Energy absorption potential of concrete floors containing secondary (shrinkage and temperature) reinforcements

Abstract

This paper experimentally investigates the energy absorption potential of two types of concrete floors, namely, normal density concrete and structural low-density concrete, containing secondary (shrinkage and temperature) reinforcements. The test program considered the following secondary reinforcements: 1) traditional welded-wire steel mesh, 2) steel fiber and 3) poly composite fiber. To estimate the extent to which crushing of floor slab materials would help absorb energy, a series of concrete penetration tests employing patch loading was undertaken on scaled down model slabs. Each concrete-secondary reinforcement combination considered slabs of 50 mm in depth with square plan dimensions ranging from 50 to 500 mm, resulting in a total of 30 test specimens. The first part of the paper discusses the test specimens, the test setup, and the test procedure. The second part of the paper presents the experimental results and establishes the energy absorption of different concrete- secondary reinforcement combinations. Sieve analysis results of the crushed specimens were used to derive a “work index” value that relates the pulverized particle size distributions to energy inputs. The work index values of concrete-secondary reinforcement systems can be used to assess the energy dissipation potential associated with such floor slabs in buildings undergoing progressive collapse. The results indicate that floors with secondary reinforcements could play an important role in helping arrest global progressive collapse.

Keywords

concrete floors / structural low-density concrete / shrinkage and temperature reinforcements / energy absorption / penetration tests / sieve analysis

Introduction

In an earlier paper by Korol and Sivakumaran [1], the issue of energy absorption potential of structural low-density concrete (LC) utilized in floor slabs of buildings experiencing post-failure response was investigated. The purpose of that study was to determine whether such slabs, devoid of any reinforcements, could offer a degree of resistance to a global collapse scenario if an extreme loading event induced wreckage from floors above impacted such slabs. In that study, it was determined that model floor slabs subjected to quasi-static patch loads had the potential to offer resistance to a possible catastrophic collapse. Based on the experiments undertaken on confined and unconfined concrete slabs, and especially for cases where the edges were restrained, we concluded that such floor systems would likely play a considerable role in dissipating kinetic energy generated during a collapse scenario.

However, even in areas where the top surfaces of floor slabs would not require steel reinforcing, it is standard practice to employ secondary reinforcements to prevent concrete cracking due to shrinkage or temperature changes. In this regard, steel mesh has traditionally been employed in practice for many decades to counter such effects, although steel fibers and others of composite materials are becoming increasingly popular in the industry. Since the early 90s guidelines for employing steel fiber reinforced concrete (SFRC) have been in existence with particular applications for ground floors in industrial buildings [2]. Falkner et al. [3] undertook a major study that showed much higher toughness values for such floors with fibers compared to those devoid of reinforcing. These conclusions were then augmented by research done at Purdue University by Chiu and Sivaram [4] which focused on fracture energy and toughness comparisons of fiber reinforced normal concrete with its unreinforced counterpart. A much later entrant into the fiber category of concrete reinforcement is polypropylene. A recent study [5] employing such fibers up to 0.10% of concrete volumes, showed improved performance under conditions of compression, splitting tensile strength and impact resistance. In the same year 2012, Widodo et al. [6] studied the effects of combining steel and polypropylene fibers in lightweight concrete, noting that such additions provide much higher modulus of rupture properties than such concrete devoid of such secondary reinforcement.

Since the focus of Korol and Sivakumaran [1] study involved energy absorption potential of unreinforced LC in floor slabs subject to patch loadings, this investigation extended the study to include both structural low-density concrete and normal density concrete with secondary reinforcements added. The latter employed three types of such reinforcements, i.e., traditional steel mesh, steel fibers and an advanced form of non-metallic macro-synthetic fiber, which is a proprietary blend of polypropylene resins. Although secondary reinforcement is traditionally prescribed to prevent cracking, it is our view that its toughness property can play a major role in preventing global collapse of high rise buildings as well. Since structural low-density concrete is desirable for minimizing the structure’s own dead loading, while normal density concrete is a more commonly used structural form in general, it was deemed important to include both types of concrete. Since concrete is known to be a brittle material offering limited tensile resistance in imminent collapse circumstances, it was our view that secondary reinforcement could provide additional robustness to a structure’s ability to resist extreme loading events.

