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Creep properties of recycled tyre rubber concrete
Construction and Building Materials 209 (2019) 126–134
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Technical note
Creep properties of recycled tyre rubber concrete D.V. Bompa ⇑, A.Y. Elghazouli Department of Civil and Environmental Engineering, Imperial College London, UK
a r t i c l e
i n f o
Article history: Received 2 January 2019 Received in revised form 26 February 2019 Accepted 11 March 2019
Keywords: Rubberised concrete Recycled rubber Creep Mechanical properties
a b s t r a c t This paper investigates the creep response and long-term strength properties of unconfined and FRPconfined concrete materials incorporating relatively high proportions of recycled tyre rubber particles. The high-strength reference conventional concrete, from which the rubberised concrete is derived, is also examined for comparison purposes. After discussing fundamental characteristics of creep behaviour, this study describes an experimental investigation in which three groups of concrete specimens are subjected to sustained uniaxial compressive stress, in the range of about 20% of the estimated strength, for a period of over a year. The test results indicate that both confined and unconfined rubberised concrete materials tend to develop higher creep coefficients by about 53% and 20%, respectively, in comparison to their reference conventional concrete. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction Creep effects may influence the overall structural response through changes in internal member actions. Hence, an in-depth understanding and reliable prediction of such effects are essential for the practical application of novel construction materials. The need for innovative applications for the reuse of waste tyre components led to a significant number of investigations focusing on the properties of rubberised concrete (RuC) materials in which a proportion of mineral aggregates is replaced with recycled rubber particles. Rubberised concrete is a relatively weak material in compression due to the softer nature of rubber particles compared to mineral aggregates, yet, in contrast, it is intrinsically more deformable than conventional concrete materials (CCM) both in terms of lateral and axial deformation [1]. Moreover, compressive strengths comparable to those of CCM can be recovered through external confinement measures, such as fibre reinforced polymer (FRP) jackets or tubular steel elements, whilst further enhancing the deformation capacity. The effectiveness of such confinement approaches depends on the thickness of the confinement and the rubber content within the concrete material [2]. Also, the energy dissipation and inherent damping typically increase for RuC in comparison to CCM [1–3]. More recent investigations also focused on the structural response of reinforced rubberised concrete (RRuC) members provided with various types of confinement (i.e. internal hoops, exter-
⇑ Corresponding author. E-mail address: [email protected] (D.V. Bompa). https://doi.org/10.1016/j.conbuildmat.2019.03.127 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.
nal steel tubes or fibre reinforced polymer sheets), and subjected to combined cyclic lateral displacement and co-existing axial load up to 20% of the squash capacity [4–6]. In comparison with specimens without external confinement, confined RRuC members tend to exhibit higher level of energy dissipation in the range of 9% for 2 layers of FRP confinement and below 10% of total aggregate replacement [4], and by a factor up to 2.7 for 3 layers of FRP and 60% rubber content [5,6]. Additionally, RRuC members develop low cyclic degradation in comparison with conventional RC specimens [6], having the potential to be integrated in isolation or energy dissipation systems to mitigate the effects of seismic loads and other dynamic actions. Creep is defined as the increase in strain in the direction of sustained stress which is obtained by removing the instantaneous, shrinkage and any thermal strain from the total strain. It occurs in most construction materials including concrete, steel, FRP and rubber as a function of their properties. In conventional concrete, the total creep is divided into basic creep and drying creep [7,8]. The former is a movement governed by an elastic structure within the paste. The latter is controlled by an oriented loss of loosely bonded hydrate water and is typically referred to as the creep exceeding the basic creep of sealed concrete [7,8]. The moisture conditions determine the level of drying creep, similarly to the concurrently developing load-independent drying shrinkage, as a response of the skeletal structure of the cement paste to capillary tensions, sorption stresses and external loads [9]. The creep level is controlled mostly by the aggregates’ stiffness, whilst the fineness of binders has a relatively minimal impact [10,11]. For the same type of aggregates used, high strength mixes
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Nomenclature fc fc0 fc,28 d fc,ti frc fcc fcrc ti Ec N
e e1 eccu
compressive strength compressive strength of reference concrete compressive strength at 28 days compressive strength at the time i (in days) unconfined rubberised concrete strength confined compressive strength confined rubberised concrete strength time modulus of elasticity, applied axial load strain compressive strain ultimate confined strain
provided with fly ash and micro-silica tend to have a denser microstructure, hence develop smaller creep deformations in comparison to conventional mixes for the same stress-to-strength ratios [11,12]. This is attributed to the formation of the new Tobermorite gel as a result of the secondary hydration between the calcium hydroxide and silica at room temperature, rather than to the fineness of the binders [13]. The primary factor influencing specific creep of conventional and fly ash concretes (20% and 40% of total binders), cured under a wide range of temperatures (20 °C, 50 °C and 90 °C) and subjected to loads within the elastic creep regime (33% fc) is the quantity and microstructure of calcium silicate hydrate (CASAH) gel [14]. For both concretes without and with low fly ash content, a higher temperature produces more C-S-H gel and a lower specific creep in comparison to curing under room temperature [14]. It is worth noting that Tobermorite is often used in thermodynamical calculations to represent the pole of the most evolved calcium silicate hydrate (CASAH) [15]. Unsurprisingly, concrete with relatively soft aggregates develops more deformation after a given time under load [8,10]. For example, for concrete specimens loaded to 30% of their axial capacity, a full replacement of coarse natural aggregate by recycled aggregate increase the creep deformation by 51% [16]. Also, the specific creep of polystyrene aggregate concrete (PAC) with an aggregate replacement of 20–80% by PAC is strongly influenced by the polystyrene content, with specific creep values between 101 and 880% of the corresponding reference concrete[17]. PAC with relatively large proportions of polystyrene aggregate is a lightweight material with good thermal and insulation properties and its use in structures can reduce the self-weight of the structure as well as the loads on the foundation, in comparison to conventional concrete [17]. Considering that for the same replacement ratio, the reduction in compressive strength of PAC is about half that of RuC, it is expected that the creep-time response of RuC would also be affected. Recycled tyre rubber typically includes about 25% carbon black and polymers in the range of 40–55%, whereas the remaining constituents are softeners and fillers [18]. As recycled rubber particles result from a variety of sources, their long-term response is dependent on the properties and ratios of their constituents. For such materials, as for concrete, creep is directly related to stress and time. The majority of creep occurs in a relatively short amount of time, minimizing its long-term effect [19]. The shape of the rubber creep-time curve resembles that of concrete, although, when plotted on log scales, for very short and for very long times it deviates from linearity [20]. Flexural creep tests on recycled guardrail posts incorporating rubber indicate that the post flexural stiffness was much lower, and the tendency to creep was much greater in comparison to conventional counterparts [21].
eccu1 eccu2 ec1,1 ec2,1 ecr ein erc1,1
u(ti)
r r/fc r/fcrc r/frc
confined ultimate axial strain confined ultimate lateral strain unconfined crushing strain unconfined lateral strain at crushing creep strain strain at applied stress unconfined crushing strain of rubberised concrete creep coefficient applied axial stress, stress to strength ratio stress-to-confined strength ratio stress-to-unconfined strength ratio
Limited tests on creep response of FRP-confined concrete indicate that due to a sealing effect produced by the FRP, the concrete core only has basic creep, whilst the drying creep and shrinkage are minimised or eliminated [22]. Creep tests on aramid FRP-wrapped cylinders under a stress-strength ratio of 30% show that axial load is primarily carried by the concrete core while the contribution from FRP is relatively negligible [23]. Although a direct comparison with unconfined concrete elements was not presented [23], the shape of the creep-time curves of the FRP-confined concretes are similar to those of unconfined concrete reported in the literature [13]. As noted above, FRP laminates generally play a beneficial role in the long-term response of confined concrete elements, yet the type of fibre (carbon, glass, aramid) may have a different contribution to creep, particularly aramid FRP, which is a material that creeps considerably [24]. Long-term effects can have a significant influence on the overall structural response through an increase in internal actions [22,25,26]. Movements due to creep and shrinkage tend to limit the dimensions of reinforced concrete buildings or long-span bridges requiring expansion joints and/or bearings. FRP-Confined RuC elements could be incorporated as discrete link elements acting as isolation or energy dissipation systems against large movements as well as vibrations in such structures. In order to avoid possible failures under service load, the superposed creep effects of the FRP-confined RuC constituent materials need to be assessed in advance [22]. Although previous studies indicate reliable structural behaviour under short-term loading, practical application of such materials and elements is constrained by the lack of studies on the long-term response of unconfined and confined RuC. To this end, this paper focuses on the creep response and the long-term strength properties of RuC with 60% rubber replacement ratio of both fine and coarse mineral aggregates by volume, with and without FRP confinement, as well as of the high-strength reference CCM from which the RuC was derived. 2. Experimental programme Eighteen cylindrical samples were prepared for assessing the response of unconfined and FRP-confined RuC as well as of the reference high-strength CCM from which the RuC was derived, under sustained stress over a period of over a year. In previous studies it was shown that a stronger influence on mechanical properties, particularly the compressive strength [3,27,28], was typically observed for small rubber content and tended to stabilise as the rubber content increased. A higher benefit in terms of deformation properties is obtained from relatively higher volumetric rubber replacement ratio (e.g. 60%, referred to as R60). The term volumetric rubber replacement ratio depicts the pro-
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portion of the fine and coarse aggregates replaced by rubber particles, by volume. From a structural point of view, such composites benefit from relatively large rubber contents (e.g. 40–60%) as these mobilise supplementary confinement and enhance the deformability levels. Additionally, as the strength reduction rate reduces with an increase in rubber content, a higher replacement ratio would also add benefits to the environment through reuse of a larger quantity of waste materials. There is very limited data on mixes in which the replacement ratio is higher than 60%, primarily since above this limit the physical characteristics are strongly governed by the rubber [1] and because the concrete workability and slump decrease drastically [28]. Hence, long-term tests on RuC with 60% replacement would represent a reliable upper bound of aggregate replacement to obtain workable concrete materials. Also, results from creep tests with 0 and 60% replacement ratios could be interpolated to obtain qualitative information regarding creep response of RuC with intermediate rubber content. On the other hand, two FRP layers (referred to as R60F2 for 2 layers in conjunction with 60% rubber content) offered a good balance between the strength recovered due to the presence of the rubber and offered significant enhancement in deformation capacity. The axial crushing strain of FRP-confined concrete was enhanced by a factor of 24.5 for R60, reaching values above 3.0% in comparison to the crushing strain of the corresponding unconfined concrete [6]. In terms of dilation, the ultimate lateral strains of FRP-confined concrete were 15.9 times higher for R60 concrete (above 2%), in comparison to the lateral strain at crushing of the unconfined concrete [6]. Hence, these material and confinement configurations were adopted for long-term testing at room temperature in a controlled environment for more than a year. Implicitly, the key observations from this study are limited to possible practical configurations that adopt similar properties to those considered herein. To expand the applicability of rubberised concrete in practice, further studies could involve a wider range of replacement ratios, as well as an in-depth investigation into the environmental and UV effects, whilst testing could be performed at low and high temperatures and their performance would be assessed after several years.
