moment resistance under bending without lateral restraint. In order to demonstrate the effectiveness of the proposed method, experimental and numerical investigations were carried out to study the flexural strength of the ECC encased HSS beams (ECC-HSS beams). Six simply supported beams fabricated with identical HSS sections but with different encasement configurations were tested until failure. Flexural resistance and failure modes of the ECC-HSS beams were compared with similar bare HSS and normal concrete (NC) encased HSS beams (NC-HSS beams). It was found that when compared with the bare HSS and NCHSS beams, a significant enhancement in flexural resistance was achieved for the ECC-HSS beams. More importantly, this study confirmed that the compressive ECC layers was crushed after the compression flanges were yielded and therefore successfully prevented the onset of lateral torsional buckling. Besides the flexural responses, the interfacial slip behaviours along the compression flange of the HSS section were also studied. Finally, a finite element (FE) model was developed and validated against the experimental results.
1. Introduction When comparing with Normal Strength Steel (NSS) which normally has a yield strength not more than 460 MPa, High Strength Steel (HSS), which specifically refers to steel with a yield strength equal to or more than 690 MPa, can offer many advantages in construction. Due to their higher yield strength, the use of HSS in structural design may reduce the cross-section dimension and thus the self-weight of the structures significantly. Such merits of HSS could alleviate the fabrication and lifting difficulties during construction. However, one of the disadvantages of using HSS sections is that their thinner wall thicknesses could lead to a higher slenderness ratio which eventually results in instability. In general, with the same full sectional plastic moment resistance, a laterally unrestrained HSS beam is more vulnerable to lateral torsional buckling (LTB) and local buckling than an NSS beam. Such unstable effects may lead to pre-mature failure before the beam’s sectional plastic moment resistance is fully utilized. In recent years, HSS had gained much research interest but most of the researches were focused on the behaviours of bare HSS sections such as columns [1–3], beams [4–6] and connections [7,8]. For steelconcrete composite structures, although some researches on concrete-
⁎
filled HSS tubular columns had been conducted [9–12], researches on HSS-concrete composite beams are relatively uncommon. Most of the researches done on HSS-concrete composite beams were focused on composite beams which are constructed by connecting concrete slab with HSS section using shear connectors. Uy and Sloane [13] studied the flexural behaviour of two composite steel-concrete T-beams and numerical results were compared with experimental results. A satisfactory agreement was observed between them when the slip between concrete and steel section was considered. However, the flexural strengths predicted by the Australian [14] and European [15] design standards were found to be slightly unconservative. Zhao and Yuan [16] tested four composite steel-concrete T-beams subjected to twopoint concentrated loads. It was found that by using HSS to replace NSS, significant improvements were observed regarding the moment capacity, longitudinal split and interfacial slip. Recently, Ban and Bradford [17] conducted an extensive parametric study on the flexural behaviours of composite beams fabricated by using different HSS grades. Results obtained from the finite element (FE) models were compared with those obtained from classical plastic analysis. It was found that while the Eurocode 4 [15] predicted the bending resistance of concreteNSS composite beams accurately, it over-predicted the flexural
Corresponding author. E-mail address: [email protected] (C.K. Lee).
https://doi.org/10.1016/j.engstruct.2020.110288 Received 19 October 2019; Received in revised form 20 December 2019; Accepted 23 January 2020 Available online 30 January 2020 0141-0296/ © 2020 Elsevier Ltd. All rights reserved.
Engineering Structures 207 (2020) 110288
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Table 1 Compositions of ECC, LWC and NC. ECC
Table 2 Section details of encased beams (all dimensions are in mm).
LWC
NC
Ingredient
Mix ratio
Ingredient
Mix ratio
Ingredient
Mix ratio
Cement Fly Ash Sand/ binder Water HRWRA PVA-fibre (vol%)
1.0 1.2 0.36
Cement Sand 8 mm LWA
1.0 1.7 0.92
Cement Sand 7 mm aggregate
1.0 1.81 1.63
0.56 0.01 2.2
14 mm LWA Water
0.92 0.4
10 mm aggregate Water Superplasticiser
1.1 0.42 0.015
Specimen
Overall size B×D
HSS Section d × bf × tw × tf
tECC,
Bare HSS beam NC100-f E50bt-f E75bt-f E25t-f E30t-p
–
150 × 75 × 6 × 8
–
–
–
– 52.5 78.75 52.5 52.5
– 52.5 78.75 – –
210 105 52.5 157.5 –
175 175 175 175 175
× × × × ×
210 210 210 210 180
comp
tECC,
ten
tLWC/NC
Legends: B = width of beam. D = depth of beam. d = depth of HSS section. bf = flange width. tf = flange thickness. tw = web thickness. tECC, comp = compression ECC layer thickness. tECC, ten = tension ECC layer thickness. tLWC/NC = LWC/NC layer thickness.
Note: LWA = Light-weight aggregate.
resistance of HSS-concrete composite beams. It appears that most recent research related to fully encased steel beam is largely limited to concrete encased NSS beams only. An exception is the study by Rana et al. [18] which showed that by replacing Normal Concrete (NC) with Engineered Cementitious Composites (ECC) and Lightweight Concrete (LWC), the flexural performances of a fully encased NSS-NC composite beams could be improved significantly. In addition, the recent work done by Kabir et al. [19] demonstrated that unstable LTB failure was prevented by using ECC to encase the compression flange of the NSS section. Numerical analyses conducted in [19] further predicted that the flexural resistance of a fully encased beam could be enhanced by substituting the NSS section with HSS section. However, no experimental study was conducted to confirm such predictions that the encasement could lead to full composite actions and prevent local buckling. Hence, in this study, a similar approach was applied to enhance the flexural resistance of HSS beams by using ECC and LWC to provide continuous lateral restraint to the compressive flange of the HSS section. The results obtained were then compared with the performances of a bare HSS beam and a control NCHSS encased beam. Furthermore, a FE modelling procedure was employed to model the bending and interfacial slip behaviours of the beams in details. It should be noted that while ECC is well-known for its superior tensile ductility and tensile strain hardening behaviour, it also has superior compression ductility (compressive strain at peak strength can easily exceed 0.5%) when comparing with normal concrete (0.23–0.28%). However, ECC’s higher tensile ductility and tensile strength are still much smaller than any grade of steel. Hence, when ECC is employed to encase the HSS section, such superior tensile performance is not very important as nearly all the tensile force at the tension side of the beam will be resisted by the steel section anyway. On the contrary, ECC should be used in the compression zone of the beam to prevent immature local buckling of steel section as its compressive strain at peak strength (0.5%) is larger than the yield strength of HSS (0.38%). In this case, it is the superior compression ductility that makes ECC a suitable material that could be used in conjunction with HSS.
2. Experimental investigation 2.1. HSS, ECC and LWC used The 3.4 m long, built-up HSS I-sections used were fabricated by using BISALLOY structural 80 steel plates which has a nominal yield strength (0.2% proof stress) of at least 690 MPa. Single pass, 6 mm continuous fillet welding were employed to join the two flange plates with the web plate. In order to minimize thermal distortion generated during welding, welding along the top and bottom flange joints was completed in an alternative manner after every 0.5 m of welding. The polyvinyl alcohol ECC (PVA-ECC) proposed by Meng et al. [20], which use of local ingredients to reduce the materials cost, was adopted in this study. It has sufficient ductility in both tension and compression to work with HSS and was found to has a slightly higher bond strength with steel section [21] than the conventional concrete. For the LWC, two different sizes (8 mm passing and 14 mm passing) of light-weight aggregate (LWA) with same mix ratio were used. For the NC, it was prepared by using two different sizes of aggregates (7 mm passing and 10 mm passing) of equal mix ratio. Table 1 lists the mix ratios for ECC, LWC and NC. 2.2. Specimen configurations Six beams were prepared and tested for studying their flexural and bond-slip behaviours. These beams include one bare HSS beam, one NC encased HSS (NC-HSS) beam and four ECC-LWC encased HSS (ECCHSS) beams. For the NC-HSS beam, the whole HSS section was encased by NC and is labelled as the NC100-f beam. The ECC-HSS beams were prepared by changing the portion of encasement of ECC/LWC and positions of the ECC layer(s). The first and second ECC-HSS beams were prepared by encasing the two flanges of the HSS sections by ECC in such
Fig. 1. Cross-sectional profiles of the tested beams. 2
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Table 3 Material properties of ECC, LWC, NC and HSS. Material/Specimen
Properties
ECC (tension side) E50bt-f, E75bt-f ECC (Compression side) E50bt-f, E75bt-f ECC (Compression side) E25t-f, E30t-p beam LWC: E50bt-f, E75bt-f LWC: E25t-f NC: NC100-f HSS section (flanges) HSS section (web)
Compressive strength (MPa)
Tensile strength (MPa)
Modulus of elasticity (MPa)
HSS yield strength, fy (MPa)
HSS ultimate strength, fu (MPa)
69.41 61.53 67.26
4.79 5.05 5.19
21,129 22,855 19,220
– – –
– – –
44.35 45.27 58.65 – –
– – – – –
26,127 25,479 32,863 197,655 198,220
– – – 748 (at εy = 0.38%) 759 (at εy = 0.38%)
– – – 842 (at εu = 10%) 856 (at εu = 10%)
Fig 2. Details of (a) instrumentation, (b), (c) positions of LVDTs for end slip measurement (d) locations of strain gauges. 3
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Fig. 3. Load-deflection curves of specimens.
a way that the ECC thicknesses are equal to 50% and 75% of beams’ depth, respectively. For the remaining parts of the beams, the HSS sections are encased by LWC to reduce the weight of beams. Therefore, these two beams are named as the E50bt-f and the E75bt-f beams, respectively. For the third ECC-HSS beam, only the compression flange is covered by a layer of ECC corresponding to 25% of the beams’ depth and remaining part of the beam is covered by LWC so that it is still fully encased. Therefore, this beam is referred as the E25t-f beam. The last ECC-HSS beam is a partially encased beam such that only the compression flange is covered by an ECC layer corresponding to 30% of the beams’ depth and is labelled as the E30t-p beam. It is the lightest ECCHSS beams tested to explore the slip characteristics between the HSS section and the ECC layer. The encasement configurations and detailed dimensions of the beams tested are shown Fig. 1 and Table 2. Note that while the NC100-f and E30t-p beams only required one casting step, the E25t-f beam and the two “bt-f” beams required two and three cast steps, respectively. Detailed information of the casting procedure of the beams is described in [19]. In order to determine the mechanical properties of the ECC, LWC and NC, three cylinders with standard dimensions of 100 mm ∅ × 200 mm H were prepared from each mix of ECC, LWC and NC [22,23]. Their Young’s Modulus and compressive strengths at 28 days were then determined in accordance with ASTM C469 [24]. Furthermore, three dog-bone shaped samples of standard dimensions [20] were tested under uniaxial tension to obtain the ECC’s tensile properties. In order to obtain the tensile properties for the HSS sections, direct tensile tests were conducted on three standard HSS coupons. The mean values of the mechanical properties obtained are given in Table 3.