The purpose of this investigation, therefore, is threefold. First, this paper presents the experimental results employing three types of secondary reinforcements in model floor slabs subjected to patch loadings, with the focus being on post-peak quasi-static loading resistance. Secondly, this study determines the degree to which such slabs exhibit resistance via a brittleness parameter known as the Bond “Work Index,” established in a classic paper on the theory of comminution [7], i.e., a measure of a material’s resistance to pulverization. And thirdly, this paper compares the results from the tests reported herein with those described in the earlier paper by Korol and Sivakumaran [1] on unreinforced structural low-density concrete (lightweight concrete) having slab ends free, and edge-supported boundary conditions.

The experimental program

The experimental program considered slabs made of two kinds of concrete namely, normal density concrete and structural low-density concrete, containing three types of shrinkage and temperature reinforcements, namely steel mesh, steel fiber and polypropylene fiber.

Concrete material details

The structural low-density concrete used in the slab tests, consisted of a light weight aggregate (bulk density 761 kg/m³), known as “Truelight” obtained courtesy of Lafarge Slag Inc. in Hamilton, Canada, together with regular Portland cement, sand and water. The actual batch mix for structural low-density concrete of a target volume of 0.11 m³ was as follows: cement – 34.5 kg, water – 21.9 kg, sand – 72.5 kg and light weight coarse aggregate – 65 kg. For the normal density concrete(NC) slabs, which had the same volume as did structural low-density concrete(LC), however, the measured bulk density of the normal weight coarse aggregate was 1483 kg/m³, the concrete batch mix was with weights of 41.0 kg, 17.1 kg, 123.5 kg, respectively, as above, and normal weight coarse aggregate of 77.5 kg.

Because steel and poly-composite fibers were employed in our investigation, we needed to add a super-plasticizer that would ensure adequate workability of concrete and prevent “balling” when the fibers are added to the material batch in the concrete mixer utilized. The amount of super-plasticizer liquid mixed in with the water, which was established based on the producer recommended dosage range, was 140 ml for the structural low-density concrete (LC) batch and 160 ml in the normal density concrete (NC) case. These batches were then divided into approximately three equal parts to allow for either of the two types of fibers mentioned above, or for an equivalent array of conventional steel mesh reinforced specimens. Standard concrete cylinder tests were undertaken to establish the 28 days compressive strengths of the concrete material with average values being 20.5 MPa for structural low-density concrete, and 40.9 MPa for normal density concrete, while their density values were 1729 and 2364 kg/m³, respectively.

Secondary reinforcement details

Photographs of the secondary reinforcement types considered are shown in Fig. 1. The welded wire steel mesh (WWM) employed as secondary reinforcement had a 50 mm by 50 mm aperture with its wire diameter being 3.18 mm (2″× 2″× 1/8″), resulting in an approximate reinforcement ratio of ρ = 0.32%, in each direction. This value is in between the minimum reinforcement (0.20%) to be used in structural floor given in ACI318 [8] and the acceptable level of reinforcement to control cracking specified by the ACI committee report 224 [9] of 0.60%. The size of mesh, noted above, determined the amount of steel fibers needed for its category of tests, made equal in weight to allow for rational comparisons, i.e., 48 kg/m³. As depicted from Fig. 1, the steel fibers under consideration had a diameter of 1.00 mm, and a length of 50 mm, thus an aspect ratio of 50. The ends of the steel fibers are mechanically deformed in order to provide better anchorage in concrete. According to the technical data sheet issued by the supplier, the tensile strength of the wire is greater than 1100 MPa and the ultimate strain smaller than 4%. For the steel fiber (SF) cases, 1.65 kg of this type of reinforcing was added to one-third of the concrete batch, while for the poly-composite fibers (PF), 0.33 kg was the dosage given to its one-third portion (recommended by the supplier, BASF). It should be noted that the poly-fibers used were of nominal 50 mm lengths with thicknesses in the micro-meter (µm) range. As can be seen in Fig. 1, the poly-fibers came with undeformed straight ends.

