Effect of nylon fibres on mechanical and thermal properties of hardened concrete for energy storage systems

Materials and Design 51 (2013) 989–997

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Effect of nylon fibres on mechanical and thermal properties of hardened concrete for energy storage systems O.B. Ozger a, F. Girardi a, G.M. Giannuzzi b, V.A. Salomoni c, C.E. Majorana c, L. Fambri d, N. Baldassino a, R. Di Maggio a,⇑ a

Department of Civil, Environmental and Mechanical Engineering, University of Trento, Italy ENEA, Agency for New Technologies, Energy and Environment, Thermodynamic Solar Project, CRE Casaccia, Rome, Italy Department of Civil, Environmental and Architectural Engineering, University of Padua, Italy d Department of Industrial Engineering, University of Trento, Italy b c

a r t i c l e

i n f o

Article history: Received 18 December 2012 Accepted 26 April 2013 Available online 17 May 2013 Keywords: Fibre-reinforced concrete Textile carpet waste Thermal energy storage device

a b s t r a c t The use of recycled materials in new concrete production is very appealing because of the low cost of the raw materials and the landfill space saving, as well as the enhancement of the resulting concrete’s properties. This paper focuses on the production of concrete for thermal energy storage systems using polyamide fibres from post-consumer textile carpet waste. We investigated the influence of a fibre content of 5 kg/m3 on the mechanical and thermal properties of fresh and hardened concrete with a w/c ratio of 0.35, comparing it with plain concrete. The nylon fibre-reinforced concrete was slightly more ductile and tougher than the plain concrete because the fibres have an important role after failure. The tensile strength, maximum load-bearing capacity and modulus of elasticity decreased slightly with the addition of nylon fibres, but drying shrinkage was lower than for plain concrete. Given that heat capacity and thermal conductivity are the most important parameters for thermal energy storage purposes, we measured these parameters for all the samples at 25 °C, and after a thermal treatment at 300 °C. The heat capacity values were 0.63 and 0.81 J g1 K1 for fibre-reinforced concrete (FC) and plain concrete (C), respectively, and the thermal conductivity values were 1.16 and 1.02 W m1 K-1. Ó 2013 Elsevier Ltd. All rights reserved.

1. Introduction So far, technology for thermal energy storage has been used largely in concentrated solar power (CSP) plants. Like conventional power plants, solar plants must cover the consumers’ energy demand around the clock, and this requirement often does not coincide with the energy input of the solar plant, which is limited by diurnal, seasonal and weather-related insolation changes [1]. Thermal energy storage (TES) systems have the potential for increasing the effective use of solar energy and facilitating large-scale switching of generation plants [2]. To balance energy supply and demand, solar plants must be built to include TES and/or fossil-fired backup systems in order to prolong their operating times, guarantee a given output, shift energy output from low-price off-peak periods to peak periods, and obtain capacity payments [1–5]. Thermal storage units have been tested mainly in the liquid state. For example, the ‘‘Solar Two’’ solar power plant in the USA, and the ‘‘Archimede’’ advanced parabolic trough plant developed by ENEA in Italy use molten salts as a heat transfer and thermal

⇑ Corresponding author. Tel.: +39 0461 282419. E-mail address: [email protected] (R. Di Maggio). 0261-3069/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.matdes.2013.04.085

storage fluid. Notably too, traditional trough plants (e.g. ‘‘Andasol’’ in Spain) distinguish the fluid transiting through the solar field (synthetic oil) from the fluid used in the storage system (molten salt). As for the use of solid materials, concrete is generally the preferred medium for thermal energy storage, as already tested at the Solar Platform in Almeria (Spain) and by DLR (Germany), where it revealed an appropriate response for this specific use [1–5]. When it comes to developing TES material, a high heat capacity (q  cp) reduces the storage volume and a high thermal conductivity increases the dynamics in the system. The coefficient of thermal expansion (CTE) of the storage material should match that of the material of the embedded metallic heat exchanger; a high stability in relation to charging–discharging cycles is also important for a long lifetime of the storage unit [2,6]. Concrete is used because it is cheap, readily available, easy to process and handle, its main ingredients are available all over the world, and its components pose no environmentally critical issues [7]. With its high specific heat, good mechanical properties (e.g. compressive strength) and a thermal expansion coefficient close to that of steel (pipes), concrete also promises to have a high mechanical resistance to cyclic thermal loading [8,9]. The concrete’s response to heating is fundamental to a prediction of its temperature and heat and mass diffusion, however, also