Fig. 4. Final failures of the specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
employed to record the mid-span (LVDT-1) and the loading points (LVDT-2 and LVDT-3) deflections. Since HSS sections were used to form the encased beams, it is expected a high shear stress would be developed between the ECC/NC-steel interfaces and slipping could occur before the beams are failed by bending. Therefore, four additional LVDTs (LVDT-4 to 7) were attached at the two ends of the encased beams to measure the interfacial slip between the ECC/LWC/NC and
2.3. Bending test procedures The specimens were loaded using four-point bending with a clear span of 3000 mm and two load points were placed at 1000 mm from the supports. The supports and loading configuration are shown in Fig. 2(a). Three linear variable displacement transducers (LVDTs) were Table 4 Bending test results. Beam
Failure load (kN)
Mid-span deflection at failure load (mm)
Enhancement of failure load (%) as compared to bare HSS beam
Weight of thebeam (kg)
Weight reduction (kg) as compared with NC100-f
Bare HSS beam NC100-f E50bt-f E75bt-f E25t-f E30t-p
139.1 175.79 192.15 191.79 214.61 204.91
54.47 43.47 56.02 54.73 76.77 148.31
– 26.37% 38.14% 37.88% 54.28% 47.31%
53.5 337 275 281 270 111
– – 62 (18.3%) 56 (16.5%) 67 (20%) 226 (67.1%)
4
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Fig. 5. Comparison of mid-span deflection measured by LVDT and DIC: (a) NC100-f, b) E50bt-f, (c) E75bt-f, (d) E25t-f, (e) E30t-p.
the HSS flanges. For LVDT-4 to 7, the base of the LVDT was attached to one end of a PVC tube while the other end of the tube was attached to the ECC/LWC/NC surface. The tip of the LVDT was placed at the flange as shown in Fig. 2(b) and (c). As a result, the PVC tubes and the LVDTs will rotate with the beam end and the measured relative horizontal displacements between the ECC/LWC/NC surfaces and the top flange will not be affected by the beam end rotation. The HSS sections were also instrumented with four strain gauges (SG-1, SG-2, SG-3 and SG-4) to measure the strain history during the tests. For the encased beams,
two strain gauges (SG-5 and SG-6) were installed at the top surfaces of beams to monitor the compressive strain of ECC/NC. The instrumentation scheme of the beams is shown in Fig. 2(a) and (d). In addition, three high-resolution digital cameras were placed in front of the beams to capture the development of crack distribution along the entire beam by using the Digital Image Correlation (DIC) technique. In all tests, displacement controlled static loading was applied by a hydraulic jack through a spreader beam. The applied force was measured by load cells placed under the supports (Fig. 2a). Loads were 5
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3. Bending test results 3.1. Flexural strengths of the specimens Fig. 3 summarizes the bending test results in the form of load-deflection curves. The full plastic moment resistance of the HSS section (170 kN), which is calculated based on the actual yield strength of the HSS coupons (Table 3), is drawn as a horizontal line in Fig. 3. The failure loads carried by the beams and the corresponding mid-span deflections are listed in Table 4. In Table 4, the failure load for the bare steel beam is the load when LTB occurred. For the fully encased beams (i.e. NC100-f, E50bt-f, E75bt-f and E25t-f), the failure load is the load when the top NC or ECC layer was crushed. For the partial encased beam E30t-p, the failure load is the highest load sustained at the end of the test. It should be mentioned that for all beams, no failure nor any crack was detected in any part of the welded joints between the HSS sections’ flanges and web at the end of the tests. 3.1.1. The bare HSS beam The bare HSS beam was failed by LTB at a mid-span deflection of 54.47 mm under an applied load of 139.1 kN. Fig. 4(a) shows that the pure bending region of the HSS section was buckled laterally at a loading about 18% lower than the sections’ full plastic moment resistance. 3.1.2. The NC100-f beam The failure load of this beam was 175.79 kN which was only 3.4% higher than the plastic moment resistance of the HSS section despite that the HSS section was fully encased by NC with a compressive strength of 58.65 MPa. Hence, the NC encasement virtual did not make contribution to the beams’ flexural resistance. This was because the top NC cover was crushed before the HSS section was fully yielded. Nonetheless, as the HSS section was laterally restrained by the NC encasement, no LTB occurred and the flexural resistance was increased by 26.67% when compared with the bare HSS beam. The failure mode of this beam was a brittle crushing of the compressive concrete layer which caused a sharp 13.5% drop in loading. At the end of the test, extensive cracking observed within the constant moment region as shown in Fig. 4(b). 3.1.3. The E50bt-f and E75bt-f beams The failure load of the E50bt-f beam was 192.15 kN at a defection of 56.02 mm. The failure load is 13% and 38.1% higher than the full plastic moment resistance of the HSS section and bare HSS beam, respectively. Since ECC has a higher compressive ductility (0.45% strain) than NC (0.28% strain), the mid-section of encased HSS section attained its plastic resistance before the top ECC layer was crushed. Therefore, the E50bt-f beam carried 9.3% higher load than the NC100-f beam. After the yielding of the flanges of HSS section, the outermost fibre of the ECC layer in the constant moment region was crushed which resulted a drop of 9.34% in loading. The failure mode of the E50bt-f beam was the yielding of HSS section first and then the crushing of compressive ECC layer. When compared with the NC100-f beam, the crushing of the compressive ECC layer was found to be much more ductile and gentler. The shape of the E50bt-f beam at the end of the test is shown in Fig. 4(c). As the only difference between the E50bt-f beam and the E75bt-f beam was the thickness of the ECC layers, the failure load, load-deflection curve and failure mode of the E75bt-f beam are all similar to the E50bt-f beam. Furthermore, Fig. 4(d) and Table 4 show that the deflected shape, flexural resistance and the weight of the E75bt-f beam are practically identical with the E50bt-f beam. Hence, this test results strongly suggested that in practice, there is no need to increase the thickness of the ECC layer beyond half of the beam’s depth.
Fig. 6. Crack analysis from DIC for the specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
applied at a rate of 0.5 mm/min until LTB occurred (for the bare steel beam) or until the compressive ECC/NC layer within the constant moment region of beams was crushed. The loading rate was increased to 1 mm/min after the compressive NC/ECC layer was failed and the tests were stopped when a 125 mm mid-span deflection was reached. Measurements from the load cells, LVDTs and strain gauges were automatically recorded by data loggers at an interval of 10 s. 6
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Fig. 7. Strain (measured at mid-section of beam) distribution along the depth of the specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30tp (tensile strain is positive).
3.1.4. The E25t-f beam Fig. 3 shows that the failure load of the E25t-f is the highest among all tested beams. This can be explained by the fact that the top flange of HSS section of E25t-f beam was encased by ECC with 10% higher compressive strength (67.26 MPa) than that of the E50bt-f/E75bt-f beams (61.53 MPa). In addition, the strain at peak compressive strength was 20% higher for E25t-f beam when compared with E50bt-f/E75bt-f beam. Therefore, its failure load is not only 26.2% higher than the HSS section’s full plastic moment but also is 54.1%, 22.1%, 11.7% and 11.9% higher than the bare HSS, the NC100-f, the E50bt-f and the
E75bt-f beams, respectively. Moreover, Fig. 3 shows that the failure load was achieved at a deflection of 76.8 mm which is greater than all other fully encased beams. Fig. 4(e) shows the deflected shape and crack patterns of this beam at failure. 3.1.5. The E30t-p beam The E30t-p beam is the lightest encased beam and has a weight equal to 32% of the NC100-f beam. However, its failure load is 47.3% and 16.6% higher than the bare HSS and NC100-f beams, respectively. The tension flange of the E30t-p beam was yielded when the applied 7
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Table 5 Beam-end slips of encased beams. Beam
NC100-f E50bt-f E75bt-f E25t-f E30t-p
Beam-end slips at different stages of testing At yielding of tension flanges of HSS section (mm)
At crushing of NC/ ECC (mm)
At final stage (mm)
0.37 0.12 0.05 0.36 1.81
0.68 0.63 0.18 0.38 –
1.14 1.67 0.37 0.39 6.87
Table 6 Initial stiffness of tested beams. Beam
Initial stiffness, kN/ mm
% increase compared to bare HSS beam
Bare HSS beam NC-100 E50bt-f E75bt-f E25t-f E30t-p
2.59 4.58 4.38 4.39 3.98 3.91
– 76.8 69.11 69.5 53.67 50.96
Fig. 9. Mesh sensitivity analysis for the E75bt-f beam.
