An experimental program was carried out by Rizk et al. (2011) to investigate the influence of the slab size on the punchingshear resistance. Five thick square slabs were tested. The slabs were 300 mm and 400 mm thick with side dimension of 2650 mm. The slabs were loaded through a small column stub 400 × 400 mm. Four slabs were cast using high strength concrete and one using normal concrete. Shear reinforcement was used in one of the high strength concrete slab. The results of the punchingshear strength showed good agreement with the values predicted using the CEB-FIP Model Code (1990). The ACI 318-11 code was found to underestimate the punchingshear strength for all slabs by an average of 17% except for one high strength concrete slab. This high strength slab had the lowest reinforcement ratio and the predicted value was overestimated by 19%. It was highlighted by the authors that the experimental shear strength was different in twoslabs. The slabs had different concrete strength, similar slab thickness and reinforcement ratio. Both slabs were designed to fail under punchingshear stresses. This finding showed that having a constant number for the size effect factor depending on the slab depth only, as used in major design codes, may not account for the size effect and it should also include the concrete compressive strength.
However, both the support including concrete column and with steel plate cannot be modelled in detail in the shell FEA (level III). Instead a simplified modelling method is needed. To investigate alternative modelling methods, two different methods were inves- tigated: (iii) the use of non-tension spring elements and (iv) a ver- tically restrained support area; see Fig. 7 . With non-tension spring elements (iii) the varying pressure over the column support can be modelled, including the possible lifting of the slab from a part of the support area. The compressive stiffness of the springs was cal- culated to equal the stiffness provided by the concrete in the sup- porting column. With vertically restrained support area (iv), fixed boundary conditions in the vertical direction to the slab was added. This approach was studied because it would be the easiest way to model the support, even though it would not reflect the column stiffness or the possible lifting of the slab from the support.
ð1Þ where P exp is the normalized failure load for the control spec- imen, P u is the test failure load for the same specimen, f cu,s is the concrete cube strength for the strengthened specimens and f cu,c is the concrete cube strength for the control specimen. The failure loads of the strengthened specimens were increased by 14–39% compared with the control specimen as shown in Table 2 . The use of a 4 mm steel plate with 16 studs of 8 mm diameter for specimen S-3 increased the failure load by 20%. Increasing the steel plate thickness to 6 mm for spec- imen S-4 did not affect the failure load of that specimen com- pared with S-3. This showed that the steel plates did not develop their full strength. On the other hand, using weaker studs, 6 mm in diameter, for specimen S-2 resulted in limiting the increase of the failure load to only 14% compared with the control specimen. This showed that the contribution of the steel plates relied mainly on the shear studs which improved the bond between the steel plates and the concrete surface due to the developing shear friction and the pull-out strength resisting the formation of the punching failure mechanism. The use of a large size steel plate with 24 studs for specimen S-5 resulted in increasing the failure load by 39% compared with the control specimen. This result further confirmed the previous conclusion. The increase in the number of shear studs enhanced the shear friction and pull-out resistance of the studs and, therefore, increased the contribution of the steel plates to the strength of the tested specimens.
Abstract. An innovative solution to the corrosion problem is the use of fiber-reinforced polymer (FRP) as an alternative reinforcing material in concrete structures. In addition to the non corrodible nature of FRP materials, they also have a high strength-to-weight ratio that makes them attractive as reinforcement for concrete structures. Extensive research programs have been carried out to investigate the flexural behavior of concrete members reinforced with FRP reinforcement. On the other hand, the shear behavior of concrete members, especially punchingshear of two-wayslabs, reinforced with FRP bars has not yet been fully explored. The existing provisions for punching of slabs in most international design standards for reinforced concrete are based on tests of steel reinforced slabs. The elastic stiffness and bonding characteristics of FRP reinforcement are sufficiently different from those of steel to affect punching strength. In the present study, the equations of existing design standards for shear capacity of FRP reinforced concrete beams have been evaluated using the large database collected. The experimental punchingshear strengths were compared with the available theoretical predictions, including the CSA S806 (CSA 2012), ACI-440.1R-15 (ACI 2015), BS 8110 (BSI 1997), JSCE (1997) a number of models proposed by some researchers in the literature. The existing design methods for FRP reinforced concrete slabs give conservative predictions for the specimens in the database. This paper also presents a simple yet improved model to calculate the punchingshear capacity of FRP- reinforced concrete slabs. The proposed model provides the accurate results in calculating the punchingshear strengths of FRP-reinforced concrete slender slabs.