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Nomenclature fcm Rcm fIf PIf fctm Em CMOD CTOD l a0 b h D0

mean cylinder-compressive strength mean cube-compressive strength first-crack strength first crack load mean tension strength mean elastic modulus crack mouth opening displacement crack tip opening displacement distance between the inferior supports in four point bending test notch depth in beam sample in four point bending test width of beam sample in four point bending test height of the beam in four point bending test ductility index in the CTOD ranges of 0–0.6 mm

taking into account that water is driven off and several reactions and transformations take place that influence the material’s strength and other physical properties, e.g. density q, specific heat capacity cp, thermal conductivity k, CTE, and cyclic stability. The compressive strength of concrete decreases by about 20% at 400 °C, and its specific heat and thermal conductivity drop in the range of 20–120 °C and 20–280 °C, respectively [2,4,5]. Fu and Chung [10] recently published room-temperature values for the specific heat of cement pastes prepared with various admixtures. For a plain cement paste with a water/cement ratio of 0.45, they found a density of 1990 kg/m3 and a specific heat of 0.703 J/g K. For a cement paste containing 15% of silica fume (w/w of the cement) and 3% of water-reducing agent, with a water/cement ratio of 0.35, they found a density of 1720 kg/m3 and a specific heat of 0.765 J/g K [10]. The corresponding thermal conductivities were 0.52 W/m K and 0.36 W/m K, respectively. In another recent work, Tamme et al. [11] presented the thermophysical properties of a high-temperature concrete with thermal conductivities in the range of 1.2–1.45 W/m K and a volumetric heat capacity of 0.64 kW h/(m3 K). Steel needles and reinforcement are generally added to the concrete to avoid cracking, thus increasing its thermal conductivity by about 15% at 100 °C and 10% at 250 °C [2,12]. Fibre reinforcement has also been recognised as an effective way to enhance the fracture resistance of concrete under all modes of failure. The risk of spalling can be reduced by increasing the concrete’s permeability and could be eliminated altogether by adding polymer fibres (e.g. polypropylene) to the mixture [13–15]. Polypropylene fibres melt at about 160 °C, leaving channels in the concrete through which moisture can escape. Polypropylene (PP) fibres have been used in concrete primarily to control shrinkage cracking and (to a more limited extent) to improve its toughness and impact resistance, increasing the concrete’s energy absorption capacity [16,17]. Much recent research has shown that adding a small volume of PP fibres clearly reduces a concrete’s plastic shrinkage cracking in early age, and these fibres can also distinctly restrain surface bleeding and the settlement of aggregate in fresh concrete, thereby preventing the formation of setting cracks [18]. Some authors have demonstrated a post-cracking activity of polyamide fibres: when a crack starts in a concrete matrix, the fibres hold its edges together [19–31]. In fresh concrete, evenly-distributed fibres help to prevent the formation of plastic shrinkage cracks, while in hardened concrete, they prevent the growth of microcracks [32–33]. Furlan, Hanai and Karahan also observed in their experiments that adding fibres reduced the fresh concrete’s workability [34,35], particularly when they used synthetic fibres with a very high aspect ratio (length/diameter) [35].

