compressive strength (f'c) according to ASTM C39. Secondly, similar ... researchers [24, 25] and was applied to remove the free water in the concrete pores and.
Residual strength of high strength concentric column-SFRC flat plate exposed to high temperatures By FARID H. ARNA'OT, AHMMAD A. ABBASS, AHMED ABBAS ABUALTEMEN, SALLAL R. ABID, and MUSTAFA OZAKCA
Abstract It became well known that steel fibers could significantly improve the mechanical properties of concrete and the ductility of flat plat-column assembly at ambient conditions. The investigation of the contribution of the steel fiber and the high strength concrete under high temperature conditions on the structural behavior of flat plates is still in need. This research presents experimental results of sixteen steel fiber reinforced slabs exposed to four levels of temperature reaching 550 °C, in addition to four control slab specimens. The residual mechanical properties, ultimate slab strength, and deformations were investigated. Relations between the concrete tensile and compressive strengths after exposing to high temperatures were introduced. For all temperature levels and steel fiber contents, the strength of most of the slabs was improved compared to the corresponding strength at room temperature. The highest percentage improvement was recorded for specimens exposed to 150 °C and reinforced with 1 % of steel fiber. The strengths of the slabs with 1% and 1.25% of steel fiber and exposed to 550 °C were almost the same as their strengths at ambient conditions, however, slabs with 0.75% showed 13% increase. Keywords: steel fiber reinforced concrete, high temperature, flat plate-column connection, punching shear
1.
Introduction
Flat plate-column assembly that is reinforced with suitable flexural reinforcement and without shear reinforcement may fail before the yielding of the flexural reinforcement. This failure type may take place in two-way shear, which is a brittle and catastrophic 1
failure mode. In general, the shear behaviors always associated with many other forces, which make the analysis processes difficult to tackle [1]. The problem is much complicated when the slab system is exposed to high temperatures, where two additional direct and indirect effects arise due to load redistribution and decay in concrete properties [2 - 4]. Many mechanical models were proposed for predicting the two-way (punching) shear strength under ambient conditions [5-12]. For all, the main two failure mechanisms are related to cracking due to diagonal tension and crashing due to compression inclined struts. Unfortunately, there are few researches available in the literature related to the concentric flat plate-column joints under high temperature conditions [2, 13-17]. Moreover, there are no provisions in the practical design codes related to both punching shear problem under fire condition and the effect of steel fiber on punching strength at both ambient [18] and at high temperature conditions. Punching shear resistance may be improved by increasing the concrete strength because of its contribution in the enhancement of the structures strength. On the contrary, in High Strength Concrete (HSC), spalling [19] and brittleness problems arise. Among the best and cheapest processes to improve the flat plate-column joint is the using of Steel Fiber Reinforced Concrete (SFRC). In spite of the discontinuity and the random distribution of steel fibers, which decreases its efficiency in sustaining the tensile stress compared to conventional reinforcing bars [20], it shows good results in controlling the propagation of the local cracks, which in turn increases the ductility and energy absorption [18, 21, 22]. The contribution of steel fibers arises by increasing the bond between the reinforcement bars and the surrounding concrete matrix and bridging the two sides of cracks, which transfers the stresses even in wide crack openings [20, 23]. Based on the premises, an experimental program was directed in this research to investigate the residual punching shear strength of SFRC plates after exposure to 2
different levels of high temperature. The combined effect of concrete strength and the presence of steel fibers under high temperatures were also investigated, where SFRC plates with compressive strength of approximately 60 MPa were tested.
2.
Experimental work
2.1 Materials properties and concrete mixture Slab panels with high strength concrete and reinforced with steel fibers and conventional steel bars were produced to conduct the experimental work of this study. The concrete mix was designed to achieve a 28-day cylinder compressive strength (f 'c) of 60 MPa, in which higher volume of fine aggregate was used. In all specimens, two types of aggregates were used, crushed stone as coarse aggregate and sand as fine aggregate. The maximum size of aggregate was 10 mm. It is clear in Table 1 that coarse aggregate composes only 37% of the total amount of aggregate, whereas the rest 63% is fine aggregate. Fig. 1 shows the grading of the coarse and fine aggregates. Portland cement type CEM II/ A-LL 42.5 R in accordance with EN 197-1 was used. In addition to the increasing of fine particles in the mix, 7% of the cement was replaced by silica fume. It is expected that increasing the fine materials, presence of steel fiber, and the approximately moderate water / (cementitious materials) would decrease the slump. Therefore, the high range water reducer Glenium 51 was used in all mixes. The slump results of the produced mixes are shown in Table 1. At the first stage of mixing, the fine and the coarse aggregates were mixed together for 10 minutes before adding the cementitious materials and then all the materials were mixed for 5 minutes. The water was then gradually added. The slump was checked in this stage, then after, the water reducer was added prior to the adding of the steel fibers. Three percentages of cold drawn glued steel fibers of 0.75, 1, and 1.25% by volume of 3
concrete were used. The characteristic of the steel fibers with hooked-ends that employed in this research are shown in Fig. 2 and Table 2. All casted specimens were kept under laboratory conditions for 24 hours. The slabs and the associated cylinders were removed from their molds, and were cured in water pools for 28 days so that both slab and cylinders are exposed to the same curing conditions. The average of three standard cylinders with the same properties and conditions was employed in the structural analysis. The experimental program includes four test series. In the first, standard cylinders of 100×200 mm subjected to uniaxial compression load were used to evaluate the concrete compressive strength (f'c) according to ASTM C39. Secondly, similar cylindrical specimens were subjected to indirect tensile load (splitting) to evaluate the concrete tensile strength (fct) according to ASTM C 496. In the third series of cylinders, the specimens were subjected to compression load to evaluate the Young’s modulus according to ASTM C469. The fourth specimen series includes the two-way slab plates, which were tested under monotonic increasing load
2.2 Specimens’ geometry All tested slab specimens are 500× 500× 60 mm square reinforced concrete plates as shown in Fig. 3. The slabs of conventional reinforced concrete under ambient conditions were designed to fail due to low flexural capacity. For which, they were reinforced orthogonally by ten deformed reinforcing bars of 5.3 mm diameter in each direction. Thus, the bar spacing was equal along the two directions where Sx = Sy = 52 mm centerto-center. There were no shear or compression reinforcements. The ultimate and yield strengths of the reinforcing bars were found to be equal to 477 MPa and 400 MPa, respectively. All bars were 90-deg hooked as shown in Fig. 3. The concrete cover was 4
15 mm, which leads to an average effective depth d = 39.7 mm. For structural analysis, the effective depth d and the reinforcement ratio ρ of the slab section are taken as the average value of the two layers of the tension reinforcement. So that d = (d1 + d2) / 2, and ρ
ρ x d1
ρy d2 2
d
, where d1 and d2 are the distances from the center of the bar in the
orthogonal directions at the tension face to the upper fiber (the compression face) of the slab, respectively, while ρx and ρy are the associated reinforcement ratios. Twenty specimens were designed to represent an interior flat plate-column assembly. These specimens are divided into four groups (A to D) based on the maximum exposure temperature (150, 300, 450, and 550 °C). Each group consists of four specimens (A1 to A4, B1 to B4, C1 to C4, and D1 to D4) with different steel fiber dosages. In addition, all mentioned groups are compared to the control group (E1 to E4) that was tested under ambient conditions (20 °C). Table 3 shows the details of the tested groups.