The test specimens

Figure 2 shows photographic images of sample specimens made of normal density concrete (identified as NC-PF20) and structural low-density concrete (identified as LC-PF20). All concrete slab specimens having a depth of 50 mm were cast in wooden forms, shaped as squares, with side wall lengths, b, ranging from 50 mm (2″) to 500 mm (20″). This range was selected in order to be consistent with the earlier study by Korol and Sivakumaran [1]. For the 50 mm × 50 mm slabs containing nominally 50 mm long fibers, some fibers formed bends and consequently localized small voids formed around the slab boundaries because of boundary restraint imposed by the small metal casting forms (50 mm × 50 mm) employed. However, in the case of the steel mesh (SM) reinforcement, the 50 mm square of steel mesh inserted at mid-depth of the 50 mm × 50 mm slabs just barely allowed embedment into all four sides at mid-height of the cube, and hence permitted inclusion in this investigation. To identify individual specimens in this paper, we denote sequentially; the type of concrete, i.e., structural low-density (LC) or normal density concrete (NC), the type of secondary reinforcements denoted by the symbols SM, SF or PF, representing steel mesh, steel fibers and poly-fibers, respectively, and the slab size dimension b, in inches, noted as: 2, 3, 5, 10 or 20 (inch units). Hence, LC-PF20 is an example of a structural low-density concrete slab employing poly-fibers having a surface dimension of 500 mm × 500 mm (20″× 20″). As a consequence of what were deemed to be acceptable specimens, 15 tests for each of NC and LC were performed on five slabs and three types of reinforcement giving a total of 30 tests. These model slabs were then subject to penetration-type loads as described below.

Loading application details

Figure 3 shows a typical specimen prior to testing in the Tinius-Olsen testing machine at McMaster’s Applied Dynamics Laboratory. It will be noted that for all tests a square 50 mm × 50 mm (2″ × 2″) steel loading block was centrally positioned on a given slab, with edges parallel to those of the specimen being tested, while LVDTs were used to monitor displacements. Quasi-static loading was then slowly applied until the steel block underwent a displacement of about 25 mm (1/2 depth), at which point loading was terminated. Such a displacement is several times greater than that associated with a specimen’s peak-load and tends to result in localized pulverization of the concrete into fine particles directly underneath the steel block, while outer areas tended simply to experience cracking and breakup into large pieces. An array of typical failure states for a group of six specimens is shown in Fig. 4. For any given specimen, the area under the load-displacement curve represents the experimental energy dissipated during that particular event. These results are presented in the section below.

Experimental energy absorption results

The patch load-penetration plots for all 30 specimens are compiled together and shown in Fig. 5, in order to convey the differences according to reinforcement type and patch-to-slab area ratios, a/A, where a is the area of the patch load (50 mm × 50 mm), and A is the surface area of the test specimen. Note that the graphs identify a maximum penetration of 25 mm with any additional data points beyond that limit ignored, thus providing consistency when comparing energy absorption values, obtained as the area under the load-penetration curve. In general, all 30 test specimens exhibited similar responses, which consisted of several distinct regions. Settling in of the loading block with its concrete surface resulted in initially low stiffness response, followed by a region of high gradient load-penetration response, after which the peak load was reached, often associated with formation of an initial crack. Table 1 identifies these peak load values obtained from the 30 tests undertaken. As noted from individual columns in the table, corresponding to different concrete-secondary reinforcement combination, the larger specimens generally reached higher peak loads than those of the corresponding smaller specimens. The peak loads associated with the normal density concrete specimens are in general higher than their corresponding structural low-density concrete specimens. It is, however, not surprising to note that the normal density specimens performed much better than their structural low-density counterparts since their respective concrete cylinder tests produced a nearly 2 to 1 ratio of maximum compressive stresses, f′_c, (40.9 to 20.5 MPa), as noted earlier.