the test, the whole ECC layer was not crushed. 3.2. Cracking patterns from DIC analysis Since the DIC technique was employed to detect the first cracking load, to measure the total crack width and to monitor the propagation of cracks, the accuracy of the DIC results were first validated by comparing the DIC mid-span deflections with the measurements by LDVT (Fig. 5). The validated DIC model was then utilized to study the cracking behaviours of the beams. For the NC100-f beam, tiny cracks were first seen at the loading level of 25.8 kN in the constant moment region. At the final stage of test, ten major cracks were observed. Their widths were ranging from 0.4 mm to 1.7 mm and their depths (measured from the bottom of the beam) were approximately 100 mm as shown in Fig. 6(a). For the E50bt-f beam, fine cracks were observed when the loading was 28.8 kN. Four major cracks with widths from 2 mm to 6.5 mm and depths up to 86 mm were found after the specimen failed (Fig. 6(b)). Apart from the major cracks, many micro cracks were observed within the constant moment region. For the E75bt-f beam, a similar crack pattern was observed. The widths of the five major cracks were ranging from 1 mm to 5.5 mm with a maximum depth of 95 mm. However, the first micro-crack appeared at a lower loading level of 14.7 kN. The final crack distribution of E75bt-f beam is shown in Fig. 6(c). The final crack pattern for the E25t-f beam was different from those
Fig. 8. Slip vs mid-span deflections of the specimens (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
load was 137.1 kN and an obvious beam end slip was detected. As a result, very little composite action was achieved during the test and the ECC layer was bent separately. As shown in Fig. 4(f), even at the end of 8
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Fig. 10. Mesh for the E75bt-f beam: (a) section view, (b) 3D view of the FE model.
visible near the end of the test. The final crack distribution of the E30t-p beam is shown in Fig. 6(e). 3.3. Strain analysis The strain variations (measured at mid-section of beams) along the depth of all beams are plotted at different loading stages in Fig. 7. Two solid and one dotted vertical lines are also added in Fig. 7, to represent the yield strain of HSS (0.38%, corresponding to the actual yield strength of the HSS) and the peak compressive strain of ECC (0.5%), respectively. It should be noted that, in Fig. 7, failure load refers the loading when ECC/NC started to crush while ultimate load refers to the loading at the final stage of test. After the failure stage, as the strain gauges (SG-5 and SG-6) were damaged, strain data at ECC/NC surfaces are not included in strain distribution. Fig. 7 indicates that both flanges of the bare HSS beam were just yielded before the beam was failed by LTB. By contrast, the tension flanges of all encased beams were started to yield when the loading was approximately 150 kN and attained at least 0.89% strain at the end of the tests (125 mm mid-span deflection). For the compression flanges of encased beams, all the top flanges of the HSS sections were yielded at the end of the tests. Note that in this plot, strain data measured at the top flanges of HSS sections (SG-1) are not included as after the crushing of the ECC/NC layer the strain gauges were damaged. For E50bt-f, E75bt-f and E30t-p beams, strain readings from SG-5 and SG-6 confirmed that the ECC crushed at a compressive strain of at least 0.42% which is higher than the yield strain (0.38%) of
Fig. 11. Tensile stress-strain curve of HSS.
of the E50bt-f and E75bt-f beams. The first fine crack was detected when the applied load was 15 kN. Since LWC has a much lower crack controlling ability when compared with ECC, thirteen smaller major cracks with widths ranging from 0.25 mm to 0.5 mm were found at end of the test. The depths of these cracks were up to 83 mm as shown in Fig. 6(d). For the E30t-p specimen, as a large amount of slip occurred when the compression flange was yielded, only a few small cracks were
Fig. 12. Material models for ECC (a) under compression and (b) under tension. 9
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Table 7 Compression parameters for ECC. Material/Specimen
σc0 (MPa)
εc0 (%)
σcl (MPa)
εcl (%)
σc (MPa)
εmax (%)
ECC (tension side) E50bt-f and E75bt-f ECC (compression side) E50bt-f and E75bt-f ECC (Compression side) E25t-f and E30t-p
70.4
0.48
35.4
0.73
24.22
3.6
59.51
0.45
16.94
0.73
10.05
3.6
67.26
0.54
21.85
0.86
15.46
2.5
Material
σt0 (MPa)
εt0 (%)
σtp (MPa)
εtp (%)
εtu (%)
ECC (Tension side): E50bt-f and E75bt-f ECC (Compression side): E50bt-f and E75bt-f ECC (Compression side): E25t-f and E30t-p
4.79 5.05 5.19
0.022 0.024 0.021
5.17 5.82 5.65
0.77 0.72 0.53
0.86 0.81 0.75
Table 8 Tension parameters for ECC.
Table 11 Validation of numerical model. Specimen
Bare HSS NC100-f E50bt-f E75bt-f E25t-f E30t-p Mean
Load carrying capacity Tests, Pu-test
FE model, Pu-FE
Pu-FE/Pu-test
139.1 175.79 192.15 191.79 214.61 204.91
148.857 181.089 194.103 201.053 205.314 191.488
1.07 1.03 1.01 1.05 0.96 0.93 1.01
Fig. 13. Compression material curve for LWC.
the HSS used. For the E25t-f beam, an even higher strain of 0.56% was recorded as the beam has the highest resistance and mid-span deflection among the encased beams.
Table 9 Compression parameters for NC and LWC. Material/Specimen
0.4f 'c (MPa)
εy (%)
f 'c (MPa)
NC:NC100-f LWC: E50bt-f and E75bt-f LWC: E25t-f
22 17.74
0.1 0.061
55 44.35
18.11
0.14
45.27
σcu (MPa)
εcu (%)
0.28 0.21
5.25 18.95
0.6 0.4
0.36
12.76
0.8
' c (%)
3.4. Bond-slip behaviours In this study, the beam-end method was used to examine the interfacial slip between the surfaces of ECC/LWC/NC and the HSS section flanges. The beam-end slips were calculated from the data gathered by LVDT-4 to LVDT-7. The slips-deflection and the load–deflection curves for all encased beams tested are plotted in Fig. 8. The final beam-end slip values obtained are summarized in Table 5. Fig. 8 and Table 5 show that the beam-end slip values were negligible (from 0.18 mm to 0.68 mm) for the NC100-f, E50bt-f, E75bt-f and E25t-f beams when the compressive surface of the NC/ECC layer was crushed (the vertical lines in the figures). However, as shown in Fig. 8(e) for the E30t-p beam, a large beam-end slip (1.81 mm) was observed when the HSS section was yielded. The slip was increased to 6.9 mm at the end of the test. Detailed numerical modelling described in the next section shown that such a large beam-end slip was due to the rapid development of shear stress within the shear spans that reached the bond strength of the ECCHSS interfaces.
Fig. 14. Steel-ECC interfaces bond-slip model.
3.5. Initial stiffness The initial stiffnesses of all beams are obtained by calculating the slopes of the linear parts of the load-deflection curves and listed in Table 6. It is observed that the NC-100 beam showed the highest stiffness. For the fully encased HSS ECC beams (E50bt-f, E75bt-f and E25t-f), as the elastic moduli of ECC and LWC are approximately twothird of normal concrete (Table 3), their stiffness were lower than the NC100-f beam but still at least 53% higher than the bare steel beam. On the other hand, the initial stiffness of the partially encased E30t-p beam
Table 10 Bond-slip parameters. Properties
Values
Bond strength, τm Slip at bond strength, Sm ultimate slip, SL
0.5 MPa 0.5 mm 15 mm
10
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Fig. 15. Comparison of test and modelling results for specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
is the lowest among all encased beams.
respectively. Loadings were applied by prescribing downward displacements at two patches of elements under the loading points. The same displacement steps used in the test were applied until a displacement of 125 mm was reached. The default *RAMP function available in ABAQUS [25] was employed to apply the prescribed displacement. During modelling, the beams were discretised by the eight-node hexahedral solid elements (C3D8R) which had been successfully employed by other researchers [18,19,26] in similar analysis. In order to determine an optimal element size, an element size sensitivity analysis was performed by using six different meshes. Fig. 9 shows the predicted load carrying capacity of the E75bt-f beam with different element sizes. From Fig. 9, it can be concluded that a mesh with 40 mm elements could predict the strength of the beam precisely with the smallest
4. Finite element modelling In order to predict the enhancement effects of ECC/LWC/NC encasement, a three-dimensional (3D) FE model which considers the geometrical and material non-linearities as well as the bond-slip behaviours between the ECC and HSS interfaces was developed. 4.1. Loading and boundary conditions and mesh convergence study The support conditions used in the tests were reproduced in the model by applying roller and pin support conditions to two patches of elements corresponding to the actual right end and left end supports, 11
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Fig. 16. Validation of final failure and cracking for the specimens.
For ECC, the compressive and tensile stress-strain relationships used by Kabir et al. [19] and Meng [20] (Fig. 12) were proved to be reliable and were adopted in this study. The average values of the parameters used are given in Tables 7 and 8. Furthermore, the damage model by Lubliner et al. [28], which had been successfully used before [19] was employed again to define the failure of the ECC under compression and tension. For NC and LWC, the compressive behaviours were described by the constitutive model suggested by Carreira and Chu [29] (Fig. 13). The values of all key parameters, which were determined from the material tests results, are listed in Table 9. For their tensile behaviours, a simple a linear relationship was considered [19] so that the peak tensile strength was assumed as 10% of the peak strength of NC/LWC under compression (f’c in Fig. 13). For the material damage models for NC and LWC, the Concrete Damage Plasticity model was employed, and the same set of parameters used in Table 8 of Ref. [19] were used.
Table 12 Beam-end slip values comparison. Beam
NC100-f E50bt-f E75bt-f E25t-f E30t-p Mean
Beam-end slip (mm) Test
FEA
FE/test
1.14 1.67 0.37 0.40 6.86
1.09 1.81 0.35 0.38 6.9
0.96 1.08 0.95 0.95 1.01 0.99
amount of computational efforts and was adopted in subsequent models. The 3D view of the mesh employed in the modelling of the E75bt-f beam is shown in Fig. 10. 4.2. Material properties used in FE modelling
4.3. Bond-slip behaviour modelling
For the material properties of HSS, the well-accepted bi-linear stress-strain model indicated in Fig. 11 [19,27] was used. Values of the materials parameters used in the model were determined by tension coupon tests and are listed in Table 3.