2.3. Application of temperature After 28 days of water curing, all slabs and standard cylinders were removed from curing tanks for testing. Excluding group E (tested at ambient conditions), the surfaces of the specimens in other groups were dried and then heated for 24 hours at 105 °C before the specified temperatures are applied. This procedure was followed by previous researchers [24, 25] and was applied to remove the free water in the concrete pores and prevent sudden damaging of the specimens due to the sudden increase in the pore water pressure [25]. This procedure would also reduce the shape effect on water evaporation during the heating process between the cylinders and the slabs. However, it is worthy to mention here that in this study, increasing the fine particles in the mix would reduce the porosity and/or reduces the link between the pores, which increases the growth of the
5
pore water pressure. Thus, the preheating to 105 oC may reduce the free water in the pores but not expel it all. Felicetti and Gambarova [26] showed that after 7 days of drying at 105 °C, approximately 40 % of the free water was remained restricted, which may lead to concrete spalling due to the sudden increase in temperature. In order to obtain a reliable comparison, each group of slab specimens and their associated standard cylinders for each temperature level were step inside a furnace of clear dimensions of 1360× 1300× 620 mm and of 1200 °C capacity, which is shown in Fig. 4a. Except the furnace floor, electrical heaters are installed on the five other sides to ensure that the heat is applied homogenously on the heated specimens. The temperature inside the furnace is measured using type-K thermocouples. After drying the specimens for 24 hours, the heating process was applied in three continues stages of heating, soaking, and cooling as depicted in Fig. 4b. For the heating stage, the temperature was increased up to the specified temperature at an average heating rate of 10 °C/min. The heating rate plays an active role when the maximum temperature exceeds 300 ºC [27]. Different heating rates were found in the literature. Purkiss [26] used 2 °C/min, Chen et al. [28] and Poon et al. [29] used 2.5 °C/min, and Lawson et al. [30] used 5 °C/min. On the other hand, Zhang et al. [27] studied the effect of heating rate on the strength of concrete. They used three different rates of 1, 3, and 10 °C/min. They found that the 10 °C/min is the most severe case, where after heating to a maximum temperature of 450 oC, the compressive and tensile strength deferred by 36 % and 60 % when the heating rate increased from 3 to 10 °C/min, respectively. The specimens of groups A, B, C, and D were exposed to 150, 300, 450, and 550 °C, respectively. These temperature levels are almost in-between the max temperature levels adopted by many researchers [26, 28-30]. It should be mentioned that the adopted
6
heating rate is lower than the ISO834. However, it was found that the thermal effects of the standard ISO834 are more severe than those induced from real fires [4]. The adopted heating process is more proper to simulate the maximum temperatures induced from accidental fire in close car parking that includes lot of flat slabs. For instance, Bamonte and Felicetti [31] showed that the maximum temperature resulting from six cars sequentially ignited in a close car parking is 550 °C. The number of cars used in their work is approximately within 98 % of the incident found in an underground car parking in Paris [32]. As a next step in the heating process, the furnace was programmed to keep the temperature constant at the specified maximum temperature for three continuous hours. The soak period (constant temperature) was adopted to allow uniform temperature distribution [28]. The same procedure was followed by Chen et al. [28] and many other previous researchers. However the soak period in this work was extended to minimize the temperature differences between the cylinder and slab specimens due to their different shapes and sizes. At the end of the soak period, the furnace was cooled down smoothly by opening the ventilation cover at the roof of the furnace (Fig. 4). The gradual cooling process aimed to prevent the effect of sudden variation in thermal strains due to the high temperature gradient produced from sudden cooling, which is dependent of the maximum reached temperature.
2.4. Test setup Both slabs and the standard cylinders of the same age and mixture properties were tested at the same time. In all testes, the load was increased monotonically until failure. For the slabs, the tests were conducted using INSTRON universal testing machine (250 kN 7
capacity) and were loaded in a displacement rate of 0.4 mm/min. For convenience in conducting the slab tests, a special steel support was constructed where the tested slabs were supported by eight steel half balls that were fixed on a steel frame as shown in Fig. 5b. Cracks initiation and propagation were monitored using a mini camera, which is fixed at the base of the testing machine and connected to a computer. The monotonic increasing load was applied from the hydraulic jack through the load cell to a steel cubic shift of 55 mm size that represents the column stub, which is not affected by load and temperature during the test. In such case of loading, the slab has some freedom to control its internal forces to distribute in boundary points [33]. The circulation of the half balls is necessary to allow the slab to rotate in two directions freely without any significant resistance. Moreover, the curvature of the half-balls was manufactured so as not to penetrate the concrete slab due to singularity of the stress in the contact area. These steel half-balls were uniformly distributed with a central angle equal to π/4 and the radii rq equal to 200 mm as detailed in Fig. 5a.
3.
Design of the specimens
The dimensions of the slabs were chosen within the relative size found in literature [2, 13-17], where the relative size in these references are as follows, c/d is ranging between 1.39 to 3.87, b/d is ranging between 10.5 to 45, and c/b is ranging between 0.085 to 0.15. As introduced previously, there are no clear provisions on SFRC punching shear strength or its residual strength at high temperature conditions. The salient aspect is that different codes of practice introduced the nominal punching shear stress at ambient conditions vn as a shear force per critical area (V / u1.d). This stress should be resisted by
8
concrete strength parameters in term of compressive strength and (if existed) with the contribution of special shear reinforcement. In general, the codes’ equation of punching shear strength is in the form of, Vn = Vc + Vs
(1)
where Vn is the nominal shear strength, Vs is the shear reinforcement contribution (which is ignored in the present work), and Vc is the characteristic concrete shear strength that is equal to; Vc = φ × τc × ξ(d) × f(ρ) × u1 × d
(2)
where φ is the safety factor, τc is a function in terms of compressive strength, ξ is the size factor in terms of effective depth d, f (ρ) is a function of tension reinforcement ratio, and u1 is the critical perimeter that distances ℓcri from the column face in term of the effective depth d. It is worthy to mention here that the parameters in Eq. 2 are different from a code to another where ℓcri = 0.5d in ACI318 [34], AS3600 [35], CSA A23.3 [36], and IS456 [37], ℓcri = 1.5d in DIN1045 [38], and ℓcri = 2d in EC2 [39]. Moreover, the effects of the flexural reinforcement ratio ρ and the size factor ξ(d) are also altering among the codes where some of codes do not include these effects. Among these codes are the ACI318 [34], AS3600 [35], CSA A23.3 [36], and IS456 [37]. The missing factors may have small effects in case of analyzing thick slabs, whereas, these effects significantly appear in thin slabs. For example, the ACI318 equation overestimates the shear stress in case of ρ < 1% and d > 200mm [40, 41]. In contrast, the advocates of this code deem that the proposed equation is never used as a shear prediction. It is intended to be interacted with proper flexural design aimed to fail in flexural shear Pflex< Vc so that full moment redistribution is allowed [42], where Pflex is the load of failure produced due to the formation of a yield-line.