2.1. Materials and mix proportions High strength cement CEM I 52, 5 conforming to BS EN 197-1 [29,30], mineral aggregates, admixtures and tap water were used in the reference normal concrete mix with a target compressive strength of 70 MPa (R00). The sand (<5 mm) and gravel (5– 10 mm) had a moisture content of about 5 and 3% of the bulk aggregate weight, respectively, while the specific gravity was 2.65 for both constituents [31]. The fineness modulus was 2.35 for fine aggregates, and 5.88 for coarse aggregates. The cement
Table 1 Concrete mixtures per cubic meter.
Cement (kg) Fly ash (kg) Microsilica (kg) Aggregates (kg) Rubber (kg)
5–10 mm 0–5 mm 0/0.5 mm 0.5/0.8 mm 1/2.5 mm 2/4 mm 4/10 mm 10/20 mm
Admixtures (l) w/c (-)
R00 425 – – 1001 820
– – 7.6 0.35
R60 350 42.5 42.5 400 328 16.5 16.5 49.5 66.0 33.0 148.5 7.6 0.35
composition included on average: 3.3% Sulfate SO3, <0.10% Chloride Cl, 0.7% Alkali Eq Na2O, 55% Tricalcium Silicate C3S, 20% Dicalcium Silicate C2S, 9.5% Tricalcium Aluminate C3A and 8% Tetracalcium Aluminoferrite C4AF [30]. From the total quantity of cement listed in Table 1 for the reference concrete mix, in the rubberised concrete, 15% was replaced by EN 450-1 fineness category S fly ash [32,33] and 15% was replaced by undensified Grade 940 microsilica with a minimum 90% SiO2 [34]. Microsilica and fly ash, primarily added to improve workability, segregation and slump, and to optimise the particle packing of the mixture, improve the strength and concrete flowability. Also, polycarboxylate superplasticizer admixtures were added to improve the workability of the mixes [35,36]. The rubberised concrete was produced using a blend of rubber particles with smaller sizes (up to 10 mm) from car tyres, and larger sizes (up to 20 mm) from truck tyres. As depicted in Fig. 1a, they were supplied in the following size ranges: 0–0.5 mm, 0.5– 0.8 mm, 1.0–2.5 mm, 2–4 mm and 4–10 mm, and were used in the concrete mix in the 5%, 5%, 15%, 20% and 10% ratios of the total added rubber content, respectively. The remaining 45% consisted of particles with sizes in the range 10–20 mm. The car and truck tyre rubber had a specific gravity of 1.1 and water absorption of 7% and 1%, respectively. The particle size distribution of mineral aggregates, rubber and blend determined following EN 933-1:2012 [37] are depicted in Fig. 1. The rubber particles were used in the condition received from the manufacturer without pre-treatment. All mineral aggregates and half of the water were mixed together for 1 min, followed by the added rubber for another 1 min. The binders, admixtures and remaining water were then mixed all together for a minimum of 3 further minutes. After being placed in plastic forms, the concrete was subsequently compacted using a vibrating table until the air content in the fresh concrete was at a minimum. Fresh concrete properties were determined by means of a slump cone. The slump was 125 mm for R00 and 50 mm for R60. To ensure appropriate curing, two days after
Fig. 1. a) Rubberised concrete constituents b) Sieve analysis for mineral aggregates and rubber blend.
D.V. Bompa, A.Y. Elghazouli / Construction and Building Materials 209 (2019) 126–134
casting, the samples were placed in a water tank located in a curing room, and these were removed from the wet environment at 26 days. Two layers of aramid FRP sheets (AFRP) were used to confine rubberised concrete samples using a hand wet lay-up procedure. To achieve this, the AFRP sheets were impregnated with a commercial moisture-tolerant EP epoxy bonding adhesive. Before applying the AFRP confinement, at about 36 h before loading, concrete surfaces were brushed and cleaned to improve the adherence between the existing concrete and the fibre sheets. The aramid fibre (of Grade S&P A120/290) was unidirectional and had 0.2 mm design thickness [38]. The ends of the cylindrical samples were ground and capped with high strength mortar slurry made of CEM I 52,5 [29,30] and tap water (1:1) to ensure flatness of the surfaces and good contact conditions during testing. The dimensions of all samples were accurately measured for appropriate assessment of stress and strain characteristics. Samples used for concrete strength assessment at 28 days were Ø100 in diameter and 196– 202 mm in length, whilst all samples used for creep assessment had a Ø102 diameter and a length which varied from 247 to 254 mm for R00 and 255–256 mm for R60. 2.2. Specimen details and testing arrangement The total number of cylindrical specimens prepared were divided into three groups. The first group of Ø100 200 mm specimens was used to assess the compressive strength at 28 days, which was required to determine the level of loading for creep testing. One-third of the remaining samples (Ø102 254 mm) were used as control specimens to record the free-standing deformations, whilst the remaining third (Ø102 254 mm) were subjected to constant sustained compression for a period of over a year. The specimen dimensions conform to the requirements of
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ISO 1920-9:2009 [39] for the determination of creep of concrete cylinders in compression. The latter were loaded to about 20% of their squash load, determined from the compressive strength assessed at 28 days. It is worth noting that for stress-to-strength (r/fc) ratios below 30% the creep response is typically linear and time-dependent only, while between 30 and 70%, the response becomes non-linear being time and stress-dependent [40]. The compressive strength tests were carried out in displacement control using a four-post displacement control Instron Satec 3500 kN machine. To avoid eccentric loading, the testing arrangement included top and bottom high strength steel transfer plates and a 3D hinge directly connected to the actuator (Fig. 2a). Three displacement transducers were employed around the specimens to record axial deformation between the machine platform and the top transfer plate. A displacement rate of 0.2 mm/min was applied from the machine to the specimens. Besides the displacement transducers used to record member deformations, the load corresponding to the applied displacement was recorded by the machine load cells. Surface strains were recorded by means of a Digital Image Correlation (DIC) system. The DIC system consists of two 5MP cameras provided with 35 mm f/2D lenses [41]. To capture crack propagation and member displacements, high contrast black/white patterns were provided to the front face of the specimens. Before the start of the recording, a calibration procedure was performed to ensure measurement accuracy as required by the manufacturer recommendations. The specimens used for creep assessment were placed in a vertical position in a loading frame, which was able to maintain the required load for the duration of the test (Fig. 2b). The load was applied through a hydraulic ram using prestressed Ø12 mm ties prepared with threaded ends, nuts and steel plates to allow positioning and load transfer, and it was recorded continuously by means of a data logger. At the top and bottom, between the two samples and the steel plates, 3D hinges were positioned to accommodate potential non-even loading. The loads were monitored throughout the testing period using a computer and were maintained throughout testing by tightening the nuts using hand-held spanners. These values never dropped below 2% between two consequent recording times. The non-loaded specimens were placed next to the loaded specimen following the same arrangement. To record long-term axial deformations, all samples were prepared with two steel discs at 150 mm distance along equally located three generatrices of the cylinder. The average deformations under sustained load were measured by means of a dial gauge immediately after loading and at established intervals during a period of over a year [39]. Deformations in non-loaded control specimens were measured concomitantly to assess the level of control shrinkage strains. A purpose-built program was developed in order to collect the data from the digital dial gauge via a COM PC port. Careful tracking of the recordings was followed in order to avoid the introduction of any spurious data. The measurements were performed three times and averaged for each testing sequence, whilst the temperature and relative humidity were found to be in the range of 24–27 °C and 50–60%, respectively.
