Effect of vacuum mixing and curing conditions on mechanical properties and porosity of reactive powder concretes

Construction and Building Materials 209 (2019) 326–339

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Effect of vacuum mixing and curing conditions on mechanical properties and porosity of reactive powder concretes Tomasz Zdeb Cracow University of Technology, Faculty of Civil Engineering, Institute of Building Materials and Structures, Warszawska 24, 31-155 Kraków, Poland

h i g h l i g h t s
The most favourable RPC mix homogenisation conditions have been determined.
Changes to RPC microstructure caused by vacuum mixing have been described.
Changes to RPC microstructure caused by hydrothermal treatment have been described.
The maximum fibres amount introduced during RPC mix production has been determined.
RPC properties with variable: fibre content, curing, vacuum mixing have been verified.

a r t i c l e

i n f o

Article history: Received 9 October 2018 Received in revised form 6 February 2019 Accepted 11 March 2019

Keywords: Reactive powder concrete Vacuum mixing Heat treatment Porosity Mechanical properties High volume steel fibre reinforced concrete

a b s t r a c t The paper contains quantitative characteristics of changes in the porosity and mechanical properties of RPCs caused by applying different pressures during concrete mixing, i.e. atmospheric pressure and pressure reduced to 500 and 40 mbar. The effect of curing conditions was determined as well: the impact of curing in water (Tmax = 20 °C), steaming (Tmax = 90 °C) and autoclaving (Tmax = 250 °C) on the porosity, compressive strength and flexural tensile strength of the materials tested. In addition, the maximum volumetric proportion of 6 and 14 mm steel micro-fibres which can be introduced in a standard manner, i.e. during the production of the RPC mix, was determined depending on their length. Besides the basic mechanical characteristics, tests of mechanical properties of RPCs containing fibres included the determination of the following parameters: rLOP (limit of proportionality), rMOR (modulus of rapture), Wf (work of fracture) and the toughness indices. Test results clearly confirm the considerable impact of pressure during RPC mixing and curing conditions on the durability and mechanical properties of the hardened composites. The maximum volumetric proportion of micro-fibres introduced during mixing was 8%. Steel fibre reinforcement was found to be more effective when longer fibres were used. Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction Despite the fact that reactive powder concretes (RPCs) have been present in the construction material market for more than 20 years, they still represent one of the most important development directions for cementitious composites. Their excellent mechanical properties alongside considerable durability have been confirmed repeatedly. However, they still exhibit certain shortcomings which limit the full use of their potential. One of them is their porosity, which is characteristic of cementitious materials. Results of numerous tests, including those carried out in the CSTB laboratory on RPC composites produced on an industrial scale, have confirmed that their total porosity amounts

E-mail address: [email protected] https://doi.org/10.1016/j.conbuildmat.2019.03.116 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

to around 10% on average [1]. The scale of possibilities offered by reducing porosity in terms of increasing compressive strength is illustrated by Balshin’s model quoted in [2–4]:

r ¼ r0  ð1  PÞn where: r – compressive strength at porosity P [MPa]; r0 – compressive strength at zero porosity [MPa]; P – porosity [–]; A – experimentally determined constant [–]. The pore size distribution is accounted for in the value of the exponent n, which ranges from 2 in the case of highly porous cementitious composites, e.g. containing light aggregates, up to 6 where the model achieves a better fit for lower-porosity composites. In the case of cement pastes, the value of the exponent may reach as high as 14. Therefore, it can be estimated that lowering

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porosity by just one percent may result in a strength increase of around a dozen MPa. In addition to the obvious treatments which are applied as standard when producing RPCs, such as limiting the amount of mixing water (even to amounts below the stoichiometric level) or introducing silica fume which contributes to reducing porosity, it is also very important to pack the dry ingredients of the material as densely as possible. So far, many concepts have been developed to reduce the void content of these fine-grained cementitious composites; these have been published, among others, by authors such as Furnas [5], Aim and Goff [6], Stovall [7], De Larrard [8], Andreasen [9] or Funk [10]. One method which allows further reduction in the amount of pores in the RPC material, and is also used on an industrial scale, is the vacuum mixing or pressing of the concrete mix during its production [2,11–14]. The studies conducted by Dils et al. [15–17] demonstrated that compared to mixing at atmospheric pressure, the application of pressures of 500 and 50 mbar when producing the mix resulted in a reduction in the amount of pores, both determined by the pressure method in the mix and by the Rapid Air test in the hardened material (by 30% and 60%, respectively). Unfortunately, the reduction in porosity was not reflected in improved mechanical properties, as compressive strength only increased by 10% while flexural tensile strength remained unchanged. A natural consequence of the typical composition of reactive powder concretes in which the binder content ranges from 50% to 60% by volume are significant shrinkage deformations. These emerge particularly in composites which are cured under natural conditions. Tam et al. [18] consider that the main reasons for high overall RPC shrinkage include the small amount of mixing water as expressed by the w/b ratio, and the amount of superplasticiser used. For w/b values in the range from 0.17 to 0.40 and the amount of SP (in relation to cement mass) ranging from 2.0% to 3.5%, linear shrinkage after 130 days of curing under laboratory conditions ranged from 400 to 1200 lm/m. It should be added here that observations of shrinkage deformation only started at the time when test specimens were demoulded. Cwirzen et al. [19] also demonstrated the effect of ambient relative humidity on the overall shrinkage value of RPCs. The tests were conducted at ambient temperature and at relative humidities of 65% and 85% during a 360-day period. The authors observed that the shrinkage of the material stored in the more humid environment was lower by half. On the other hand, Van Tuan et al. [20] studied exclusively the autogenous shrinkage of RPCs by determining changes in linear dimensions of the specimen without mass exchange with the environment, in accordance with the procedure described in ASTM C1698. Material which had a water-binder ratio of 0.18 and whose composition included 20% of silica fume exhibited autogenous shrinkage as high as 2000 lm/m. Since the order of magnitude of the thermal expansion coefficient of cementitious composites is virtually the same as that of shrinkage deformations, some authors predicate that it could be possible to reduce RPC shrinkage by applying the frequently used hydrothermal treatment [21,22]. According to Staquet [22], raising the ambient temperature to 42 °C during the 100 initial hours of curing reduces shrinkage deformations by around half. Another serious shortcoming of RPCs is their brittleness. Owing to the microstructure of RPC composites, they exhibit virtually no plastic deformation. As a result of reducing the porosity of the composite in all possible known ways, decreasing the amount of mixing water to below the stoichiometric amount and the presence of pozzolanic additives which amount to 25% of cement mass on average, the resulting composite matrix is composed mostly of hydrated calcium silicates with a very densely packed structure [23]. The result is a stiffer matrix and its deformability values approach that of the micro-aggregate, and this in turn boosts the elastic modulus value of the entire composite, which typically

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ranges from 50 to 70 GPa [24,25]. Additionally, since the maximum grain size generally does not exceed 500 lm, the distribution of actual stresses in loaded RPC material is very homogeneous in comparison with ordinary concrete. As a consequence, loaded material exhibits virtually no cracking almost up to the level of its strength. Therefore, stress-strain relationships both for tensile and compressive strength tests are almost linear throughout the entire range [25]. In order to endow composites with pseudoplastic fracture characteristics steel fibres have become the standard component. 1.1. Research significance In the presented studies, the main purpose was to mitigate the aforementioned basic RPC shortcomings. As described in the Introduction, to date many studies have been carried out on the impact of various technological treatments during the manufacture of RPC composites on their mechanical properties or durability [11–17]. However, the results presented in this article in the first part deal with both mechanical and durability properties of RPCs without fibres subjected to both vacuum mixing and hydrothermal curing. Thus the principle objectives of this part of the study were as follow:
determination of the simultaneous effect of vacuum mixing at variable pressure and hydrothermal treatment (steaming and autoclaving) on the total porosity and porosity distribution determined by the helium pycnometry and MIP methods respectively of RPC without fibres,
determination of the simultaneous effect of vacuum mixing as well as hydrothermal treatment on the compressive strength and the tensile strength at bending of RPC without fibres.
influence of the basic mixing parameter, i.e. rotational speed, on the efficiency of obtaining the proper consistency of the RPC mix has been also verified. Steel fibres are usually incorporated to the RPC matrix mainly in order to enhance its ductility. They are added to concrete mix in amounts ranging from 2% to 3% and generally not exceeding 5% by volume or in the case of SIFCON technology (Slurry Infiltrated Fibre Concrete) their proportion may reach from 10% to even 25% by volume [2,26–29]. Therefore, the main purposes of the second part of the research conducted were:
determination of the maximum volumetric proportion of steel micro-fibres of varying lengths which can be introduced into RPCs when producing the composite mix along with mechanical properties of hardened composites,
application of technological treatments that have the most intense impact on the reduction of porosity and improvement of mechanical properties of the RPC reinforced with steel micro-fibres. 2. Materials and the experiment 2.1. Materials The ingredients used for the production of RPC specimens included: CEM I 52.5 Portland cement, silica fume, 0/0.2 mm ground quartz, 0/0.5 mm quartz sand, polycarboxylate superplasticiser and 6 and 14 mm long straight steel micro-fibres. Details of individual components are summarised in Tables 1–4. The component proportions of RPC matrix without micro-fibres presented in Table 5 were determined in such a manner as to obtain the appropriate consistency of the concrete mix as well as to ensure that the hardened material is as compacted as possible and exhibits minimum porosity as a result. The method for designing the composition of the material is described in detail in [30]. In turn, the maximum volumetric proportion of steel fibres with lengths of 6 and 14 mm, which amounted to Vf,6 = 8% by volume and Vf,14 = 6% by volume, respectively, was determined so