9
The estimated flexural capacity of the designed slab (Pflex) is obtained from Eq. 3, based on the expected yield line patterns as shown in Fig. 6 [41]. 8 m b2 (b c) (c 2 /4) b c b b2 2 (c b1 )
flex
where, mR is the section moment capacity, while the other parameters are illustrated in Fig. 6. m
fy
d2 (1
ρ
0 5ρ
fy ) f c
Table 4 shows both the flexural capacity Pflex according to Eq. 3, and the punching shear strength of the slab based on different codes’ provisions. In the present work, the specimens under ambient conditions were reinforced with conventional deformed steel bars and were designed to fail due to low punching shear strength (Pflex > Vc). This is to recognize the effects of the interaction of high strength concrete with the steel fiber dosages. Where Vc is the non-factored concrete punching shear strength, which obtained from different codes’ equations [34, 38, 39, 43]. Moreover, the designed sections were also checked by employing the failure mode criterion (Eq. 4) that was proposed by Gesund and Kaushik [44]. ρ2 fy d2 104 16 c
b
√f
(4)
c
where f 'c, and fy are the compressive strength of concrete and the yield strength of the reinforcement bars, respectively, ρ is the reinforcement ratio, d is the effective depth, and b and c are illustrated in Fig. 6. If the dimensions are expressed in pounds and inches, the types of failure can be characterizing based on Q ranges as follows, For Q < 2, yield line will form, i.e. the specimens fail in flexure.
10
For 2 ≤ Q ≤ 4, both flexure and shear failure types have an opportunity to take place. For Q > 4, punching shear failure is dominant. For the current designed slabs, Q was 2.66. Therefore, either flexure or shear failure type is expected. Knowing that this criterion is conservative in shear, however in flexural it may not [22].
4. Experimental results and discussions 4.1. Mechanical properties of fresh and hardened concrete 4.1.1. Slump test The main problem in fibrous concrete is the tendency to produce balling of fibers in the fresh mix, which is affected by many factors such as the maximum size of aggregate and its overall gradation, in addition to the fiber three main parameters (aspect ratio, volume fraction, and fiber shape). It was found that as the maximum size of aggregate and the aspect ratio of the steel fiber increase less volume fraction of fibers should be used (ACI 544.1). Table 1 and Fig. 7 show the slump results of the concrete mixes with the adopted fiber contents. From Fig. 7, it is can be recognized that the slump decreases smoothly with increasing the steel fiber dosage. The adding of 0.75% of SF decreased the slump by approximately 30%, while using 1.25% of SF decreased the slump by only additional 10%. The results are compared with slump tests from the literature [25], in which steel fiber having lf /Df = 80 was used. With almost the same content of fine aggregate, the slump of this study was higher as the content of SF exceeds 0.75%. This can be attributed to the effect of the incorporation of silica fume in this study, which has higher surface area.
11
4.1.2. Concrete compressive strength After completion of heating and cooling processes, the diameters of the cylinders were checked to ensure that there are no differences in diameters more than 2% according to ASTM C39. The top contact surface with compression machine was capped using melting sulfur mortars. The results of the tested specimens are listed in Table 3 and shown in Fig. 8. From Fig. 8a, it is shown that the compressive strength (f'c) of the cylinders exposed to 150 °C was increased by 10% to 19% compared to its strength at ambient conditions, no matter what amount of steel fiber was used. The aforementioned improvements were decreased to about 3% to 8% for the cylinders exposed to 300 °C. Increasing of compressive strength after exposure to temperatures up to 300 °C agrees with findings of previous researches [19, 45]. Many physical and chemical factors related to both cementitious materials (cement + silica fume) and aggregate lead to the increase of compressive strength of concrete exposed to temperatures up to 300 °C. Recalling the mix design of the current work, the percentage volume of fine aggregate is much higher than of coarse aggregate, which increases the surface area that absorbs the free water uniformly and effectively. Thus, produces a matrix with more homogenous particle distribution. Moreover, increasing the graded fine particles reduces the pores and envelopes the coarse aggregate actively, which enhances the bond between the steel fibers and the coarse aggregate. Consequently, such factors lead to better distribution of stresses and hence reduce the stress localization. In this study, this activity was more obvious after exposure to 150 ºC as shown in Fig. 8a. Slow-rate decay in compressive strength was noticed for specimens exposed to temperatures above 300 ºC where exposure to temperatures of 450 °C and 550 °C decreased the compressive strength compared to that at ambient conditions. The 12
percentage decrease defers according to the steel fiber content. The cylinders with 0.75 % showed the lowest decay percentages (about 10% at 450 °C and 23% at 550 °C). The compressive strength of cylinders with 1% decreased by approximately 22% and 47% at 450 °C and 550 °C, respectively, while the percentage decay values at the same temperatures were 17% and 49%, respectively, for cylinders with 1.25% of steel fibers. Many factors control the decay of compressive strength due to exposure to high temperatures. In addition to the effect of mix proportions, these factors include the heating and cooling rates, maximum exposed temperature, initial compressive strength, and moisture content. Inclusion of steel fiber decreases the rate of decay because of the bridging effect of these fibers. However, due to the dehydration of the calcium hydroxide between 400 and 550 ºC [46], the bridging effect was significantly reduced. Nonlinear regressions of second and third orders were obtained for the effect of high temperatures on f 'c, which are represented by Eq. 5 and Fig. 8b. f 'c(T)0 = f 'c × [(-3.94 × 10-6 × T2) + (0.00115 × T) + 1]
(5a)
f 'c(T)0.75 = f 'c × [(7.9 × 10-9 × T3) - (1 × 10-5 × T2) + (0.0028 × T) + 0.95]
(5b)
f 'c(T)1.0 = f 'c × [(6.12 × 10-9 × T3) - (1.005 × 10-5 × T2) + (0.0028 × T) + 0.946]
(5c)
f 'c(T)1.25 = f 'c × [(- 4.25 × 10-6 × T2) + (0.00156 × T) + 0.965]
(5d)
where f 'c(T) is the compressive strength of plain concrete and SFRC that is exposed to temperature T. The subscripts 0, 0.75, 1, and 1.25 represent the steel fiber volume fractions. Eq. 5a of plain concrete that exposed to high temperature was compared with other references [28, 47-49] as shown Fig. 9. For the standard cylinders exposed to temperatures between 450 °C and 550 °C, the results showed that the rates of decay in compressive strength (f 'c) are within the formulae in the mentioned references. For the
13
standard cylinders exposed to temperatures less than 300 °C, there was no degradation in f 'c of plain concrete. Oppositely, there was an increasing of about 19% for cylinders exposed to 150 °C.