groups of specimens were assessed in this study. The first group (A) involves three solid slab specimens, while the second group (B) includes three slab specimens with an opening in the shear zone. Two variables were included in these experiments, namely, the effect of strengthening by CFRP laminate and the effect of high strength self-compacting concrete. The outcomes were discussed based on cracking load, ultimate punchingshear capacity, crack patterns, and load-deflection response. The results showed that strengthening by CFRP enhanced the ultimate punchingshear capacity by (23%-65%) and increased the deflection by (33%-79.5%).
Structural failures of highway bridge structures are not common under static loading. However, in highway bridges, beam and slab failure usually occur in two common forms, direct flexural and/or punchingshear. The direct flexural failure typically occurs in beam or slab members and is associated with overall bending. This type of failure arises from the formation of diagonal tension cracks in the region of maximum bending moment and extends across the entire width of the member. However, punchingshear failure is a more localised effect associated with thin slabs or two-way slab-column members when subjected to a highly concentrated load. Punchingshear failure occurs when the principal stress across the critical surface of the section exceeds the tensile strength of the concrete due to applied loading and failure occurs with limited warning. An example of this type of complete failure is seen in Figure 1. Failure occurs with the potential diagonal crack following the surface of a truncated cone around the load. The failure surface extends from the bottom of the member diagonally upward to the top surface. For a normal concrete slab, the angle of inclination of the failure surface ranges approximately from 20 to 45 degrees depending on the amount of shear reinforcement . However, very little information on this parameter is available for UHPFRC.
They used different CFRP systems to increase the flexural capacity of twoway simply supported slabs , (strips of laminate plates with Cold cured adhesive bonding, prestressing strips of laminate plates, wet lay-up ply of Fiber laminate sheets, near surface mounted strips of laminate bars (NSM) ). They found that CFRP increased the flexural strength between 63% to 145% and remarkably reduced the deflections and crack widths, especially the prestressing CFRP system. Two modes of failure were observed. Delamination occurred in the cases CFRP with cold cured adhesive and prestressing CFRP while rupture of the CFRP reinforcement was observed in the other cases of the CFRP system.
Previous studies (Sagaseta et al. 2011; Matesˇan et al. 2012; Fall et al. 2014) that conducted two-way bending tests used a pointed load at the center of the slab. However, with a pointed load, an unexpected punchingshear failure could occur because of the concentration of shear stress around the loading point. Therefore, the application of multiple loading conditions was considered to provide better results, partic- ularly in the case of voided slabs: voided slabs are vulner- able to the shear strength deterioration because they use of less concrete in the slab web and therefore less concrete is available to resist shear. For this reason, Ibrahim et al. (2013) Fig. 1 Details of the donut type void former.
In the 1990ies, a new system was invented by Jorgen Breuing, eliminating the above problems, the so called BubbleDeck technology. The Bubble Deck technology uses spheres made of recycled industrial plastic to create air voids while providing strength through arch action. As a result, this allows the hollow slab to act as a normal monolithic two-way spanning concrete slab. These bubbles can decrease the dead weight up to 35% and can increase the capacity by almost 100% with the same thickness. As a result, bubble deck slabs can be lighter, stronger, and thinner than regular reinforced concrete slabs (3).