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T. Zdeb / Construction and Building Materials 209 (2019) 326–339 2.2. Specimen preparation and curing conditions

Table 1 Properties, chemical and phase composition of CEM I 52.5 Portland cement. Properties Initial setting time [min] Final setting time [min] Specific surface area [cm2/g] Compressive strength after 2 days [MPa] Compressive strength after 28 days [MPa]

2.2.1. Mix production Concrete ingredients were mixed in a horizontal mixer, which was equipped with a three-phase 1.1 kW electric motor that allowed the speed to be regulated by means of an inverter, and also with a vacuum pump which made it possible to set the desired reduced pressure value in the mixing chamber. Maintaining a constant value of pressure in the mixing chamber was possible owing to the choke installed around the axis between the chamber and the clutch as well as an additional regulating valve placed on the mixer’s lid (see Fig. 1). The volume of the mixing chamber was 5 dm3, and the mix was prepared in portions of a constant volume of 2 dm3, i.e. it constituted 40% of the volume of the mixing chamber. The mixing procedure was carried out in consecutive stages:

130 220 4100 34.5 70.8

Chemical composition [wt%] CaO – 63.7; SiO2 – 20.8; Al2O3 – 4.1; Fe2O3 – 5.6; SO3 – 2.9; MgO – 0.9; Na2Oe – 0.51; Cl- – 0.009 Phase composition [wt%] C3S – 59.09; C2S – 17.97; C3A – 8.12; C4AF – 6.38

Table 2 Particle size distribution and chemical composition of ground quartz and quartz sand. Properties

Ground quartz

Quartz sand

Dmax [mm] D50 [mm] Specific surface BET [m2/g] Density [g/cm3] Polymorphic modification SiO2 [%] Al2O3 [%] Fe2O3 [%]

200 16 0.8 2.65 b-quartz 99.0 0.3 0.05

500 110 0.04

98.5 0.8 0.03

2.2.2. Specimen moulding and curing Immediately after the mixing process had been completed, series of three specimens with dimensions of 40  40  160 mm each were moulded. Since the RPC mix has thixotropic properties, its formation, especially in the presence of a highvolume fibres fraction, requires the use of high frequency vibrations. For this purpose, a Vebe vibrating table was used, which made it possible to precisely fill the molds with the concrete mix. The specimens were pre-cured under conditions which limited the evaporation of water, i.e. in a chamber with a relative humidity of 98% for the time shown in Fig. 2. Subsequently, individual specimen series were cured under three different sets of conditions:

Table 3 Properties and chemical composition of silica fume. Properties Specific surface area [m2/g] Density [g/cm3]

22.4 2.23

Chemical composition [wt%] SiO2 – 94.06; Al2O3 – 0.74; Fe2O3 – 0.78; CaO – 0.06; MgO – 0.49; Na2Oe – 1.43; SO3 – 0.63; LOI – 0.74


curing in water at 20 °C for 28 days, (W)
steaming at 90 °C, according to the cycle shown in Fig. 2a, (S)
autoclaving at 250 °C and at a pressure of 40 bar, according to the cycle shown in Fig. 2b, (A)

Table 4 Steel micro-fibre properties. Length [mm] Diameter [mm] Modulus of elasticity [GPa] Tensile strength [MPa] Density [g/cm3]


pre-mixing of dry components for 1 min;
adding the entire assumed amount of water with an admixture of superplasticiser;
setting the pressure in the mixing chamber at one of three levels: atmospheric pressure (P1013) and pressure reduced to 500 mbar (P500) and 40 mbar (P40);
further mixing until proper consistency was obtained after a time which depended on the mixer’s rotational speed. The rotational speeds (controlled by an inverter) were 70, 140, 175, 210 and 280 rpm.
in the case of mixtures with steel micro-fibres, after obtaining a suitable consistency of a mixture containing only mineral constituents together with water and superplasticizer, in order to avoid the clumping problem dosing of fibres was carried out in small portions with simultaneous slow mixing to disperse them initially in the concrete mix. This procedure was repeated until the maximum assumed amount of fibres was introduced. Subsequently, mixing was continued for 1 min at the established pressure and rotational speed values.

6 175 210 2200 7.76

14 200

that the resulting mix had a proper consistency to mould the specimens using the vibrating table. The proportions of RPC matrix constituents remained constant so increasing amount of micro-fibres caused decrease in amount of other components and the same time decrease mix consistency but still allowing to keep the microfibres homogeneously dispersed in the entire volume of the composite. Steel fibre reinforcement was added in steps of 2% by volume and hence its proportion by mass was respectively from 0 to 155, 311, 466 and 621 kg/m3. All the compositions of manufactured RPCs are presented in Table 5.

The aforementioned hydrothermal treatment parameters were adopted as optimal based on the test results presented in [31]. Mechanical properties of RPCs cured in water were tested after 28 days, while the specimens subjected to hydrothermal treatment were tested immediately after the material had been cooled to ambient temperature. In the test results presented below, symbols similar to the following one have been adopted, consisting of four elements: RPM70/P500/W/F6-2. The symbol indicates specimens produced from a mix homogenised at a rotational speed of 70 rpm, at a pressure of 500 mbar in the mixer chamber, cured in water and containing steel micro-fibres with a length of 6 mm, which amounted to 2% by volume. 2.3. Research methods 2.3.1. Testing of fresh concrete During the production of the RPC mix, the total energy consumed by the mixer in kWh was measured, which depended on the rotational speed of the mixer and on the mixing time required until the proper consistency of the RPC mix was achieved. The flow diameter of the mix was measured using a flow table for mortars according to EN 1015-3, while air content was determined using the pressure method according to EN 1015-7.

Table 5 Composition of the tested RPCs [kg/m3]. Component

Mass fraction

CEM I 52.5 R cement Silica fume 0/0.20 mm ground quartz 0/0.50 mm quartz sand Water Superplasticiser Steel micro-fibre 6 or 14 mm Steel micro-fibre volume fraction Vf,6;14 [%]

903 181 312 729 217 19.4 0 0

885 177 306 714 212 19.0 155 2

867 173 300 699 208 18.6 311 4

849 170 294 685 204 18.0 466 6

831 166 287 670 199 18.0 621 8

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Fig. 1. Mixer equipped with an inverter which enables rotational speed to be adjusted in the range from 0 to 280 rpm and with a vacuum pump which enables pressure to be adjusted in the range from 40 mbar to atmospheric pressure.

Fig. 2. Hydrothermal treatment cycles applied during composite curing: a) steaming, b) autoclaving.