4.1.3. Concrete tensile strength Cylinder splitting test was conducted to evaluate the tensile strength. The test results are summarized in Table 3. The residual tensile strength was lower than the residual compressive strength, which means that the exposure to high temperatures is more severe on tensile strength than on compressive strength. The results of this study agree with results of splitting tensile tests conducted by Faiyadh and Al-Ausi [50]. The inclusion of steel fiber was found to improve the concrete tensile strength compared to plain concrete specimens under the same exposure conditions, which is attributed to the bridging activity of the steel fibers. The maximum increase in tensile strength was recorded at 150 oC for specimens with 0.75% of steel fiber as shown in Fig. 10, which was approximately 35% compared to unheated specimens (ambient conditions). Fig. 10 also shows that the tensile strength of specimens with 1.25% of steel fiber was increased by 14%, 17%, and 14% for specimens exposed to 150, 300, and 450 °C, respectively, compared to their corresponding unheated specimens. Fig. 11a shows the comparison of the residual tensile strength of plain concrete with that adopted by Eurocode 2 [48] and Faiyadh and Al-Ausi [50]. As shown in the figure, up to 200 °C, the decay rate seems to be similar to that of Faiyadh and Al-Ausi. In this work, the tensile strength decreased continuously with a semi-constant slow rate up to 550 °C. Beyond 100 oC, Eurocode 2 proposes linear decrease, yet with higher decay rate than that obtained in the current study. On the other hand, the results obtained by Faiyadh and
14
Al-Ausi show that the tensile strength-temperature relationship is multilinear with strength recovery at 350 oC. Although the rate of strength decay is almost identical with that of the current study up to 200 °C, the rate of decay beyond 350 oC is clearly higher and is comparable to that of Eurocode 2. The gap between the three curves is apparently the highest at 350 °C. The residual tensile strength in the current study after exposure to 350 °C is 66% compared to 50% proposed by Eurocode 2 and 87% introduced by Faiyadh and Al-Ausi [50]. These differences are attributed to the variation in heating and cooling rates, mix proportions, type and maximum size of aggregate, cooling process, and age of specimens. In their work, Faiyadh and Al-Ausi [50] soaked their specimens for 90 minutes, which differs from that adopted in this study. Moreover, the water/cement ratio in their study was 0.55, which is higher than that used in current work. As shown in Fig. 11b, the behavior of specimens reinforced with 1% of steel fiber is quite similar to that of their corresponding specimens tested by Faiyadh and Al-Ausi [50], which is multilinear with partial strength recovery between 300 to 350 ºC. This behavior can be attributed to the same affecting factors discussed earlier in section 4.1.2.
4.1.4. Compressive-related concrete tensile strength As mentioned in section 3, codes of practice have different provisions related to Eq. 2. Among the parameters included in this equation is the compressive strength f'c. The provisions of ACI318 [34], AS3600 [35], CSA A23.3 [36], and IS456 [37] adopt the square root of compressive strength (√ impose to use the cubic root (√
), while Eurocode 2 [39] and DIN 1045-1 [38]
). However, SNiP 2.03.01-84 [51] and TS500 [43]
suggest using the concrete tensile strength.
15
The relationship between concrete tensile and compressive strengths can take the form of, ki
fct
(6)
a
√f c
where fct and f'c are the splitting tensile and compressive strengths, respectively, ki and a are constants acquired from the experimental results. In the literature, there are many relationships between SFRC compressive and tensile strengths for specimens treated in room temperature, among which are those introduced by Shaaban and Gesund [22] and De Hanai and Holanda [18] in Eq. 7 and Eq. 8, respectively, 196 25 wc 0 15
f
f
0 567
(7)
0 51
(8)
where k1 is related to the square root of compressive strength, while wc is the plain concrete unit weight. Eq. 7 and Eq. 8 were applied to the test results under ambient conditions as shown in Fig. 12. It is shown that Eq. 7 and Eq. 8 underestimate the obtained experimental results of the current study by approximately 67% and 80%, respectively. Based on the same previous two approaches, k1 can be introduced within the limits of this study as follows, f
0 98
(9)
In the same manner, it is shown that the splitting tensile strength of SFRC exposed to high temperatures is better represented by the cubic root of the compressive strength √
as shown in Eq. 10. The test results of the current study showed that the behavior of
ki with temperature is noticeably affected by the steel fiber content. fct
(10)
3
√f c
16
Fig. 13 shows the relationship between k2 and temperature for different steel fiber contents. The experimental data are represented by polynomial curves with R2 = 0.99. In order to produce a consistent relationship, cylinders with 0.75% of steel fiber were modified (dashed line in Fig. 13). This modification is conservatively considered. It is clearly shown that k2 decreases consistently with increasing the temperature. Gradually, all curves approach the same target at 550 oC. It was found that the variation of k2 could be expressed for different steel fiber contents with the growth of temperature according to Eq. 11, k2(0.75 %) = (-17 × 10-5 × T2) + (0.061 × T) + 50.814
(11a)
k2 (1%) = (-19×10-5 ×T2) + (0.0409 × T) + 67.607
(11b)
k2 (1.25%) = (-22 × 10-5 ×T2) + (0.0693 × T) + 59.241
(11c)
4.1.5. Concrete modulus of elasticity The static modulus of elasticity was measured based on ASTM C469, where the compressive strengths of the 100×200 mm cylinders presented in section 4.1.2 were considered. The specimens were tested for modulus of elasticity using a compressometer with two Linear Variable Differential Transformers (LVDT). The experimental results are shown in Table 3. Moreover, the decay in the modulus of elasticity (Ec,(T)) of the plain concrete were compared to that proposed by Li and Purkiss [52] and Bastami et al. [49] as shown in Fig. 14. The comparison of the three curves shows that the rate of decay is higher in the current work for temperatures above 300 °C. Fig. 15 shows that the Ec,(T) decreases with the increase of temperature, the decay is more severe than the corresponding decay in compressive and tensile strengths, which agrees with results of tests conducted by Chen et al. [28]. The observation of Fig. 15 shows that at 150 °C the inclusion of 1.25 % of steel fiber enhanced the modulus of 17
elasticity slightly, while 1.0% of steel fiber decreased the decay compared to plain concrete specimens. On the other, there was no significant effect of low steel fiber content (0.75%) on the modulus of elasticity at the same temperature. After exposure to 300 °C and higher, it is obvious in the figure that all steel fiber contents exhibited close percentage decreases at each temperature. Thus, according to the results of the current research, it can be concluded that steel fiber inclusion has no significant effect on modulus of elasticity of concrete after exposure to temperatures higher than 150 oC.
