Effects of high temperature on residual punching strength of slab-column connections after cooling and enhanced post-punching load resistance

Engineering Structures 199 (2019) 109580

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Effects of high temperature on residual punching strength of slab-column connections after cooling and enhanced post-punching load resistance

T

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Chunyu Zhanga, Wenchen Mab, Xiang Liub, Ying Tianb, , Sarah L. Ortonc a

GCW, Inc., 1555 S. Rainbow Boulevard, Las Vegas, NV 89146, USA Department of Civil and Environmental Engineering and Construction, University of Nevada, Las Vegas, 4505 S. Maryland Parkway, Las Vegas, NV 89154, USA c Department of Civil and Environmental Engineering, University of Missouri – Columbia, Columbia, MO 65211, USA b

A R T I C LE I N FO

A B S T R A C T

Keywords: Flat plate Flat slab Fire High temperature Punching Post-punching

Large-scale experiments were conducted on six isolated slab-column connection specimens. Each specimen contained a 300 mm square center column and a 150 mm thick slab. The objectives of the experiments were to study the effects of fire-induced high temperatures on the residual punching shear strength of reinforced concrete flat-plate structures after cooling and to examine the effectiveness of a detailing approach for enhancing the post-punching load-carrying capacity. Ceramic fiber heating panels were used to apply high temperatures to slab shear-critical regions at the compressive face. The test results indicate that the high temperatures up to 800 °C did not seriously impact connection punching shear strength. Moreover, the use of crossties can effectively engage slab tensile reinforcement in resisting post-punching loads with a loading capacity close to or even greater than the punching failure load. Complementary to experiments, finite element simulations were conducted. The numerical simulations predicted the punching failure load of slab-column connections at room temperature with a good accuracy; however, the simulations underestimated the post-heating punching strength of the cooled slab-column connections by up to 11%.

1. Introduction Due to low construction cost and architectural versatility, reinforced concrete flat plate, consisting of a slab with uniform thickness supported directly on the columns, is a popular structural system for residential and office buildings as well as parking garages. The main structural concern regarding flat plates is slab punching failure near the columns due to highly concentrated bending moment and shear. Although a punching failure happens locally, it can trigger a chain reaction of failure over the entire floor, causing partial or even complete collapse of the building [1–3]. Uncontrolled fire threatens the structural safety of flat plates due to combined loading effects of flexure and shear on the risk of punching failure. On one hand, even if fire-induced elevated temperature causes little change in slab shear force transferred to the columns, it drastically increases slab bending moment near the columns because the thermal-induced slab rotational deformation is restrained by the columns [4,5]. On the other hand, the strength degradation of slab concrete and reinforcement may decrease slab twoway shear strength. In 2004, a flat-plate underground parking garage in Switzerland collapsed due to punching failure propagation after a fire lasted for 90 min, causing the death of seven firefighters [6,7].

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Past experimental studies focused mainly on the fire endurance of flat-plate systems [8] or slab-column connections [7,9,10] subjected to a constant gravity load; if no failure occurred, a few test specimens were tested with increased load in the heated condition until failure. The post-fire punching strength without cooling identified in this manner by Annerel et al. [7] was about 60% of the punching strength in room temperature. Such a strength reduction can be partially attributed to the loss of nearly 1/3 of slab depth after severe concrete spalling due to a high concrete moisture content of 3.5%. Additionally, according to ACI 216 [11] and Knaack et al. [12], the concrete strength after cooling from a high temperature is lower than that found in the hot condition. For instance, the compressive strength of calcareous concrete at 500 °C is about 80% of the initial strength at room temperature, whereas this ratio is reduced to approximately 40% after cooling [11]. However, little is known to date regarding the residual punching strength of slabcolumn connections after cooling from fire-induced high temperatures. Not only is post-fire capacity important, but post-punching capacity needs evaluation. Following the punching failure due to abnormal conditions, such as faulty design or construction error, the ability of a flat-plate structure to resist progressive collapse depends on the postpunching load-carrying capacity of slab-column connections. Compared

Corresponding author. E-mail address: [email protected] (Y. Tian).

https://doi.org/10.1016/j.engstruct.2019.109580 Received 21 June 2019; Received in revised form 30 July 2019; Accepted 22 August 2019 Available online 28 August 2019 0141-0296/ © 2019 Elsevier Ltd. All rights reserved.