Table 2 provides energy absorption values computed by numerical integration for each of the plotted curves shown in Fig. 5 having a penetration range from 0 to 25 mm. It is noteworthy that in every case without exception, for which the same reinforcing type was employed for a given concrete, the amount of absorbed energy increased as the slab size increased. For example the 500 mm × 500 mm NC – SF specimen absorbed eight times more energy compared to its 50 mm × 50 mm counterpart of the same material combination. Since the structural low-density concrete was of much lower strength than the normal one, it is not surprising to note that normal density concrete energy values were greater than their structural low-density counterparts (same slab size and reinforcing type) except for the 125 mm × 125 mm size employing SF (1136 vs. 1229 J, respectively). The observation of relation between the size of the specimens and the peak load and the energy values indicate the importance of restraint offered by slab material external to the loading block, i.e., a/A << 1 results in higher peak loads and greater energy dissipation than for cases of a/A nearer to or equal to unity.

Energy density can be established as energy per unit volume expressed in Joules/cm³ (J/ml) for a specimen. It was found that the energy density decreases with increasing size of specimen for a given patch load area. For example, the energy density of the 50 mm x 50 mm normal density concrete containing polypropylene fibers (NC-PF2) was 4.32 J/ml, whereas those of 125 mm × 125 mm (NC-PF5) and 500 mm x 500 mm (NC-PF20) were 1.75 and 0.36 J/ml, respectively. The reason for such precipitous changes is due to the a/A ratio, i.e., all specimens were subjected to the same patch load of 50 mm × 50 mm at the middle of the specimen, and as such the energy absorbed by the materials remote from the patch load location would be less than the material right underneath the patch load.

The influence of secondary reinforcements can be quantified from an energy absorption point of view. Using the energy values given in Table 2, the normal density concrete specimens with steel fibers(NC-SF) absorbed between 66% and 121% (average of different sizes 95%) of the energy expended by the same size normal density concrete slab containing steel mesh. The corresponding comparison for structural low-density concrete-steel fiber (LC-SF) was between 44% and 188% (average of different sizes 83%) of the energy absorption by structural low-density concrete-steel mesh specimens (LC-SM). The normal density concrete with polypropylene fibers (NC-PF) absorbed between 59% and 148% (average of different sizes 108%) of the energy compared to the same concrete containing steel mesh. Similar results for structural low-density concrete with polypropylene fibers (LC-PF) were between 51% and 204% (average of different sizes 106%) of LC-SM specimens. Based on energy values obtained, it is seen that there is considerable scatter in the results and as such, other factors such as costs of material and labor would need to be considered before recommending one type of secondary reinforcement over another.

In the case of structural low-density concrete tests, the energy values as shown in Table 2 are in between the energy values for unconfined light weight concrete specimens and edges confined test specimens reported in the earlier study by Korol and Sivakumaran [1]. In that investigation, the strength of the low-density concrete used was 19.8 MPa, only slightly less than the value reported here for structural low-density concrete of 20.5 MPa. As such, a comparison between the energy dissipation values for the unconfined 50 mm cubes tested without any reinforcement was made with the three test results from this study. Noting the results from the top row of Table 2 for the structural low-density concrete cases having the three secondary reinforcements, the energy values range from 140 to 291 Joules with an average of 225 Joules, whereas for the two unconfined light-weight specimens of 50mm, listed as 2A1 and 2A2 in the previous investigation [1], gave an average result of 174 Joules or 23% less than those possessing secondary reinforcements considered herein.