4.2. Flat plate specimens 4.2.1. Ultimate slab strength Twenty slab specimens of 500×500 mm and with 60 mm thickness were exposed to different temperatures. All specimens were tested under monotonically increased concentric loading until failure, which was applied through a 55 mm square steel plate. Table 5 shows the test results of the flat plate specimens. The ultimate strength (Pu) was compared with different code equations as shown in Fig. 16. Except the Turkish standard TS500 [43], the predicted ultimate strength values of all other codes are too conservative with high safety factor reaching 2.4, whereas the safety factor introduced by all codes ranges from 0.65 to 0.75. The missing bars in Fig. 16 refer to the specimens were the compressive strength exceeded the upper limit adopted in that code. For example, √f
c
of specimen A3 is more than 8.3 MPa, so that, it does not satisfy the
conditions of both ACI318 [34] and CSA A23.3 [36]. Practically, Eq. 2 can be adopted to predict the punching shear capacity of slabs without transvers shear reinforcement. Since the specimens in the present work are of the same dimensions and reinforcement ratio, therefore the characterized parameter in Eq. 2 is the function of compressive strength τc. Eurocode 2 has adopted the cubic root of 18
compressive strength (f'c)1/3. The effect of high temperatures on the specimens’ ultimate loads in term of steel fiber dosage and temperature Pu(Vf, T) are drawn in Fig. 17a. Using the results of the ultimate failure load of the slab specimens and their corresponding compressive strength results, second-degree polynomial regressions in terms of the target temperature were conducted as presented in Eq. 12 Pu(0%, T) = (f 'c)1/3 × [(- 11 × 10-5 × T2) + (0.057 × T) + 19.6]
(12a)
Pu(0.75%, T) = (f 'c)1/3 × [(- 6.9 × 10-5 ×T2) + (0.043 × T) + 21]
(12b)
Pu(1%, T) = (f 'c)1/3 × [(- 9 × 10-5 ×T2) + (0.049 × T) + 17.8]
(12c)
Pu(1.25%, T) = (f 'c)1/3 × [(- 5 × 10-5 ×T2) + (0.0276 × T) + 23.5]
(12d)
From Fig. 17a, it is clear that the exposure to 300 °C increased the ultimate slab resistance for all steel fiber contents. Moreover, the specimens with Vf = 0.75 % and 1.25 % showed higher residual strengths compared to the plain concrete specimens and those reinforced with 1 % steel fibers. This result is consistent with the cylinder tests results illustrated in Fig. 8a and Fig. 10. However, Fig.18b, shows that the rate of variation in the ultimate strength with respect to temperature (
for the
specimens of plain concrete and the specimens with 1% steel fiber was higher than for other specimens, which affected the results of Eq. 12. Therefore, Eq. 12 can be modified by including the effect of the (
, which can be done by equating Eq. 12 with Eq. 13
that was obtained from the regression of the data in Fig. 17b. Pu(0%, T) = Pu(0%, 20) × [(- 5 × 10-6 × T2) + (0.0026 × T) + 0.93]
(13a)
Pu(0.75%, T) = Pu(0.75%, 20)× [(- 1.3 × 10-6× T2) + (0.001 × T) + 0.98]
(13b)
Pu(1%, T) = Pu(1%, 20) × [(- 5 × 10-6 × T2) + (0.0027 × T) + 0.98]
(13c)
Pu(1.25%, T) = Pu(1.25%, 20) × [(- 2 × 10-6× T2) + (0.0011 × T) + 0.98]
(13d)
19
where the term Pu(Vf %, 20) is the ultimate punching strength, which can be substituted by Eq. 9 for specimens of plain concrete, and by Eq. 11 for SFRC. Many factors affect the improvement of the ultimate resistance of the flat-plate column connection that exposed to high temperature based on the crack pattern (discussed later in section 4.2.3). In addition to the factors discussed previously, the mechanical behavior of the flat plate plays the major role. Improving the stress distribution (increasing the fine particles) leads to the dispersing of the stress away from the compression strut that transfers the load to the supports. This is obvious by the multi-cracks shown in the tension face of the slabs. Many researches [11] assumed that the failure in flat plate column connection occurs either due to crushing, when the compression stress at internal fictitious column capital exceeds the compression capacity, and/or due to micro cracking of the soften concrete at the slab-column interface. The temperature gradient between the tension and compression zones makes the slab behaves as a composite structure of two materials with different strengths. This improves the peak cracking and the crashing resistance due to rearranging of the neutral axis gradually with temperature rising.
4.2.2. Load-deflection behavior Two LVDTs were attached to each specimen to record the deflection. The first LVDT was positioned at the mid-span of the slab, directly beneath the column, while the second LVDT was located at 110 mm from the center of slab (2d from the column face). The deflections obtained from the LVDTs on the aforementioned locations are symbolized by δ1 and δ2, respectively. Fig. 18 depicts the typical deflections for specimens exposed to 450 °C along the slab half span. The deflections at the center, δ1, due to the applied load for different steel fiber contents and under different temperatures 20
are presented in Fig. 19. It is shown that the steel fiber has significantly reduced the specimens’ deflection and increased the deflection rate
), which is fully agrees with
the findings of Swamy and Ali [21]. For all groups (all temperature levels), the specimens with 1.25% showed the highest
) values. This means that the enhancing
of the ductility by steel fibers continues under the effect of high temperatures and that the specimens with 1.25% steel fiber can sustain the highest deformations in addition to increasing the failure load. As shown in Fig, 19, the specimens with 0.75% and 1% showed almost the same behaviors under temperatures of 300 °C and 450 °C. It is worthy to mention here that specimen B1 (specimen without steel fiber) at 300 °C showed unexpected behavior after 50% of the failure load. This can be attributed to the significant spalling noticed for this specimen, knowing that the specimens in group B showed flexural failure.
4.2.3. Failure mode After the heat process is completed, there was no spalling and were no visible cracks except few small cracks shown for slabs exposed to 550 ºC. At the end of the mechanical test, the failed specimens were removed from the test machine and the type of failure was examined. The actual critical radius was measured and the type of failure was then visually classified. The observed failure types of the slab specimens are listed in Table 5. Red marker indicated the small cracks in some specimens as it is shown in Figs. 20 through 24. In spite that the slabs were designed to fail in punching, many types of failure could be recognized due to the effect of the steel fiber used and the level of applied temperatures. Fig. 20 shows the observed failure modes of the specimens of group E (under ambient conditions). In general, four out of five specimens of plain 21
concrete (in different temperature levels) failed in punching shear with elliptical failure surfaces and showed extensive spalling. This is because of the intersection of the flexural and punching cracks, which confirms the design purposes. The observed angles of shear cracks ranged between 17° and 22°. These angles increased slightly with increasing the temperature level. The same behavior was observed for E2 and E4 (0.75% and 1.25% steel fiber), but with incomplete punching behavior that allowed the propagation of flexural cracks to reach the edges of the slab. Pure punching failure was observed for all specimens of groups A and C, as shown in Figs. 21 and 23, respectively, with significant differences in hair cracks density. The higher temperature of group C increased the hair cracks and the spalling layer thickness. Different behaviors were observed for group B that exposed to 300 °C (Fig. 22). Pure flexural failure was recognized for specimens B1 of plain concrete and B4 of 1.25% steel fiber. Moreover, incomplete punching cracks, and small spalling thickness were associated with the flexural cracks of specimens B2 and B3 that include 0.75% and 1 % of steel fiber, respectively. In group D, where the specimens were exposed to 550 °C, excessive spalling was shown for the plain concrete specimen D1. The color of concrete in the exterior peels changed to light brown for more than 20 mm thickness, while wide parts along the thickness could be easily removed by hand. The inclusion of different dosages of steel fiber (in specimens D2, D3, and D4) prevented the deteriorated parts from falling down in spite of the high thermal and load stresses that led to such cracking. It can be concluded that the inclusion of steel fiber did not significantly affect the failure shape, which agrees with the conclusion of Shaaban and Gesund [22]. The inclination of cracks with the horizontal axis of the tested specimens under ambient conditions (group E) was 17° to 22°, which also agrees with findings of Gardner [53] and Shaaban and 22
Gesund [22]. However, the crack inclinations of the SFRC specimens with 0.75% steel fiber that exposed to high temperatures reached 37°.