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with the numerous experimental studies of the punching strength of flat plates, fewer data of the post-punching resistance of slab-column connections exist. Slab bottom reinforcement passing through the columns is crucial for developing a tensile membrane action and thus postpunching load resistance [13]. Immediately after the punching failure at a slab-column connection, the slab reinforcing bars crossing the shear cracks become inclined. The vertical force component in the slab bottom bars is the major source of post-punching gravity load resistance. If the post-punching load-carrying capacity is high, less load will be redistributed to the surrounding slab-column connections, thereby reducing the risk of punching failure propagation and progressive collapse. Accordingly, the slab bottom bars anchored into the columns are taken as structural integrity reinforcement in the ACI 318 [14] and CSA [15] design codes. The role of slab bottom reinforcement can be confirmed by the experiments conducted on multi-panel flatplate systems [16,17] and isolated slab-column specimens [18–22]. A mechanical model [23] and a finite element model [24] have been developed to simulate the post-punching load-deformation response of slab-column connections. Once punching failure occurs at a slab-column connection, the slab top (tensile) bars designed to resist negative bending moment progressively break the concrete cover outside the punching cone (Fig. 1a), making the bars ineffective to carry vertical loads. Moreover, the slab bottom (compressive) bars crossing the inclined shear crack apply large vertical stresses to the bearing concrete and thus can damage slab concrete at the lower toe [23] and column concrete at the upper toe [25]. Bending up a portion of slab bottom reinforcement into the columns can enhance both punching and post-punching strength [26,27]. However, in either case of using straight or bent-up slab bottom bars, the post-punching loading capacity not only depends on the amount and strength of slab compression bars, but also is limited by the bearing strength of concrete surrounding the compressive bars. It follows that, if the slab bottom face has been exposed to fire, the concrete strength degradation due to high temperatures would lead to a reduced postpunching strength of slab-column connections. This paper presents the results of a series of large-scale experiments conducted to (1) explore the feasibility of a simple detailing approach to enhance the post-punching loading capacity of slab-column connections with and without fire damage, and (2) identify the residual punching shear strength of slab-column connections after cooling from fire-induced high temperatures. The proposed detailing approach intends to engage the slab tensile bars passing the column in resisting post-punching loads. As shown in Fig. 1b, a number of crossties are added to the slab outside the punching failure cone. The upper end of each crosstie hooks with a slab tension bar passing through the column to restrain its splitting from slab concrete, while the lower side is tied with slab compression bars. The concrete between slab tension and compression bars provide anchorage for the crossties. Following a

Fig. 2. Specimen dimension and reinforcement layout.

connection punching failure, the ripping of slab tensile bars is arrested by the crossties so that a large vertical force component can be developed in the slab tension bars to carry post-punching loads.

Fig. 1. Post-punching load-resisting mechanism for slab-column connections provided by (a) slab compression reinforcement only, and (b) both slab tension and compression reinforcement.

2. Specimen properties Six 3/4-scale specimens simulating the interior slab-column 2

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Fig. 3. Crossties installed with slab longitudinal reinforcement.

900 No. 3

600

100

No. 4

500 400

80 60

No. 3 bars: fy = 520 MPa fu = 682 MPa

300 200

40

No. 4 bars: fy = 494 MPa fu = 663 MPa

100

Stress (ksi)

700

Srtess (MPa)

an average slab tensile reinforcement ratio of ρ = 0.65% and 1.12%, respectively; however, the slab compressive (top) reinforcement had identical layout. Each group had a control specimen (A-1 and B-1) without experiencing high temperature and deploying crossties. All other specimens (A-2, A-3, B-2, and B-3) were reinforced with crossties. Specimens A-2 and B-2 were subjected to high temperatures. All the specimens were intended to simulate slab-column connections without shear reinforcement. Therefore, the cross-ties in Specimen A-2, A-3, B2, and B-3 were designed to only affect the post-punching loading capacity by situating them outside the potential shear crack region in the slab. As shown in Fig. 2, two slab bottom bars in Group A specimens and three bars in Group B specimens passed through the column in both directions. These bars were installed with crossties at each side of the column. The distance from the crossties to column face was chosen as 203 mm and 356 mm for Group A and B specimens, respectively. Concrete moisture content affects both thermal conductivity and the likelihood of spalling. According to EC2 [28], spalling is not a concern if moisture content is less than 3%. Moreover, the typical concrete moisture content of a floor slab is about 2% [29]. The specimens tested by Liao et al. [10] using normal strength concrete did not show spalling because the tests were not conducted until the specimens reached an age of at least 250 days, which allowed concrete moisture content to be reduced significantly. To achieve the typical concrete moisture content, the specimens tested in this study were stored in the laboratory, which was air-conditioned and had a low humidity, for nearly two years. During this period, the slab moisture content was continuously monitored using a meter measuring moisture based on electrical impedance. The moisture content was reduced to 1.0% for Specimen A-2 and 1.8% for Specimen B-2 by the time of testing. The concrete cylinder compressive strength in room temperature was measured as 39.0 MPa for the Group A specimens and 41.2 MPa for the Group B specimens at time of testing.

120

800

20

0

0 0

0.04

0.08

0.12

0.16

0.2

Strain Fig. 4. Stress-strain relationship of slab reinforcement.

connections of a prototype flat-plate building were tested under quasistatic concentric loading. Fig. 2 shows specimen dimension and reinforcement layout. The specimens were constructed using normal strength, normal weight concrete and tested upside down relative to the actual position in a building. Each specimen consisted of a slab and a center column stub. The slab was 152 mm (6 in.) thick and 1880 mm (74 in.) square. The column stub had a 305 mm (12 in.) square cross section, extending 152 mm (6 in.) beyond slab lower face and 140 mm (5.5 in.) beyond the upper face. Fig. 3 shows the crossties installed with slab reinforcing bars prior to concrete casting. The center column was reinforced by four No. 7 longitudinal bars (22.2 mm diameter) and five closed hoops. No. 3 bars (9.53 mm diameter) with a yield strength of fy = 520 MPa were used for slab top reinforcement, crossties, and column transverse reinforcement. No. 4 bars (12.7 mm diameter) with fy = 494 MPa were used for slab bottom reinforcement. As shown in Fig. 4, the slab reinforcing bars had clearly defined yielding in the measured stress-strain relationship. The thickness of concrete clear cover was 12.7 mm (0.5 in.) for the slab and 25.4 mm (1 in.) for the column. As shown in Table 1, the test variables included slab tensile (bottom) reinforcement ratio, temperature applied to slab top face near the column, and the use of crossties. According to reinforcement ratio, the tests were divided into two groups, A and B, each consisting of three specimens numbered from one to three. The two groups of specimens had identical concrete mix design and were cast two months apart. Calcareous (limestone) aggregates, having a maximum size of 10 mm, were used for the concrete mix. The specimens in Groups A and B had