Sieve analysis to ascertain particle distributions

The importance of ascertaining the particle sizes and quantities directly relates to the susceptibility of a brittle material to fragment under methods used to pulverize materials. In this case, we are dealing with crushing of concrete containing various types of secondary reinforcement meant to prevent unsightly cracks from shrinkage and temperature changes. Sieve analysis allows one to assess the tendency of a material to pulverize into smaller sized particles and to ascertain its potential to dissipate energy in extreme loadings events. After each patch load test, the small fragments and large pieces of the failed test specimen were carefully collected and the weights of the large pieces (>> 20 mm) were recorded. During the crush tests, all of the 50 mm × 50 mm specimens were reduced to bits less than 20 mm, whereas, the larger slabs typically broke up into a full array of large pieces down to small dust particles (see Fig. 4). Primarily, fine particle pulverization occurred in the region under or very near to the patch load, while the area beyond the loading block tended to break up into larger pieces. The remnants that were less than 20 mm in size were subjected to a standard five minutes 8-sieve analysis in order to establish the particle size distribution and the associated weights. Table 3 shows the corresponding results of a dozen of the mid-range sized slabs (i.e., 125 mm and 250 mm). This group of NC and LC specimens was chosen to represent patch load-to-slab area ratios that may embrace a practical range of a/A ratios when postulating upper storey floor system assembly collisions with floor slabs below in a building collapse scenario. The table lists the slab weight prior to testing, observed number of large pieces (> 20 mm size), and the percentage weight of the particles retained in sieve sizes ranging between 20 mm and 60 μm. Additional details related to the 30 sieve analysis tests discussed herein can be found in the thesis by Fan [10].

Analysis of results

Our objective, then, is to employ a method by which we can estimate the amount of energy associated with pulverizing a concrete slab containing secondary (shrinkage and temperature) reinforcements, an attribute that is important when tallying up a structure’s resistances to collapse. As a matter of fact, input energy estimates are important to industries that employ processes that involve grinding, crushing or blasting to break down materials to reduced sizes. In those mining and milling industries, despite the passing of several decades, the Bond’s approach [7] is widely used to calculate such energy demands. Our interest, of course, is to identify secondary reinforced concrete systems that maximize the energy of comminution (breakdown into particles), whereas mining and milling industries aim to minimize such energy. This paper utilizes the method of Bond [7] to establish the so-called Bond Work Index value for a concrete slab containing secondary (shrinkage and temperature) reinforcements. The approach is similar to what was used in an earlier paper by Korol and Sivakumaran [1]. The kernel of the method proposed by Bond [7] is a formula given as follows:

(1)E=\left\{\mathrm{10}\left[\frac{\mathrm{1}}{\sqrt{x_f}}-\frac{\mathrm{1}}{\sqrt{x_i}}\right]\right\}{W}_i

where E is the energy per unit mass to pulverize it from the initial to its final particle size state, whereby x_i and x_f represent the initial and final average linear dimensions, respectively, expressed in μm. The factor 10 in Eq. (1) is actually √100 μm to provide dimensional consistency for size units. In experiments such as the ones we are describing, a slab’s breakup will results in assorted final particle size values, determined from a sieve analysis procedure such as described earlier. One can compute the Bond Work Index, W_i for any material from Eq. (1), when E and the particle size distributions are known. In mining and milling industries, traditionally, E and W_i are both established in units of kilowatt hours per ton (kW-hr/t). However, in order to use E in units of Joules/kg, a convenient unit for our purposes, while maintaining the Bond Work Index W_i in units of kW-hr/t, we need to multiply the right hand side of the Eq. (1) by 3,600. One additional correction to the Eq. (1) should be made, however. It has been shown by Eloranta [11], that there is a lesser degree of efficiency, by a factor of 3.4, in the comminution of materials, if direct crushing is used, rather than breakdown predominantly by a sliding action, or by grinding. Since crushing dominates that part of slab directly below the steel loading block used in our tests, we therefore propose to modify the left hand side of Eq. (1) to account for Eloranta’s observation by a factor {[3.4 a + 1(A-a)]/A}, where, a is the area of the patch load (50 mm × 50 mm), and A is the surface area of the test specimen. Essentially, had it not been direct crushing failure the corresponding experimentally obtained energy would be less. Hence, providing a correction for the parts of the specimen failed under direct crushing, the energy-comminution formula becomes:

(2)\frac{E_{\mathrm{c}}}{\mathrm{1}+\mathrm{2.4}a/A}=\frac{\mathrm{10}\times \mathrm{3600}}{\mathrm{sqrt}\mathrm{of}x}{W}_{\mathrm{i}},

where E_c is the energy of combined methods of pulverization in units of Joules/kg, here, the experimental energy. It may be pointed out that had this correction was not made the calculated work index would be higher, which would lead to unconservative estimates of energy absorption.