2.3.2. Porosity measurement During the examination of physical properties of the hardened composite, in order to determine the total porosity of the materials, their bulk (envelope) density qbulk was determined with the powder picnometry method using the Micrometrics GeoPyc 1360 device and the true (skeleton) density qtrue was determined using a helium pycnometer Quantachrome Ultrapycnometer 1200e. The total porosity Ptot was calculated according to the following relationship:

 Ptot ¼



1

qbulk  100½%v ol: qtrue

Pore size distribution in the material was determined with the mercury intrusion porosimetry (MIP) method using the Quantachrome Poremaster 60. This device allows mercury to be introduced into the specimen within the pressure range from 0.1 to 400 N/mm2 and thus allows the volumetric proportion of pores to be determined in the range from 3.75 nm to around 0.25 mm. 2.3.3. Mechanical properties measurement Basic mechanical properties were tested as follows: flexural strength (ff) using the 3-point method (see Fig. 3) on three 40  40  160 mm prisms, and compressive strength (fc) was tested on six 40  40  40 mm cubes cut from prism halves after bending. In addition to the basic mechanical properties, in the case of composites which contained micro-fibres, additional mechanical parameters were determined on the basis of the recorded load-deflection relationship. As in the case of [32], coefficient values were determined according to standards [33,34]. Thus, stress (r), deflection (d) and toughness (T) values were calculated for characteristic points: LOP (limit of proportionality – assigned to the first crack of the matrix) and MOR (modulus of rapture – assigned to the point when material softening is observed after LOP). Additionally, the following coefficient values were calculated: d5 at deflection of 3dLOP, d10 for d = 5.5dLOP and d20 for d = 10.5dLOP. The dimensionless d5–d20 coefficients were calculated as toughness at beam deflections ranging from 3dLOP to 10.5dLOP, respectively, with reference to the TLOP value. In addition, the total work of fracture Wf was determined in deflection range from 0 to dMAX = 5 mm for 6 mm fibres and from 0 to dMAX = 10 mm for 14 mm fibres. All mechanical properties of RPCs were tested using the Zwick/Roell Z100 (see Fig. 3) and Zwick/Roell Z1600 machines, which enabled testing with a constant increase in strain over time and were additionally equipped with an extensometer that enabled measurements accurate to 0.1 lm. Microstructure observations were conducted using a Zeiss EVO10M scanning microscope.

Fig. 3. Universal testing machine for testing strength at 3-point bending, equipped with an extensometer to measure beam deflection.

3. Test results and analysis 3.1. Mixing parameters and mix properties In order to determine the most favourable homogenisation parameters (mixing time, mixer speed and pressure in the mixing chamber), mixes with a volume of 2 dm3 without the addition of steel micro-fibres were tested. Initially, the mixing took place at atmospheric pressure. After the best combination of mixer speed

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and mixing time had been determined, the effect of pressure on mix properties was established. On the basis of the observations conducted, a virtually linear relationship was found between the rotational mixing speed and the time required to obtain the desired consistency (flow diameter around 25 cm). Doubling the speed caused the mixing time to be reduced by around half in the speed range from 70 to 280 rpm. Interestingly, energy consumption tended to decrease when rotational speed increased up to 210 rpm. Exceeding this value resulted in significantly greater energy consumption by the mixer – despite the short homogenisation time at 280 rpm, the amount of energy consumed rose by more than 50% in relation to the minimum value, which was obtained when mixing at 210 rpm (see Table 6). When the ingredients were mixed at the lowest speed of 70 rpm, this slightly lowered the flow diameter of the mix, which resulted in its limited tendency to deaerate spontaneously. This explains the increased number of technological pores in this case, which amounted to 2.6% by volume. A further increase in speed to 140 rpm increased the flow diameter of the mix and significantly (by almost 30%) reduced the amount of trapped air. Unfortunately, the presence of large amounts of superplasticizer, which reduces the surface tension of mixing water, results in the RPC mix turning into foam at higher mixing speeds, and an increase in porosity occurs despite a simultaneous increase in liquidity. Additionally, the aforementioned increase in liquidity to values above 28 cm may prove inappropriate for composites containing steel micro-fibres, since high steel density may result in their sedimentation during the moulding of specimens. Summing up, on the basis of test data presented in Table 6, the most favourable mixing parameters are present when the mixer operates at a rotational speed of 175 rpm for 3 min, as mixing under these conditions consumes the least energy while at the same time making it possible to obtain mixes with a minimum air content and satisfactory consistency. Vacuum mixing at 500 bar allowed the content of air trapped in the mix to be reduced by 40%. When the pressure was reduced to 40 mbar, the air content was so low that it became virtually unmeasurable using the method described in EN 1015-7. Further, mixes containing different amounts of the two types of fibres were tested. Those mixes were homogenized for a fixed rotational speed of 175 rpm, while the total mixing time was about 5 min. The variable factor was the pressure in the mixing chamber: first atmospheric, and then 500 and finally 40 mbar. In the mixer used, the increasing resistance during mixing, which was caused by both vacuum mixing and by the presence of fibres, resulted in

a negligibly small increase in the amount of energy consumed by the mixer motor (see Table 6). The presence of steel micro-fibres resulted in a clear increase in the aeration of the concrete mix, which was at a similar level for both fibre lengths of 6 and 14 mm, while longer fibres decreased the flow diameter of the mix considerably more. When mixes including 6 mm fibres were produced, it was possible to introduce those fibres in amounts of up to 8% by volume, which reduced the flow diameter to 17 cm. For 14 mm fibres, it was possible to introduce a slightly smaller amount (6% by volume) with flow diameter decreased to 13 cm. From the point of view of the amount of micro-fibres successfully introduced into the RPC matrix and the verified homogeneity of their distribution in the material, the reinforced RPC is similar to SIFCON composites [13]. Exemplary pictures taken while the flow measurement of mixtures without and with a maximum amount of 6 and 14 mm fibers are presented in Fig. 4. In addition, the Fig. 5 shows highly contrasted images of the cross-sectional area (40  40 mm) of all hardened RPC composites containing steel micro-fibres in the entire tested range of volume fraction in both length cases. 3.2. Density, total porosity and pore size distribution RPCs without micro-fibres added were subject to density and total porosity tests. The mix was homogenised for 3 min at a constant mixer rotational speed of 175 rpm. The variable was the pressure in the mixing chamber: atmospheric or reduced to either 500 or 40 mbar. The specimens formed from mixes produced under different conditions were cured in water or under low pressure steaming or autoclaving conditions (see Section 2.2.2). All the aforementioned RPC variants were also subject to mercury intrusion porosimetry (MIP) tests, but the analysis only covered the composites produced under extreme pressure conditions during mixing, i.e. at atmospheric pressure and at 40 mbar. True (helium) density of materials cured under the same conditions remained virtually unchanged irrespective of the pressure used during mixing, only differing by around 0.5–1.5%. Very small changes were also observed between materials cured in water and using steaming, but with an indistinct tendency for the helium density to decrease as the curing temperature increased. As can be seen from the bar chart in Fig. 6, the autoclaving process reduces the helium density value irrespective of mixing conditions. This effect can be attributed to an increase in the degree of cement hydration under these hydrothermal conditions. This fact has been confirmed by many researchers, especially in initial stages of

Table 6 Mixing parameters and technological properties of the RPC mixes tested.

RPM70/P1013/F0-0 RPM140/P1013/F0-0 RPM175/P1013/F0-0 RPM210/P1013/F0-0 RPM280/P1013/F0-0 RPM175/P500/F0-0 RPM175/P40/F0-0 RPM175/P1013/F6-2 RPM175/P1013/F6-4 RPM175/P1013/F6-6 RPM175/P1013/F6-8 RPM175/P500/F6-8 RPM175/P40/F6-8 RPM175/P1013/F14-2 RPM175/P1013/F14-4 RPM175/P1013/F14-6 RPM175/P500/F14-6 RPM175/P40/F14-6

Mixing time [min]

Energy consumed [kWh]

Air content [%] vol.

Flow table [cm]

10 5 3 2.5 1.5 3

0.20 0.12 0.08 0.07 0.11 0.08

5

0.13

2.6 1.9 1.7 3.6 4.2 1.1 0 3 3.5 3.5 3.7 2.1 0 2 2.5 3.7 2.9 0

23 25 25 28 29 25 23 21 20 18 17 17 16 21 18 13 13 12

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Fig. 4. RPC mixture flow according to EN 1015-3 a) without fibres, b) with 6 mm fibres 8% vol. and c) with 14 mm fibres 6% vol.