3. Test results 3.1. Strength properties
Fig. 2. Testing arrangement: a) compression tests, b) creep tests.
Fig. 3 depicts the complete axial and lateral stress–strain (r-e) curves for the materials tested at 28 days after casting, whilst Fig. 4 illustrates the damage kinematics for unconfined and confined RuC. For the unconfined samples (R00, R60), the r-e responses are characterised by an increase in stress with strain until crushing, followed by a softening region. As expected, the response of the
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Fig. 3. Stress–strain response at 28 days a) complete response, b) detail of axial compressive response.
Fig. 4. Damage kinematics for: a) R60, b) R60F2.
reference CCM (R00) was brittle with a sudden drop in stress after the maximum capacity was reached. In contrast, the unconfined RuC (R60) exhibited a softer post-crushing behaviour which typically leads to similar or higher energy dissipation in the postpeak regime as shown previously in other investigations [1,3,21]. Unlike the average test measurements for R00, in which a distinct peak lateral strain is observed, for R60 the peak lateral strain is followed by a much smoother softening. The average test compressive strengths at 28 days (fc,28d), the elastic modulus (Ec), the applied axial load (N) and the corresponding stress to strength ratio (r/fc), the applied axial stress (r) and the strain at the applied stress (ein), are listed in Table 2. The stress values r depicted in Table 2 were determined by dividing the applied axial load (N) by the specimen cross-section area (Ac), whist the strain (ein) was assessed by using Hooke’s law and the
modulus of elasticity (Ec) obtained from the tests. In terms of strength, as observed in Fig. 3 and Table 2, the compressive strength of the reference concrete R00 material fc0 = 79.5 MPa is about 10 times that of the rubberised R60 counterpart (frc = 8.07 MPa), indicating a significant influence from the rubber particles (Table 2). The pre-peak behaviour of the concrete was strongly influenced by the volumetric replacement ratio of the mineral aggregates as shown in a previous study [1]. Naturally, the rubber aggregates are softer and lighter in weight compared to mineral aggregates. Hence, for the RuC, the reduction in stiffness is a function of the stiffness ratio between the cement matrix and inclusions (rubber particles and voids) [1]. For the R60 material investigated in this paper, the elastic modulus is less than a fifth of the corresponding R00 concrete. As observed in Fig. 3a, the axial r-e curve of FRP-confined RuC can be characterised by four (r,e) pairs: (i) zero stress, (ii) a proportionality limit between zero stress and the end of the linear elastic branch, (iii) a transition zone between the proportionality limit and the beginning of the hardening regime, (iv) a hardening regime, and (v) the ultimate condition (fcc,eccu). As illustrated in Fig. 3b, up to the proportionality limit, the stiffness of the confined RuC (R60F2) is similar to that of the unconfined concrete Ec. As the concrete gradually crushes inside the wrapping, the r-e response softens. Beyond this limit, the response under axial compression is governed by the jacket properties. The ultimate state of FRPconfined concrete occurs due to extensive lateral expansion and ultimately fracture of the FRP sheets in the transverse direction. Fig. 3 shows that two layers of external AFRP confinement, added to R60 cylinders, enhanced the confined rubberised concrete compressive strength fcrc = 38.8 MPa to about 5 times the unconfined R60 (i.e. frc = 8.07 MPa), which recovers a significant proportion of the strength of its reference conventional concrete (fc0 = 79.5 MPa). The complete recovery of the compressive strength may be attained by increasing the jacket thickness, yet as the jacket becomes thicker the deformation capacity would reduce [6]. For two layers of confinement, the axial deformation capacity of R60F2 was about 12 times higher than R00, and 24 times higher than R60. On the other hand, the lateral deformation increased by a factor of about 15 in comparison to both R00 and R60. This shows that external confinement, in the form of AFRP,
Table 2 Material and loading details. Concrete
fc,28 d (MPa)
Ec (MPa)
N (kN)
r/fc (-)
r (MPa)
ein (%)
R00 R60 R60F2
79.5 ± 2.58 8.07 ± 0.76 38.8 ± 1.45
41,380 7550 7200
130 16 65
0.21 0.25 0.21
16.5 2.04 8.28
0.038 0.027 0.115
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Fig. 5. Stress–strain response after removal from the creep frame: a) R00, R60, R60F2 non-loaded/control specimens, b) R00, R60, R60F2 loaded specimens; c) details of R60 and R60F2 non-loaded/control specimens, d) details of R60 and R60F2 loaded specimens.
may compensate for the strength loss due to the presence of rubber, whilst enhancing the axial and lateral deformation capacity. One-third of the specimens were tested to failure at 28 days, whilst the remaining two-thirds were tested after the removal from the creep frame. The r-e curve of the R00, R60 and R60F2 control non-loaded specimens and those loaded for a year to 20% axial load are illustrated in Fig. 5, whilst the full details are given in Table 3. As mentioned before, one third of the remaining samples (Ø102 254 mm) were used as control specimens (referred to as ‘control non-loaded in Table 3) to record the free-standing deformations, whilst the remaining third (Ø102 254 mm) were subjected to constant sustained compression for a period of over a year. The latter are referred to as ‘loaded to 0.2 fc’ in Table 3. From the results in the figures and table, it can be observed that the unconfined compressive strength of non-loaded specimens after removal from the creep frame was about 10% and 18% higher than at 28 days for R00 and R60 respectively. These results are largely similar to those available in the literature indicating that creep has a minor influence on the concrete strength for r/fc < 0.7 [40]. To this end, Fig. 6 illustrates the relationship between fc,ti/fc,28 d
and time for non-loaded unconfined R00 and R60 concrete, as obtained from testing throughout a period of two years as part of a wider research programme [3,6,42,43]. Additional compressive strengths for the corresponding non-loaded R00 and R60 specimens described in this paper are also illustrated in Fig. 6. The term fc,ti depicts the compressive strength at the time i (in days), whilst fc,28 d is the strength at 28 days. As observed, the logarithmic trend lines for R00 and R60 are overlapping, indicating a similar increase in unconfined rubberised concrete compressive strength frc with time and suggests that methods typically used for CCM may be directly applied for RuC. As indicated in Table 3, for the FRP-confined specimens there was a slight reduction in compressive strength with time. The reduction in strength was 4.5% for the control specimens and about 9% for the loaded samples. The ultimate deformation capacity was also 13% lower for the control samples and 35% lower for the loaded samples, in comparison to the values at 28 days. It is worth noting that one of the loaded samples had an initial drop in capacity before ultimate deformation was reached due to gradual failure of the wrapping; hence, the strength of this sample is not
Table 3 Strength and strain properties.
Sample
28 days
After removal from the creep frame
Initial
Control non-loaded
fc (MPa)
ec1,1 or eccu1 (%)
ec2,1 or eccu2 (%)
fc (MPa)
ec1,1 or eccu1 (%)
Loaded to 0.2 fc
ec2,1 or eccu2 (%)
fc (MPa)
ec1,1 or eccu1 (%)
R00
a b avg
81.3 77.6 79.5
0.26 0.26 0.26
0.14 0.12 0.13
85.7 89.7 87.7
0.27 0.27 0.27
0.08 0.14 0.12
92.1 84.4 88.2
0.27 0.26 0.26
0.13 0.14 0.13
R60
a b avg
8.6 7.5 8.1
0.13 0.14 0.13
0.16 0.12 0.14
9.4 9.7 9.5
0.12 0.13 0.12
0.13 0.12 0.12
9.7 9.4 9.6
0.15 0.17 0.16
0.11 0.13 0.12
R60F2
a b avg
37.5 41.7 39.6
2.92 3.76 3.34
2.14 2.01 2.08
38.6 37.0 37.8
3.26 2.56 2.91
1.61 1.01 1.31
36.0 30.5 33.2
2.18 1.81 2.00
1.42 1.11 1.26
Notes: a, b – two identical specimens, avg – average, axial strains are depicted with positive sign, lateral strains are depicted with negative sign.
ec2,1 or eccu2 (%)
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Fig. 6. Compressive strength increase versus time (fc,ti/fc,28d – t) ratio for R00 and R60.