Based on the particle size distribution for a given slab (sample noted in Table 3), together with an estimate of the average size of pieces > 20 mm, and the energy values shown in Table 2, the “Work Index” value W_i was established for the 30 test specimens under consideration. To illustrate the “Work Index” calculation method, consider the slab specimen LC-PF10. As noted in Table 3, its mass was 5.622 kg, with 4 large pieces constituting 92.79% of the mass. Even though this specimen dimensions were specified as 250 mm × 250 mm × 50 mm, the measured volume, established based on measured dimensions of the actual specimen before test, was 3460,000 mm³ [10]. Therefore, a side dimension of an equivalent cube representing the large pieces after test can be computed as (3460,000 × 0.9279/4)^1/3 = 92.934 mm = 92,934 µm. For the concrete pieces having size more than 20 mm, this value is substituted into Eq. (1) for x_f, while x_i, the original size, is taken as an equivalent cube of (3460,000)^1/3 = 151.249 mm = 151,249 µm. The result of substituting x_f and x_i into Eq. (1) and accounting for the mass involved (92.79% of 5.622 kg) results in an energy quantity of 133.15 W_i Joules. Similar calculations were performed for the other nine fractions of materials (20 mm down to 60µ) resulting in a grand total of 667.64 W_i Joules. Since, 50 mm × 50 mm (2″ square) block is penetrating a 250 mm × 250 mm (10″× 10″) slab, therefore a/A = 0.04, which means that the 3.4 factor will apply to 4% of the area, while 96% retains the value of unity. Equating the corresponding experimental energy values listed in Table 2 to E_c, we can now calculate a Work Index value for specimen LC-PF10 as W_i = {1614/(1 + 2.4 × 0.04)}/667.64 = 2.21 kWh/t.

Table 4 lists the calculated Work Index values W_i for all 30 tests, collated based on concrete-secondary reinforcement combination. For readers convenience, Table 4 also lists the patch load to specimen surface area ratios. Accordingly, the average work index values for normal density concrete floor slabs containing steel mesh, steel fiber and poly composite fiber secondary reinforcements are 2.65, 2.69, and 2.94, respectively. The corresponding values for structural low-density concrete floor slabs containing corresponding secondary reinforcements are 4.29, 2.11, and 2.57, respectively. The earlier work on this subject by Korol and Sivakumaran [1] produced a work index value of W_i of 0.90 kW-hr/t for unconfined plain lightweight concrete, which is well below of the average values LC- SM, LC-SF and LC-PF type floor systems. It is noted that the work index for structural low-density concrete floor slabs containing traditional steel mesh is 4.29, which is very much in line with that of the confined lightweight concrete, a value of 4.14 obtained in the earlier study [1]. Somewhat unexpectedly perhaps, is the observation that the NC-SM had lower Work Index value W_i than its LC counterpart. The NC-SM average W_i was 2.65, or 62% of the LC-SM average W_i. The overall average work index for normal density concrete floor slabs was 2.76, whereas the overall average work index for structural low-density concrete floor slabs was 2.99. The rationale for this observation may be due to micro voids that exist in light weight aggregate that allows for a more ductile response to loadings, than the more dense aggregates that are used in standard concrete.