D1 feq(0–0.6) feq(0.6–3) U1

ductility index in the CTOD ranges of 0.6–3 mm equivalent strength in the CTOD ranges of 0–0.6 mm equivalent strength in the CTOD ranges of 0.6–3 mm area under the load-CTOD in the CTOD ranges of 0– 0.6 mm U2 area under the load-CTOD in the CTOD ranges of 0.6– 3 mm CTE coefficient of thermal expansion k thermal conductivity Cp specific heat capacity Dl/l0 and e strain Dm mass variation

The present paper addresses the preparation of fibre-reinforced concrete and the identification of its properties – such as compressive strength, toughness, specific heat capacity, thermal conductivity, thermal expansion, and hygrometric shrinkage – crucial to prefiguring the concrete’s behaviour before and after working in a TES unit. 2. Experimental activities 2.1. Materials We used crushed natural dolomite aggregates to prepare a plain concrete mix serving as a control (C), considering three grades of aggregate, i.e. gravel with a maximum particle size of 25 mm, fine gravel with a maximum size of 12 mm, and limestone sand with a fineness modulus of 3.5. We used a commercial type of Portland limestone cement (CEM II/A-LL 42.5 N [36], made by Titan Cement Co.), adding recycled nylon PA66 fibres in proportions of 0.5% v/v (almost 5 kg/m3) to prepare a fibre-reinforced concrete mix (FC) with the natural aggregate. The fibres were obtained by recycling post-consumer carpet waste. Table 1 shows the mix proportions for the concretes investigated. We optimised a specific procedure for the dispersion of the short fibres in the concrete mixture (blowing the fibre wisps with compressed air at a pressure of 8–10 bar proved very effective for this purpose). We further distributed the fibres in the water and fine aggregates with the aid of a hand mixer driven by an electric drill, then we added the fibre-sand mixture to the other concrete ingredients for batching. We mixed the concrete with a 100 l rotary drum mixer. The fine and coarse aggregates were dry mixed for 5 min beforehand.

Table 1 Mix proportions of the concretes investigated.

a

Mixture name

C

FC

Mixing water (kg/m3) Cement (kg/m3) Gravel (12/25) (kg/m3)a Fine gravel (8/12) (kg/m3)a Sand (kg/m3)a Fibres (kg/m3) w/c ratio Specific weight (kg/m3) Slump (cm)

97 280 400 200 1000 – 0.35 2451 0

97 280 400 200 995 5 0.35 2440 0

Aggregates are used in the condition of saturated surface dry.

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2.2. Characterisation methods and instruments We used a Philips XL-30 environmental scanning electron microscope (ESEM) to obtain information on the size, morphology and surface topography of the fibres, aggregates and concretes after 28 days of wet-curing and after heat treatment. We also conducted energy-dispersive X-ray spectroscopy (EDXS) analyses to investigate the chemical composition of the materials. We measured the fibres’ tensile properties by means of a singlefibre tension test performed on a Seiko Exstar TMA/SS 6000 machine fitted with a special fibre fixture. We performed the tests at a constant temperature of 20 °C and a loading rate of 100 mN/ min, corresponding to a deformation rate of about 2 mm/min. We recorded the fibres’ tensile modulus, ultimate strength, and ultimate elongation as the means obtained by measuring five fibre specimens (mean length 5 mm and equivalent diameter of 41 lm). We measured the fibres’ absorption at 25 ± 1 °C. We dried the fibres in an oven at 100 °C, then placed them in a humid static chamber with a relative humidity of 100 ± 2%. We measured the increase in their mass every hour for about 12 h, using an electronic balance (Toledo Mettler AG245) with a precision of 0.01 mg, and their weight became constant within 24 h. We performed differential scanning calorimetry (DSC) on the fibres with a DSC 92 Setaram over the range of 20–300 °C at a rate of 10 °C/ min in both N2 and air, using aluminium (Al) pans. We conducted thermogravimetric (TG) and high-temperature differential scanning calorimetry (HT-DSC) analyses simultaneously on aggregates and concretes on a TG-DTA/DSC Setaram, using alumina pans at a heating rate of 10 °C/min in air, under a flow of 100 ml/min. The cooling rate was 50 °C/min. Before conducting any thermal analyses, we ground the specimens to a powder using a Fritsch Pulverisette 0 vibratory ball mill. We weighed the samples before and after the DSC test using a Mettler AE 240 microbalance with an accuracy of ± 10 lg to ascertain the percentage weight change. We examined the crystallographic structure and chemical composition of the aggregates using powder X-ray diffraction analysis (XRD) and an imaging plate diffractometer (Italstructure) in pure reflection geometry, with a Cu Ka anode (40 kV–22.5 mA), a Si multilayer monochromator on the incident beam, and a Ni filter on the diffracted beam. We characterised the fresh concrete in accordance with the European Standards. To measure drying shrinkage, in compliance with the Italian standard UNI 11307: 2008 [37], we sealed three prisms obtained from each mix (10  10  35 cm) in a large plastic box and kept them at a constant relative humidity of 55 ± 1% and temperature of 20 ± 2 °C. We measured the variation in their length with a digital comparator affording a precision of 0.001 mm at 1, 7, 28, 60 and 90 days. 2.3. Mechanical characterisation of the concrete after 28 days of wet curing For each mix, we cast six 100  200 mm cylinders and three 150  150 mm cubes in PVC moulds, and three 150  150  600 mm beams in wooden moulds, compacting all the moulds on a vibrating table. We wet-cured the specimens at a relative humidity of 95% and a temperature of 20 °C for 28 days, de-moulding them one day before testing them. In compliance with UNI 11039-2:2003 [38], we cut a 45 mm deep and 3–5 mm wide Vshaped notch in the middle of each beam’s span (but the crack tip shape was rectangular in this study). We performed compressive strength tests on the cubes with a 3000 kN Controls Universal Testing Machine. We established the elastic modulus and performed compressive strength tests on the cylinder with a 1000 kN Metrocom universal testing machine 28 days after batching. We used an HBM WA Series displacement