5. Conclusions In present work, the effect of exposure to high temperatures on the punching shear resistance of steel fiber reinforced concrete was investigated using small slab specimens. The slab specimens were reinforced with moderate percentage of conventional reinforcing bars in order to fail in punching shear. Three different steel contents of 0.75, 1.0, and 1.25% were considered in addition to specimens without steel fiber. The slabs were exposed to different levels of temperatures of 150, 300, 450 and 550 oC. In addition to the slab specimens, cylinder specimens from the same mixes and exposed to the same conditions were prepared and tested. The cylinder specimens were used to evaluate the effect of high temperatures on the compressive strength, tensile strength, and modulus of elasticity of steel fiber reinforced concrete. Based on the test results of the current research, the following summarizes the most important conclusions:
1. Exposure to temperatures between 150 °C and 300 °C increases the SFRC mechanical properties, where the compressive strength was increased by 18%, while the tensile strength showed lower percentage of increase. For temperatures above 300 °C, all the mechanical properties were decreased but with lower rate compared to the plain concrete. The SFRC with 0.75 % showed the highest residual compressive and tensile strength values after exposure to 550 °C compared to the specimens with the other steel fiber contents. The residual compressive and tensile strengths after exposure to 550 °C for SFRC with 0.75 % were 77% and 74%. 23
2. Polynomial regression formulas were derived from the experimental results to represent the decay of the tensile strength of concrete reinforced with steel fibers after exposure to high temperatures. These polynomial formulas are dependent of the cubic root of compressive strength. On the other hand, for plain concrete, the regression formulas were represented using the square root of compressive strength instead of the cubic root. 3. Most of the available codes of practice (ACI318, AS3600, CSA A23.3, IS456,
Eurocode 2 and DIN 1045-1) underestimate the punching shear strength for concentric SFRC flat plate-column connections. The punching shear strength values estimated from these codes for the experimental slabs of this study were in general less than 41% of the obtained experimental strengths. On the other hand, the TS500, which adopts the concrete tensile strength to evaluate the punching shear strength, overestimates the strength of some of the current experimental slabs. 4. The incorporation of steel fibers enhanced the behavior of the tested flat plates
both before and after exposure to high temperatures. The use of steel fibers increased the ultimate failure load, decreased the associated deflection, and improved the capability of the slabs to withstand more sever cracking. However, there is no evidence of significant effect of steel fibers on the failure type of the tested specimens.
24
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[14] E. Annerel, L. Lu, L. Taerwe, Punching shear tests on flat concrete slabs exposed to fire, Fire Saf. J. 57 (2013) pp. 83-95. [15] J.S. Liao, F.P. Cheng, C.C. Chen, Fire Resistance of Concrete Slabs in Punching Shear, ASCE, J. Struct. Eng. ISSN 0733-9445/04013025(9) (2013) 9p. [16] H.K.M. Smith, T.J. Stratford , L.A. Bisby, Punching shear of reinforced concrete slabs under fire conditions: experiment vs. design, in the First International Conference on Struct. Saf. Fire & Blast. CONFAB Univ. Edinburgh, (2015). [17] M. Ghoreishi, A. Bagchi, M.A. Sultan, Punching shear behavior of concrete flat slabs in elevated temperature and fire, Adv. Struct. Eng. J. 18 (5) (2015) pp. 659674. [18] J.B. De Hanai, K.M.A. Holanda, Similarities between punching and shear strength of steel fiber reinforced concrete (SFRC) slabs and beams, Struct. Mater. J. 1 (1) (2008) pp. 1-16. [19] N. Yermak , P. Pliya , A. L. Beaucour, A. Simon, A. Noumowé, Influence of steel and/or polypropylene fibres on the behaviour of concrete at high temperature: spalling, transfer and mechanical properties, Constr. Build Mater. 132 (2017) pp. 240-250. [20] J. Michels, D. Waldmann, S. Maas, A. Zürbes, Steel fibers as only reinforcement for flat slab construction – experimental investigation and design, Constr.Build Mater. 26 (2012) pp. 145-155. [21] R.N. Swamy, S.A.R. Ali, Punching shear behavior of reinforced slab-column connections made with steel fiber concrete, ACI J. (1982) pp. 392-406. [22] A.M. Shaaban, H. Gesund, Splitting tensile strength of steel fiber reinforced concrete cylinders consolidated by rodding or vibrating, ACI Mater. J. 90 (4) (1993) pp. 366-369. [23] R. Serrano, A. Cobo, M.I. Prieto, M.N. González, Analysis of fire resistance of concrete with polypropylene or steel fibers, Constr. Build Mater. 122 (2016) pp. 302-309. [24] A. Lau, M. Anson, Effect of High Temperatures on High Performance Steel Fibre Reinforced Concrete, Cem. Conc. Res. 36 (2006) pp. 1698–1707. [25] O Düğenci, T Haktanir, Experimental research for the effect of high temperature on the mechanical properties of steel fiber-reinforced concrete, Constr. Build Mater. 75 (2015) pp. 82-88. 26
[26] R. Felicetti, P.G. Gambarova, Effects of high temperature on the- residual compressive strength of high-strength siliceous concretes, ACI Mater. J. 95 (4) (1998). [27] B. Zhang, N. Bicanic, C.J. Pearce, G. Balabanic, J.A. Purkiss, Discussion: Residual Fracture Properties of Normal- and High-Strength Concrete Subject to Elevated Temperatures, Mag. Concr. Res. 52 (2) (2000) 123-136. [28] G.M. Chen, Y.H. He, H. Yang, J.F. Chen, Y.C. Guo, Compressive behavior of steel fiber reinforced recycled aggregate concrete after exposure to elevated temperatures, Constr. Build. Mater. 71 (2014) pp. 1-15. [29] C.S. Poon, Z.H. Shui, L. Lam, Compressive Behavior of Fiber Reinforced HighPerformance Concrete Subjected to Elevated Temperatures, Cem. Conc. Research. 34 (2004), pp. 2215–2222. [30] J.R. Lawson, T. Phanl, F. Davis, Mechanical Properties of High Performance Concrete after Exposure to Elevated Temperatures, National Institute of Standard and Technology, Gaithersburg, NISTIR 6475, 2000. [31] P. Bamonte, R. Felicetti, Fire scenario and structural behaviour of underground parking lots exposed to fire, in Proceeding of International Conferance Applications of structural Fire Engineering, session 2 Fire Modelling, Prague, 2009, pp. 60-65. [32] Y. Li, M. J. Spearpoint, Analysis of vehicle fire statistics in New Zealand parking buildings, Fire Tech, 43 (2)(2007), pp. 93-106. [33] S.D.B Alexander, H. Simmonds, Shear-moment transfer in column-slab connections, Civil Eng. Univ.Alberta, SEI-Rep 141 (1986) 95p. [34] ACI318-14, Building code requirements for structural concrete, American Concrete Institute, ACI committee 318; 2014. [35] Australian Standards, Concrete structures, Sydney, Australia, (AS3600), 2009. [36] Canadian standards association, Design of concrete structures, Canada, (CSA A23.3), 2004. [37] Bureau of Indian standards, Plain and reinforced code of practice, New Delhi, India, (IS456-Fourth Revision), 2000. [38] German Standards, Concrete, reinforced and prestressed concrete structures- Part 1: Design and construction, Deutsches Institut für Normung, Berlin, Germany, (DIN1045-1), 2008.