Fig. 5. Cross-section of all hardened RPC composites containing 6 and 14 mm micro-fibres in the entire tested range of volume fractions Vf.

hydration [35–37]. Since the specific density of clinker grains is typically around 3.15 g/cm3, while that of hydrated calcium silicates oscillates around 2.0 g/cm3 [38], the true density of the RPC composite subjected to autoclaving is reduced compared to the density of the composite cured in water as well as cured under steaming conditions. The trend of changes in bulk density depending on curing conditions remains analogous to that of changes in helium density. However, a clear increase in qbulk was observed as pressure during mixing decreased. When specimens produced

Fig. 6. Impact of curing conditions and vacuum mixing on bulk density, specific density and porosity.

from mixes at atmospheric pressure and at 40 mbar were compared, this increase amounted to around 5%. On the basis of the test results presented, it can be concluded that irrespective of the mixing pressure applied, an increase in curing temperature reduces the total porosity of the RPC composite. In the case of steaming, this reduction is small and amounts to at most 7%. However, the autoclaving process has the excellent effect of reducing porosity by 50%. Reasons for this phenomenon may include the increased amount of binder hydration products, also contributed to by the increase in silica solubility (regardless of whether it is in crystalline or amorphous form) along with the increase in temperature [4]. Further, and most importantly, vacuum mixing results in a very significant reduction in the total porosity of the hardened composite. For materials cured in water, the application of pressure reduced to 500 and 40 mbar instead of atmospheric pressure results in a reduction in porosity by 30% and 60% respectively, for steamed materials porosity is reduced by 40% and 60%, respectively, and after autoclaving, a reduction by 50% and by almost 80% was observed, respectively. Mercury intrusion porosity (MIP) tests made it possible to determine changes with respect to the total porosity of the RPC, but also with respect to pore size distribution. Total porosity determined by the mercury porosimetry method is generally lower than that determined using the method described above. This is related to the type of medium which fills voids in the composite. As can be seen in Fig. 7, the total porosity of composites cured in water and subjected to steaming is similar, with a slight tendency to decrease as the curing temperature increases. It is only the autoclaving process that clearly reduces this value by around 25% for both anal-

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ysed mixing pressures, i.e. atmospheric pressure and 40 mbar. On the other hand, lowering the mixing pressure causes a reduction in porosity of around 60% irrespective of subsequent curing conditions. The application of vacuum mixing during the production of the RPC results mainly in eliminating the air trapped in large pores whose size cannot be measured using the MIP method. However, a composite subjected to the same curing conditions at the same time is characterized by a constant binder hydration degree. Therefore, the same amount of cement, silica fume and water reaction products will have to fill pores in the vacuumed material with a smaller diameter. Analysing porosity distribution according to the classification given in [39] and [40], i.e. gel micropores <4.5 nm, meso-pores 4.5–50 nm, middle capillary pores 50– 100 nm and large capillary pores >100 nm, it can be concluded that the prevailing RPC pore fractions are meso-pores (ranging from 47% to 68%) and gel micropores in the range from 21% to 37%. However, vacuum mixing results in a reduction in the amount of mesopores and capillary pores by 6% to 12% depending on the curing method, which in turn increases the share of gel pores by the same amount. As mentioned above, the increase in temperature during binder hydration also results in an increase in silica solubility. Thus, the increasing amount of hydration products further tightens the structure of the material, which also explains the increase in gel pore and meso-pore fractions at the expense of capillary pores as hydrothermal treatment temperature increases (see Fig. 8). On the other hand, the reduction in gel pore content at the same mixing pressure where the steaming and autoclaving processes are used can be explained by an increase in the tendency for the CS-H phase to crystallise, which is observed in particular during high-pressure hydrothermal treatment. Moreover, according to [4] precipitation of the C-S-H phase at elevated temperature may results in decrease in the amount of bounded water. This can be also confirmed by the results obtained by [41,42], where the authors describe changes in the C-S-H phase morphology at elevated temperature. Under such conditions, the C-S-H globules may agglomerate and reduce or even destroy the interlayer space. 3.3. Mechanical characteristics The flexural strength test was carried out in each case on three 40  40  160 mm specimens and the load-deflection relationship was recorded continuously. In turn, compressive strength was determined as the average of six measurements on 40  40  40 mm cubes cut from prism halves after the flexural strength test. RPC composites that did not contain micro-fibres but were subject to vacuum mixing during production were tested. Three pressure levels were applied, i.e. atmospheric pressure, 500 and 40 mbar, and three curing methods were used: in water (W), during steaming (S) and autoclaving (A). In addition, the impact of micro-fibre proportion and length as well as of curing conditions on flexural strength was determined alongside the determination of toughness values at characteristic points, work of fracture and compressive strength. Values of the aforementioned mechanical properties for all RPC configurations are presented in Table 7. 3.3.1. Flexural and compressive strengths Fig. 9 presents the variability in compressive and flexural strength values of composites without fibres, produced at different pressure values during mixing and subjected to curing in water, steaming and autoclaving. The results obtained were highly homogeneous as the coefficient of variation of compressive strength did not exceed 6%, while for flexural strength it did not exceed 14%. Irrespective of conditions under which the mix is prepared, both compressive and flexural strengths of the hardened composite increase both after steaming and autoclaving. In the case of low-

pressure hydrothermal treatment, the increase in fc compared to the material cured in water ranged from 1% to 8% depending on the conditions under which the mix was produced. On the other hand, the increase in flexural strength is more pronounced, ranging from 4% to 34%. The autoclaving process had a much better effect on the improvement in RPC mechanical properties than lowpressure steaming. In the case of compressive strength, an increase ranging from 20% to almost 40% was measured, while flexural strength increases from 75% to more than 100% were recorded where the mix had been prepared at a pressure of 40 mbar. The reduction in pressure during mix production from atmospheric levels to 500 mbar results in clear advantages in terms of mechanical properties, since irrespective of the subsequent curing conditions, an increase in compressive strength of around 10% is observed, while that in flexural strength reaches up to 50%. On the other hand, a reduction in mixing pressure to 40 mbar results in a further increase in compressive strength – by about 25% on average. At the same time, flexural strength increases by 6% in the case of steaming and by up to 70% when autoclaving is used. The obtained mechanical properties correspond well with the results of reduced capillary porosity (above 50 nm) in favour of meso-pores and gel pores (below 50 nm) due to the use of hydrothermal treatment. The larger the pore diameter, the stronger the reduction in mechanical properties can be expected. In addition, as reported in [43], the impact of pores with a diameter below 20 nm on the strength of cement composites is negligible. When analysing the compressive strength of the above composites in the context of their total porosity, it can generally be assumed that Balshin’s model describes this relationship well. However, the equation in the graph (see Fig. 10) demonstrates that the exponent value at n = 9.2 exceeds those found for ordinary cementitious composites which contain aggregate. Reasons for this include significantly reduced porosity as a result of vacuum mixing and the modification of the RPC composite interfacial transition zone as a result of hydrothermal treatment. This applies in particular to autoclaved composites where the surface dissolution of quartz grains, which are inert under normal conditions, occurs. The analysis of basic mechanical characteristics was also conducted for materials containing variable amounts of micro-fibres (from 0 to 8% by volume) of various lengths (6 and 14 mm). The impact of the minimum mixing pressure, i.e. 40 mbar, on those properties was analysed as well. Variation coefficient values were slightly higher than in the case of composites without fibres added and were 6% and 19% for compressive strength and flexural strength, respectively. The results obtained are shown in the bar chart in Fig. 11. As concerns compressive strength, the effectiveness of the lowpressure steaming process is at a similar level irrespective of the length of the micro-fibres used and does not exceed 10% compared to composites being cured in water. However, in the case of autoclaving, this effectiveness is much more pronounced, especially for composites containing longer fibres. The positive effects of hydrothermal treatment on mechanical properties of RPC mention many authors, among others [25,31,36,44–48]. Where composites with 6 mm micro-fibres which have been cured in water and autoclaved are compared, the increase in compressive strength oscillates around 50%. When 14 mm micro-fibres are used, this increase reaches up to 70%, especially in the Vf range from 2% to 4% by volume (155–310 kg/m3). It should be noted that in general, the increase in fibre content, regardless of fibre length, has a positive effect on compressive strength for all curing conditions. Moreover, longer fibres prove more effective compared to 6 mm fibres. While the increase in strength is similar for maximum contents of both micro-fibre lengths, both the number and the mass proportion of 14 mm fibres are much lower (466 versus 621 kg/m3). Materials cured in water and subjected to steaming at maximum

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Fig. 7. Cumulative distribution of RPC pore sizes for curing under different hydrothermal conditions: a) homogenisation at atmospheric pressure; b) homogenisation at a pressure of 40 mbar.