Fig. 7. Test creep strain–time curves.
considered in the comparisons. Considering that both fcrc and ecrcu reduced with time, it is suggested that along with inherent experimental variations, the hardening of the epoxy resin may have increased the stiffness of the jacket with time, allowing less dilation under axial loading and resulting in a reduction in the deformation capacity. 3.2. Creep measurements As noted above, creep effects were determined from an average of three experimental measurements on a pair of loaded specimens from which average control/shrinkage strains were removed. Creep-time test results depicted in Fig. 7 indicate that creep strains (ecr) increase rapidly after loading in a non-linear manner and tend to become gradually linear after a period of time, with most of the creep largely developing within the first three months. At a similar r/fc = 0.20–0.25, the creep response of the high-strength conventional concrete (CCM) is lower that of the rubberised concrete (RuC). Although ecr are generally similar at early age, the RuC continues to deform under sustained axial stresses, whilst the CCM tends to settle. Prior to removal from the creep frame, ecr values for RuC were about 55% higher than for CCM, indicating that, as for the case of polystyrene aggregate concrete (PAC) described in Section 1, the presence of rubber increased ecr as a function of r/ fc and level of mineral aggregate replacement. The creep of confined rubberised concrete was found to be significantly of a higher magnitude in comparison to the corresponding unconfined materials at the same stress-to-strength ratio. Importantly, the creep strains of FRP-confined RuC were higher than of RuC by a factor of 2.80. As the stress-to-confined strength
ratio was r/fcrc 0.20 corresponding to a stress-to-unconfined strength ratio of r/frc 1.00, the strain levels are within the crushing ranges of the R60 unconfined material. For r/frc 1.00, the concrete creep was in the non-linear range, which for unconfined concrete materials may result in creep rupture [44]. As the strain of the FRP-confined RuC reached erc1,1 = 0.115%, a relatively ductile unstable crushing propagation I throughout the concrete core. This is illustrated in the panel located at the right-hand side of Fig. 7, which shows that strains in the range of e1 = 0.115–0.170% (r/frc 1.00) correspond to r/fcrc 0.20 for R60F2. The above observations suggest that besides the strain from initial loading (e1 = 0.115%) of R60F2, the concrete core gradually crushed inside the FRP confinement after being loaded, adding to the total mechanical strain about De0.05%. After deducting both the initial, gradual crushing and shrinkage deformations, the creep deformations of FRP-confined RuC, at one day after loading, were within the same values as for the unconfined RuC (R60). After a year, the creep deformations of R60F2 were 77% higher than of R60. It is noted that when the r/fcrc (for confined RuC) is around or above the r/frc 1.00 of the corresponding unconfined RuC, and implicitly the stress range is within the transition region of confined r-e curve, the creep deformations are intertwined with those resulting from the unstable crushing propagation within the RuC core. As mentioned above, the transition zone of the confined r-e curve is delimited by the proportionality limit (zero stress and the end of the linear elastic branch) and the beginning of the hardening regime. Another possibility to depict the creep response under sustained stress is through the creep compliance parameter, defined as the creep strain (ecr) divided by the applied stress (r). Close inspection of the creep compliance-time curves indicates improved performance for the confined RuC in comparison to the corresponding unconfined material, in contrast to the creep strain–time (ecr-ti) curves, which show the highest creep for the FRP-confined RuC. However, in all situations, the best creep performance, or lowest ranges of ecr and ecr/r, were observed for the reference highstrength CCM. In practice, to enable the assessment of creep response of concrete materials under various levels of stresses, a creep coefficient u(ti) is determined. This includes both the drying and basic creep components. The u(ti) assessed from the ratio between the creep strains and instantaneous strains (ecr/ein) is shown in Fig. 8. For a similar r/fc, the high-strength CCM has the lowest u(ti) in comparison to the unconfined and confined RuC. The u(ti) prior to specimen removal from the creep frame was about 17% and 106% higher for R60F2 and R60, respectively, in comparison to R00. The experimental results described above permit a direct comparison with predictions of existing codified provisions. Fig. 9 depicts an assessment of the u(ti) using Eurocode 2 provisions [45] for the three cases investigated in this paper.
Fig. 8. Test creep coefficient versus time curves.
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Fig. 9. Comparison between test and predicted creep coefficient vs time curves: a) R00, b) R60, c) R60F2.
Considering that r/fc at loading was below 45% of the 28 days strength, the creep deformation is directly proportional to the applied stress. Hence, the linear creep model was considered. It is worth noting that the mean coefficient of variation between the Eurocode 2 creep test database and codified expressions, is in the order of 20%. Alongside the predicted average values represented by continuous black curves, the minimum and maximum ranges of the creep dataset are depicted by continuous grey curves. It is shown that for R00 and R60F2, the average predicted curves are slightly above the test data, indicating generally conservative predictions, yet within the code test database ranges. In contrast, for R60, the test results from this paper are outside of the codified range with overly-conservative predictions of u(ti). As the range of applicability of codified expressions is limited to concrete strengths above 20 MPa, predictions of u (ti) below this range would be unreliable. An assessment of the u(ti) using fc = 20 MPa, together with the test geometries and conditions described in Section 2, shows that the predicted values overlap the test data for R60. The observations noted above indicate that Eurocode 2 offers relatively conservative estimates of u(ti) for r/fc within the linear creep range for high-strength CCM, yet within reasonable margins. This suggests that the codified provisions can be reliably used for creep assessments. In contrast, the limitations of the model do not permit reliable assessments of creep response for rubberised concrete with relatively high rubber proportions as compressive strengths of these materials would be outside of the specified ranges. Importantly, u(ti) for the FRP-confined rubberised concrete are reliably estimated, since the material may possibly be regarded as an unconfined normal-strength CCM. Nonetheless, further studies on rubberised concrete materials with lower replacement ratios and other confinement types and levels are necessary for wider practical application in structural elements. Additionally, further detailed experimental assessments incorporating a broader range of r/fc for unconfined RuC, as well as r/fc ratios within the transition and hardening zones of FRP-confined RuC materials, needs to be considered in order to support the development of full design expressions. 4. Concluding remarks This study investigated the creep response of unconfined and FRP-confined rubberised concrete materials provided with relatively high rubber content, as well as of high-strength conventional concrete from which the rubberised concrete was derived. A full account of creep tests carried out under sustained axial compression, and of material tests focusing on the complete constitutive
response of the materials investigated, was given. The key observations are outlined below. The compressive strength of the considered rubberised concrete material with 60% volumetric rubber replacement was about one-tenth of the reference conventional concrete from which it was derived, yet it exhibited enhanced energy dissipation and comparatively ductile response. External confinement measures by means of FRP recovered a significant proportion of the strength lost due to the presence of rubber and enhanced the deformation characteristics further. The compressive strength-time curves for both unconfined rubberised concrete and its reference concrete followed similar logarithmic trend lines, with 18% and 10% increase in strength over one year after casting, respectively. For confined elements, along with inherent experimental variations, the hardening of the epoxy resin may have increased the stiffness of the jacket with time, allowing less dilation under axial loading leading to a reduction in strength and deformation. For two layers of FRP-confinement, the confined-to-unconfined compressive strength ratio of the rubberised concrete with 60% rubber replacement was 4.9 at 28 days. After a year from casting, these ratios were 4.0 for non-loaded control specimens and 3.5 for specimens subjected to an axial load ratio of 20%. At stress-to-strength ratios within the linear creep range, about a year after loading, the creep strains of the high-strength conventional concrete were 35% lower than that of the rubberised concrete elements. For the same stress-to-strength ratio, creep strains of FRP-confined rubberised concrete were higher than those of unconfined rubberised concrete by a factor of 2.80. Although the stress-to-confined strength ratio was 0.20, for the FRP-confined concrete, the applied stress corresponded to a strain level within the crushing and non-linear creep ranges of the unconfined material, leading to an unstable crushing propagation throughout the concrete core, which increased the creep strains for the confined elements. Comparative assessment indicated that Eurocode 2 offers relatively conservative estimates of the creep coefficient for stress-to-strength ratios within the linear creep range for high-strength conventional concrete and confined rubberised concrete. However, it provides overly-conservative estimates for rubberised concrete, particularly since its concrete strength is outside of the codified model ranges. Conflict of interest None.