Discussion

Even though the experimental program employed a wide range of a/A ratios, the a/A of 0.01 to 0.16 domain perhaps represents the most likely crush load area to floor area ratio that would be experienced in practice during a progressive collapse event related to building floors. With beams and girders and even open web steel joist spacing taken into account, the contact areas offered by bottom flanges or chords would not likely exceed 0.16 times the floor area below, and this could only happen when their full lengths made contact. In many areas of a floor area, one might expect only a 1% contact ratio when excessive bending of collapsing members would make a direct hit, while other portions would remain suspended above floor surfaces. The possibilities are endless, of course, but our best guesses of what lower floors would experience suggests that high a/A contact areas are unlikely. Hence, one may focus on a/A ratios that are likely more representative of reality in catastrophic events that are poised to cause total global collapse of a typical hi-rise steel framed building. Therefore, the average work index values W_i for specimens possessing a/A = 0.01, 0.04, 0.16 (Table 4 - upper three rows for each group) are given special focus. The corresponding averages for NC-SM, NC-SF, and NC-PF are 3.87, 3.86, and 4.04, respectively, with an overall average of 3.92. Similarly, averages for LC-SM, LC-SF, and LC-PF are 6.75, 3.18, and 3.79, respectively, gave an overall average of 4.57.

To make a recommendation of what value of a Work Index is suitable to ascribe to such concrete secondary reinforcement systems is not possible, but it certainly does seem that W_i values in realm of about 4.0 seems to us to be practical regardless of the type of secondary reinforcement employed. Nonetheless, it seems that the lightweight concrete has the edge over normal density concrete when factoring in both reducing dead loading of a structure and in absorbing energy in catastrophic events that could cause its total collapse.

Conclusions

Based on the qausi-static penetration tests conducted on normal density concrete and structural low-density concrete slab models that utilized one of – welded mesh steel, steel fibers or poly-composite fibers, the applied patch loadings were effectively resisted well beyond the yield strain limit of the material, suggesting a high degree of toughness compared to totally unreinforced and unrestrained slab boundary edges. It is evident indeed that whatever secondary reinforcement is employed will help provide considerable energy dissipation capacity that contributes to arresting a progressive collapse event. As such, we are of the opinion that structural low-density concrete floor slabs, having a prescribed content of shrinkage steel in the form of mesh or fibers or, an alternative type of fiber made of macro-synthetic material based on polypropylene resins do indeed offer a significant degree of resistance to motion caused by gravity driven forces that could arise from an extreme loading event.

References

[1] Korol R M, Sivakumaran K S. Energy Absorption Potential of Light Weight Concrete Floors. Canadian Journal of Civil Engineering, 2012, 39(11): 1193-1201

[2] Concrete Society. Concrete Industrial Ground Floors: A Guide to their Design and Construction – TR34 Report. 2nd ed. General document and Appendix F (Slab Design with Steel Fibers), 1995

[3] Falkner H, Huang Z, Teutsh M. Comparative Study of Plain and Steel Fiber Reinforced Concrete Ground Slabs – ACI. Concrete International, 1995

[4] Chiu Y C, Sivaram A. Analysis of Fracture Energy: Comparative Study of PCC & FRC. Report dated 4/27/2011, Purdue University, USA, 2011

[5] Widodo S. Fresh and Hardened Properties of Polypropylene Fiber-Added Self Consolidating Concrete. Int J Civ and Struct Eng, 2012, 3(1): 85-93

[6] Widodo S, Satyarno I, Tudjono S. Effects of Hybrid Polypropylene-Steel Fiber Addition on Some Hardened Properties of Lightweight Concrete with Pumice Breccia Aggregate. ISRN Civil Engineering, Article ID 475751, 2012, 8

[7] Bond F C. The Third Theory of Comminution. Mining Engineering Journal. <month>May</month> 1952, 484-494

[8] American Concrete Institute. Building Code Requirements for Structural Concrete and Commentary (ACI 318–08), Farmington Hills, MI, USA, 2008

[9] American Concrete Institute. Control of Cracking in Concrete Structures. ACI Committee report 224. Farmington Hills, MI, USA, 2001

[10] Fan X. Characteristics of concrete slabs employing shrinkage and temperature reinforcements, Master Thesis. Hamilton, Ontario, Canada: Thesis, McMaster University, 2013, xiii: 127

[11] Eloranta J. The Efficiency of Blasting Versus Crushing and Grinding. In: Proceedings of the 23rd Conference of Explosives and Blasting Technique. Las Vegas, Nevada February. 1997, 1-7

RIGHTS & PERMISSIONS

Higher Education Press and Springer-Verlag Berlin Heidelberg