Fig. 8. Distribution of the volumetric proportion of pores in RPCs subjected to vacuum mixing and curing under different hydrothermal conditions.

micro-fibre content showed an increase in compressive strength by slightly over 20% in relation to composites without fibres. In the case of autoclaving, this increase is greater and varies depending on micro-fibre length. The maximum content of shorter fibres causes fc to rise by 30% (to 324 MPa), while longer fibres (at a maximum content) increase it by up to 40% (to 345 MPa). In addition, it should be mentioned that the increase in compressive strength under all curing conditions is virtually linear as a function of the volumetric proportion of micro-fibres (see Fig. 11). Linear regression analysis demonstrated that the value of the R2 determination coefficient for the linear model was not lower than 0.83 for all curing conditions and micro-fibre lengths considered. Generally, it can be concluded that for 6 mm fibres and with specimens which are cured in water and steamed, when fibre amount rises by 1% by volume, compressive strength increases by around 5 MPa. For autoclaving, this increase is twice as large. On the other hand, an increase in the proportion of 14 mm fibres by 1% results in an fc increase of more than 7 MPa for curing in water and steaming, while in the case of autoclaving the increase is more than 15 MPa. The effect of vacuum mixing under 40 mbar with a maximum micro-fibre proportion, i.e. 8% or 6% by volume, is unclear, since depending on curing conditions, compressive strength increases vary from 3 to 7% irrespective of the length of the fibres. In the case of flexural strength, an increase in this characteristic is also observed after low-pressure steaming. In general, this increase varies from a few to around 10% for the entire analysed

range of volumetric proportions of reinforcing fibres. Just as in the case of compressive strength, the autoclaving process has a much greater influence on the ff value. However, no effect of changing micro-fibre length is visible here. In other words, when we compare composites with the same amounts and lengths of fibres which are cured in water and subjected to autoclaving, the increase in strength is about 20%. Similarly as for compressive strength, the increase in fibre content, regardless of fibre length, has a positive effect on flexural strength for all curing conditions. Interestingly, the most pronounced increase in flexural strength attributable to the increasing proportion of fibres was observed for composites cured in water. In the case of short fibres at their maximum proportion, i.e. 8% by volume, this increase was almost fourfold compared to the same material without fibres. When the maximum amount of 14 mm fibres was added, an increase of more than five times was recorded. A similar trend could be observed for materials subject to hydrothermal treatment. Both after steaming and autoclaving, 6 mm fibres at maximum proportions resulted in an approximately threefold increase in flexural strength compared to materials without fibres. The use of 14 mm fibres resulted in a more than fourfold increase. The obtained test results correspond well with the results described in [49,50], where a pullout test of single fibres from, among other things, the RPC matrix was carried out. As the authors claim, the increase in fibre embedment length in the RPC matrix results in an increase in the pull-out peak load value. This confirms more favourable distribution of stresses before material cracking, and hence an increase in flexural strength. Similarly to compressive strength, flexural strength is also virtually linearly related to the volumetric proportion of steel micro-fibres for all curing conditions. Linear regression analysis of the relationship between the volumetric proportion of fibres and flexural strength demonstrated that the value of the R2 determination coefficient was not lower than 0.88 for all cases considered. Generally, it can be concluded that irrespective of curing conditions, an increase in the amount of fibres with a length of 6 mm by 1% by volume results in a flexural strength increase ranging from around 3 to 4 MPa. On the other hand, an increase in the proportion of 14 mm fibres by 1% results in a strength increase of around 8% to 9%. Vacuum mixing at 40 mbar with maximum proportions of short and long fibres, i.e. 6% or 8% by volume, respectively, does not yield a significant increase in flexural strength. For all curing conditions and irrespective of micro-fibre length, the increase in flexural strength does not exceed a few percent, which is less than the natural range of variation for this characteristic. The exception are RPCs subjected to steaming with a maximum proportion of 14 mm fibres where a decrease in flexural strength was recorded in fact.

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Table 7 List of mechanical properties of the RPCs tested.

RPM175/P1013/W/F0-0 RPM175/P500/W/F0-0 RPM175/P40/W/F0-0 RPM175/P1013/W/F6-2 RPM175/P1013/W/F6-4 RPM175/P1013/W/F6-6 RPM175/P1013/W/F6-8 RPM175/P40/W/F6-8 RPM175/P1013/W/F142 RPM175/P1013/W/F144 RPM175/P1013/W/F146 RPM175/P40/W/F14-6 RPM175/P1013/S/F0-0 RPM175/P500/S/F0-0 RPM175/P40/S/F0-0 RPM175/P1013/S/F6-2 RPM175/P1013/S/F6-4 RPM175/P1013/S/F6-6 RPM175/P1013/S/F6-8 RPM175/P40/S/F6-8 RPM175/P1013/S/F14-2 RPM175/P1013/S/F14-4 RPM175/P1013/S/F14-6 RPM175/P40/S/F14-6 RPM175/P1013/A/F0-0 RPM175/P500/A/F0-0 RPM175/P40/A/F0-0 RPM175/P1013/A/F6-2 RPM175/P1013/A/F6-4 RPM175/P1013/A/F6-6 RPM175/P1013/A/F6-8 RPM175/P40/A/F6-8 RPM175/P1013/A/F14-2 RPM175/P1013/A/F14-4 RPM175/P1013/A/F14-6 RPM175/P40/A/F14-6

fc [MPa]

ff [MPa]

LOP

MOR

rLOP [MPa]

dLOP [mm]

TLOP [Nm]

rMOR [MPa]

dMOR [mm]

TMOR [Nm]

182 202 239 178 199 202 226 231 184

10.6 13.6 14.7 20.4 31.5 37.5 40.8 41.1 41.1

– – – 20.4 23.5 35.0 39.6 37.0 29.1

– – – 0.33 0.30 0.44 0.51 0.50 0.35

– – – 1.2 1.3 2.9 3.9 3.6 2.4

– – – 16.9 31.5 37.5 40.8 41.1 41.1

– – – 0.33 0.46 0.52 0.55 0.64 0.83

197

55.4

42.9

0.52

4.6

55.4

227

59.7

50.5

0.63

6.5

235 190 218 241 198 202 205 231 242 203 214 235 251 250 273 289 265 308 309 324 336 310 335 345 354

61.7 14.3 14.8 15.2 24.7 34.3 35.9 45.1 45.6 41.6 55.8 60.1 58.7 18.6 28.6 31.6 27.3 40.4 46.5 52.2 52.7 44.5 67.2 75.1 76.0

51.9 – – – 21.0 32.6 34.9 43.7 42.4 33.9 47.6 53.0 48.5 – – – 23.9 39.2 45.3 45.8 50.9 38.8 57.8 63.2 70.1

0.60 – – – 0.32 0.46 0.42 0.51 0.51 0.51 0.65 0.70 0.69 – – – 0.37 0.53 0.57 0.57 0.68 0.58 0.73 0.89 0.86

6.9 – – – 1.2 2.8 2.9 4.3 4.1 3.9 6.8 8.2 7.2 – – – 1.6 3.8 4.7 4.7 6.4 5.0 8.8 11.7 12.2

Fig. 9. Impact of curing conditions and vacuum mixing on compressive and flexural strengths of RPC specimens without fibres added.