3. Test setup and instrumentation Fig. 5 shows the setup for testing punching and post-punching strengths of slab-column connections. The specimen was simply supported by eight steel columns made of W-flange sections and evenly distributed at a radius of 962 mm from the center of column stub (Fig. 5a). The distance between steel column center and slab edge was 51 mm and the distance between the centers of two steel columns along a slab edge was 736 mm. A 203-mm long, 25.4-mm wide, and 12.7-mm thick steel strip was placed between each steel column and the slab. The steel strips were oiled at the top and bottom faces to decrease friction 3

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Table 1 Specimen properties. Specimen

Tensile reinforcement ratio (%)

39.0 39.0 39.0 41.2 41.2 41.2

0.65 0.65 0.65 1.12 1.12 1.12

Maximum temperature (°C)

Room 800 °C Room Room 800 °C Room

Use of crossties Quantity

Distance to column (mm)

– 8 8 – 12 12

– 203 203 – 356 356

Punching failure load (kN)

Post-punching strength (kN)

314 308 293 495 459 468

200 243 333 250 347 411

Support

Support

A-1 A-2 A-3 B-1 B-2 B-3

Concrete compression strength (MPa)

Heating Area

51 mm

572 mm

736 mm

572 mm

(a)

Fig. 6. Test setup for heating slabs: (a) position of ceramic fiber heating panels; (b) ceramic fiber blankets for insulation.

deformation was determined. Specimens A-2 and B-2 were exposed to high temperatures prior to structural loading. As shown in Fig. 6a, the thermal loading was applied by ceramic fiber heating panels supported by 10-mm high ceramic spacers on the slab top face. Each heater was 305-mm square (same size as the center column stub) and had an operating temperature up to 1200 °C. Totally 8 heaters were used to cover slab shear-critical regions surrounding the column, as shown in Fig. 5a. Other portions of the slab were not heated due to the limitation of electricity supply; however, the ratio of heated area to slab thickness was comparable to that considered by Liao et al. [10] using electric furnace to test the fire endurance of slab-column connections. During thermal loading, the heating panels were covered with 51mm thick ceramic fiber blankets to ensure insulation (Fig. 6b). Slab temperatures were measured at a distance of 150 mm from the column face by K-type thermocouples. Vertical holes with a diameter identical

(b) Fig. 5. Setup for structural loading tests: (a) plan view and (b) 3D view.

and thus permit slab in-plane deformation. Load was applied downward on the upper column stub by a hydraulic jack bearing against a reaction frame. The hydraulic jack had a loading capacity of 136 tons and a maximum stoke of 254 mm. The applied load was measured by a load cell placed between the hydraulic jack and the upper column stub. Linear variable differential transformers (LVDTs) were placed underneath a specimen to measure slab vertical displacements at the center and quarter span. Additional LVDTs measured slab horizontal displacements at two opposite edges, from which the slab in-plane 4

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900

900 Top

A-2

800

700

6 mm from top

Temperature (ÛC)

Temperature (ÛC)

700 600 500 400

Mid-depth

300

Top

B-2

800

200

600 500 400 Mid-depth

300 200

Bottom

100

Bottom

100

0

0 0

1

2

3

4

5

6

7

0

1

Time (hour)

2

3

4

5

6

7

Time (hour) Fig. 7. Slab temperature history.

connection specimens: three were heated to a target furnace temperature of 800 °C in unstressed condition and another three to 600 °C. The cylinder specimens, with a 102-mm diameter and 203-mm height, were prepared from the same batch of concrete used for the slab-column connection specimens. In addition to furnace temperature, temperature inside each concrete cylinder was measured at center and quarter diameter by thermocouples embedded to the mid-height. Because of the relatively high heating rate (up to 8 °C/minute), the temperature was not uniformly distributed in a cylinder. To reduce temperature gradient in the cylinders, the electric power was turned off several times during heating. Once the target temperature was reached, the power was completely turned off and the cylinders were allowed to further increase temperature for about 20 min until the inner temperatures became stabilized. Fig. 10a and b show the temperature history of concrete cylinders (Group B specimens) at different locations with a target furnace temperature of 600 °C and 800 °C, respectively. The temperature was the average value obtained from a group of three cylinders put together into the furnace. After the cylinders were naturally cooled, compression tests were conducted. Fig. 10c plots the normalized residual compressive strength (the ratio of residual strength after cooling to that in room temperature) as a function of the average temperature evaluated based on the maximum temperatures the cylinders experienced at the surface, quarter-diameter, and center. Because of the relatively high heating rate (up to 8 °C/minute), the temperature was not uniformly distributed in a cylinder. However, the extent of strength loss based on average value was comparable to the data available in literature. For instance, the tests performed by Lee et al. [30] with a heating rate of 5 °C/minute indicated that the exposure of normal strength concrete to 600 °C and 800 °C resulted in a normalized residual strength of 50% and 15%, respectively.