Since bond-slip was observed in the test of the E30t-p beam, an accurate bond-slip model between the HSS-ECC interfaces is essential to predict the encased beams’ behaviours. In ABAQUS, bond-slip between 12
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HSS [19]. Finally, as no bond-slip was observed between the LWC-ECC interfaces, all interfaces between LWC and ECC were tied together in the models. 4.4. Validation of model against experimental results 4.4.1. Load-deflection responses and failure mode predictions All six beams tested were modelled by the proposed FE model. Comparisons between the FE modelling failure load, Pu-FE, and the test failure load, Pu-test, are presented in Table 11. The comparisons of load vs mid-span displacement curves are presented in Fig. 15. Fig. 15 shows that all key stages of the bending tests including initial linear responses, yielding of HSS sections, flexural resistance (i.e. failure load) and crushing of the top NC/ECC layer as well as the post-failure behaviours, were accurately predicted by the FE model. Regarding the failure mode, crushing of the compressive NC/ECC layer and bottom cracking distribution, Fig. 16 indicates that the FE model again produced good predictions which well agreed with the test results. 4.4.2. Bond-slip predictions and analysis The reliability of the model is further verified by comparing the measured bond-slips at the two support ends with the FE predictions. The maximum measured and FE predicted bond-slips for different beams are summarized in Table 12. From Table 12, it is clear that the proposed FE model predicted the final beam-end slips accurately. Furthermore, in order to confirm that the surface-based cohesive model could predict both the initiation and development of slip, the interfacial slip histories for the E75bt-f beam (which represents those fully encased beams) and for the E30t-p beam (which represents the partially encased beam) were compared with the FE predictions in Fig. 17. Fig. 17 indicates that the surface-based cohesive behaviours model predicted the slip histories for both cases accurately. The validated FE model was employed to examine the advancement of interfacial bond stress and slip along the length of the ECC-HSS encased beams. The results obtained for the four ECC-HSS encased beams at five different loading levels, namely at 50 kN, 100 kN, 150 kN and at the peak (failure) load the beams, are presented in Figs. 18–21. From Figs. 18–21, it is observed that the bond stress between the HSS sections and ECC surfaces increased as the applied load increased. As expected, the bond stress curves and the contour plots at the peak load indicate that high shear bond stress was generated along the shear spans. In addition, when comparing Fig. 21 with Figs. 18–20, one can see that as the applied load increased, bond stress was developed much more rapidly for the E30t-p beam than all other fully encased beams. For the E30t-p beam, the bond stress reached the maximum slip strength quickly throughout the shear spans when the applied load is 100 kN which is well below the HSS section’s full plastic moment resistance (170 kN).
Fig. 17. Comparison of beam-end slip vs mid-span deflection between FE and test results for the (a) E75bt-f and (b) E30t-p beams.
the HSS and ECC interfaces can be modelled by prescribing their surface-based cohesive behaviours [18,19,25]. In this model, a linear elastic traction-separation behaviour is assumed prior to the maximum traction (bond strength) τm at a slip of Sm. The bond is failed once the ultimate slip, SL is reached (Fig. 14). Such approach is particularly attractive and accurate for negligibly small thickness of interfaces [30] and when reliable bond-slip data are available from carefully designed push tests [21]. In this study, the data obtained by Rana et al. [21] were used to define the surface-based cohesive behaviours between the HSS and ECC interfaces (Table 10). As no slip was detected between the HSS and LWC interfaces, a simpler Coulomb friction model with a frictional coefficient of 0.25 is found to be accurate enough to model the interfaces between LWC and
Fig. 18. Shear (Bond) stress distribution for the E50bt-f specimen. 13
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Fig. 19. Shear (Bond) stress distribution for the E75bt-f specimen.
Fig. 20. Shear (Bond) stress distribution for the E25t-f specimen.
Fig. 21. Shear (Bond) stress distribution for the E30t-p specimen.
5. Conclusions
encased beams tested, only a very small amount of slip between the ECC encasement and HSS flanges occurred even after the whole HSS section was yielded. As a result, the flexural resistance of the fully encased beams was significantly higher than the HSS section’s full plastic moment resistance and the flexural strength of a similar normal concrete encased beam in which the top concrete layer was crushed before the HSS section was fully yielded. However, for the partially encased ECCHSS beam, noticeable amount of interfacial slip was measured when the HSS flange was yielded. Nevertheless, the ECC encasement was still able to prevent the occurrence of LTB. As it was found that the ECC encasements were crushed after the HSS sections were fully yielded, one of the potential extensions of the proposed enhancement approach is to apply similar encasement to Class 4 slender HSS sections [31,32] by preventing local buckling of the slender flange and web so that the full plastic moment resistance of the slender section could be achieved.
In this study, the effectiveness of using Engineered Cementitious Composites (ECC) encasement to enhance the flexural strength of welded High Strength Steel (HSS) beam was investigated experimentally and numerically. A 3D finite element model was proposed to simulate the flexural responses and interfacial slip of the encased beams. The results obtained confirmed the predictions from previous study [19] that ECC encasement can significantly enhanced the flexural resistance of a bare HSS beam by preventing the onset of lateral torsional buckling (LTB). More importantly, the use of ECC prevented the brittle crushing of the beam’s top layer at peak load. The strain gauge measurements and numerical modelling results also confirmed that so long as the compressive strain at the peak strength of the ECC is greater than the yield strain of the HSS, failure of the ECC layer at the compression side of the beam would occur after the HSS flange is fully yielded. The experimental and numerical results further shown that for those fully 14
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Declaration of Competing Interest
2015;141(7): 04014166. [11] Kim CS, Park HG, Chung KS, Choi IR. Eccentric axial load capacity of high-strength steel–concrete composite columns of various sectional shapes. J Struct Eng 2014;140(4):04013091. [12] Uy B. Stability and ductility of high performance steel sections with concrete infill. J Constr Steel Res 2008;64:748–54. [13] Uy B, Sloane RJ. Behaviour of composite tee beams constructed with high strength steel. J Constr Steel Res 1998;46:203–4. [14] Standards Australia. AS 2327.1-1996-Australian Standard, Composite Structures Part 1: Simply Supported Beams. Standards Australia; 1996. [15] Eurocode 4. ENV 1994-1-1 Design of composite steel and concrete structures, Part 1.1, General Rules and Rules for Buildings, British Standards Institution”. Eurocode 4; 1994. [16] Zhao H, Yuan Y. Experimental studies on composite beams with high-strength steel and concrete. Steel Compos Struct 2010;10:373–83. [17] Ban H, Bradford MA. Flexural strength of high-strength steel-concrete composite beams with varying steel grades. Compos Constr Steel Concr VII 2016:120–30. [18] Rana MM, Lee CK, Al-Deen S, Zhang YX. Flexural behaviour of steel composite beams encased by engineered cementitious composites. J Constr Steel Res 2018;143:279–90. [19] Kabir MI, Lee CK, Rana MM, Zhang YX. Flexural and bond-slip behaviours of engineered cementitious composites encased steel composite beams. J Constr Steel Res 2019;57:229–44. [20] Meng D, Lee CK, Zhang YX. Flexural and shear behaviours of plain and reinforced polyvinyl alcohol-engineered cementitious composite beams. Eng Struct 2017;151:261–72. [21] Rana MM, Lee CK, Al-Deen S. A study on the bond stress-slip behaviour between engineered cementitious composites and structural steel sections. In: 8th European Conference on Steel and Composite Structures, Copenhagen, Denmark; 2017. [22] AS 1012. Methods of testing concrete. SAI Global, Australia; 2014. [23] ACI Committee. Building code requirements for structural concrete (ACI 318–14) and commentary. American Concrete Institute; 2014. [24] ASTM C469. Standard test method for static modulus of elasticity and Poisson’s ratio of concrete in compression. West Conshohocken, Pennsylvania; 2002. [25] ABAQUS, ABAQUS Standard User’s Manual, Version 6.14, Rhode Island, USA: Dassault Systemes Simulia Corp; 2014. [26] Zhong T, Wang Z-B, Yu Q. Finite element modelling of concrete-filled steel stub columns under axial compression. J Constr Steel Res 2013;89:121–31. [27] Ban H, Bradford M. Flexural behaviour of composite beams with high strength steel. Eng Struct 2013;56:1130–41. [28] Lubliner J, Olive S, Onate E. A plastic-damage model for concrete. Int J Solid Struct 1989;25(3):299–326. [29] Carreira D, Chu K. Stress-strain relationship for plain concrete in compression. ACI J 1985, Title no.82-72:797-804. [30] Tekle BH, Khennane A, Kayali O. Bond behaviour of GFRP reinforcement in alkali activated cement concrete. Constr Build Mater 2017;154:972–82. [31] Lee CK, Chiew SP. An efficient modified flange only method for plate girder bending resistance calculation. J Constr Steel Res 2013;89:98–106. [32] Lee CK, Chiew SP. A review on class 4 slender section properties calculation for hot finished sections according to EC3. Adv Steel Constr 2019;15:259–66.