27
[39] European Committee for Standardization. Eurocode 2: Design of concrete structures - Part 1-1: General rules and rules for buildings, (EN1992-1-1), Brussels, 2004. [40] Widianto, O. Bayrak , J.O. Jirsa , Two-way shear strength of slab-column connections: reexamination of ACI 318 provisions, ACI Struct. J. 106 (2), (2009) pp. 160-170. [41] S. Guandalini, O.L. Burdet, A. Muttoni, Punching tests of slabs with low reinforcement ratios, ACI Struct. J. 106 (1) (2009) pp. 87-95. [42] C.E. Ospina, N.M. Hawkins, Addressing punching failure, Struct. Mag. (2013) pp. 14-17. [43] TS500, Requirements for design and construction of reinforced concrete structures. Ankara, TUBITAK, 2000 (in English). [44] H. Gesund, Y. P. Kaushik, Analysis of punching shear failures in slabs, International Association for Bridge and Structu. Eng. J. 30-I (1970) pp. 41-60. [45] J.A. Purkiss, Steel fibre reinforced concrete at elevated temperature, Int. J. Cem. Com. Lightweight Conc. 6 (3) (1984) pp. 170-184. [46] Q. Zhang, G. Ye, Dehydration kinetics of portland cement paste at high temperature, J Therm Anal Calorim, 110 (2012) pp. 153–158. [47] ACI216R-89, Guide for determining the fire endurance of concrete elements, American Concrete Institute, ACI committee 216; 1989 Reapproved 1994. [48] European Committee for Standardization. Eurocode 2: Design of concrete structures- part 1-2: General rules—structural fire design, (EN1992-1-2), Brussels, 2004. [49] M. Bastami, F. Aslani, M.E. Omaran, High-temperature mechanical properties of concrete, Int. J. Civil Eng. 8 (4) (2010) pp. 337-351. [50] F.I. Faiyadh, M.A. Al-Ausi, Effect of elevated temperature on splitting tensile strength of fibre concrete, Int. J. Cem. Compos. and Lwt. Conc., 11 (3), (1989), pp. 175-178. [51] SNIP2.03.01-84, Concrete and reinforced concrete structures, SNIP 2.03.01-84, 1997 Edition. [52] L.Y. Li, J. Purkiss, Stress–strain constitutive equations of concrete material at elevated temperatures, Fire Saf. J. 40 (7) (2005) pp. 669-686.
28
[53] N.J. Gardner, Relationship of the punching shear capacity of reinforced concrete slabs with concrete strength, ACI Struct. J., 87 (1) (1990) pp. 66-71.
29
Table 1 Concrete mix slump Mix Steel fiber W/C W C G S S.fm SP 3 3 3 3 3 3 3 code Vf % kg/m kg/m kg/m kg/m kg/m kg/m kg/m mm 216 97 S0 0.43 465 680 1170 35 6.6 0.43 216 68 S0.75 0.75 58.9 465 680 1170 35 6.6 0.43 216 62 S1 1.00 78.5 465 680 1170 35 6.6 0.43 216 58 S1.25 1.25 98.2 465 680 1170 35 6.6 W; water, C; cement, G; crashing stone, S; river sand, S.fm; silica fume, WR; high water reducer
Table 2 Steel fibers characteristic Type Geometry
KMX 55/30 BG
Crimpedends
Density kg/m3 7850
Table 3 Specimens’ mechanical properties Steel Age at Spec. Max.Temp. fiber day of No. (°C) Vf (%) test A1 0 46 A2 0.75 45 150 A3 1 48 A4 1.25 44 B1 0 46 B2 0.75 43 300 B3 1 47 B4 1.25 44 C1 0 32 C2 0.75 33 450 C3 1 32 C4 1.25 33 D1 0 48 D2 0.75 45 550 D3 1 47 D4 1.25 43 E1 0 34 E2 Ambient 0.75 37 E3 conditions 1 34 E4 1.25 37
lf mm 30
Df mm 0.55
f 'c (MPa)
fct (MPa)
63 60 71 65 51 55 64 61 34 46 47 49 24 39 32 30 53 51 60 59
6.2 9.2 8.6 8.0 5.3 6.6 8.8 8.3 4.0 6.9 6.5 8.0 2.9 5.0 5.5 4.5 7.1 6.8 11.5 7.0
30
lf/Df 55
Moisture content (%) 0.777 1.353 0.915 1.472 1.476 1.688 0.745 1.037 0.929 1.039 1.104 1.008 0.738 0.743 0.949 1.246 _ _ _ _
Tensile strength MPa 1500
Ec (GPa) 33.21 31.22 34.57 38.37 23.82 20.93 21.50 23.78 10.73 9.64 9.98 10.53 5.39 5.86 4.32 4.07 38.13 37.27 37.32 37.41
Table 4 Designed slab parameters, flexural capacity, and punching shear strength Spec.
E1 (1) (2) (3)
d
f 'c
fy
ρ
(mm)
(MPa)
(MPa)
(%)
Q(1)
Pflex(2) (kN)
Vc (3)(kN) ACI
EC2
DIN
TS
39.7 65 400 1.1 2.66 65.2 40.3 42 40 61.9 failure mode criterion according to Eq. 4 section flexural capacity according to Eq. 3 concrete punching shear capacity according to ACI318 [34], EC2 [39], DIN1045-1 [38], and TS500 [43].
Table 5 Experimental characteristics of the slab specimens Spec. Pu δ1(1) Failure(2) Punching u u (T) type shape flex u (20)
r0, test(3)
Angle of failure (deg.) 17-18 20-27 21 19 19-22 27-29 21 19-21 22 27 20 19 17-22 18 30-33 18
(kN) (mm) (mm) A1 91.00 4.0 1.39 1.14 P Ellipse 218 A2 98.95 3.8 1.52 1.26 P circle 170 A3 95.80 3.5 1.46 1.34 P circle 180 A4 82.86 2.8 1.27 0.89 P circle 195 B1 103.23 3.8 1.59 1.30 F B2 92.45 4.0 1.42 1.17 F B3 93.66 4.3 1.43 1.31 P+F B4 106.88 4.8 1.64 1.15 F C1 88.50 5.8 1.40 1.11 P Ellipse 200 C2 92.48 5.9 1.44 1.17 P circle 160 1.15 C3 82.08 4.9 1.27 P circle 180 1.05 C4 98.06 5.7 1.52 P Ellipse 200 0.82 D1 64.99 4.7 1.06 P circle 180 1.13 D2 89.0 6.2 1.40 P square 180 D3 70.65 3.2 1.12 0.99 P circle 190 D4 93.10 4.8 1.49 1.00 P+F circle 200 E1 79.63 3.9 1.23 1.00 P circle 218 E2 78.80 3.7 1.22 1.00 P circle 188 E3 71.4 1.09 1.00 P+F square 131 E4 93.28 4.0 1.43 1.00 P circle 179 (1) δ1 is the displacement at the center corresponding to the ultimate load Pu, test (2) P refers to the punching shear failure and F refers to the flexural failure (3) radius of the critical punching crack taken as the longest size in ellipse and square
31
Passing particles (%)
100 River sand Crushed stone
80 60 40 20 0 12.5
9.5
4.75
2
1.18
0.6
Sieve opening (mm)
Fig. 1. Sieve analysis of the river sand and the crushed stone
Fig. 2. Steel fibers with hooked-ends
32
0.3
0.15 0.075
Fig. 3. Slab geometry and reinforcement details
33
(a)
(b)
(c) Fig. 4. (a) The used furnace, (b) typical stages of heating process, and (c) actual temperaturetime curves