to that of the thermocouples were drilled into the slabs so that the thermocouples were inserted from slab bottom to measure concrete temperature at 6 mm below the top (heated) face of Specimen A-2 and slab mid-depth of both Specimens A-2 and B-2. The holes were sealed by cement paste. Temperature was also measured at the slab top and bottom faces by thermocouples. LVDTs measured slab upward displacement at the center column stub and in-plane expansion. 4. Test results of thermal loading 4.1. Temperature history and slab damage Fig. 7 shows the measured temperature history at different slab depths for Specimens A-2 and B-2. Likely due to higher moisture content, temperature increased at a lower rate in Specimen B-2 than in Specimen A-2. It took 1.1 h and 1.3 h to reach 500 °C at the slab top face in A-2 and B-2, respectively. The rate of temperature increase was significantly reduced in A-2 after 700 °C. As a result, the thermal loading was terminated at 800 °C after 5.15 h’s heating despite an initial target of 1000 °C. For consistency, the heating for Specimen B-2 was also stopped at 800 °C after 6.25 h’s heating. Because the heating caused slab to camber upward, the vertical position of the thermocouples measuring slab temperatures was accordingly adjusted at time t = 3.6 h for Specimen A-2. This led to a sharp increase in temperature measured at 6 mm below slab top surface, where temperature reached 780 °C by the end of heading. The position of thermocouples for Specimens B-2 was adjusted more frequently. For both specimens, the temperature at slab mid-depth stayed at 100 °C for about 30 min, indicating the effects of water evaporation on temperature. Moreover, the evaporated free water in concrete penetrated the slab and condensed at the bottom face. After heating, the specimens were naturally cooled and the slab cracks developed mainly along the slab top reinforcement during heating were marked. Fig. 8 shows the damage condition of Specimen A-2 after the heating panels were removed (ceramic spacers are shown in Fig. 8a). The concrete color was changed to white and pink in the heated region. Some scaling of the cement paste occurred, but no severe spalling of concrete cover was observed.

5. Results of punching and post-punching tests 5.1. Damage pattern The specimens failed in punching failure at a center deflection of δ = 10 to 23 mm. At each side of the column, the punching cone always intersected slab tension face along a bottom bar. The distance from the punching perimeter to column face was 76 mm for Group A specimens (ρ = 0.65%) and 114 mm for Group B specimens (ρ = 1.12%). Given that the slab effective height was 127 mm, the angle of inclined shear cracks relative to slab plane was greater than 45°. Even if the Specimens A-2 and B-2 were exposed to high temperatures at the top face, the temperature at slab bottom face never exceeded 85 °C, causing

4.2. Residual concrete strength after cooling To examine the residual strength of concrete cylinders after experiencing high temperatures, an electric furnace was assembled using six pieces of ceramic heating panels (Fig. 9) and covered with ceramic fiber blankets during heating. For each group of the specimens, totally six concrete cylinders were tested at the same age of the slab-column 5

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(a)

(b)

Fig. 8. Damage due to elevated temperature on slab concrete (Specimen A-2): (a) slab top face after cooling, and (b) crack map.

Fig. 9. Furnace built by ceramic fiber heating panels.

Fig. 11. Post-punching connection damage in Specimen A-1 at center deflection of δ = 54 mm.

negligible effects on concrete mechanical properties. Accordingly, the location of slab punching perimeter and thus the orientation of inclined shear cracks in Specimens A-2 and B-2 were identical to the specimens without experiencing high temperatures. As shown in Fig. 11 for Specimen A-1, the slab bottom reinforcement broke the concrete cover and ripped out of the slab following the punching failure of the control specimens, a classical damage pattern of slab-column connections. In contrast, the post-punching behavior of all other specimens was altered due to the use of crossties to vertically

anchor slab tensile (bottom) bars. Fig. 12 shows the damage pattern of Specimen A-3 after being loaded beyond the punching failure that occurred at a center deflection of δ = 15 mm. At δ = 38 mm, the slab bottom bars were still able to break the concrete cover (Fig. 12a); however, starting from δ = 46 mm, the crossties located 203 mm away from column face arrested the separation of bottom reinforcement from the surrounding concrete. The damage condition of Specimen A-3 at δ = 98 mm is shown in Fig. 12b, which clearly indicates the slab bottom bars were able to develop a hanging mechanism to carry the vertical

(a)

Temperature (ÛC)

800

(b)

800

700

700

600

600

500

500

400

400

300

300 Furnace Quarter-diameter Center

200 100

Normialized Residual Strength (%)

900

900

Furnace Quarter-diameter Center

200 100 0

0 0

0.5

1

1.5

Time (hour)

2

2.5

0

0.5

1

1.5

Time (hour)

2

2.5

50

(c) 40 30 20 10 0 500

600

700

800

900

Temperature (ÛC)

Fig. 10. Testing residual concrete compressive strength: (a) temperature history of heating to 600 °C; (b) temperature history of heating to 800 °C; (c) normalized residual concrete strength. 6

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Fig. 12. Post-punching connection damage in Specimen A-3 at center deflection of (a) δ = 38 mm and (b) δ = 98 mm.