The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. Acknowledgements The authors greatly acknowledge BISALLOY, Australia for providing the high strength steel plates used in this study. However, the findings and conclusions drawn in this publication are solely the view of the authors only and not necessarily the view of BISALLOY, Australia. Appendix A. Supplementary material Supplementary data to this article can be found online at https:// doi.org/10.1016/j.engstruct.2020.110288. References [1] Ban H, Shi G, Shi Y, Wang Y. Overall buckling behavior of 460 N/mm2 high strength steel columns: experimental investigation and design method. J Constr Steel Res 2012;74:140–50. [2] Ban H, Shi G, Shi Y, Bradford MA. Experimental investigation of the overall buckling behaviour of 960 N/mm2 high strength steel columns. J Constr Steel Res 2013;88:256–66. [3] Wang YB, Li GQ, Chen SW, Sun FF. Experimental and numerical study on the behavior of axially compressed high strength steel box-columns. Eng Struct 2014;58:79–91. [4] Lee CH, Han KH, Uang CM, Kim DK, Park CH, Kim JH. Flexural strength, rotation capacity of I-shaped beams fabricated from 800 MPa steel. J Struct Eng 2012;139:1043–58. [5] Bradford MA, Lui X. Flexural-torsional buckling of high-strength steel beams. J Constr Steel Res 2016;124:122–31. [6] Ma JL, Chan TM, Young B. Experimental investigation of cold-formed high strength steel tubular beams. Eng Struct 2016;126:200–9. [7] Puthli R, Fleischer O. Investigations on bolted connections for high strength steel members. J Constr Steel Res 2001;57:313–26. [8] Coelho AMG, Bijlaard FS. Experimental behaviour of high strength steel end-plate connections. J Constr Steel Res 2007;63:1228–40. [9] Mashiri FR, Uy B, Tao Z, Wang ZB. Concrete-filled VHS-to-steel fabricated section stub columns subjected to axial compression. J Constr Steel Res 2014;95:141–61. [10] Hsiao PC, Hayashi KK, Nishi R, Lin XC, Nakashima M. Investigation of concretefilled double-skin steel tubular columns with ultrahigh-strength steel. J Struct Eng
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Strength enhancement of high strength steel beams by engineered cementitious composites encasement
T
Md. Imran Kabira, C.K. Leea, , Mohammad M. Ranaa, Y.X. Zhangb ⁎
a b
School of Engineering and Information Technology, The University of New South Wales, Canberra, ACT 2600, Australia School of Computing, Engineering and Mathematics, Western Sydney University, NSW 2751 Australia
ARTICLE INFO
ABSTRACT
Keywords: High strength steel Engineered cementitious composite encasement Flexural resistance Bond-slip behaviours Finite element modelling
This study proposes a method of using Polyvinyl Alcohol Engineered Cementitious Composites (PVA-ECC) encasement to provide continuous restraints along the compression flange of High Strength Steel (HSS) section so that it will reach its sectional plastic moment resistance under bending without lateral restraint. In order to demonstrate the effectiveness of the proposed method, experimental and numerical investigations were carried out to study the flexural strength of the ECC encased HSS beams (ECC-HSS beams). Six simply supported beams fabricated with identical HSS sections but with different encasement configurations were tested until failure. Flexural resistance and failure modes of the ECC-HSS beams were compared with similar bare HSS and normal concrete (NC) encased HSS beams (NC-HSS beams). It was found that when compared with the bare HSS and NCHSS beams, a significant enhancement in flexural resistance was achieved for the ECC-HSS beams. More importantly, this study confirmed that the compressive ECC layers was crushed after the compression flanges were yielded and therefore successfully prevented the onset of lateral torsional buckling. Besides the flexural responses, the interfacial slip behaviours along the compression flange of the HSS section were also studied. Finally, a finite element (FE) model was developed and validated against the experimental results.
1. Introduction When comparing with Normal Strength Steel (NSS) which normally has a yield strength not more than 460 MPa, High Strength Steel (HSS), which specifically refers to steel with a yield strength equal to or more than 690 MPa, can offer many advantages in construction. Due to their higher yield strength, the use of HSS in structural design may reduce the cross-section dimension and thus the self-weight of the structures significantly. Such merits of HSS could alleviate the fabrication and lifting difficulties during construction. However, one of the disadvantages of using HSS sections is that their thinner wall thicknesses could lead to a higher slenderness ratio which eventually results in instability. In general, with the same full sectional plastic moment resistance, a laterally unrestrained HSS beam is more vulnerable to lateral torsional buckling (LTB) and local buckling than an NSS beam. Such unstable effects may lead to pre-mature failure before the beam’s sectional plastic moment resistance is fully utilized. In recent years, HSS had gained much research interest but most of the researches were focused on the behaviours of bare HSS sections such as columns [1–3], beams [4–6] and connections [7,8]. For steelconcrete composite structures, although some researches on concrete-
⁎
filled HSS tubular columns had been conducted [9–12], researches on HSS-concrete composite beams are relatively uncommon. Most of the researches done on HSS-concrete composite beams were focused on composite beams which are constructed by connecting concrete slab with HSS section using shear connectors. Uy and Sloane [13] studied the flexural behaviour of two composite steel-concrete T-beams and numerical results were compared with experimental results. A satisfactory agreement was observed between them when the slip between concrete and steel section was considered. However, the flexural strengths predicted by the Australian [14] and European [15] design standards were found to be slightly unconservative. Zhao and Yuan [16] tested four composite steel-concrete T-beams subjected to twopoint concentrated loads. It was found that by using HSS to replace NSS, significant improvements were observed regarding the moment capacity, longitudinal split and interfacial slip. Recently, Ban and Bradford [17] conducted an extensive parametric study on the flexural behaviours of composite beams fabricated by using different HSS grades. Results obtained from the finite element (FE) models were compared with those obtained from classical plastic analysis. It was found that while the Eurocode 4 [15] predicted the bending resistance of concreteNSS composite beams accurately, it over-predicted the flexural
Corresponding author. E-mail address: [email protected] (C.K. Lee).
https://doi.org/10.1016/j.engstruct.2020.110288 Received 19 October 2019; Received in revised form 20 December 2019; Accepted 23 January 2020 Available online 30 January 2020 0141-0296/ © 2020 Elsevier Ltd. All rights reserved.
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Table 1 Compositions of ECC, LWC and NC. ECC
Table 2 Section details of encased beams (all dimensions are in mm).
LWC
NC
Ingredient
Mix ratio
Ingredient
Mix ratio
Ingredient
Mix ratio
Cement Fly Ash Sand/ binder Water HRWRA PVA-fibre (vol%)
1.0 1.2 0.36
Cement Sand 8 mm LWA
1.0 1.7 0.92
Cement Sand 7 mm aggregate
1.0 1.81 1.63
0.56 0.01 2.2
14 mm LWA Water
0.92 0.4
10 mm aggregate Water Superplasticiser
1.1 0.42 0.015
Specimen
Overall size B×D
HSS Section d × bf × tw × tf
tECC,
Bare HSS beam NC100-f E50bt-f E75bt-f E25t-f E30t-p
–
150 × 75 × 6 × 8
–
–
–
– 52.5 78.75 52.5 52.5
– 52.5 78.75 – –
210 105 52.5 157.5 –
175 175 175 175 175
× × × × ×
210 210 210 210 180
comp
tECC,
ten
tLWC/NC
Legends: B = width of beam. D = depth of beam. d = depth of HSS section. bf = flange width. tf = flange thickness. tw = web thickness. tECC, comp = compression ECC layer thickness. tECC, ten = tension ECC layer thickness. tLWC/NC = LWC/NC layer thickness.
Note: LWA = Light-weight aggregate.
resistance of HSS-concrete composite beams. It appears that most recent research related to fully encased steel beam is largely limited to concrete encased NSS beams only. An exception is the study by Rana et al. [18] which showed that by replacing Normal Concrete (NC) with Engineered Cementitious Composites (ECC) and Lightweight Concrete (LWC), the flexural performances of a fully encased NSS-NC composite beams could be improved significantly. In addition, the recent work done by Kabir et al. [19] demonstrated that unstable LTB failure was prevented by using ECC to encase the compression flange of the NSS section. Numerical analyses conducted in [19] further predicted that the flexural resistance of a fully encased beam could be enhanced by substituting the NSS section with HSS section. However, no experimental study was conducted to confirm such predictions that the encasement could lead to full composite actions and prevent local buckling. Hence, in this study, a similar approach was applied to enhance the flexural resistance of HSS beams by using ECC and LWC to provide continuous lateral restraint to the compressive flange of the HSS section. The results obtained were then compared with the performances of a bare HSS beam and a control NCHSS encased beam. Furthermore, a FE modelling procedure was employed to model the bending and interfacial slip behaviours of the beams in details. It should be noted that while ECC is well-known for its superior tensile ductility and tensile strain hardening behaviour, it also has superior compression ductility (compressive strain at peak strength can easily exceed 0.5%) when comparing with normal concrete (0.23–0.28%). However, ECC’s higher tensile ductility and tensile strength are still much smaller than any grade of steel. Hence, when ECC is employed to encase the HSS section, such superior tensile performance is not very important as nearly all the tensile force at the tension side of the beam will be resisted by the steel section anyway. On the contrary, ECC should be used in the compression zone of the beam to prevent immature local buckling of steel section as its compressive strain at peak strength (0.5%) is larger than the yield strength of HSS (0.38%). In this case, it is the superior compression ductility that makes ECC a suitable material that could be used in conjunction with HSS.