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Acknowledgements The authors acknowledge the financial support of the European Union Seventh Framework Programme FP7/2007-2013 under grant agreement No 603722 within the project ‘Anagennisi: Innovative Use of all Tyre Components in Concrete’ for the tests described in this paper. References [1] D.V. Bompa, A.Y. Elghazouli, B. Xu, P.J. Stafford, A.M. Ruiz-Teran, Experimental assessment and constitutive modelling of rubberised concrete materials, Constr. Build. Mater. 137 (2017) 246–260, https://doi.org/10.1016/ j.conbuildmat.2017.01.086. [2] S. Raffoul, R. Garcia, D. Escolano-Margarit, M. Guadagnini, I. Hajirasouliha, K. Pilakoutas, Behaviour of unconfined and FRP-confined rubberised concrete in axial compression, Constr. Build. Mater. 147 (2017) 388–397, https://doi.org/ 10.1016/j.conbuildmat.2017.04.175. [3] A.O. Atahan, A.Ö. Yücel, Crumb rubber in concrete: static and dynamic evaluation, Constr. Build. Mater. 36 (2012) 617–622, https://doi.org/10.1016/ j.conbuildmat.2012.04.068. [4] O. Youssf, M.A. ElGawady, J.E. Mills, Static cyclic behaviour of FRP-confined crumb rubber concrete columns, Eng. Struct. 113 (2016) 371–387, https://doi. org/10.1016/j.engstruct.2016.01.033. [5] A.Y. Elghazouli, D.V. Bompa, B. Xu, P.J. Stafford, A.M. Ruiz-Teran, Inelastic Behaviour of RC Members Incorporating High Deformability Concrete, in: High Tech Concrete: Where Technology and Engineering Meet – Proceedings of the 2017 Fib Symposium, Springer International Publishing, 2017, pp. 2399–2406, https://doi.org/10.1007/978-3-319-59471-2_273. [6] A.Y. Elghazouli, D.V. Bompa, B. Xu, A.M. Ruiz-Teran, P.J. Stafford, Performance of rubberised reinforced concrete members under cyclic loading, Eng. Struct. 166 (2018) 526–545, https://doi.org/10.1016/j.engstruct.2018.03.090. [7] L.J. Parrott, Some observations on the components of creep in concrete, Mag. Concr. Res. 22 (72) (1970) 143–148, https://doi.org/10.1680/macr.1970. 22.72.143. [8] J.J. Brooks, 30-year creep and shrinkage of concrete, Mag. Concr. Res. 57 (9) (2005) 545–556, https://doi.org/10.1680/macr.2005.57.9.545. [9] B.R. Gamble, L.J. Parrott, Creep of concrete in compression during drying and wetting, Mag. Concr. Res. 30 (104) (1978) 129–138, https://doi.org/ 10.1680/macr.1978.30.104.129. [10] A.M. Neville, Creep of concrete as a function of its cement paste content, Mag. Concr. Res. 16 (46) (1964) 21–30, https://doi.org/10.1680/macr.1964.16.46.21. [11] E.W. Bennett, D.R. Loat, Shrinkage and creep of concrete as affected by the fineness of Portland cement, Mag. Concr. Res. 22 (71) (1970) 69–78, https:// doi.org/10.1680/macr.1970.22.71.69. [12] L. Maia, J. Figueiras, Early-age creep deformation of a high strength selfcompacting concrete, Constr. Build. Mater. 34 (2012) 602–610, https://doi.org/ 10.1016/j.conbuildmat.2012.02.083. [13] H. Marzouk, Creep of high-strength concrete and normal-strength concrete, Mag. Concr. Res. 43 (155) (1991) 121–126, https://doi.org/10.1680/macr.1991. 43.155.121. [14] Q. Zhao, X. Liu, J. Jiang, Effect of curing temperature on creep behaviour of fly ash concrete, Constr. Build. Mater. 96 (2015) 326–333, https://doi.org/ 10.1016/j.conbuildmat.2015.08.030. [15] S. Brunauer Tobermorite gel—the heart of concrete. American Scientist. 1962 Mar 1;50(1):210-29. https://www.jstor.org/stable/27838347? seq=1#metadata_info_tab_contents [Accessed 25 February 2019]. [16] A. Domingo-Cabo, C. Lázaro, F. López-Gayarre, M.A. Serrano-López, P. Serna, J. O. Castaño-Tabares, Creep and shrinkage of recycled aggregate concrete, Constr. Build. Mater. 23 (2009) 2545–2553, https://doi.org/10.1016/ j.conbuildmat.2009.02.018. [17] W.C. Tang, H.Z. Cui, M. Wu, Creep and creep recovery properties of polystyrene aggregate concrete, Constr. Build. Mater. 51 (2014) 338–343, https://doi.org/ 10.1016/j.conbuildmat.2013.10.093. [18] CONICA, Recycled Rubber from Used Tyres, http://www.conica.com/ en/commodities/black-granules/rubber-granules/recycled-tyre-rubber/ [Accessed 2 December 2017]. [19] S.H. Hahn, I. Gazdik, The creep of natural and synthetic rubber compounds in shear, Rubber Chem. Technol. 14 (2) (1941) 433–443, https://doi.org/10.5254/ 1.3540039. [20] A.N. Gent, Relaxation processes in vulcanized rubber. I. relation among stress relaxation, creep, recovery, and hysteresis, J. Appl. Polym. Sci. 6 (22) (1962) 433–441, https://doi.org/10.1002/app.1962.070062207.
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Technical note
Creep properties of recycled tyre rubber concrete D.V. Bompa ⇑, A.Y. Elghazouli Department of Civil and Environmental Engineering, Imperial College London, UK
a r t i c l e
i n f o
Article history: Received 2 January 2019 Received in revised form 26 February 2019 Accepted 11 March 2019
Keywords: Rubberised concrete Recycled rubber Creep Mechanical properties
a b s t r a c t This paper investigates the creep response and long-term strength properties of unconfined and FRPconfined concrete materials incorporating relatively high proportions of recycled tyre rubber particles. The high-strength reference conventional concrete, from which the rubberised concrete is derived, is also examined for comparison purposes. After discussing fundamental characteristics of creep behaviour, this study describes an experimental investigation in which three groups of concrete specimens are subjected to sustained uniaxial compressive stress, in the range of about 20% of the estimated strength, for a period of over a year. The test results indicate that both confined and unconfined rubberised concrete materials tend to develop higher creep coefficients by about 53% and 20%, respectively, in comparison to their reference conventional concrete. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction Creep effects may influence the overall structural response through changes in internal member actions. Hence, an in-depth understanding and reliable prediction of such effects are essential for the practical application of novel construction materials. The need for innovative applications for the reuse of waste tyre components led to a significant number of investigations focusing on the properties of rubberised concrete (RuC) materials in which a proportion of mineral aggregates is replaced with recycled rubber particles. Rubberised concrete is a relatively weak material in compression due to the softer nature of rubber particles compared to mineral aggregates, yet, in contrast, it is intrinsically more deformable than conventional concrete materials (CCM) both in terms of lateral and axial deformation [1]. Moreover, compressive strengths comparable to those of CCM can be recovered through external confinement measures, such as fibre reinforced polymer (FRP) jackets or tubular steel elements, whilst further enhancing the deformation capacity. The effectiveness of such confinement approaches depends on the thickness of the confinement and the rubber content within the concrete material [2]. Also, the energy dissipation and inherent damping typically increase for RuC in comparison to CCM [1–3]. More recent investigations also focused on the structural response of reinforced rubberised concrete (RRuC) members provided with various types of confinement (i.e. internal hoops, exter-
⇑ Corresponding author. E-mail address: [email protected] (D.V. Bompa). https://doi.org/10.1016/j.conbuildmat.2019.03.127 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.