3.3.2. Toughness The following analysis was carried out on the basis of the load– deflection relationships obtained for RPC composites cured in water, subjected to steaming and autoclaving, with variable volumetric proportions of micro-fibres of different lengths, which were also additionally subjected to vacuum mixing during mix production (P = 40 mbar). The test results presented in Fig. 12 are

d5 [–]

d10 []

d20 []

Wf [Nm]

– – – 1.2 2.9 3.9 4.6 5.8 8.9

– – – 3.5 5.4 4.1 3.9 4.3 5.1

– – – 4.8 7.8 5.6 5.4 6.0 9.4

– – – 5.5 9.5 6.5 – 7.2 14.9

0.3 0.6 0.7 6.7 12.6 18.9 24.7 26.2 51.6

0.87

11.2

5.1

8.9

13.3

74.8

59.7

0.97

14.0

4.3

6.3

7.9

55.9

61.7 – – – 24.7 34.3 35.9 45.1 45.6 41.6 55.8 60.1 58.7 – – – 27.3 40.4 46.5 52.2 52.7 44.5 67.2 75.1 76.0

0.87 – – – 0.47 0.55 0.48 0.57 0.61 0.74 0.97 0.91 1.26 – – – 0.45 0.58 0.65 0.75 0.74 0.78 1.08 1.27 1.08

12.7 – – – 2.5 3.8 3.5 5.3 5.8 7.1 13.1 13.5 18.6 – – – 2.4 4.5 6.1 8.2 7.7 8.5 16.9 21.1 18.4

4.2 – – – 4.9 3.7 3.9 3.6 4.2 4.7 4.1 3.9 4.8 – – – 3.7 3.7 3.5 4.3 3.8 4.1 4.5 4.6 4.0

6.7 – – – 7.5 4.7 5.1 4.8 5.5 8.6 6.0 5.9 7.8 – – – 5.1 4.7 4.4 5.8 5.2 6.4 6.8 7.1 5.7

9.4 – – – 8.9 5.2 5.7 – – 13.1 7.7 7.8 10.6 – – – 5.5 – – – – 8.7 8.5 9.1 7.1

73.0 0.7 0.7 0.7 11.1 14.5 16.6 24.3 26.3 64.1 56.4 68.3 82.2 0.9 1.4 1.5 9.2 19.4 21.9 30.2 37.2 49.5 81.6 107.9 91.8

Fig. 10. Relationship between the compression strength of RPCs without fibres added and their total porosity as determined by powder and helium pycnometry.

strength-deflection curves of the samples that represent this relationship best, while all parameter values calculated are averaged and listed in Table 7. It may be stated in general that the rLOP parameter increases together with the amount of micro-fibres used, analogously to flexural strength ff. This confirms that at greater proportions of steel micro-fibres, tensile stress is distributed more favourably through these fibres before the composite cracks. This effect is explained by

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Fig. 11. Impact of mixing pressure and of the length and volumetric proportion of steel micro-fibres on RPC compressive and flexural strengths: a) curing in water; b) steaming; c) autoclaving.

the much higher elastic modulus of the reinforcement in relation to the matrix. Hydrothermal treatment also positively affects this parameter, because low-pressure steaming results in an increase in rLOP by up to 38%, while autoclaving raises it by up to 66%. These results confirm the effect of curing conditions on the mechanical properties of RPCs without steel fibre reinforcement, which were discussed in the previous chapter. The deformability of the composites tested until the LOP point is very similar. In other words, the slope angle of the load–deflection curve within the linear range and thus the elastic modulus under bending are similar irrespective of curing conditions and the amount and length of the micro-fibres added. Thus, the other parameters (dLOP and TLOP)

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which describe the behaviour of the composite in this range change in proportion to rLOP. The basic criterion of suitably selected steel fibre reinforcement parameters, i.e. that the rMOR/ rLOP ratio be greater than 1, is met for virtually all tested fibre quantities and lengths. The only exception is the RPM175/P1013/ W/F6-2 composite where the value of this ratio is 0.83. Interestingly, the maximum reinforcement of composite after matrix cracking was achieved using 14 mm micro-fibres at 2% by volume. This relatively small amount of fibres was insufficient to result in the distribution of tensile stresses in the composite before cracking, but it significantly increased post-critical stress values. Therefore, the rMOR/rLOP ratio was the highest for precisely these composites, and it varied depending on curing conditions (curing in water – 1.41, steaming – 1.23, autoclaving – 1.15). The d5, d10 and d20 parameters are calculated as the proportion of work which must be performed until the assumed deflection of a prism reinforced with fibres is achieved in relation to the work which results in the composite cracking. From the experiments conducted, it follows that increasing the volumetric proportion of micro-fibres (Vf) results in a proportional increase of the TLOP parameter and also of the work required to pull the fibres from the RPC matrix until the assumed deflection is reached. Hence the d5, d10 and d20 parameters remain virtually constant at a constant fibre length irrespective of the curing conditions applied. For 6-millimeter fibres, d5 generally oscillated between 3.0 and 4.0, and d10 ranged from 4.0 to 6.0. For most composites tested, especially those containing high proportions of fibres, the d20 value could not be determined owing to the increase in the dLOP parameter – according to standards, after being multiplied by 10.5, this parameter exceeded the value of deflection measurable for those composites, i.e. 5 mm. On the other hand, composites containing longer (14 mm) fibres showed slightly higher values of the d parameters. For d5, these varied from around 4.0–5.0, and for d10 – from 6.0 to 9.0, while d20 could be determined in all cases analysed and ranged from around 7.0–15.0. The parameter that changed in the most pronounced manner depending on the length and amount of micro-fibres was the work of fracture Wf (see Fig. 13). Compared to a composite without steel fibre reinforcement, short fibres already at a level of 2% resulted in the work of fracture increasing by 10–20 times depending on curing conditions. For composites containing 6 mm fibres, increasing their volumetric proportion within the Vf range from 2% to 8% resulted in virtually linear increases in the work of fracture. For each percent of fibre added, Wf rose by 4–6 Nm depending on curing conditions. The most rapid increase was observed in the case of autoclaving. In the case of 14 mm fibres, a general increase in Wf is also observed as the volumetric proportion of fibres rises, but it is difficult to describe this relationship as linear for any given curing conditions. The increase in toughness along with the increasing fibre embedment length and matrix strength has been confirmed, among others, in [49,51] during the single-fibre pull-out study. The authors observed this relationship regardless of the setting time in water and the type of smooth or hooked fibre used. This is clearly reflected in significantly higher values of post-critical stress and thus Wf when RPC contains longer micro-fibres. In addition, it should be noted that the Wf parameter proved to be more sensitive to the vacuum mixing process applied than flexural strength ff. The production of the RPC mix at a pressure of 40 mbar resulted in an increase in the average value of this parameter which ranged from 6% to 23% depending on curing conditions for short fibres. In contrast, the vacuum mixing of the material which contained the maximum amount of 14 mm fibres (6%) resulted in an increase in the Wf parameter which ranged from 20% to 30%. An exception here was the autoclaved composite where a surprising decrease in the work of fracture was noted.

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Fig. 12. Load-deflection relationships for RPCs reinforced with 6 mm micro-fibres: a) cured in water; b) subjected to steaming; c) subjected to autoclaving; and reinforced with 14 mm micro-fibres: d) cured in water; e) subjected to steaming; f) subjected to autoclaving.

3.4. SEM observations Before the analysis of the microstructure of the RPCs tested is presented, significant macro-scale changes in the composites produced should be noted. Namely, the process of vacuum mixing of fibre-free RPC composites at a pressure of 40 mbar resulted in the removal of virtually all technological pores with sizes that could be observed with the naked eye. The difference in the texture of the composite can also be easily observed at minimum magnifications during SEM observation. Fig. 14 shows fracture surfaces of RPM175/P1013/W/F0-0 and RPM175/P40/W/F0-0 specimens. As a result of intensive mixing at atmospheric pressure and the high viscosity of RPC mixes, during their moulding the technological pores trapped in the material mostly take the form of spherical pores with

diameters not exceeding 1 mm (see Fig. 14a). In a material which was vacuum mixed during production, technological pores are practically absent, and the texture of the material at 100x magnification (see Fig. 14b) is practically free of pores of this type. The observation of the microstructure of the amorphous C-S-H phase demonstrates that irrespective of the subsequent hydrothermal treatments, this phase, which is formed during initial curing at ambient conditions, remains virtually unchanged. All its properties indicate that it is of type IV according to Diamond’s classification [52], i.e. it occurs in the form of spherical agglomerates presenting as compact gel under a scanning microscope [4]. The morphology of the C-S-H phase, cured in water and subjected to autoclaving, at 10,000 magnification is shown in Fig. 15. The photo in Fig. 15b additionally shows an edge of a defect which under auto-

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Fig. 13. Work of fracture for RPCs cured under variable hydrothermal conditions with variable proportions of reinforcing micro-fibres.

claving conditions is filled with whiskers of the C-S-H phase which crystallises inside, mainly in the form of tobermorite and xonotlite. Numerous cross-sections of RPC composites prepared after flexural strength tests in order to obtain cubes for compressive strength tests made it possible to observe a homogeneous distribu-

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tion of steel micro-fibre reinforcement in the entire volume of the material irrespective of the amount and length of the fibres used (see Fig. 5). Additionally, the macro-scale observations carried out during and after the flexural strength testing of composites containing steel micro-fibre reinforcement made it possible to conclude that virtually all possible fracture mechanisms described by Zollo, i.e. mainly fibre pull-out, fibre bridging and fibre/matrix debonding, were present [53]. However, even in the most favourable case the 3 mm anchoring of short fibres and 7 mm anchoring of long fibres proved insufficient to result in fibre rupturing. The adhesion of the composite matrix to steel micro-fibres was excellent irrespective of the curing conditions applied. Within the ITZ (interfacial transition zone), there was no debonding caused by shrinkage or portlandite crystallisation (see Fig. 16). The good adhesion between the matrix and the steel inclusion is also evidenced by numerous traces of the C-S-H phase on the surface of the fibres pulled out of the composite. However, the observed excellent mechanical properties of RPC composites, especially those that have been subjected to autoclaving, among other things, result from increased fibre adhesion to the RPC matrix. According to Beglarigale and Yazıcı [49], the C-S-H phase in the ITZ is congested as a result of its crystallization in the form of tobermorite whiskers.