2. Experimental investigation 2.1. HSS, ECC and LWC used The 3.4 m long, built-up HSS I-sections used were fabricated by using BISALLOY structural 80 steel plates which has a nominal yield strength (0.2% proof stress) of at least 690 MPa. Single pass, 6 mm continuous fillet welding were employed to join the two flange plates with the web plate. In order to minimize thermal distortion generated during welding, welding along the top and bottom flange joints was completed in an alternative manner after every 0.5 m of welding. The polyvinyl alcohol ECC (PVA-ECC) proposed by Meng et al. [20], which use of local ingredients to reduce the materials cost, was adopted in this study. It has sufficient ductility in both tension and compression to work with HSS and was found to has a slightly higher bond strength with steel section [21] than the conventional concrete. For the LWC, two different sizes (8 mm passing and 14 mm passing) of light-weight aggregate (LWA) with same mix ratio were used. For the NC, it was prepared by using two different sizes of aggregates (7 mm passing and 10 mm passing) of equal mix ratio. Table 1 lists the mix ratios for ECC, LWC and NC. 2.2. Specimen configurations Six beams were prepared and tested for studying their flexural and bond-slip behaviours. These beams include one bare HSS beam, one NC encased HSS (NC-HSS) beam and four ECC-LWC encased HSS (ECCHSS) beams. For the NC-HSS beam, the whole HSS section was encased by NC and is labelled as the NC100-f beam. The ECC-HSS beams were prepared by changing the portion of encasement of ECC/LWC and positions of the ECC layer(s). The first and second ECC-HSS beams were prepared by encasing the two flanges of the HSS sections by ECC in such
Fig. 1. Cross-sectional profiles of the tested beams. 2
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Table 3 Material properties of ECC, LWC, NC and HSS. Material/Specimen
Properties
ECC (tension side) E50bt-f, E75bt-f ECC (Compression side) E50bt-f, E75bt-f ECC (Compression side) E25t-f, E30t-p beam LWC: E50bt-f, E75bt-f LWC: E25t-f NC: NC100-f HSS section (flanges) HSS section (web)
Compressive strength (MPa)
Tensile strength (MPa)
Modulus of elasticity (MPa)
HSS yield strength, fy (MPa)
HSS ultimate strength, fu (MPa)
69.41 61.53 67.26
4.79 5.05 5.19
21,129 22,855 19,220
– – –
– – –
44.35 45.27 58.65 – –
– – – – –
26,127 25,479 32,863 197,655 198,220
– – – 748 (at εy = 0.38%) 759 (at εy = 0.38%)
– – – 842 (at εu = 10%) 856 (at εu = 10%)
Fig 2. Details of (a) instrumentation, (b), (c) positions of LVDTs for end slip measurement (d) locations of strain gauges. 3
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Fig. 3. Load-deflection curves of specimens.
a way that the ECC thicknesses are equal to 50% and 75% of beams’ depth, respectively. For the remaining parts of the beams, the HSS sections are encased by LWC to reduce the weight of beams. Therefore, these two beams are named as the E50bt-f and the E75bt-f beams, respectively. For the third ECC-HSS beam, only the compression flange is covered by a layer of ECC corresponding to 25% of the beams’ depth and remaining part of the beam is covered by LWC so that it is still fully encased. Therefore, this beam is referred as the E25t-f beam. The last ECC-HSS beam is a partially encased beam such that only the compression flange is covered by an ECC layer corresponding to 30% of the beams’ depth and is labelled as the E30t-p beam. It is the lightest ECCHSS beams tested to explore the slip characteristics between the HSS section and the ECC layer. The encasement configurations and detailed dimensions of the beams tested are shown Fig. 1 and Table 2. Note that while the NC100-f and E30t-p beams only required one casting step, the E25t-f beam and the two “bt-f” beams required two and three cast steps, respectively. Detailed information of the casting procedure of the beams is described in [19]. In order to determine the mechanical properties of the ECC, LWC and NC, three cylinders with standard dimensions of 100 mm ∅ × 200 mm H were prepared from each mix of ECC, LWC and NC [22,23]. Their Young’s Modulus and compressive strengths at 28 days were then determined in accordance with ASTM C469 [24]. Furthermore, three dog-bone shaped samples of standard dimensions [20] were tested under uniaxial tension to obtain the ECC’s tensile properties. In order to obtain the tensile properties for the HSS sections, direct tensile tests were conducted on three standard HSS coupons. The mean values of the mechanical properties obtained are given in Table 3.
Fig. 4. Final failures of the specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
employed to record the mid-span (LVDT-1) and the loading points (LVDT-2 and LVDT-3) deflections. Since HSS sections were used to form the encased beams, it is expected a high shear stress would be developed between the ECC/NC-steel interfaces and slipping could occur before the beams are failed by bending. Therefore, four additional LVDTs (LVDT-4 to 7) were attached at the two ends of the encased beams to measure the interfacial slip between the ECC/LWC/NC and
2.3. Bending test procedures The specimens were loaded using four-point bending with a clear span of 3000 mm and two load points were placed at 1000 mm from the supports. The supports and loading configuration are shown in Fig. 2(a). Three linear variable displacement transducers (LVDTs) were Table 4 Bending test results. Beam
Failure load (kN)
Mid-span deflection at failure load (mm)
Enhancement of failure load (%) as compared to bare HSS beam
Weight of thebeam (kg)
Weight reduction (kg) as compared with NC100-f
Bare HSS beam NC100-f E50bt-f E75bt-f E25t-f E30t-p
139.1 175.79 192.15 191.79 214.61 204.91
54.47 43.47 56.02 54.73 76.77 148.31
– 26.37% 38.14% 37.88% 54.28% 47.31%
53.5 337 275 281 270 111
– – 62 (18.3%) 56 (16.5%) 67 (20%) 226 (67.1%)
4
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Fig. 5. Comparison of mid-span deflection measured by LVDT and DIC: (a) NC100-f, b) E50bt-f, (c) E75bt-f, (d) E25t-f, (e) E30t-p.
the HSS flanges. For LVDT-4 to 7, the base of the LVDT was attached to one end of a PVC tube while the other end of the tube was attached to the ECC/LWC/NC surface. The tip of the LVDT was placed at the flange as shown in Fig. 2(b) and (c). As a result, the PVC tubes and the LVDTs will rotate with the beam end and the measured relative horizontal displacements between the ECC/LWC/NC surfaces and the top flange will not be affected by the beam end rotation. The HSS sections were also instrumented with four strain gauges (SG-1, SG-2, SG-3 and SG-4) to measure the strain history during the tests. For the encased beams,
two strain gauges (SG-5 and SG-6) were installed at the top surfaces of beams to monitor the compressive strain of ECC/NC. The instrumentation scheme of the beams is shown in Fig. 2(a) and (d). In addition, three high-resolution digital cameras were placed in front of the beams to capture the development of crack distribution along the entire beam by using the Digital Image Correlation (DIC) technique. In all tests, displacement controlled static loading was applied by a hydraulic jack through a spreader beam. The applied force was measured by load cells placed under the supports (Fig. 2a). Loads were 5
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3. Bending test results 3.1. Flexural strengths of the specimens Fig. 3 summarizes the bending test results in the form of load-deflection curves. The full plastic moment resistance of the HSS section (170 kN), which is calculated based on the actual yield strength of the HSS coupons (Table 3), is drawn as a horizontal line in Fig. 3. The failure loads carried by the beams and the corresponding mid-span deflections are listed in Table 4. In Table 4, the failure load for the bare steel beam is the load when LTB occurred. For the fully encased beams (i.e. NC100-f, E50bt-f, E75bt-f and E25t-f), the failure load is the load when the top NC or ECC layer was crushed. For the partial encased beam E30t-p, the failure load is the highest load sustained at the end of the test. It should be mentioned that for all beams, no failure nor any crack was detected in any part of the welded joints between the HSS sections’ flanges and web at the end of the tests. 3.1.1. The bare HSS beam The bare HSS beam was failed by LTB at a mid-span deflection of 54.47 mm under an applied load of 139.1 kN. Fig. 4(a) shows that the pure bending region of the HSS section was buckled laterally at a loading about 18% lower than the sections’ full plastic moment resistance. 3.1.2. The NC100-f beam The failure load of this beam was 175.79 kN which was only 3.4% higher than the plastic moment resistance of the HSS section despite that the HSS section was fully encased by NC with a compressive strength of 58.65 MPa. Hence, the NC encasement virtual did not make contribution to the beams’ flexural resistance. This was because the top NC cover was crushed before the HSS section was fully yielded. Nonetheless, as the HSS section was laterally restrained by the NC encasement, no LTB occurred and the flexural resistance was increased by 26.67% when compared with the bare HSS beam. The failure mode of this beam was a brittle crushing of the compressive concrete layer which caused a sharp 13.5% drop in loading. At the end of the test, extensive cracking observed within the constant moment region as shown in Fig. 4(b). 3.1.3. The E50bt-f and E75bt-f beams The failure load of the E50bt-f beam was 192.15 kN at a defection of 56.02 mm. The failure load is 13% and 38.1% higher than the full plastic moment resistance of the HSS section and bare HSS beam, respectively. Since ECC has a higher compressive ductility (0.45% strain) than NC (0.28% strain), the mid-section of encased HSS section attained its plastic resistance before the top ECC layer was crushed. Therefore, the E50bt-f beam carried 9.3% higher load than the NC100-f beam. After the yielding of the flanges of HSS section, the outermost fibre of the ECC layer in the constant moment region was crushed which resulted a drop of 9.34% in loading. The failure mode of the E50bt-f beam was the yielding of HSS section first and then the crushing of compressive ECC layer. When compared with the NC100-f beam, the crushing of the compressive ECC layer was found to be much more ductile and gentler. The shape of the E50bt-f beam at the end of the test is shown in Fig. 4(c). As the only difference between the E50bt-f beam and the E75bt-f beam was the thickness of the ECC layers, the failure load, load-deflection curve and failure mode of the E75bt-f beam are all similar to the E50bt-f beam. Furthermore, Fig. 4(d) and Table 4 show that the deflected shape, flexural resistance and the weight of the E75bt-f beam are practically identical with the E50bt-f beam. Hence, this test results strongly suggested that in practice, there is no need to increase the thickness of the ECC layer beyond half of the beam’s depth.
Fig. 6. Crack analysis from DIC for the specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
applied at a rate of 0.5 mm/min until LTB occurred (for the bare steel beam) or until the compressive ECC/NC layer within the constant moment region of beams was crushed. The loading rate was increased to 1 mm/min after the compressive NC/ECC layer was failed and the tests were stopped when a 125 mm mid-span deflection was reached. Measurements from the load cells, LVDTs and strain gauges were automatically recorded by data loggers at an interval of 10 s. 6
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Fig. 7. Strain (measured at mid-section of beam) distribution along the depth of the specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30tp (tensile strain is positive).