nal steel tubes or fibre reinforced polymer sheets), and subjected to combined cyclic lateral displacement and co-existing axial load up to 20% of the squash capacity [4–6]. In comparison with specimens without external confinement, confined RRuC members tend to exhibit higher level of energy dissipation in the range of 9% for 2 layers of FRP confinement and below 10% of total aggregate replacement [4], and by a factor up to 2.7 for 3 layers of FRP and 60% rubber content [5,6]. Additionally, RRuC members develop low cyclic degradation in comparison with conventional RC specimens [6], having the potential to be integrated in isolation or energy dissipation systems to mitigate the effects of seismic loads and other dynamic actions. Creep is defined as the increase in strain in the direction of sustained stress which is obtained by removing the instantaneous, shrinkage and any thermal strain from the total strain. It occurs in most construction materials including concrete, steel, FRP and rubber as a function of their properties. In conventional concrete, the total creep is divided into basic creep and drying creep [7,8]. The former is a movement governed by an elastic structure within the paste. The latter is controlled by an oriented loss of loosely bonded hydrate water and is typically referred to as the creep exceeding the basic creep of sealed concrete [7,8]. The moisture conditions determine the level of drying creep, similarly to the concurrently developing load-independent drying shrinkage, as a response of the skeletal structure of the cement paste to capillary tensions, sorption stresses and external loads [9]. The creep level is controlled mostly by the aggregates’ stiffness, whilst the fineness of binders has a relatively minimal impact [10,11]. For the same type of aggregates used, high strength mixes
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Nomenclature fc fc0 fc,28 d fc,ti frc fcc fcrc ti Ec N
e e1 eccu
compressive strength compressive strength of reference concrete compressive strength at 28 days compressive strength at the time i (in days) unconfined rubberised concrete strength confined compressive strength confined rubberised concrete strength time modulus of elasticity, applied axial load strain compressive strain ultimate confined strain
provided with fly ash and micro-silica tend to have a denser microstructure, hence develop smaller creep deformations in comparison to conventional mixes for the same stress-to-strength ratios [11,12]. This is attributed to the formation of the new Tobermorite gel as a result of the secondary hydration between the calcium hydroxide and silica at room temperature, rather than to the fineness of the binders [13]. The primary factor influencing specific creep of conventional and fly ash concretes (20% and 40% of total binders), cured under a wide range of temperatures (20 °C, 50 °C and 90 °C) and subjected to loads within the elastic creep regime (33% fc) is the quantity and microstructure of calcium silicate hydrate (CASAH) gel [14]. For both concretes without and with low fly ash content, a higher temperature produces more C-S-H gel and a lower specific creep in comparison to curing under room temperature [14]. It is worth noting that Tobermorite is often used in thermodynamical calculations to represent the pole of the most evolved calcium silicate hydrate (CASAH) [15]. Unsurprisingly, concrete with relatively soft aggregates develops more deformation after a given time under load [8,10]. For example, for concrete specimens loaded to 30% of their axial capacity, a full replacement of coarse natural aggregate by recycled aggregate increase the creep deformation by 51% [16]. Also, the specific creep of polystyrene aggregate concrete (PAC) with an aggregate replacement of 20–80% by PAC is strongly influenced by the polystyrene content, with specific creep values between 101 and 880% of the corresponding reference concrete[17]. PAC with relatively large proportions of polystyrene aggregate is a lightweight material with good thermal and insulation properties and its use in structures can reduce the self-weight of the structure as well as the loads on the foundation, in comparison to conventional concrete [17]. Considering that for the same replacement ratio, the reduction in compressive strength of PAC is about half that of RuC, it is expected that the creep-time response of RuC would also be affected. Recycled tyre rubber typically includes about 25% carbon black and polymers in the range of 40–55%, whereas the remaining constituents are softeners and fillers [18]. As recycled rubber particles result from a variety of sources, their long-term response is dependent on the properties and ratios of their constituents. For such materials, as for concrete, creep is directly related to stress and time. The majority of creep occurs in a relatively short amount of time, minimizing its long-term effect [19]. The shape of the rubber creep-time curve resembles that of concrete, although, when plotted on log scales, for very short and for very long times it deviates from linearity [20]. Flexural creep tests on recycled guardrail posts incorporating rubber indicate that the post flexural stiffness was much lower, and the tendency to creep was much greater in comparison to conventional counterparts [21].
eccu1 eccu2 ec1,1 ec2,1 ecr ein erc1,1
u(ti)
r r/fc r/fcrc r/frc
confined ultimate axial strain confined ultimate lateral strain unconfined crushing strain unconfined lateral strain at crushing creep strain strain at applied stress unconfined crushing strain of rubberised concrete creep coefficient applied axial stress, stress to strength ratio stress-to-confined strength ratio stress-to-unconfined strength ratio
Limited tests on creep response of FRP-confined concrete indicate that due to a sealing effect produced by the FRP, the concrete core only has basic creep, whilst the drying creep and shrinkage are minimised or eliminated [22]. Creep tests on aramid FRP-wrapped cylinders under a stress-strength ratio of 30% show that axial load is primarily carried by the concrete core while the contribution from FRP is relatively negligible [23]. Although a direct comparison with unconfined concrete elements was not presented [23], the shape of the creep-time curves of the FRP-confined concretes are similar to those of unconfined concrete reported in the literature [13]. As noted above, FRP laminates generally play a beneficial role in the long-term response of confined concrete elements, yet the type of fibre (carbon, glass, aramid) may have a different contribution to creep, particularly aramid FRP, which is a material that creeps considerably [24]. Long-term effects can have a significant influence on the overall structural response through an increase in internal actions [22,25,26]. Movements due to creep and shrinkage tend to limit the dimensions of reinforced concrete buildings or long-span bridges requiring expansion joints and/or bearings. FRP-Confined RuC elements could be incorporated as discrete link elements acting as isolation or energy dissipation systems against large movements as well as vibrations in such structures. In order to avoid possible failures under service load, the superposed creep effects of the FRP-confined RuC constituent materials need to be assessed in advance [22]. Although previous studies indicate reliable structural behaviour under short-term loading, practical application of such materials and elements is constrained by the lack of studies on the long-term response of unconfined and confined RuC. To this end, this paper focuses on the creep response and the long-term strength properties of RuC with 60% rubber replacement ratio of both fine and coarse mineral aggregates by volume, with and without FRP confinement, as well as of the high-strength reference CCM from which the RuC was derived. 2. Experimental programme Eighteen cylindrical samples were prepared for assessing the response of unconfined and FRP-confined RuC as well as of the reference high-strength CCM from which the RuC was derived, under sustained stress over a period of over a year. In previous studies it was shown that a stronger influence on mechanical properties, particularly the compressive strength [3,27,28], was typically observed for small rubber content and tended to stabilise as the rubber content increased. A higher benefit in terms of deformation properties is obtained from relatively higher volumetric rubber replacement ratio (e.g. 60%, referred to as R60). The term volumetric rubber replacement ratio depicts the pro-
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portion of the fine and coarse aggregates replaced by rubber particles, by volume. From a structural point of view, such composites benefit from relatively large rubber contents (e.g. 40–60%) as these mobilise supplementary confinement and enhance the deformability levels. Additionally, as the strength reduction rate reduces with an increase in rubber content, a higher replacement ratio would also add benefits to the environment through reuse of a larger quantity of waste materials. There is very limited data on mixes in which the replacement ratio is higher than 60%, primarily since above this limit the physical characteristics are strongly governed by the rubber [1] and because the concrete workability and slump decrease drastically [28]. Hence, long-term tests on RuC with 60% replacement would represent a reliable upper bound of aggregate replacement to obtain workable concrete materials. Also, results from creep tests with 0 and 60% replacement ratios could be interpolated to obtain qualitative information regarding creep response of RuC with intermediate rubber content. On the other hand, two FRP layers (referred to as R60F2 for 2 layers in conjunction with 60% rubber content) offered a good balance between the strength recovered due to the presence of the rubber and offered significant enhancement in deformation capacity. The axial crushing strain of FRP-confined concrete was enhanced by a factor of 24.5 for R60, reaching values above 3.0% in comparison to the crushing strain of the corresponding unconfined concrete [6]. In terms of dilation, the ultimate lateral strains of FRP-confined concrete were 15.9 times higher for R60 concrete (above 2%), in comparison to the lateral strain at crushing of the unconfined concrete [6]. Hence, these material and confinement configurations were adopted for long-term testing at room temperature in a controlled environment for more than a year. Implicitly, the key observations from this study are limited to possible practical configurations that adopt similar properties to those considered herein. To expand the applicability of rubberised concrete in practice, further studies could involve a wider range of replacement ratios, as well as an in-depth investigation into the environmental and UV effects, whilst testing could be performed at low and high temperatures and their performance would be assessed after several years.
2.1. Materials and mix proportions High strength cement CEM I 52, 5 conforming to BS EN 197-1 [29,30], mineral aggregates, admixtures and tap water were used in the reference normal concrete mix with a target compressive strength of 70 MPa (R00). The sand (<5 mm) and gravel (5– 10 mm) had a moisture content of about 5 and 3% of the bulk aggregate weight, respectively, while the specific gravity was 2.65 for both constituents [31]. The fineness modulus was 2.35 for fine aggregates, and 5.88 for coarse aggregates. The cement
Table 1 Concrete mixtures per cubic meter.
Cement (kg) Fly ash (kg) Microsilica (kg) Aggregates (kg) Rubber (kg)
5–10 mm 0–5 mm 0/0.5 mm 0.5/0.8 mm 1/2.5 mm 2/4 mm 4/10 mm 10/20 mm
Admixtures (l) w/c (-)
R00 425 – – 1001 820
– – 7.6 0.35
R60 350 42.5 42.5 400 328 16.5 16.5 49.5 66.0 33.0 148.5 7.6 0.35
composition included on average: 3.3% Sulfate SO3, <0.10% Chloride Cl, 0.7% Alkali Eq Na2O, 55% Tricalcium Silicate C3S, 20% Dicalcium Silicate C2S, 9.5% Tricalcium Aluminate C3A and 8% Tetracalcium Aluminoferrite C4AF [30]. From the total quantity of cement listed in Table 1 for the reference concrete mix, in the rubberised concrete, 15% was replaced by EN 450-1 fineness category S fly ash [32,33] and 15% was replaced by undensified Grade 940 microsilica with a minimum 90% SiO2 [34]. Microsilica and fly ash, primarily added to improve workability, segregation and slump, and to optimise the particle packing of the mixture, improve the strength and concrete flowability. Also, polycarboxylate superplasticizer admixtures were added to improve the workability of the mixes [35,36]. The rubberised concrete was produced using a blend of rubber particles with smaller sizes (up to 10 mm) from car tyres, and larger sizes (up to 20 mm) from truck tyres. As depicted in Fig. 1a, they were supplied in the following size ranges: 0–0.5 mm, 0.5– 0.8 mm, 1.0–2.5 mm, 2–4 mm and 4–10 mm, and were used in the concrete mix in the 5%, 5%, 15%, 20% and 10% ratios of the total added rubber content, respectively. The remaining 45% consisted of particles with sizes in the range 10–20 mm. The car and truck tyre rubber had a specific gravity of 1.1 and water absorption of 7% and 1%, respectively. The particle size distribution of mineral aggregates, rubber and blend determined following EN 933-1:2012 [37] are depicted in Fig. 1. The rubber particles were used in the condition received from the manufacturer without pre-treatment. All mineral aggregates and half of the water were mixed together for 1 min, followed by the added rubber for another 1 min. The binders, admixtures and remaining water were then mixed all together for a minimum of 3 further minutes. After being placed in plastic forms, the concrete was subsequently compacted using a vibrating table until the air content in the fresh concrete was at a minimum. Fresh concrete properties were determined by means of a slump cone. The slump was 125 mm for R00 and 50 mm for R60. To ensure appropriate curing, two days after
Fig. 1. a) Rubberised concrete constituents b) Sieve analysis for mineral aggregates and rubber blend.