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transducer (maximum displacement 100 mm, class 0.1, sensitivity 80 mV/V) to measure the vertical displacement of the test samples. Similarly, we used three full model DD1 HBM displacement transducers (measurement base 50 mm, maximum strain + 2.5 ram, class 0.1, sensitivity 1 mV/V/mm) to measure strain in the test samples. The elastic modulus and compression tests complied with the Italian Standard UNI 6556:1976 [39] and the European Standard UNI EN 12390-3:2009 [40], respectively. For each mix, we tested three specimens under compression alone, and another three to assess both the elastic modulus and the compressive strength. The loading rate was approximately 1.5–2.1 kN s1 and we conventionally stopped the compression tests when either the load reached around 1/10–1/20 of the peak value (for the FC), or the sample failed abruptly (for the C). We recorded loads and displacements using a Spider8 multi-channel electronic PC measuring unit. For each mix, we tested three notched beams to ascertain the first-crack strength and the ductility indexes of the fibre-reinforced concrete mix (in accordance with Italian Standard UNI 110391:2003 [41]). We conducted these tests on a 100 kN Galdabini mechanical universal testing machine, using a four-point configuration. We measured the crack mouth opening displacement (CMOD) and crack tip opening displacement (CTOD) with a model DD1 HBM displacement transducer, recording top loads and corresponding mid-span crack widths automatically during the test, and taking five readings per second. To comply with UNI 11039-2:2003 [38], we had to run the test under a controlled CMOD, with a CMOD increment rate of 0.05 ± 0.01 mm/min. For CMOD values greater than 0.65 mm, we could increase the displacement rate gradually to 0.5 ± 0.02 mm/ min. Based on past experience, we set the testing machine’s displacement increment rate in the range of 0.3–0.4 mm/min so as to meet the standard requirements. 2.4. Thermal characterisation of the concrete We completed the isothermal treatments in air using a Nabertherm N 30/85 Ha oven, set at four different temperatures, i.e. 150 °C, 250 °C, 350 °C, 450 °C, and checking the mass every 30 min for a total of 4 h using a Mettler AE 240 microbalance with an accuracy of ±10 lg. We recorded thermal expansion using a Seiko Exstar TMA/SS 6000 dilatometer over the range of 20–550 °C at a rate of 3 °C/ min in air with an alumina probe, repeating the measurement two more times. We measured heat capacity using a DSC 92 Setaram differential scanning calorimeter over the range of 25–350 °C at a rate of 3 °C/ min in N2 using Al pans. We analysed thermal conductivity with a k-meter EP500 by Lambda-Messtechnik GmbH at 25 °C, before and after an isothermal treatment at 300 °C in air. 3. Results and discussion 3.1. Recycled nylon PA66 fibres from waste carpets We used polyamide fibres from post-consumer carpets in this study because of their potential suitability for use in the concrete industry, primarily due to their low cost and local availability. Instead of their disposal in landfills or recycling as PA 66 polymer by compounding and pelleting, they can also be used as short fibres and converted into useful products, including special concretes. We chose various types of nylon PA66 fibre derived from recycled carpet waste in different colours (white, brown, light blue) to ensure a representative composition of typical BCF (bulk continuous