3.1.4. The E25t-f beam Fig. 3 shows that the failure load of the E25t-f is the highest among all tested beams. This can be explained by the fact that the top flange of HSS section of E25t-f beam was encased by ECC with 10% higher compressive strength (67.26 MPa) than that of the E50bt-f/E75bt-f beams (61.53 MPa). In addition, the strain at peak compressive strength was 20% higher for E25t-f beam when compared with E50bt-f/E75bt-f beam. Therefore, its failure load is not only 26.2% higher than the HSS section’s full plastic moment but also is 54.1%, 22.1%, 11.7% and 11.9% higher than the bare HSS, the NC100-f, the E50bt-f and the
E75bt-f beams, respectively. Moreover, Fig. 3 shows that the failure load was achieved at a deflection of 76.8 mm which is greater than all other fully encased beams. Fig. 4(e) shows the deflected shape and crack patterns of this beam at failure. 3.1.5. The E30t-p beam The E30t-p beam is the lightest encased beam and has a weight equal to 32% of the NC100-f beam. However, its failure load is 47.3% and 16.6% higher than the bare HSS and NC100-f beams, respectively. The tension flange of the E30t-p beam was yielded when the applied 7
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Table 5 Beam-end slips of encased beams. Beam
NC100-f E50bt-f E75bt-f E25t-f E30t-p
Beam-end slips at different stages of testing At yielding of tension flanges of HSS section (mm)
At crushing of NC/ ECC (mm)
At final stage (mm)
0.37 0.12 0.05 0.36 1.81
0.68 0.63 0.18 0.38 –
1.14 1.67 0.37 0.39 6.87
Table 6 Initial stiffness of tested beams. Beam
Initial stiffness, kN/ mm
% increase compared to bare HSS beam
Bare HSS beam NC-100 E50bt-f E75bt-f E25t-f E30t-p
2.59 4.58 4.38 4.39 3.98 3.91
– 76.8 69.11 69.5 53.67 50.96
Fig. 9. Mesh sensitivity analysis for the E75bt-f beam.
the test, the whole ECC layer was not crushed. 3.2. Cracking patterns from DIC analysis Since the DIC technique was employed to detect the first cracking load, to measure the total crack width and to monitor the propagation of cracks, the accuracy of the DIC results were first validated by comparing the DIC mid-span deflections with the measurements by LDVT (Fig. 5). The validated DIC model was then utilized to study the cracking behaviours of the beams. For the NC100-f beam, tiny cracks were first seen at the loading level of 25.8 kN in the constant moment region. At the final stage of test, ten major cracks were observed. Their widths were ranging from 0.4 mm to 1.7 mm and their depths (measured from the bottom of the beam) were approximately 100 mm as shown in Fig. 6(a). For the E50bt-f beam, fine cracks were observed when the loading was 28.8 kN. Four major cracks with widths from 2 mm to 6.5 mm and depths up to 86 mm were found after the specimen failed (Fig. 6(b)). Apart from the major cracks, many micro cracks were observed within the constant moment region. For the E75bt-f beam, a similar crack pattern was observed. The widths of the five major cracks were ranging from 1 mm to 5.5 mm with a maximum depth of 95 mm. However, the first micro-crack appeared at a lower loading level of 14.7 kN. The final crack distribution of E75bt-f beam is shown in Fig. 6(c). The final crack pattern for the E25t-f beam was different from those
Fig. 8. Slip vs mid-span deflections of the specimens (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
load was 137.1 kN and an obvious beam end slip was detected. As a result, very little composite action was achieved during the test and the ECC layer was bent separately. As shown in Fig. 4(f), even at the end of 8
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Fig. 10. Mesh for the E75bt-f beam: (a) section view, (b) 3D view of the FE model.
visible near the end of the test. The final crack distribution of the E30t-p beam is shown in Fig. 6(e). 3.3. Strain analysis The strain variations (measured at mid-section of beams) along the depth of all beams are plotted at different loading stages in Fig. 7. Two solid and one dotted vertical lines are also added in Fig. 7, to represent the yield strain of HSS (0.38%, corresponding to the actual yield strength of the HSS) and the peak compressive strain of ECC (0.5%), respectively. It should be noted that, in Fig. 7, failure load refers the loading when ECC/NC started to crush while ultimate load refers to the loading at the final stage of test. After the failure stage, as the strain gauges (SG-5 and SG-6) were damaged, strain data at ECC/NC surfaces are not included in strain distribution. Fig. 7 indicates that both flanges of the bare HSS beam were just yielded before the beam was failed by LTB. By contrast, the tension flanges of all encased beams were started to yield when the loading was approximately 150 kN and attained at least 0.89% strain at the end of the tests (125 mm mid-span deflection). For the compression flanges of encased beams, all the top flanges of the HSS sections were yielded at the end of the tests. Note that in this plot, strain data measured at the top flanges of HSS sections (SG-1) are not included as after the crushing of the ECC/NC layer the strain gauges were damaged. For E50bt-f, E75bt-f and E30t-p beams, strain readings from SG-5 and SG-6 confirmed that the ECC crushed at a compressive strain of at least 0.42% which is higher than the yield strain (0.38%) of
Fig. 11. Tensile stress-strain curve of HSS.
of the E50bt-f and E75bt-f beams. The first fine crack was detected when the applied load was 15 kN. Since LWC has a much lower crack controlling ability when compared with ECC, thirteen smaller major cracks with widths ranging from 0.25 mm to 0.5 mm were found at end of the test. The depths of these cracks were up to 83 mm as shown in Fig. 6(d). For the E30t-p specimen, as a large amount of slip occurred when the compression flange was yielded, only a few small cracks were
Fig. 12. Material models for ECC (a) under compression and (b) under tension. 9
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Table 7 Compression parameters for ECC. Material/Specimen
σc0 (MPa)
εc0 (%)
σcl (MPa)
εcl (%)
σc (MPa)
εmax (%)
ECC (tension side) E50bt-f and E75bt-f ECC (compression side) E50bt-f and E75bt-f ECC (Compression side) E25t-f and E30t-p
70.4
0.48
35.4
0.73
24.22
3.6
59.51
0.45
16.94
0.73
10.05
3.6
67.26
0.54
21.85
0.86
15.46
2.5
Material
σt0 (MPa)
εt0 (%)
σtp (MPa)
εtp (%)
εtu (%)
ECC (Tension side): E50bt-f and E75bt-f ECC (Compression side): E50bt-f and E75bt-f ECC (Compression side): E25t-f and E30t-p
4.79 5.05 5.19
0.022 0.024 0.021
5.17 5.82 5.65
0.77 0.72 0.53
0.86 0.81 0.75
Table 8 Tension parameters for ECC.
Table 11 Validation of numerical model. Specimen
Bare HSS NC100-f E50bt-f E75bt-f E25t-f E30t-p Mean
Load carrying capacity Tests, Pu-test
FE model, Pu-FE
Pu-FE/Pu-test
139.1 175.79 192.15 191.79 214.61 204.91
148.857 181.089 194.103 201.053 205.314 191.488
1.07 1.03 1.01 1.05 0.96 0.93 1.01
Fig. 13. Compression material curve for LWC.
the HSS used. For the E25t-f beam, an even higher strain of 0.56% was recorded as the beam has the highest resistance and mid-span deflection among the encased beams.
Table 9 Compression parameters for NC and LWC. Material/Specimen
0.4f 'c (MPa)
εy (%)
f 'c (MPa)
NC:NC100-f LWC: E50bt-f and E75bt-f LWC: E25t-f
22 17.74
0.1 0.061
55 44.35
18.11
0.14
45.27
σcu (MPa)
εcu (%)
0.28 0.21
5.25 18.95
0.6 0.4
0.36
12.76
0.8
' c (%)
3.4. Bond-slip behaviours In this study, the beam-end method was used to examine the interfacial slip between the surfaces of ECC/LWC/NC and the HSS section flanges. The beam-end slips were calculated from the data gathered by LVDT-4 to LVDT-7. The slips-deflection and the load–deflection curves for all encased beams tested are plotted in Fig. 8. The final beam-end slip values obtained are summarized in Table 5. Fig. 8 and Table 5 show that the beam-end slip values were negligible (from 0.18 mm to 0.68 mm) for the NC100-f, E50bt-f, E75bt-f and E25t-f beams when the compressive surface of the NC/ECC layer was crushed (the vertical lines in the figures). However, as shown in Fig. 8(e) for the E30t-p beam, a large beam-end slip (1.81 mm) was observed when the HSS section was yielded. The slip was increased to 6.9 mm at the end of the test. Detailed numerical modelling described in the next section shown that such a large beam-end slip was due to the rapid development of shear stress within the shear spans that reached the bond strength of the ECCHSS interfaces.
Fig. 14. Steel-ECC interfaces bond-slip model.
3.5. Initial stiffness The initial stiffnesses of all beams are obtained by calculating the slopes of the linear parts of the load-deflection curves and listed in Table 6. It is observed that the NC-100 beam showed the highest stiffness. For the fully encased HSS ECC beams (E50bt-f, E75bt-f and E25t-f), as the elastic moduli of ECC and LWC are approximately twothird of normal concrete (Table 3), their stiffness were lower than the NC100-f beam but still at least 53% higher than the bare steel beam. On the other hand, the initial stiffness of the partially encased E30t-p beam
Table 10 Bond-slip parameters. Properties
Values
Bond strength, τm Slip at bond strength, Sm ultimate slip, SL
0.5 MPa 0.5 mm 15 mm
10
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Fig. 15. Comparison of test and modelling results for specimens: (a) bare HSS, (b) NC100-f, (c) E50bt-f, (d) E75bt-f, (e) E25t-f, (f) E30t-p.
is the lowest among all encased beams.
respectively. Loadings were applied by prescribing downward displacements at two patches of elements under the loading points. The same displacement steps used in the test were applied until a displacement of 125 mm was reached. The default *RAMP function available in ABAQUS [25] was employed to apply the prescribed displacement. During modelling, the beams were discretised by the eight-node hexahedral solid elements (C3D8R) which had been successfully employed by other researchers [18,19,26] in similar analysis. In order to determine an optimal element size, an element size sensitivity analysis was performed by using six different meshes. Fig. 9 shows the predicted load carrying capacity of the E75bt-f beam with different element sizes. From Fig. 9, it can be concluded that a mesh with 40 mm elements could predict the strength of the beam precisely with the smallest
4. Finite element modelling In order to predict the enhancement effects of ECC/LWC/NC encasement, a three-dimensional (3D) FE model which considers the geometrical and material non-linearities as well as the bond-slip behaviours between the ECC and HSS interfaces was developed. 4.1. Loading and boundary conditions and mesh convergence study The support conditions used in the tests were reproduced in the model by applying roller and pin support conditions to two patches of elements corresponding to the actual right end and left end supports, 11
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Fig. 16. Validation of final failure and cracking for the specimens.