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casting, the samples were placed in a water tank located in a curing room, and these were removed from the wet environment at 26 days. Two layers of aramid FRP sheets (AFRP) were used to confine rubberised concrete samples using a hand wet lay-up procedure. To achieve this, the AFRP sheets were impregnated with a commercial moisture-tolerant EP epoxy bonding adhesive. Before applying the AFRP confinement, at about 36 h before loading, concrete surfaces were brushed and cleaned to improve the adherence between the existing concrete and the fibre sheets. The aramid fibre (of Grade S&P A120/290) was unidirectional and had 0.2 mm design thickness [38]. The ends of the cylindrical samples were ground and capped with high strength mortar slurry made of CEM I 52,5 [29,30] and tap water (1:1) to ensure flatness of the surfaces and good contact conditions during testing. The dimensions of all samples were accurately measured for appropriate assessment of stress and strain characteristics. Samples used for concrete strength assessment at 28 days were Ø100 in diameter and 196– 202 mm in length, whilst all samples used for creep assessment had a Ø102 diameter and a length which varied from 247 to 254 mm for R00 and 255–256 mm for R60. 2.2. Specimen details and testing arrangement The total number of cylindrical specimens prepared were divided into three groups. The first group of Ø100 200 mm specimens was used to assess the compressive strength at 28 days, which was required to determine the level of loading for creep testing. One-third of the remaining samples (Ø102 254 mm) were used as control specimens to record the free-standing deformations, whilst the remaining third (Ø102 254 mm) were subjected to constant sustained compression for a period of over a year. The specimen dimensions conform to the requirements of
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ISO 1920-9:2009 [39] for the determination of creep of concrete cylinders in compression. The latter were loaded to about 20% of their squash load, determined from the compressive strength assessed at 28 days. It is worth noting that for stress-to-strength (r/fc) ratios below 30% the creep response is typically linear and time-dependent only, while between 30 and 70%, the response becomes non-linear being time and stress-dependent [40]. The compressive strength tests were carried out in displacement control using a four-post displacement control Instron Satec 3500 kN machine. To avoid eccentric loading, the testing arrangement included top and bottom high strength steel transfer plates and a 3D hinge directly connected to the actuator (Fig. 2a). Three displacement transducers were employed around the specimens to record axial deformation between the machine platform and the top transfer plate. A displacement rate of 0.2 mm/min was applied from the machine to the specimens. Besides the displacement transducers used to record member deformations, the load corresponding to the applied displacement was recorded by the machine load cells. Surface strains were recorded by means of a Digital Image Correlation (DIC) system. The DIC system consists of two 5MP cameras provided with 35 mm f/2D lenses [41]. To capture crack propagation and member displacements, high contrast black/white patterns were provided to the front face of the specimens. Before the start of the recording, a calibration procedure was performed to ensure measurement accuracy as required by the manufacturer recommendations. The specimens used for creep assessment were placed in a vertical position in a loading frame, which was able to maintain the required load for the duration of the test (Fig. 2b). The load was applied through a hydraulic ram using prestressed Ø12 mm ties prepared with threaded ends, nuts and steel plates to allow positioning and load transfer, and it was recorded continuously by means of a data logger. At the top and bottom, between the two samples and the steel plates, 3D hinges were positioned to accommodate potential non-even loading. The loads were monitored throughout the testing period using a computer and were maintained throughout testing by tightening the nuts using hand-held spanners. These values never dropped below 2% between two consequent recording times. The non-loaded specimens were placed next to the loaded specimen following the same arrangement. To record long-term axial deformations, all samples were prepared with two steel discs at 150 mm distance along equally located three generatrices of the cylinder. The average deformations under sustained load were measured by means of a dial gauge immediately after loading and at established intervals during a period of over a year [39]. Deformations in non-loaded control specimens were measured concomitantly to assess the level of control shrinkage strains. A purpose-built program was developed in order to collect the data from the digital dial gauge via a COM PC port. Careful tracking of the recordings was followed in order to avoid the introduction of any spurious data. The measurements were performed three times and averaged for each testing sequence, whilst the temperature and relative humidity were found to be in the range of 24–27 °C and 50–60%, respectively.
3. Test results 3.1. Strength properties
Fig. 2. Testing arrangement: a) compression tests, b) creep tests.
Fig. 3 depicts the complete axial and lateral stress–strain (r-e) curves for the materials tested at 28 days after casting, whilst Fig. 4 illustrates the damage kinematics for unconfined and confined RuC. For the unconfined samples (R00, R60), the r-e responses are characterised by an increase in stress with strain until crushing, followed by a softening region. As expected, the response of the
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Fig. 3. Stress–strain response at 28 days a) complete response, b) detail of axial compressive response.
Fig. 4. Damage kinematics for: a) R60, b) R60F2.
reference CCM (R00) was brittle with a sudden drop in stress after the maximum capacity was reached. In contrast, the unconfined RuC (R60) exhibited a softer post-crushing behaviour which typically leads to similar or higher energy dissipation in the postpeak regime as shown previously in other investigations [1,3,21]. Unlike the average test measurements for R00, in which a distinct peak lateral strain is observed, for R60 the peak lateral strain is followed by a much smoother softening. The average test compressive strengths at 28 days (fc,28d), the elastic modulus (Ec), the applied axial load (N) and the corresponding stress to strength ratio (r/fc), the applied axial stress (r) and the strain at the applied stress (ein), are listed in Table 2. The stress values r depicted in Table 2 were determined by dividing the applied axial load (N) by the specimen cross-section area (Ac), whist the strain (ein) was assessed by using Hooke’s law and the
modulus of elasticity (Ec) obtained from the tests. In terms of strength, as observed in Fig. 3 and Table 2, the compressive strength of the reference concrete R00 material fc0 = 79.5 MPa is about 10 times that of the rubberised R60 counterpart (frc = 8.07 MPa), indicating a significant influence from the rubber particles (Table 2). The pre-peak behaviour of the concrete was strongly influenced by the volumetric replacement ratio of the mineral aggregates as shown in a previous study [1]. Naturally, the rubber aggregates are softer and lighter in weight compared to mineral aggregates. Hence, for the RuC, the reduction in stiffness is a function of the stiffness ratio between the cement matrix and inclusions (rubber particles and voids) [1]. For the R60 material investigated in this paper, the elastic modulus is less than a fifth of the corresponding R00 concrete. As observed in Fig. 3a, the axial r-e curve of FRP-confined RuC can be characterised by four (r,e) pairs: (i) zero stress, (ii) a proportionality limit between zero stress and the end of the linear elastic branch, (iii) a transition zone between the proportionality limit and the beginning of the hardening regime, (iv) a hardening regime, and (v) the ultimate condition (fcc,eccu). As illustrated in Fig. 3b, up to the proportionality limit, the stiffness of the confined RuC (R60F2) is similar to that of the unconfined concrete Ec. As the concrete gradually crushes inside the wrapping, the r-e response softens. Beyond this limit, the response under axial compression is governed by the jacket properties. The ultimate state of FRPconfined concrete occurs due to extensive lateral expansion and ultimately fracture of the FRP sheets in the transverse direction. Fig. 3 shows that two layers of external AFRP confinement, added to R60 cylinders, enhanced the confined rubberised concrete compressive strength fcrc = 38.8 MPa to about 5 times the unconfined R60 (i.e. frc = 8.07 MPa), which recovers a significant proportion of the strength of its reference conventional concrete (fc0 = 79.5 MPa). The complete recovery of the compressive strength may be attained by increasing the jacket thickness, yet as the jacket becomes thicker the deformation capacity would reduce [6]. For two layers of confinement, the axial deformation capacity of R60F2 was about 12 times higher than R00, and 24 times higher than R60. On the other hand, the lateral deformation increased by a factor of about 15 in comparison to both R00 and R60. This shows that external confinement, in the form of AFRP,
Table 2 Material and loading details. Concrete
fc,28 d (MPa)
Ec (MPa)
N (kN)
r/fc (-)
r (MPa)
ein (%)
R00 R60 R60F2
79.5 ± 2.58 8.07 ± 0.76 38.8 ± 1.45
41,380 7550 7200
130 16 65
0.21 0.25 0.21
16.5 2.04 8.28
0.038 0.027 0.115
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Fig. 5. Stress–strain response after removal from the creep frame: a) R00, R60, R60F2 non-loaded/control specimens, b) R00, R60, R60F2 loaded specimens; c) details of R60 and R60F2 non-loaded/control specimens, d) details of R60 and R60F2 loaded specimens.