deflection curves. The concrete compressive strength in room temperature of Group B specimens was only 5.6% greater than that of Group A specimens; however, the average punching failure load was 308 kN and 481 kN for the unheated Group A specimens (ρ = 0.65%) and Group B specimens (ρ = 1.12%), respectively. The large difference in punching load indicated the strong influence of slab flexural strength dictated by tensile reinforcement ratio. As shown in Fig. 8, the previous heating in Specimens A-2 and B-2 caused extensive slab cracking thus they presented a lower initial stiffness during the structural loading; however, the post-cracking stiffness was similar to that of the unheated counterparts. More significantly, the effect of high temperature (up to 800 °C) on connection post-heating punching strength was not substantial. The punching failure load, Pu, of Specimen A-2 (308 kN) was only 1.9% lower than the maximum Pu of the unheated Specimens A-1 and A-3 (314 kN). However, Specimen A-2 also demonstrated much improved ductility, as shown by the longer general yielding plateau in the load-deflection response prior to punching failure. For Specimen B-2, the punching failure load (Pu = 458 kN) was 7.5% lower than maximum Pu of Specimens B-1 and B-3 (495 kN). As described previously, a 40% punching shear strength reduction was observed in the experiments conducted by Annerel et al. [7] when the specimens were loaded to failure in the hot condition. However, the degree of strength reduction for the specimens tested in the present study after cooling (A-2 and B-2) was much lower due to the effects of slab flexural behavior on punching resistance. Because slab concrete did not spall in Specimens A-2 and B-2 during the heating, they did not loss effective height for the slab bottom bars to resist bending moment. Moreover, slab flexural strength may not be seriously impacted by the decreased concrete compressive strength due to the prior heating. This simulation can be analogous to the experiments conducted in room temperatures. For instance, two specimens tested by Elstner and

load. Moreover, the slab bottom bars split the bearing concrete in the punching cone due to high local stress. Consequently, the distance from the bottom bars to the lower face of punching cone far exceeded the initial concrete cover thickness of 13 mm (Fig. 12b). Fig. 13 shows the post-punching damage pattern in Specimen B-3 at δ = 25 mm and 114 mm. Following the punching failure at δ = 10 mm, the concrete cover for slab bottom bars spalled to a higher extent than in Specimen A-3 (Fig. 13a) because the crossties were situated 242 mm away from the punching perimeter. When the specimen was displaced to δ = 114 mm, the two bottom bars located immediately outside the column were completed ripped out of the slab. However, the crossties at each side held the three slab bottom bars passing through the column, thereby allowing the slab-column connection to develop a tensile membrane action. As shown in Fig. 13b, during the post-punching loading, the slab bottom bars destroyed the concrete of punching cone outside the column, indicating the post-punching loading capacity is a function of local concrete strength. The punching and post-punching behaviors of Specimens A-2 and B-2 observed from slab bottom side were similar to those of the unheated Specimens A-3 and B-3. The slab bottom bars passing through the column and restrained by the crossties also successfully functioned as structural integrity reinforcement to carry the post-punching loads. No fracture of slab bottom bars was observed. However, after loading Specimen B-2, the slab top concrete cover was removed and one slab top bar bearing against another top bar oriented in the orthogonal direction was found to be fractured. 5.2. Load-deflection response Fig. 14 shows the load versus slab center deflection response of the six test specimens. The cracking load for the specimens without experiencing high temperatures (A-1, A-3, B-1, and B-3) was about 75 kN determined based on the significant change in the slope of load-

Fig. 13. Post-punching connection damage in Specimen B-3 at center deflection of (a) δ = 25 mm and (b) δ = 114 mm. 7

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Center Deflection (in.) 0

1

2

3

4

top bars or greater extent of splitting failure between the slab bottom bars and concrete.

5

600

where ∑ Asb = total area of slab bottom bars passing through the column, and fy = yield strength of slab bottom bars.

100 400

A-3

80

A-2

60

300 200

Load (kips)

Load (kN)

Pp = 0.5 ∑ Asb f y

120

500

Pp =

where ft = tensile strength of slab bottom bars, and ψt = 25° per Ruiz et al. [27]. Due to the use of crossties, the initial concrete cover spalling did not cause a significant load decrease in all other specimens even if the rate of load increase was reduced. When these specimens were loaded beyond δ = 50 mm, load drop occurred presumably due to concrete failure under high bearing stresses; however, the loading capacity was largely regained during the further loading. Compared with the control specimens, the post-punching loading capacity, Pp, was substantially enhanced. Specimens A-3 and B-3 reached Pp = 333 kN and 409 kN, both approximately 60% greater than the Pp of their counterparts (A-1 and B-1). It is noteworthy that the Pp of A-3 exceeded its punching failure load (Pu = 291 kN) by 14%. The Pp of B-3 was 0.87Pu; however, greater Pp can be anticipated if the crossties in B-3 were deployed at the same distance from column as in A-3. Immediately following a punching failure, the load is carried by both slab top and bottom bars. Because the slab bottom temperatures in Specimens A-2 and B-2 remained low during heating, the high temperatures applied to slab top face should not impact the post-punching load resisted by the slab bottom bars. Accordingly, the difference in post-punching load between the specimens with and without heating reflected the effect of elevated temperature on the strength of concrete bearing against the slab top bars. This effect was examined by the load immediately prior to the noticeable load drop marked by the solid dot in the load-deflection curve of a specimen in Fig. 14. The load defined in this way was 237 kN, 306 kN, 294 kN, and 358 kN for Specimens A-2, A-3, B-2, and B-3, respectively. It was estimated from these data that the 800 °C temperature applied to the slab top face decreased the postpunching load resistance contributed from the slab top reinforcement by 23% for Group A specimens and 18% for Group B specimens. Specimens A-2 and B-2 obtained a peak post-punching load of Pp = 243 kN and 356 kN, respectively. As shown in Fig. 14, the post-punching strength of each heated specimen was greater than that of the corresponding unheated control specimen without using crossties. The punching failure load and the post-punching strength are summarized in Table 1 for each specimen.