2. The U-Boot Beton® formworks are situated utilizing the sidelong spacers joints to put them at the coveted focus separate that will decide the shaft width. Because of the conic lift foot, theU-Boot Beton® formworks will be lifted from the surface, making it feasible for the lower section to be framed. Assuming twofold or triple components are utilized, these components should first be amassed, which will be provided on particular beds in the yard. 3. The situating of the reinforconcretes is finished by putting over the U-Boot Beton® formwork the upper bars in the two headings and in addition the support for shear and punching where vital, as per the plan.
In the laboratory, tests were planned to determine the tensile strength, modulus of elasticity and ultimate stain. The test was carried out by preparing the specimen which is a representative of the batch being tested. The length of the specimen is the full length of the test section and the lengths of the anchoring sections. The total length was 50 mm, 10 mm is the length of the test section, and 20 mm is the length of each anchoring section at both end of each specimen Figure 3-3. The number of test specimens for each diameter and type was five. All GFRP bars specimens were stored in the standard laboratory atmosphere before testing (23 ± 3 °C and 50 ± 10% relative humidity). Next, the specimen was mounted on the testing machine with care to ensure that the longitudinal axis of the specimen matches with the line joining the two anchorages fitted to the testing machine (Figure 3-4). Then, the data acquisition system was connected before starting the load. The rate of the load was kept constant increments during the test (5 kN with a rate of 0.03 mm/sec) in such a way that the specimen failed within approximately five minutes (Figure 3-5). The load was increased until tensile failure occurred, whereas, the strain measurements were recorded up to 50% of the expected tensile capacity.
A combined bending and shearpunching test series, called X, is carried out at VTT. The target structure is a two-way simply supported concrete plate with a span of 2 m and a thickness of 25 cm. The deformable projectile is made of stainless steel tube with a shallow spherical dome nose and its mass is 50 kg. In the first two tests, X1 and X2, the outer diameter of the missile was 253 mm and the wall thickness was 3 mm. The impact velocity was 166 m/s. No clear shearpunching occurred in these tests. In order to achieve shearpunching, the missile was modified for test X3. The diameter of the missile was 219 mm and the wall thickness was 6.35 mm. For test X4 the impact velocity was increased from 144.7 m/s to 168.6 m/s. Capabilities of different calculation methods in assessing both global bending deformation and local shear deformation and possible shearpunching are studied. The models and methods comprise a two-degree-of-freedom (TDOF) model CEB (1989) and two finite element (FE) programs, an in-house code and a commercial general purpose code Abaqus (2014). Additionally, some semi-empirical formulae are used for comparisons. Tests X1 and X2 have been analysed in Borgerhoff et al. (2013).
This paper introduces two new solutions for a best predic- tion of punching capacities. The ﬁrst is a new empirical model which is a modiﬁcation of the El-Gamal et al.  equation. The second solution is using the Artiﬁcial Neural Networks Tech- nique. Each of them contains two new parameters, never used before; which are the effects of the elastic or ﬂexural stiffness of the main bottom FRP reinforcement, and the effect of the continuity of slabs in the longitudinal and/or in the transverse direction on punching capacity. The paper also examines the validity of existing shear design recommendations (interna- tional design codes & models) for slabs with internal FRP bars reinforcement.
As previously stated, the calibrated FEM is based on the Concrete Damaged Plasticity Model available in ABAQUS. This material model accounts for concrete cracking through a smeared crack approach, making the model results mesh size-dependent . As mentioned previously, a global mesh size of 20 mm was found to result in the best correlation between the experimental and numerical results. In order to improve the mesh uniformity in the L-shaped column study, the slab dimensions, column dimensions and loading points were modified as summarized in Tab. 1, and Fig. 3b. The impact of these modifications for the Hawkins et al.  slabs was minor as shown in Tab. 1. The modified slab dimensions, column height, effective depth and loading point locations are used in the L-shaped column study, with equal loading applied on each point, as shown in Fig. 3b.