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filament) polyamide. These short fibres were a mean 8 ± 3 mm long and were not straight, but rather curly or coiled due to the previous texturing process. ESEM showed a trilobal cross-section, and we calculated an equivalent diameter in the range of 38–41 lm using image analysis. The corresponding titre or linear density of the PA66 fibres used here was in the range of 13–15 dtex (ASTM: D861-07) [42–43]. The aspect ratio, calculated by dividing the fibre’s length by its equivalent diameter, was in the range of 122–289. DSC analysis (not discussed here) showed that the glass transition region of the fibres was centred at about 50 °C; the melting point and enthalpy were 258 °C and 62 J/g, respectively; and their degradation started at higher temperatures. As regards this last phenomenon, oxidative degradation in air started at about 300 °C, while in nitrogen pyrolytic degradation began at about 360 °C. The ultimate strength and initial elastic modulus were 286 ± 38 MPa and 5.0 ± 0.4 GPa, respectively. The ultimate elongation was about 19%. Water absorption by the fibres at saturation, i.e. exposed to an environment with a relative humidity of 100% (in compliance with UNI EN ISO 62:2008 [44]), was about 15% [45,46]. The trilobal cross-section of the fibres explains their high water absorption. 3.2. Compressive strength and modulus of elasticity The concretes’ average compressive strength at 28 days is shown in Table 2 for the cube and cylinder specimens. The FC was as strong as the plain concrete. Albeit within the error bar, there was even a very slight increment observable in the fcm of the FC samples, probably due to lower w/c ratio of the cement paste induced by the fibres adsorbing a certain amount of water. There was no such small increment in the compressive strength Rcm due to the different resolution of the loading machines used to test the cubes or cylinders. When the compressive load on the FC specimens increased, the fibres played a significant part in the concrete’s lateral tensile strength and, by helping to increase its lateral tensile strength, they could prevent or postpone crack enlargement. When compressive loading begins and increases slowly, microcracking begins too, but debonding at the fibre–matrix interface interferes with the cracks’ further growth. The growing cracks need to form a mesh to make the concrete fail, and this demands a greater compressive load before the concrete fails. The good compressive strength of fibre-reinforced concrete is explained by the fibres inhibiting cracking in the concrete.

3.3. Flexural and tensile strength

Table 3 Flexural test results.

fIf (MPa) U1 (Nmm) U2 (Nmm) D0 D1

Table 2 Mechanical test results where Rcm and fcm are the average compressive strength for cubes and cylinders, respectively, Em is the average elastic modulus and fctm is the average tensile strength.

Rcm (MPa) fcm (MPa) Em (GPa) fctm (MPa)

FC

C

57 ± 2 39 ± 2 36 ± 2 3.9 ± 0.2

57 ± 1 38 ± 1 37 ± 1 4.4 ± 0.4

Indirect method

5.5 ± 0.2 10792 ± 1083 10775 ± 3249 0.84 ± 0.06 0.25 ± 0.05

5.3 ± 0.3 10787 ± 1044 10856 ± 3281 0.86 ± 0.25 0.25 ± 0.06

the ductility indexes. Fig. 1 shows the load-CTOD curve for all specimens. For the indirect method, we used the CTOD0 value of 0.0025 lm (as suggested in the standard). We measured the average maximum load PIf for the FC specimens as 21502 N using the direct method, while it was 20787 N using the indirect method. Clearly the difference associated with the PIf values between the direct and indirect methods was approximately 3%. The structural synthetic fibres have a low modulus of elasticity, so the concrete’s flexural strength dropped considerably once the first crack had formed. The tests showed that, with further bending after the first crack had developed, the load increased for the FC as well as for the control concrete. We calculated the fIf as 5.5 MPa using the direct method and as 5.3 MPa with the indirect method. We calculated the first-crack strength fIf using the recommended equation:

fIf ¼

P If  l b  ðh  a0 Þ2

ðMPaÞ

ð1Þ

where PIf is the first-crack load, l is the distance between the supports, a0 is the notch depth, b is the beam base, and h is the beam height. We also calculated D0 and D1 (the average ductility indexes), using (2): they were 0.84 and 0.25, respectively, for the direct method, and 0.86 and 0.25 for the indirect method.

8 < D0 ¼ feq ð00:6Þ f If

ð2Þ

: D ¼ feq ð0:63Þ 1 feq ð00:6Þ

feq(0–0.6) and feq(0.6–3) (3) represent the equivalent strength for CTOD in the ranges of 0–0.6 mm and 0.6–3 mm, respectively.

8 l < feqð00:6Þ ¼ bðha

0Þ

l : feqð0:63Þ ¼ bðha

2

U1  0:6

2

U2  2:4

0Þ

ð3Þ

U1 and U2 (4) are the area under the load-CTOD curve in kJ, within the CTOD range of (0–0.6) mm and (0.6–3) mm, respectively.

( The flexural strength values of the prism-shaped samples at 28 days are given in Table 3, where fIf represents the first-crack strength under flexural loading and U1/U2 the areas under the load-CTOD curve, while we calculated D0/D1 as the ductility indexes in compliance with the Italian Standard UNI 11039-2:2003 [38]. We used both the direct and the indirect methods to compare

Direct method

U1 ¼ U2 ¼

R 0:6 0

PðCTODÞ  dðCTODÞ

0:6

PðCTODÞ  dðCTODÞ

R3

ð4Þ

We also calculated the tension resistance (fctm) for both types of sample (Table 2) using the following equation:

fcfm ¼

Pmax  l bh

2

ð5Þ

The mean tensile strength of the control concrete was higher than that of the FC (see Table 2), being 4.4 MPa for the former 3.9 MPa (12% less) for the latter. The load-CTOD curves likewise showed a mean ultimate load-bearing capacity of 22003 and 24808 N for the FC and C samples, respectively. The plain concrete exhibited a brittle behaviour, however, whereas the samples reinforced with nylon fibres were more ductile. According to Italian Standard UNI 11039-2:2003 [38], the analysis performed is unable to indicate a precise type of ductility behaviour (softening, plastic

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Fig. 1. Load-CTOD curve plotted for the four-point flexural tests (a); and an enlargement of the curve at 0.00–0.10 mm (b).

or hardening), because D1 was always below the minimum required value (P0.3). This is not to say that the standard (intended for concrete reinforced with steel fibres) is not applicable to concrete reinforced with fibres with a low elastic modulus, but the polymer fibres used can produce an effective crack-bridging effect because their modulus of elasticity comes closer to that of the concrete matrix. Also, in the light of the investigation conducted by Roesler et al. [47], this can increase the fracture energy at the bottom of the specimen, where the notch is located and the fracture starts. 3.4. Characterisation of aggregates and concretes For the three types of aggregate (i.e. coarse gravel, fine gravel and sand), the thermogravimetric and heat flow curves (not shown here) indicated that no relevant thermal events occurred up to 500 °C. Beyond this point, the typical peak corresponding to decomposition of the carbonates (i.e. dolomite and calcite) was visible [48]. Although the thermal analyses did not show the alpha/ beta transition of quartz at 575 °C, XRD analyses indicated that the fine gravel and sand contained a certain amount of quartz (see Table 4), which may induce internal stress in concrete at temperatures beyond 500 °C, because a 4.5% change in volume is associated with its transition. The thermogravimetric and heat flow curves (not shown here) for the C and FC concretes were quite similar, with signs of free water evaporation up to 100 °C and dehydration of C–S–H, ettringite and calcium aluminate hydrate starting at 140 °C. The main peaks were attributable to the decomposition of Ca(OH)2 at 430 °C, and of magnesium and calcium carbonates at around 790 °C [49]. Due to the small percentage of fibres in the concrete, there was no evidence of glass transition or melting in the thermal analyses. An important parameter for a thermal energy storage device is heat capacity. Fig. 2a and b shows the average heat capacity vs. temperature curves for the FC and C samples, recorded during Table 4 XRD quantitative results (vol%, err. ±0.1%).