34
(b)
(a)
(c) Fig. 5. (a) Circumferential load distribution, (b) rigid support of eight half-balls, and (c) test setup in the testing machine
35
Fig. 6. Expected yield line patterns (reproduced from Guandalini et al.) [41]
140 Present work
Slump (mm)
120
Dügenci et al., (2015)
100 80 60 40 20 0 0
0.5
1 Vf (%)
Fig. 7. Effect of steel fiber content on concrete slump
36
1.5
1.40 150 °C
f 'c (T)/f 'c (20)
1.20 20 °C
1.00 0.80 0.60 0.40 0.20 0
0.25
0.5
0.75
1
1.25
Steel fiber Vf (%) (a) 1.40
f 'c (T)/f 'c (20)
1.20 1.00 0.80 0.60
Vf = 0% Vf = 0.75 % Vf = 1% Vf = 1.25 %
0.40 0.20 0
100
200
300
400
500
600
Temperature (˚C)
(b) Fig. 8. Variation of the normalized concrete compressive strength due to (a) variation in the steel fiber volume fraction and (b) increasing the temperature
37
1.4
f 'c (T)/f 'c (20)
1.2 1 0.8 Present work Chen et al., (2014) Bastami et al., (2010) EN1992-1-2, (2004) ACI216-02, unstressed
0.6 0.4 0.2 0
100
200
300
400
500
600
Temperature (˚C) Fig. 9. Comparison of normalized compressive strength decay of plain concrete of the current study due to exposing to high temperatures with results from other researches [28, 47-49]
1.40
fct(T) /fct(20)
1.20 20 °C
1.00 0.80 0.60 0.40 0.20 0
0.25
0.5
0.75
1
1.25
Steel fiber Vf (%) Fig. 10. Effect of steel fiber content on tensile strength of standard cylinders exposed to high temperatures
38
1.4 EN1992-1-2, (2004) Faiyadh and AL-Ausi, (1989) Current work
fct(T)/fct(20)
1.2 1 0.8 0.6 0.4
Plain Concrete 0.2 0
100
200
300
400
500
600
500
600
Tempreture (˚C)
(a) 1.4 Faiyadh and AL-Ausi, (1989) Current work
fct(T)/fct(20)
1.2 1 0.8 0.6 0.4
1 % Steel Fibre 0.2 0
100
200
300
400
Tempreture (˚C)
(b) Fig. 11. Comparison of normalized splitting tensile strength decay of the current study due to exposing to high temperatures with two references [44, 46]: (a) for plain concrete and (b) for steel fiber content of 1 %
39
1.4
Current work Eq. 9 De Hanai and Holanda, (2008) Shaaban and Gesund, (1993)
fct /√f 'c
1.2 1 0.8 0.6 0.4 0.2 0
0.25
0.5
0.75
1
1.25
Steel fiber Vf (%) Fig. 12. k1 steel fiber content relationship: comparison between experimental results and proposed model of the current study with some proposed relationships found in literature
80 70 60 50
k2 40 30
Vf = 0.75 % Vf = 1% Vf = 1.25 % modified k(0.75%)
20 10 0 0
100
200
300
400
500
600
Tempreture (˚C) Fig. 13. Variation of k2 in Eq. 11 due to exposing to high temperatures for different contents of steel fiber
40
1.20 Current work Li and Purkiss, 2005 Bastami et al., 2010
Ec (T) / Ec (20)
1.00 0.80 0.60 0.40 0.20 0.00 0
200
400
600
800
1000
Tempreture (˚C) Fig. 14. Comparison between the experimental results of the plain concrete in current study with proposed decay models of the modulus of elasticity found in the literature [49, 52]
1.20
Ec (T) / Ec (20)
1.00 0.80 0.60 0.40
Vf = 0% Vf = 0.75 % Vf = 1 % Vf = 1.25 %
0.20 0.00 0
100
200
300
400
500
600
Tempreture (˚C) Fig. 15. Decay of the concrete modulus of elasticity due to exposing to high temperatures for different steel fiber contents
41
4.00 EC2
3.50
ACI
DIN
AS
CSA
TS
Pu, Test /Pu, Code
3.00 2.50 2.00 1.50 1.00 0.50 A1 A2 A3 A4 B1 B2 B3 B4 C1 C2 C3 C4 D1 D2 D3 D4 E1 E2 E3 E4
0.00
Specimens code Fig. 16. Experimental ultimate slab strength compared to different codes’ proposed equations 30
Pu,test /3√f 'c
25 20 15 10
Vf = 0 % Vf = 0.75 % Vf = 1 % Vf = 1.25 %
5 0 0
100
200 300 400 Temperature (°C) (a)
500
600
200
500
600
1.5
Pu,(T) /Pu,(20)
1.25 1 0.75 0.5
Vf = 0% Vf = 0.75% Vf = 1% Vf = 1.25%
0.25 0 0
100
300
400
Temperature (°C)
(b) 42
Fig. 17. Effect of high temperature on the ultimate loads of the SFRC concentric flat
plat-column connections: (a) normalized by cubic root of compressive strength and (b) normalized by the ultimate load of slabs at ambient condition
Distance from slab center (mm) 0
50
100
150
200
250
0
at T = 450 °C
Deflection (mm)
-1 -2 -3 -4
Vf = 0% Vf = 0.75% Vf = 1% Vf = 1.25%
-5 -6 -7
Fig. 18. Typical deflections along the slab width 120
Max. T = 20 °C
Load (kN)
100 80 60 40 Vf = 0% Vf = 0.75% Vf = 1.25%
20 0 0
1
2
3
Deflection (mm)
(a)
43
4
5
120
Max. T = 150 °C
Load (kN)
100 80 60 40
Vf Vf Vf Vf
20
= 0% = 0.75% = 1% = 1.25%
0 0
1
2
3
4
5
Deflection (mm)
(b) 120
Max. T = 300
Load (kN)
100 80 60 40
Vf Vf Vf Vf
20
= 0% = 0.75% = 1% = 1.25%
0 0
1
2
3
4
5
6
Deflection (mm)
(c) 120
Max. T = 450 °C
Load (kN)
100 80 60 40
Vf Vf Vf Vf
20
= 0% = 0.75% = 1% = 1.25%
0 0
1
2
3
4
Deflection (mm)
(d) 44
5
6
7
120
Max. T = 550 °C
Load (kN)
100 80 60 40
Vf Vf Vf Vf
20
= 0% = 0.75% = 1% = 1.25%
0 0
1
2
3
4
5
Deflection (mm)
(e) Fig. 19. Load-deflections curves for all exposed temperatures
45
6
7
E1, Normal Concrete
E2, SFRC, Vf = 0.75%
E3, SFRC, Vf = 1%
E4, SFRC, Vf = 1.25%
Group E, 20 °C Fig. 20. Failure pattern of slabs in group E
A1, Normal Concrete
A2, SFRC, Vf = 0.75%
A3, SFRC, Vf = 1%
A4, SFRC, Vf = 1.25%
Group A, 150 °C Fig. 21. Failure pattern of slabs in group A
B1, Normal Concrete
B2, SFRC, Vf = 0.75%
B3, SFRC, Vf = 1%
Group B, 300 °C Fig. 22. Failure pattern of slabs in group B
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B4, SFRC, Vf = 1.25%
C1, Normal Concrete
C2, SFRC, Vf = 0.75%
C3, SFRC, Vf = 1%
C4, SFRC, Vf = 1.25%
Group C, 450 °C Fig. 23. Failure pattern of slabs in group C
D1, Normal Concrete
D2, SFRC, Vf = 0.75%
D3, SFRC, Vf = 1%
Group D, 550 °C Fig. 24. Failure pattern of slabs in group D
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D4, SFRC, Vf = 1.25%