40

A-1

100

20

Group A 0

0

0

25

50

75

100

125

150

Center Deflection (mm) Center Deflection (in.) 0

1

2

3

4

5

600

120 100

B-3

400

80 300

B-2

60

B-1

200

Load (kips)

Load (kN)

500

40

100

20

Group B 0

0

0

25

50

75

100

125

∑ Asb ft sin ψt

150

Center Deflection (mm) Fig. 14. Load versus slab center deflection response.

Hognestad [31] were identically reinforced with a tensile reinforcement ratio of 0.50%. The punching failure load for the specimen with lower concrete strength (14.2 MPa) was only 11% lower than that with much higher concrete strength (42 MPa). Following punching failure, the load immediately dropped to an average of 147 kN and 160 kN in the Group A and Group B specimens, respectively. The specimens then regained loading capacity by developing tensile forces in the structural integrity reinforcement. However, the concrete cover of slab bottom bars spalled subsequently between δ = 33 mm and 45 mm. The concrete cover spalling resulted in load drops in the control specimens A-1 and B-1 and the load at spalling was unable to be recovered in the further loading. The post-punching loading capacity, Pp, was about 200 kN in A-1 and 250 kN in B-1 when the center deflection δ became greater than 50 mm. This is about 40–60% of the punching strength and similar to the post-punching capacity found in other isolated slab-column connection tests [21,22]. The post-punching strength of Specimens A-1 and B-1 was estimated using Eqs. (1) and (2) suggested for design purpose by CSA A23.3-14 [15] and MC 2010 [32], respectively. Each equation assumes only the slab bottom bars to be effective. The reversely positioned specimens had identical slab top reinforcement layout and thus the same predicted post-punching strength. According to the material properties of slab top reinforcement shown in Fig. 4, CSA A23.3-14 and MC 2010 predict a post-punching load resistance of Pp = 148 kN and 164 kN, respectively. The predictions are at least 20% lower than the measured Pp for A-1 and B-1 (200 kN and 250 kN). However, if the specimens were loaded further to larger deflections, the measured post-punching loads may be closer to the predictions because the post-punching load resistance may degrade due to more damage to the concrete bearing against the slab

5.3. Slab in-plane deformation response Fig. 15 shows the response of slab in-plane deformation of the specimens as a function of increasing vertical load up to slab punching failure. The in-plane deformation was determined from the slab horizontal displacements measured at the mid-depth of the centers of two opposite slab edges of a specimen. Positive and negative signs are given to slab expansion and shortening, respectively. As the load exceeded 100 kN, the slabs of the four unheated Specimens A-1, A-3, B-1, and B-3 consistently expanded until the punching failures occurred. This behavior and the amount of lateral expansion were comparable to other isolated slab-column connection tests [22]. The slab expansion was primarily initiated after the cracking load of 75 kN was reached and increased almost linearly with the load. Thus, the slab expansion was caused mainly by slab cracking. In addition, Specimens A-1 and A-3 with lower slab tensile reinforcement ratio experienced greater rate of expansion. The expansion indicates that the slab near the column in both specimens was subjected to compressive membrane action prior to punching failure. The heated specimens, A-2 and B-2, were initially shortened but the shortening started to reduce when the punching 8

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500

Load (kN)

400

boundary condition in the simulations. All unheated specimen surfaces, including slab bottom and column faces, were assigned with a heat transfer coefficient of 20 J/(m2·s·K) [35] to account for the heat dissipation from concrete at slab surface to the air due to convection. Fig. 16 compares the simulated and measured time-temperature responses for Specimens A-2 and B-2 at different locations along slab depth. The simulated temperatures were determined based on a linear interpolation of the nodal temperature output. In general, the simulation overestimated the temperature of slab concrete during the heating process. The predicted temperature at slab mid-depth is 18% and 23% greater than the measured temperature for A-2 and B-2, respectively. The relative difference is reduced above slab mid-depth where the concrete can be more critical for the shear strength of a slab-column connection. One possible reason for the temperature overestimation is that the concrete thermal properties defined in the modeling according to ASCE [34] does not account for the effects of water evaporation. Accordingly, the simulation cannot capture the 100 °C temperature plateau occurring at slab mid-depth. By the end of heating, the simulation overestimated the temperature by 50–59 °C at slab mid-depth and 30–31 °C at bottom face. Fig. 17 shows the predicted slab temperature profile in Specimen A2 when the slab top face in the heated region reaches 800 °C. It is seen that the high temperature applied on slab top face propagates predominantly along the vertical direction. The concrete of a quarter of slab depth (38 mm) and the concrete of a portion of column reaches a temperature above 500 °C, a temperature causing more than 50% reduction of the residual compressive strength for calcareous concrete according to ACI 216 [11]. Given that the concrete in these regions would bear against the slab top bars after punching failure, a significant decrease in the post-punching resistance carried by the slab top bars can also be expected based on the heat transfer analysis results.