Dolomite Calcite Quartz Albite

Sand

Fine gravel 8–12

Coarse gravel 12–25

60.0 20.9 9.3 9.8

77.0 20.6 2.4 –

78.3 21.7 – –

two successive scans. The two curves resemble the heat capacity trends during the first two concrete storage system charging steps. As reported by other authors [2,3], the curves of all the samples, recorded during the first scan, show a maximum at 100 °C due to evaporation of the water adsorbed in porosities. The FC samples show a higher heat capacity beyond 250 °C, due to the fibres’ contribution. Unlike the first scan, the second shows the curve of the FC samples with a constant negative slope and lower values. In fact, the samples lost not only water, but fibres too, which could no longer contribute to heat capacity in the second scan. If we disregard the water evaporation occurring in the first scan, the two heat capacity curves for the C samples almost overlap beyond 300 °C. We measured the concretes’ thermal conductivity using the transient hot disk method, conducting the analysis at room temperature on two samples, before and after thermal treatment, up to 300 °C. The fibre-reinforced concrete had a thermal conductivity of 1.30 ± 0.02 W/m K at 25 °C and 1.16 ± 0.02 W/m K at 300 °C. The control samples had a slightly lower conductivity, i.e. 1.12 ± 0.02 W/m K at 25 °C and 1.02 ± 0.02 W/m K at 300 °C. All these results are consistent with data in the literature [2]. The higher conductivity of the FC samples is attributable to the fibres at the lower temperature measured, and to the greater interconnected porosity left by the fibres burning at the higher temperature. Variations in the density of materials with temperature depend on the coefficient of thermal expansion (CTE), which is important in the thermomechanical design of a thermal storage device. The CTE of a material for use in thermal storage should coincide with that of the embedded metallic heat exchanger pipe. Differences in the thermal expansion of the two types of concrete were mainly observable in the first scan, shown in Fig. 3: the C samples expanded more than the FC. Heating promoted vaporisation of the liquid pore water and the vapour migrated out of the medium into the atmosphere or towards lower-pressure regions. This migration occurred more easily through the channels where fibres were embedded in the FC. In the absence of fibres (as in the C samples), the pore pressure rose abruptly and caused further expansion, especially in the temperature range between 50° and 100 °C. The thermal strain on FC samples at low temperatures is also considerably reduced thanks to the greater compressive strain on the fibres before they melt. In the other two scans the curves (not shown here) became very similar. It is worth noting that the averaged CTE, 17.2  106 K1, of the FC in the same temperature range was very similar to that of stainless steel, 17.5  106 K1 [50], used commercially for the piping.

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Fig. 2. Specific heat plotted vs. temperature for the FC (a) and C (b) samples.

3.5. Isothermal treatments

Fig. 3. Thermal expansion of FC and C samples during the first scan.

The concrete samples underwent isothermal treatment in air at different temperatures. As shown in Fig. 4a and b for the FC and C samples, respectively, the mass loss increased with time due to dehydration processes, which took place mainly during the first hour of treatment, reaching equilibrium a few hours later. The mass loss increased as the temperature rose, as first the water adsorbed in porosities, then the water chemically combined in the cement hydration products was gradually released. At 150 °C, the initial mass losses of the FC and C samples were 1.2% and 1.1%, respectively. After 4 h, the mass losses settled at about 1.3% and 1.5%. Given that only free water is released up to 150 °C, the hygroscopic nature of the fibres (which absorb a certain amount of water that is only released more slowly) accounted for the lower mass loss recorded for FC samples. At higher treatment temperatures, the FC samples lost more mass than the C samples, but the latter showed signs of spalling. The greater mass loss from FC is due both to the residual water that had been absorbed by the fibres and to their decomposition after softening and melting. The greater mass loss of the FC samples also suggests a greater permeability to gases [11–16] due to the additional porosity left by the fibres’ decomposition. The ESEM findings discussed in the next paragraph confirmed this impression.