Fig. 14. Texture of RPC cured in water and mixed at atmospheric pressure (a); at P = 40 mbar (b).

Fig. 15. Morphology of the C-S-H phase cured in water (a) and after autoclaving (b).

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Fig. 16. (a) Surface of the fracture of an RPC specimen subjected to bending: steel micro-fibre tightly surrounded by the RPC matrix, (b) a polished cross-section within the ITZ (interfacial transition zone) between the micro-fibre and the matrix: the excellent adhesion of the C-S-H phase with a structure similar to that occurring in bulk material is visible.

4. Conclusions

Compliance with ethical standards

The production of the RPC mix requires intensive mixing in order to obtain a proper consistency. Under the conditions described, the optimal rotational speed of the mixer from the point of view of the amount of energy consumed, the aeration of the mix and its flow diameter turned out to be 175 rpm with a mixing time of 3 min. Both a decrease and an increase in mixer rotational speed had an adverse amount on the amount of air trapped in the mix as well as the consumption of energy during mixing. Curing RPCs under hydrothermal conditions significantly affects the microstructure of the composites tested. In particular, the autoclaving process lowers both specific and bulk densities of the material, mainly as a result of the increase in the cement hydration degree. The fact that observed bulk and specific density values are similar leads to the conclusion that the total porosity of autoclaved composites is minimal. The reduction in total porosity as a result of autoclaving compared to materials cured in water may reach up to 50%. The vacuum mixing of fibre-free RPCs also significantly affects the total porosity of the hardened composite. The application of pressure reduced to 500 and 40 mbar results in a similar reduction in total porosity both when curing in water and after steaming, which amounts to approximately 40% and 60%, respectively. For autoclaving, this effect is even more pronounced because porosity may be reduced by even 50% and 80% at mixing pressures of 500 and 40 mbar, respectively. Lower pressure during the production of the RPC mix reduces not only the amount of air trapped in the mix, but also the number of capillary pores and meso-pores in the hardened composite. The maximum amount of steel micro-fibres which can be introduced into an RPC composite during mixing depends on their length. For 6 mm fibres, this reached 8% by volume, while 14 mm fibres could only be added in amounts not exceeding 6% by volume. When comparing the properties of materials with the same volumetric proportions of short and long fibres, it should be noted that the effectiveness of 14 mm fibres is clearly greater with respect to all analysed mechanical properties of RPC composites. The increasing volumetric proportion of both shorter and longer micro-fibres results in linear increases of compressive and flexural strengths of RPCs in all cases of considered curing conditions.

Conflict of interest: The author declare that he has no conflict of interest. Acknowledgments The results and paper preparation were supported with the funds of Cracow University of Technology within the framework of the L1/87/2018/DS research project. References [1] CSTB, Béton fibré à Ultra Hautes Avertissement, 2009. [2] C. Shi, Z. Wu, J. Xiao, D. Wang, Z. Huang, Z. Fang, A review on ultra high performance concrete: Part I. Raw materials and mixture design, Constr. Build. Mater. 101 (2015) 741–751, https://doi.org/10.1016/ j.conbuildmat.2015.10.088. [3] S.N. Ghosh (Ed.), Advances in Cement Technology: Critical Reviews and Case Studies on Manufacturing, Quality Control, Optimization and Use, Elsevier, 2014. [4] W. Kurdowski, Cement and Concrete Chemistry, Springer, Netherlands, 2014. doi: 10.1007/978-94-007-7945-7. [5] C.C. Furnas, Grading the aggregates I – Mathematical relations for beds of broken solids of maximum density, Ind. Eng. Chem. 23 (1931) 1052–1058. [6] R. Aim, P. Goff, Effet de paroi dans les empilements désordonnés de sphères et application à la porosité de mélanges binaires, Powder Technol. 1 (1968) 281– 290. [7] T. Stovall, F. De Larrard, M. Buil, Linear packing density model of grain mixtures, Powder Technol. 48 (1986) 1–12. [8] F. De Larrard, T. Sedran, Mixture-proportioning of high-performance concrete, Cem. Concr. Res. 32 (2002) 1699–1704, https://doi.org/10.1016/S0008-8846 (02)00861-X. [9] A.H.M. Andreasen, J. Andersen, Ueber die Beziehung zwischen Kornabstufung und Zwischenraum in Produkten aus losen Körnern (mit einigen Experimenten), Kolloid-Zeitschrift 50 (1930) 217–228. [10] J.E. Funk, D.R. Dinger, Introduction to predictive process control, in: Predict. Process Control Crowded Part. Suspens., 1994, pp. 1–16. [11] A. Naaman, K. Wille, The path to Ultra-High Performance Fiber Reinforced Concrete (UHP-FRC): five decades of progress, in: M. Schmidt, E. Fehling (Eds.), 3rd Int. Symp. UHPC Nanotechnol. High Perform. Construction Mater., Kassel, Germany, 2012, pp. 3–15. [12] H. Detlef, L. Urbonas, T. Gerlicher, Effect of heat treatment method on the properties of UHPC, in: M. Schmidt, E. Fehling (Eds.), 3rd Int. Symp. UHPC Nanotechnol. High Perform. Construction Mater., Kassel, Germany, 2012, pp. 283–290. [13] M. Ipek, M. Aksu, M. Uysal, K. Yilmaz, I. Vural, Effect of pre-setting pressure applied flexure strength and fracture toughness of new SIFCON + RPC composite during setting phase, Constr. Build. Mater. 79 (2015) 90–96, https://doi.org/10.1016/j.conbuildmat.2015.01.023.