Table 2.1 Typical properties of various reinforced fibres 9 Table 2.2 Typical properties of various polymer matrix resins 11 Table 2.3 Typical average properties of various core materials 13 Table 3.1 Material properties of E-glass/vinylester of unidirectional lamina 43 Table 3.2 Material properties of vinylester/SLG of PFR core 44 Table 5.1 Summary of 3D FEA results for the first ply failure prediction 118 Table 5.2 Summary of 3D FEA results for PFR cracking load prediction 124 Table 5.3 First natural frequency of the two-way FRP slab 128
warning that the structure is about to fail so that there will be sufficient time to react. The aim of study is to verify the influence of steel reinforcement on the modulus of elasticity of reinforced concrete members and also to compare the parameters such as deflection, strains and crack width between the slabs that are laid by using complete natural coarse aggregate and 50% of recycled coarse aggregate.
24 removed each year before the snows melt, and then reinstalled once the danger of flooding has passed. Light weight was thus a significant factor in the design, and the two 12.5-m spans were lifted into place by helicopter (Burgoyne 1999, Keller 1999, Hollaway and Head 2001, Keller 2002). Light weight was also one of the main reasons for choosing a fibre composite solution for the Bonds Mill Bridge – the world’s first drawbridge for vehicular traffic on the Stroudwater navigation canal in England, since this meant that both lifting tower and counterweight could be eliminated. The bridge was erected in 1994 and assembled from the same modular system as used for the Aberfeldy Bridge (Åström 1997, Hollaway 2000, Hollaway and Head 2001). Recently an all-composite full-service bridge designed by the University of California at San Diego (UCSD) was proposed to provide a link between the east and west part of the campus, which are separated by a ten-lane interstate freeway (Schwartz 1997b, Hollaway and Head 2001). The bridge is designed as a 137 m long dual plane fan-type cable-stayed bridge supported by a 58 m high A-frame pylon with circular legs filled with normal concrete (Figure 2.10). The bridge deck design utilizes a new FRP modular system – Hybrid Tube System (HTS) developed at the University of California, San Diego. Details of this new FRP composite bridge system can be found in Karbhari and Seible (2000), Karbhari et al. (2000), Hollaway and Head (2001).
The design of most reinforced concrete (RC) structures is typi- cally governed by ultimate limit state performance of the various structural elements when subjected to static loading, e.g. dead (or permanent) loads and live (or imposed or variable) loads. Whilst the former are typically static in nature (e.g. structure ’ s self-weight, ﬁ nishes etc.), the magnitude of live loading tends to be variable with time (e.g. pedestrian or vehicular traf ﬁ c loading). However, in most cases, such loadings can be idealised as quasi-static, since the rate at which this loading is applied, typically described by the strain-rate, _ 3 , is of a very small magnitude.
From equation 2, it is obvious that the above ACI 318-95 code equation totally ignores the influence of tension flexural reinforcement when calculating the concrete shear resistance and depends heavily on concrete strength. This is not too unreasonable in the case of steel reinforcement, since with its high modulus of elasticity, the dominant factor determining concrete shear resistance will be the area of concrete in compression, which remains constant as the neutral axis depth does not vary much with normal steel reinforcement ratios. Hence, in case of steel, the ACI 318-95 equation gives good predictions though conservative at low levels of reinforcement. However, when using FRP reinforcement with low modulus of elasticity, the concrete shear resistance becomes more sensitive to the reinforcement stiffness, as the neutral axis depth reduces significantly with low reinforcement ratios. In such cases, the ACI 318-95 prediction becomes too unconservative.
with flat slabs with no beams and no enlarged column heads .These slabs are subjected to punchingshear failure of slab-column connections. Load concentration around the column head generally leads to increased stresses which cannot be absorbed solely in thin slab thicknesses. The punchingshear strength is an extremely significant parameter for the design of flat slabs. Solution of architecture often imposes the necessity of openings near the columns .In this paper will examine the effect of openings in different positions and with different dimensions in the punchingshear strength. The calculations performed in the punchingshear application are based on the standard EN 1992-1-1:2004.