B-2 B-3

B-1 A-1

A-2

300

A-3 200 100 0 -0.5

0

0.5

1

1.5

2

Slab Expansion (mm) Fig. 15. Load versus slab in-plane deformation response up to punching failure.

failure was approached. This dimension change indicates that, the slabcolumn connections with previous heating most likely were also under compressive membrane action, which helped to reduce the negative effects of concrete strength degradation due to high temperature on loading capacity. It is interesting to note that the plateau in the load versus in-plane deformation response of Specimen A-2 coincided with the ductile load-deflection response of this specimen shown in Fig. 14. The exact reason, however, is still unclear. 6. Finite element simulations 6.1. Heat transfer analyses Finite element simulations were performed using Abaqus [33] to simulate the behavior of specimens during the heating and punching tests. Concrete was simulated using 8-node thermally coupled brick elements with trilinear displacement and temperature. Heat transfer analyses were conducted first for Specimens A-2 and B-2. The slab mesh size for heat transfer analyses was 19 mm in the vertical direction so that eight elements were used along slab depth, and 25 mm in two horizontal directions. Because the presence of reinforcement has negligible impact on heat transfer characteristics of a RC slab [5] and the slab top reinforcement ratio was low, the reinforcement bars were omitted from the heat transfer analyses. Neglecting free water evaporation, a constant mass density of concrete was assumed. The specific heat and thermal conductivity for calcareous concrete used in the experiments of this study were defined based on ASCE [34]. The temperature history measured at the heated slab top face was applied as a

6.2. Simulations of punching tests Based on symmetry, one quarter of a specimen with boundary restraints simulating the actual support condition in the experiments was modeled for structural loading. Displacement-driven analyses using explicit algorithm were adopted. The element type of concrete and the mesh size were identical to those in the heat transfer analyses. The slab reinforcement was modeled using embedded truss elements assigned with a bilinear stress-strain relationship and a 1% strain hardening ratio. The crossties used in Specimens A-2, A-3, B-2, and B-3 were not simulated in the modeling because they were located outside the punching failure regions and thus should not impact the punching failure load. Concrete under triaxial state of compressive stresses was modeled using Concrete Damaged Plasticity [36,37]. Due to the lack of 900

900 Top

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100

0

0 0

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2

3

4

5

6

7

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Time (hour)

1

2

3

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Time (hour)

Fig. 16. Comparison between simulated and measured temperature history. 9

5

6

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Temperature (ÛC) Residual concrete strength / Initial strength 0.27 0.39 0.65 0.80 0.90 0.96 1.00 1.00

Fig. 17. Predicted temperature profile in Specimen B-2 when slab top face is heated to 800 °C. 400

45 24ÛC

40

Group A

217ÛC

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300

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Stress (MPa)

356ÛC 30 451ÛC

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0.015

0.02

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0

Strain

5

10

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Center Deflection (mm)

Fig. 18. Stress-strain relationship for Specimen A-2 concrete after cooling from different temperatures.

600

Group B 500

experimental data, the five parameters needed to construct this model were assumed as temperature-independent and defined according to the suggestions by Genikomsou and Polak [38]: (1) the dilation angle was defined as 40˚; (2) the flow potential eccentricity was defined as 0.1; (3) the ratio of the second stress invariant on the tensile meridian to that on the compressive meridian was defined as 0.667; (4) the viscosity parameter was taken as zero; (5) the ratio of initial equibiaxial compressive yield stress to initial uniaxial compressive yield stress was defined as 1.16. Prior to reaching 50% of concrete compressive strength fc’, the uniaxial compressive stress-strain (σ-ε) response of concrete in room temperature was assumed to be linear elastic with a Young’s modulus defined based on ACI 318 [14] for normal weight concrete. Beyond elastic limit, the uniaxial σ-ε was defined using Hognestad’s model [39]. The σ-ε response of concrete after cooling from different temperatures was defined based on the experimental data given by Lee et al. [30] in terms of the effects of temperature on the residual compressive strength, Young’s modulus, and the ultimate stress and strain. The temperatures inside the slab during the cooling phase were not measured in the experiments. After terminating the thermal loading, the slab inner temperature could still increase. However, it can be expected that the increase should not be significant due to the much lower heating rate and thus lower temperature gradient in the experiments than those caused by a standard fire. As shown in Fig. 10a, even if the 6 in. diameter concrete cylinders were heated at a greater rate than the slabs, the center temperature increased only 20 °C by the time when the temperature increase rate became nearly zero. Therefore, the effects of cooling were not considered, and the temperature profiles identified from the heat transfer analyses at the time of reaching 800 °C on the slab top face were used to estimate the peak temperatures that slab

Load (kN)

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200 100 0 0

5

10

15

20

25

Center Deflection (mm) Fig. 19. Comparison between simulated and measured load-center deflection response.