T. Zdeb / Construction and Building Materials 209 (2019) 326–339 [14] M. Ipek, K. Yilmaz, M. Sümer, M. Saribiyik, Effect of pre-setting pressure applied to mechanical behaviours of reactive powder concrete during setting phase, Constr. Build. Mater. 25 (2011) 61–68, https://doi.org/10.1016/ j.conbuildmat.2010.06.056. [15] J. Dils, G. De Schutter, V. Boel, Vacuum mixing technology to improve the mechanical properties of ultra-high performance concrete, Mater. Struct. 48 (2015) 3485–3501. [16] J. Dils, G. De Schutter, V. Boel, Influence of mixing procedure and mixer type on fresh and hardened properties of concrete: a review, Mater. Struct. 45 (2012) 1673–1683. [17] J. Dils, G. De Schutter, V. Boel, E. Braem, Influence of vacuum mixing on the mechanical properties of UHPC, in: M. Schmidt, E. Fehling (Eds.), 3rd Int. Symp. UHPC Nanotechnol. High Perform. Construction Mater., Kassel, Germany, 2012, pp. 241–248. [18] C.M. Tam, V.W.Y. Tam, K.M. Ng, Assessing drying shrinkage and water permeability of reactive powder concrete produced in Hong Kong, Constr. Build. Mater. 26 (2012) 79–89, https://doi.org/10.1016/ j.conbuildmat.2011.05.006. [19] A. Cwirzen, V. Penttala, C. Vornanen, Reactive powder based concretes: Mechanical properties, durability and hybrid use with OPC, Cem. Concr. Res. 38 (2008) 1217–1226, https://doi.org/10.1016/j.cemconres.2008.03.013. [20] N. Van Tuan, G. Ye, K. Van Breugel, Mitigation of early age shrinkage of Ultra High Performance Concrete by using Rice Husk Ash Nguyen, in: 3rd Int. Symp. UHPC Nanotechnol. High Perform. Construction Mater., Kassel, Germany, 2012, pp. 341–348. [21] P. Acker, M. Behloul, Ductal technology: a large spectrum of properties, a wide range of applications, in: M. Schmidt, E. Fehling, C. Geisenhansluke (Eds.), Int. Symp. Ultra High Perform. Concr., Kassel, Germany, 2004, pp. 11–23. [22] S. Staquet, B. Espion, Early-age autogenous shrinkage of UHPC incorporating very fine fly ash or metakaolin in replacement of silica fume, in: M. Schmidt, E. Fehling, C. Geisenhansluke (Eds.), Int. Symp. Ultra High Perform. Concr., Kassel, Germany, 2004, pp. 587–599. [23] A. Korpa, T. Kowald, R. Trettin, Phase development in normal and ultra high performance cementitious systems by quantitative X-ray analysis and thermoanalytical methods, Cem. Concr. Res. 39 (2009) 69–76, https://doi. org/10.1016/j.cemconres.2008.11.003. [24] J. Dugat, N. Roux, G. Bernier, Mechanical properties of reactive powder concretes, Mater. Struct. 29 (1996) 233–240, https://doi.org/10.1007/ BF02485945. [25] T. Zdeb, Ultra-high performance concrete – properties and technology, Bull. Polish Acad. Sci. Tech. Sci. 61 (2013) 183–193, https://doi.org/10.2478/bpasts2013-0017. [26] B. Abdollahi, M. Bakhshi, Z. Mirzaee, M. Shekarchi, M. Motavalli, SIFCON strengthening of concrete cylinders in comparison with conventional GFRP confinement method, Constr. Build. Mater. 36 (2012) 765–778, https://doi.org/ 10.1016/j.conbuildmat.2012.06.021. [27] A. Beglarigale, C. Yalçinkaya, H. Yiʇiter, H. YazIcI, Flexural performance of SIFCON composites subjected to high temperature, Constr. Build. Mater. 104 (2016) 99–108, https://doi.org/10.1016/j.conbuildmat.2015.12.034. [28] R. Yu, P. Spiesz, H.J.H. Brouwers, Mix design and properties assessment of Ultra-High Performance Fibre Reinforced Concrete (UHPFRC), Cem. Concr. Res. 56 (2014) 29–39, https://doi.org/10.1016/j.cemconres.2013.11.002. [29] D. Wang, C. Shi, Z. Wu, J. Xiao, Z. Huang, Z. Fang, A review on ultra high performance concrete: Part II. Hydration, microstructure and properties, Constr. Build. Mater. 96 (2015) 368–377, https://doi.org/10.1016/ j.conbuildmat.2015.08.095. [30] T. Zdeb, The Impact of Composition and Technology on Selected Properties of Reactive Powder Concretes, Cracow University of Technology, 2009. [31] T. Zdeb, An analysis of the steam curing and autoclaving process parameters for reactive powder concretes, Constr. Build. Mater. 131 (2016) 758–766, https://doi.org/10.1016/j.conbuildmat.2016.11.026. [32] D.J. Kim, A.E. Naaman, S. El-Tawil, Comparative flexural behavior of four fiber reinforced cementitious composites, Cem. Concr. Compos. 30 (2008) 917–928, https://doi.org/10.1016/j.cemconcomp.2008.08.002.

339

[33] ASTM C1018 Standard Test Method for Flexural Toughness and First-Crack Strength of Fiber-Reinforced Concrete (Using Beam With Third-Point Loading), 1997. [34] ASTM C1609 Standard Test Method for Flexural Performance of FiberReinforced Concrete (Using Beam With Third-Point Loading), 2012. [35] H. Zanni, M. Cheyrezy, V. Maret, S. Philippot, P. Nieto, Investigation of hydration and pozzolanic reaction in reactive powder concrete (RPC) using 29Si NMR, Cem. Concr. Res. 26 (1996) 93–100. [36] A. Cwirzen, The effect of the heat-treatment regime on the properties of reactive powder concrete, Adv. Cem. Res. 19 (2007) 25–33, https://doi.org/ 10.1680/adcr.2007.19.1.25. [37] J.I. Escalante-Garcia, Nonevaporable water from neat OPC and replacement materials in composite cements hydrated at different temperatures, Cem. Concr. Res. 33 (2003) 1883–1888, https://doi.org/10.1016/S0008-8846(03) 00208-4. [38] Y. Tanaka, T. Saeki, K. Sasaki, Y. Suda, Fundamental study on density of C-S-H, Cem. Sci. Concr. Technol. 1 (2009) 70–76, https://doi.org/10.14250/ cement.63.354. [39] X. Chen, S. Wu, J. Zhou, Experimental study and analytical model for pore structure of hydrated cement paste, Appl. Clay Sci. 101 (2014) 159–167, https://doi.org/10.1016/j.clay.2014.07.031. [40] Q. Zeng, K. Li, T. Fen-Chong, P. Dangla, Pore structure characterization of cement pastes blended with high-volume fly-ash, Cem. Concr. Res. 42 (2012) 194–204, https://doi.org/10.1016/j.cemconres.2011.09.012. [41] I. Maruyama, H. Sasano, Y. Nishioka, G. Igarashi, Strength and Young’s modulus change in concrete due to long-term drying and heating up to 90 °C, Cem. Concr. Res. 66 (2014) 48–63, https://doi.org/10.1016/j. cemconres.2014.07.016. [42] I. Maruyama, Y. Nishioka, G. Igarashi, K. Matsui, Microstructural and bulk property changes in hardened cement paste during the first drying process, Cem. Concr. Res. 58 (2014) 20–34, https://doi.org/10.1016/j. cemconres.2014.01.007. [43] A.M. Neville, Włas´ciwos´ci betonu, fourth ed., Polski Cement, 2000. [44] C. Tam, V.W. Tam, Microstructural behaviour of reactive powder concrete under different heating regimes, Mag. Concr. Res. 64 (2012) 259–267. [45] H. Yazıcı, E. Deniz, B. Baradan, The effect of autoclave pressure, temperature and duration time on mechanical properties of reactive powder concrete, Constr. Build. Mater. 42 (2013) 53–63, https://doi.org/10.1016/ j.conbuildmat.2013.01.003. [46] H. Yazıcı, M.Y. Yardımcı, S. Aydın, A.S ß. Karabulut, Mechanical properties of reactive powder concrete containing mineral admixtures under different curing regimes, Constr. Build. Mater. 23 (2009) 1223–1231, https://doi.org/ 10.1016/j.conbuildmat.2008.08.003. [47] C.M. Tam, K.M. Ng, V.W.Y. Tam, Optimal conditions for producing reactive powder concrete, Mag. Concr. Res. 62 (2010) 701–716, https://doi.org/ 10.1680/macr.2010.62.10.701. [48] T. Zdeb, Autogenous healing effect of Ultra-High Performance Cementitious composites, J. Adv. Concr. Technol. 16 (2018) 549–562, https://doi.org/ 10.3151/jact.16.549. [49] A. Beglarigale, H. Yazici, Pull-out behavior of steel fiber embedded in flowable RPC and ordinary mortar, Constr. Build. Mater. 75 (2015) 255–265, https://doi. org/10.1016/j.conbuildmat.2014.11.037. [50] V.M.C.F. Cunha, J.A.O. Barros, J.M. Sena-Cruz, Pullout behavior of steel fibers in self-compacting concrete, J. Mater. Civil Eng. 22 (2010) 1–9, https://doi.org/ 10.1061/(ASCE)MT.1943-5533.0000001. [51] T. Abu-Lebdeh, S. Hamoush, W. Heard, B. Zornig, Effect of matrix strength on pullout behavior of steel fiber reinforced very-high strength concrete composites, Constr. Build. Mater. 25 (2011) 39–46, https://doi.org/10.1016/ j.conbuildmat.2010.06.059. [52] S. Diamond, E.E. Lachowski, On the morphology of type III C-S-H gel, Cem. Concr. Res. 10 (1980) 703–705. [53] R.F. Zollo, Fiber-reinforced concrete: an overview after 30 years of development, Cem. Concr. Compos. 19 (1997) 107–122, https://doi.org/ 10.1016/S0958-9465(96)00046-7.