For ECC, the compressive and tensile stress-strain relationships used by Kabir et al. [19] and Meng [20] (Fig. 12) were proved to be reliable and were adopted in this study. The average values of the parameters used are given in Tables 7 and 8. Furthermore, the damage model by Lubliner et al. [28], which had been successfully used before [19] was employed again to define the failure of the ECC under compression and tension. For NC and LWC, the compressive behaviours were described by the constitutive model suggested by Carreira and Chu [29] (Fig. 13). The values of all key parameters, which were determined from the material tests results, are listed in Table 9. For their tensile behaviours, a simple a linear relationship was considered [19] so that the peak tensile strength was assumed as 10% of the peak strength of NC/LWC under compression (f’c in Fig. 13). For the material damage models for NC and LWC, the Concrete Damage Plasticity model was employed, and the same set of parameters used in Table 8 of Ref. [19] were used.
Table 12 Beam-end slip values comparison. Beam
NC100-f E50bt-f E75bt-f E25t-f E30t-p Mean
Beam-end slip (mm) Test
FEA
FE/test
1.14 1.67 0.37 0.40 6.86
1.09 1.81 0.35 0.38 6.9
0.96 1.08 0.95 0.95 1.01 0.99
amount of computational efforts and was adopted in subsequent models. The 3D view of the mesh employed in the modelling of the E75bt-f beam is shown in Fig. 10. 4.2. Material properties used in FE modelling
4.3. Bond-slip behaviour modelling
For the material properties of HSS, the well-accepted bi-linear stress-strain model indicated in Fig. 11 [19,27] was used. Values of the materials parameters used in the model were determined by tension coupon tests and are listed in Table 3.
Since bond-slip was observed in the test of the E30t-p beam, an accurate bond-slip model between the HSS-ECC interfaces is essential to predict the encased beams’ behaviours. In ABAQUS, bond-slip between 12
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HSS [19]. Finally, as no bond-slip was observed between the LWC-ECC interfaces, all interfaces between LWC and ECC were tied together in the models. 4.4. Validation of model against experimental results 4.4.1. Load-deflection responses and failure mode predictions All six beams tested were modelled by the proposed FE model. Comparisons between the FE modelling failure load, Pu-FE, and the test failure load, Pu-test, are presented in Table 11. The comparisons of load vs mid-span displacement curves are presented in Fig. 15. Fig. 15 shows that all key stages of the bending tests including initial linear responses, yielding of HSS sections, flexural resistance (i.e. failure load) and crushing of the top NC/ECC layer as well as the post-failure behaviours, were accurately predicted by the FE model. Regarding the failure mode, crushing of the compressive NC/ECC layer and bottom cracking distribution, Fig. 16 indicates that the FE model again produced good predictions which well agreed with the test results. 4.4.2. Bond-slip predictions and analysis The reliability of the model is further verified by comparing the measured bond-slips at the two support ends with the FE predictions. The maximum measured and FE predicted bond-slips for different beams are summarized in Table 12. From Table 12, it is clear that the proposed FE model predicted the final beam-end slips accurately. Furthermore, in order to confirm that the surface-based cohesive model could predict both the initiation and development of slip, the interfacial slip histories for the E75bt-f beam (which represents those fully encased beams) and for the E30t-p beam (which represents the partially encased beam) were compared with the FE predictions in Fig. 17. Fig. 17 indicates that the surface-based cohesive behaviours model predicted the slip histories for both cases accurately. The validated FE model was employed to examine the advancement of interfacial bond stress and slip along the length of the ECC-HSS encased beams. The results obtained for the four ECC-HSS encased beams at five different loading levels, namely at 50 kN, 100 kN, 150 kN and at the peak (failure) load the beams, are presented in Figs. 18–21. From Figs. 18–21, it is observed that the bond stress between the HSS sections and ECC surfaces increased as the applied load increased. As expected, the bond stress curves and the contour plots at the peak load indicate that high shear bond stress was generated along the shear spans. In addition, when comparing Fig. 21 with Figs. 18–20, one can see that as the applied load increased, bond stress was developed much more rapidly for the E30t-p beam than all other fully encased beams. For the E30t-p beam, the bond stress reached the maximum slip strength quickly throughout the shear spans when the applied load is 100 kN which is well below the HSS section’s full plastic moment resistance (170 kN).
Fig. 17. Comparison of beam-end slip vs mid-span deflection between FE and test results for the (a) E75bt-f and (b) E30t-p beams.
the HSS and ECC interfaces can be modelled by prescribing their surface-based cohesive behaviours [18,19,25]. In this model, a linear elastic traction-separation behaviour is assumed prior to the maximum traction (bond strength) τm at a slip of Sm. The bond is failed once the ultimate slip, SL is reached (Fig. 14). Such approach is particularly attractive and accurate for negligibly small thickness of interfaces [30] and when reliable bond-slip data are available from carefully designed push tests [21]. In this study, the data obtained by Rana et al. [21] were used to define the surface-based cohesive behaviours between the HSS and ECC interfaces (Table 10). As no slip was detected between the HSS and LWC interfaces, a simpler Coulomb friction model with a frictional coefficient of 0.25 is found to be accurate enough to model the interfaces between LWC and
Fig. 18. Shear (Bond) stress distribution for the E50bt-f specimen. 13
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Fig. 19. Shear (Bond) stress distribution for the E75bt-f specimen.
Fig. 20. Shear (Bond) stress distribution for the E25t-f specimen.
Fig. 21. Shear (Bond) stress distribution for the E30t-p specimen.
5. Conclusions
encased beams tested, only a very small amount of slip between the ECC encasement and HSS flanges occurred even after the whole HSS section was yielded. As a result, the flexural resistance of the fully encased beams was significantly higher than the HSS section’s full plastic moment resistance and the flexural strength of a similar normal concrete encased beam in which the top concrete layer was crushed before the HSS section was fully yielded. However, for the partially encased ECCHSS beam, noticeable amount of interfacial slip was measured when the HSS flange was yielded. Nevertheless, the ECC encasement was still able to prevent the occurrence of LTB. As it was found that the ECC encasements were crushed after the HSS sections were fully yielded, one of the potential extensions of the proposed enhancement approach is to apply similar encasement to Class 4 slender HSS sections [31,32] by preventing local buckling of the slender flange and web so that the full plastic moment resistance of the slender section could be achieved.
In this study, the effectiveness of using Engineered Cementitious Composites (ECC) encasement to enhance the flexural strength of welded High Strength Steel (HSS) beam was investigated experimentally and numerically. A 3D finite element model was proposed to simulate the flexural responses and interfacial slip of the encased beams. The results obtained confirmed the predictions from previous study [19] that ECC encasement can significantly enhanced the flexural resistance of a bare HSS beam by preventing the onset of lateral torsional buckling (LTB). More importantly, the use of ECC prevented the brittle crushing of the beam’s top layer at peak load. The strain gauge measurements and numerical modelling results also confirmed that so long as the compressive strain at the peak strength of the ECC is greater than the yield strain of the HSS, failure of the ECC layer at the compression side of the beam would occur after the HSS flange is fully yielded. The experimental and numerical results further shown that for those fully 14
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Declaration of Competing Interest
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The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. Acknowledgements The authors greatly acknowledge BISALLOY, Australia for providing the high strength steel plates used in this study. However, the findings and conclusions drawn in this publication are solely the view of the authors only and not necessarily the view of BISALLOY, Australia. Appendix A. Supplementary material Supplementary data to this article can be found online at https:// doi.org/10.1016/j.engstruct.2020.110288. References [1] Ban H, Shi G, Shi Y, Wang Y. Overall buckling behavior of 460 N/mm2 high strength steel columns: experimental investigation and design method. J Constr Steel Res 2012;74:140–50. [2] Ban H, Shi G, Shi Y, Bradford MA. Experimental investigation of the overall buckling behaviour of 960 N/mm2 high strength steel columns. J Constr Steel Res 2013;88:256–66. [3] Wang YB, Li GQ, Chen SW, Sun FF. Experimental and numerical study on the behavior of axially compressed high strength steel box-columns. Eng Struct 2014;58:79–91. [4] Lee CH, Han KH, Uang CM, Kim DK, Park CH, Kim JH. Flexural strength, rotation capacity of I-shaped beams fabricated from 800 MPa steel. J Struct Eng 2012;139:1043–58. [5] Bradford MA, Lui X. Flexural-torsional buckling of high-strength steel beams. J Constr Steel Res 2016;124:122–31. [6] Ma JL, Chan TM, Young B. Experimental investigation of cold-formed high strength steel tubular beams. Eng Struct 2016;126:200–9. [7] Puthli R, Fleischer O. Investigations on bolted connections for high strength steel members. J Constr Steel Res 2001;57:313–26. [8] Coelho AMG, Bijlaard FS. Experimental behaviour of high strength steel end-plate connections. J Constr Steel Res 2007;63:1228–40. [9] Mashiri FR, Uy B, Tao Z, Wang ZB. Concrete-filled VHS-to-steel fabricated section stub columns subjected to axial compression. J Constr Steel Res 2014;95:141–61. [10] Hsiao PC, Hayashi KK, Nishi R, Lin XC, Nakashima M. Investigation of concretefilled double-skin steel tubular columns with ultrahigh-strength steel. J Struct Eng
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