Fig. 4. Mass variation under thermal treatment at various temperatures for FC (a) and C (b).

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Fig. 5. ESEM micrographs of the FC sample after 28 days of curing (a–c), and after thermal treatment at 250 °C (d), 350 °C (e) and 450 °C (f).

We examined the concrete specimens under the ESEM after 28 days of wet curing, before and after heat treatments up to various temperatures. All the mixtures revealed a very dense

microstructure. Fig. 5a and b are ESEM micrographs of the fibrereinforced sample before any thermal treatments. In the first, the trilobal cross-section of the fibres appears randomly distributed

Fig. 6. Drying shrinkage of FC (a) and C (b) samples.

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in the cement matrix, whereas the cross-section of two fibres is clearly visible in the second (taken at a higher magnification). To see how thermal treatment affected the microstructure of the concrete around the fibres in the FC, micrographs of a specimen left untreated or treated at 250, 350 and 450 °C are compared in Fig. 5c– f. There was still evidence of fibres in the concrete after a thermal treatment up to 250 °C (Fig. 5d), but the fibres melted on exposure to higher temperatures, leaving a porosity in the bulk of the sample (Fig. 5e and f clearly show the spaces left by the fibres). Microcracks also appeared in the vicinity of the aggregate particles on heating beyond 250 °C, attributable mainly to the thermal cycling adopted for the purpose of measuring mass changes. The porosity left by the fibres prevented the cracks from propagating. 3.6. Drying shrinkage We measured drying shrinkage in compliance with the Italian standard UNI 11307:2008 [37], recording deformations in the prisms at 1, 7, 14, 28, 60 and 90 days; the results are given in Fig. 6. We included a measurement at 14 days, although this is not required by the standard. There was slightly less shrinkage in the FC specimens than in the control concrete [51,52]. It is worth emphasizing that fibres play an important part in containing shrinkage, particularly in the early curing stage. It is also important to note that a lower drying shrinkage is a key factor when it comes to the production of large thermal energy storage devices. 4. Conclusions For this study on the effects of synthetic fibre reinforcement, we manufactured concrete specimens with a 0.35 w/c ratio and investigated their mechanical and thermal properties. We obtained a high-performance concrete thermal storage material using Portland limestone cement, recycled nylon fibres, and dolomite aggregates as raw materials. Our findings indicate that large thermal energy storage (TES) devices for large-scale applications are feasible using this type of concrete, for both technological and economic reasons. The fibre-reinforced concrete suffered less drying shrinkage than the plain concrete, and this is always beneficial because it coincides with a more limited formation of microcracks in the cement paste and a more durable material. The results of our mechanical analyses showed that the added fibres did not significantly affect the concrete’s compressive strength or elastic modulus, nor did they have any negative fallout on other properties. The behaviour of the FC appeared to be slightly more ductile than the plain concrete, due to debonding and pull-out at the fibre–matrix interface lying on the path of crack propagation. The heat capacity and conductivity at 300 °C of the FC were 0.63 J/g K and 1.16 W/m K, respectively (as opposed to 0.81 J/g K and 1.02 W/m K for the reference concrete). When first heated to between 50° and 150 °C, the C samples expanded more than the FC samples. The rising pressure of the vapour from free liquid water in porosities can account for this phenomenon, which was more accentuated in the C samples than in the FC, where the channels in which the fibres are embedded facilitate water migration. The channels, left when the fibres soften above their glass transition temperature, also enable the concrete to bear the thermal strain without forming cracks. At higher temperatures, the fibres flow and deteriorate, leaving further porosities (as shown by ESEM). The better thermal behaviour of the fibre-reinforced concrete was thanks to the hygroscopic nature of the fibres. The FC exhibited a good thermal stability up to 450 °C, with no signs of spalling (observable to some degree in the plain concrete). The porosity left by the fibres melting seems to help prevent the propagation of thermal cracks.

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