concrete may experience at different locations of the slab. Fig. 17 shows the ratios of residual compressive strength to the initial strength at room temperature for different layers of slab concrete near the column of Specimen B-2 based on predicted temperatures. Fig. 18 shows as an example the σ-ε responses defined for slab concrete in Specimen A-2 after cooling from different temperatures determined from the heat transfer analyses at six locations along slab depth. For simplicity, a bilinear tensile behavior with tension softening was employed for the concrete. The failure stress of concrete in tension, ft, representing the onset of micro-cracking, was defined as 0.03fc’. Beyond ft, the stress-strain curve softens to reflect the formation of micro-cracks and the tensile stress is reduced to zero stress at a strain of 0.0065 and 10

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Fig. 20. Cracking pattern in tension surface at the punching failure of unheated specimens: (a) Group A, and (b) Group B.

conclusions were reached:

0.0072 for concrete of Groups A and B specimens, respectively. The fracture energy of concrete was taken as temperature independent and defined based on Genikomsou and Polak’s suggestions [38]. Fig. 19 compares the simulated and measured load-deflection response of the test specimens. The numerical simulations overestimated the initial stiffness and the post-cracking load for Group A specimens tested in room temperature. However, such discrepancies were considerably reduced for Group B specimens having a higher slab tensile reinforcement ratio. In general, the predicted responses in room temperature were close to the test data as the punching failure load was approached. The predicted failure load was 6.5% and 5.1% greater than the average failure load of the two unheated specimens for Group A and Group B, respectively. According to the Concrete Damage Plasticity model [36,37], when the maximum principle plastic strain becomes positive, concrete cracking occurs. The cracking pattern of the unheated Groups A and B specimens can be visualized by the maximum principle strains shown in Fig. 20, in which the slabs are reversely placed to display the tension side. For the numerical simulations of heated specimens, focus is suggested to be given to the prediction of failure load only. This is because the effects of cracking during heating in the experiments on the initial stiffness were not considered in the modeling and thus the predicted stiffness was much higher than the test result. It is seen from Fig. 19 that the post-heating punching strength was underpredicted by 8.8% and 11% for Specimens A-2 and B-2, respectively. The simulations indicated the maximum stress of slab tensile bars in Specimens A-2 and B2 at punching failure is still 3% below the yield stress; however, all tensile bars cross the column and two bars outside each side of the column develop yielding in the unheated specimens at punching failure. In other words, the finite element simulations predicted premature punching failures. This can be explained partially by the overestimation of slab temperature (thus underestimated the slab concrete residual strength) and partially by modeling uncertainty caused by the assumed temperature-independency in some material properties, such as tension fracture energy and dilation angle that impacts the prediction of a punching failure.

1. The residual punching strength of slab-column connections after cooling from a high temperature up to 800 °C applied on the compression face of slab was 1.9–7.5% lower than the slab-column connections without experiencing high temperatures. 2. It was estimated from test data that the 800 °C temperature applied to the specimens decreased the post-punching load resistance contributed from the slab compressive reinforcement by 18–23%. 3. The use of crossties outside the punching cone can effectively engage slab tensile reinforcement in resisting post-punching loads, lead to a post-punching strength close to or even greater than the punching failure load. 4. Finite element simulation has the potential of accurately predicting the punching failure load of slab-column connections in room temperature. However, the numerical simulations predicted a premature punching failure and thus underestimated the residual punching strength of slab-column connections with previously applied high temperatures by 8.8–11%. Declaration of Competing Interest The authors declared that there is no conflict of interest. Acknowledgments The authors would like to thank for the Faculty Opportunity Award provided by the University of Nevada, Las Vegas that funded this research. The help provided by undergraduate student Matthew Maler during specimen fabrication is acknowledged. References [1] Park TW. Inspection of collapse cause of Sampoong Department Store. Forensic Sci Int 2012;217:119–26. [2] Peng Z, Orton SL, Liu J, Tian Y. Experimental study of dynamic progressive collapse in flat-plate buildings subjected to an exterior column removal. J Struct Eng 2017;143(9):04017125. [3] Peng Z, Orton SL, Liu J, Tian Y. Experimental study of dynamic progressive collapse in flat-plate buildings subjected to an interior column removal. J Struct Eng 2018;144(8):04018094. [4] Moss PJ, Dhakal RP, Wang G, Buchanan AH. The fire behaviour of multi-bay, twoway reinforced concrete slabs. Eng Struct 2008;30(12):3566–73. [5] George SJ, Tian Y. Structural performance of reinforced concrete flat plate buildings subjected to fire. Int J Concr Struct Mater 2012;6(2):111–21. [6] Ruiz MF, Muttoni A, Kunz J. Strengthening of flat slabs against punching shear using post-installed shear reinforcement. ACI Struct J 2010;107(4):434–42. [7] Annerel E, Lu L, Taerwe L. Punching shear tests on flat concrete slabs exposed to fire. Fire Saf J 2013;57:83–95. [8] Bailey C. Holistic behavior of concrete buildings in fire. Struct Build

7. Conclusions Experiments were conducted on large-scale isolated slab-column connection specimens with an aim to examine the effects of fire-induced high temperatures on the residual punching shear strength of reinforced concrete flat-plate structures and the effectiveness of a detailing approach to enhance the post-punching load resistance. Finite element analyses were performed to simulate the behavior of slabcolumn connections up to punching failure. The following major 11

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