may compensate for the strength loss due to the presence of rubber, whilst enhancing the axial and lateral deformation capacity. One-third of the specimens were tested to failure at 28 days, whilst the remaining two-thirds were tested after the removal from the creep frame. The r-e curve of the R00, R60 and R60F2 control non-loaded specimens and those loaded for a year to 20% axial load are illustrated in Fig. 5, whilst the full details are given in Table 3. As mentioned before, one third of the remaining samples (Ø102 254 mm) were used as control specimens (referred to as ‘control non-loaded in Table 3) to record the free-standing deformations, whilst the remaining third (Ø102 254 mm) were subjected to constant sustained compression for a period of over a year. The latter are referred to as ‘loaded to 0.2 fc’ in Table 3. From the results in the figures and table, it can be observed that the unconfined compressive strength of non-loaded specimens after removal from the creep frame was about 10% and 18% higher than at 28 days for R00 and R60 respectively. These results are largely similar to those available in the literature indicating that creep has a minor influence on the concrete strength for r/fc < 0.7 [40]. To this end, Fig. 6 illustrates the relationship between fc,ti/fc,28 d
and time for non-loaded unconfined R00 and R60 concrete, as obtained from testing throughout a period of two years as part of a wider research programme [3,6,42,43]. Additional compressive strengths for the corresponding non-loaded R00 and R60 specimens described in this paper are also illustrated in Fig. 6. The term fc,ti depicts the compressive strength at the time i (in days), whilst fc,28 d is the strength at 28 days. As observed, the logarithmic trend lines for R00 and R60 are overlapping, indicating a similar increase in unconfined rubberised concrete compressive strength frc with time and suggests that methods typically used for CCM may be directly applied for RuC. As indicated in Table 3, for the FRP-confined specimens there was a slight reduction in compressive strength with time. The reduction in strength was 4.5% for the control specimens and about 9% for the loaded samples. The ultimate deformation capacity was also 13% lower for the control samples and 35% lower for the loaded samples, in comparison to the values at 28 days. It is worth noting that one of the loaded samples had an initial drop in capacity before ultimate deformation was reached due to gradual failure of the wrapping; hence, the strength of this sample is not
Table 3 Strength and strain properties.
Sample
28 days
After removal from the creep frame
Initial
Control non-loaded
fc (MPa)
ec1,1 or eccu1 (%)
ec2,1 or eccu2 (%)
fc (MPa)
ec1,1 or eccu1 (%)
Loaded to 0.2 fc
ec2,1 or eccu2 (%)
fc (MPa)
ec1,1 or eccu1 (%)
R00
a b avg
81.3 77.6 79.5
0.26 0.26 0.26
0.14 0.12 0.13
85.7 89.7 87.7
0.27 0.27 0.27
0.08 0.14 0.12
92.1 84.4 88.2
0.27 0.26 0.26
0.13 0.14 0.13
R60
a b avg
8.6 7.5 8.1
0.13 0.14 0.13
0.16 0.12 0.14
9.4 9.7 9.5
0.12 0.13 0.12
0.13 0.12 0.12
9.7 9.4 9.6
0.15 0.17 0.16
0.11 0.13 0.12
R60F2
a b avg
37.5 41.7 39.6
2.92 3.76 3.34
2.14 2.01 2.08
38.6 37.0 37.8
3.26 2.56 2.91
1.61 1.01 1.31
36.0 30.5 33.2
2.18 1.81 2.00
1.42 1.11 1.26
Notes: a, b – two identical specimens, avg – average, axial strains are depicted with positive sign, lateral strains are depicted with negative sign.
ec2,1 or eccu2 (%)
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Fig. 6. Compressive strength increase versus time (fc,ti/fc,28d – t) ratio for R00 and R60.
Fig. 7. Test creep strain–time curves.
considered in the comparisons. Considering that both fcrc and ecrcu reduced with time, it is suggested that along with inherent experimental variations, the hardening of the epoxy resin may have increased the stiffness of the jacket with time, allowing less dilation under axial loading and resulting in a reduction in the deformation capacity. 3.2. Creep measurements As noted above, creep effects were determined from an average of three experimental measurements on a pair of loaded specimens from which average control/shrinkage strains were removed. Creep-time test results depicted in Fig. 7 indicate that creep strains (ecr) increase rapidly after loading in a non-linear manner and tend to become gradually linear after a period of time, with most of the creep largely developing within the first three months. At a similar r/fc = 0.20–0.25, the creep response of the high-strength conventional concrete (CCM) is lower that of the rubberised concrete (RuC). Although ecr are generally similar at early age, the RuC continues to deform under sustained axial stresses, whilst the CCM tends to settle. Prior to removal from the creep frame, ecr values for RuC were about 55% higher than for CCM, indicating that, as for the case of polystyrene aggregate concrete (PAC) described in Section 1, the presence of rubber increased ecr as a function of r/ fc and level of mineral aggregate replacement. The creep of confined rubberised concrete was found to be significantly of a higher magnitude in comparison to the corresponding unconfined materials at the same stress-to-strength ratio. Importantly, the creep strains of FRP-confined RuC were higher than of RuC by a factor of 2.80. As the stress-to-confined strength
ratio was r/fcrc 0.20 corresponding to a stress-to-unconfined strength ratio of r/frc 1.00, the strain levels are within the crushing ranges of the R60 unconfined material. For r/frc 1.00, the concrete creep was in the non-linear range, which for unconfined concrete materials may result in creep rupture [44]. As the strain of the FRP-confined RuC reached erc1,1 = 0.115%, a relatively ductile unstable crushing propagation I throughout the concrete core. This is illustrated in the panel located at the right-hand side of Fig. 7, which shows that strains in the range of e1 = 0.115–0.170% (r/frc 1.00) correspond to r/fcrc 0.20 for R60F2. The above observations suggest that besides the strain from initial loading (e1 = 0.115%) of R60F2, the concrete core gradually crushed inside the FRP confinement after being loaded, adding to the total mechanical strain about De0.05%. After deducting both the initial, gradual crushing and shrinkage deformations, the creep deformations of FRP-confined RuC, at one day after loading, were within the same values as for the unconfined RuC (R60). After a year, the creep deformations of R60F2 were 77% higher than of R60. It is noted that when the r/fcrc (for confined RuC) is around or above the r/frc 1.00 of the corresponding unconfined RuC, and implicitly the stress range is within the transition region of confined r-e curve, the creep deformations are intertwined with those resulting from the unstable crushing propagation within the RuC core. As mentioned above, the transition zone of the confined r-e curve is delimited by the proportionality limit (zero stress and the end of the linear elastic branch) and the beginning of the hardening regime. Another possibility to depict the creep response under sustained stress is through the creep compliance parameter, defined as the creep strain (ecr) divided by the applied stress (r). Close inspection of the creep compliance-time curves indicates improved performance for the confined RuC in comparison to the corresponding unconfined material, in contrast to the creep strain–time (ecr-ti) curves, which show the highest creep for the FRP-confined RuC. However, in all situations, the best creep performance, or lowest ranges of ecr and ecr/r, were observed for the reference highstrength CCM. In practice, to enable the assessment of creep response of concrete materials under various levels of stresses, a creep coefficient u(ti) is determined. This includes both the drying and basic creep components. The u(ti) assessed from the ratio between the creep strains and instantaneous strains (ecr/ein) is shown in Fig. 8. For a similar r/fc, the high-strength CCM has the lowest u(ti) in comparison to the unconfined and confined RuC. The u(ti) prior to specimen removal from the creep frame was about 17% and 106% higher for R60F2 and R60, respectively, in comparison to R00. The experimental results described above permit a direct comparison with predictions of existing codified provisions. Fig. 9 depicts an assessment of the u(ti) using Eurocode 2 provisions [45] for the three cases investigated in this paper.
Fig. 8. Test creep coefficient versus time curves.
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Fig. 9. Comparison between test and predicted creep coefficient vs time curves: a) R00, b) R60, c) R60F2.
Considering that r/fc at loading was below 45% of the 28 days strength, the creep deformation is directly proportional to the applied stress. Hence, the linear creep model was considered. It is worth noting that the mean coefficient of variation between the Eurocode 2 creep test database and codified expressions, is in the order of 20%. Alongside the predicted average values represented by continuous black curves, the minimum and maximum ranges of the creep dataset are depicted by continuous grey curves. It is shown that for R00 and R60F2, the average predicted curves are slightly above the test data, indicating generally conservative predictions, yet within the code test database ranges. In contrast, for R60, the test results from this paper are outside of the codified range with overly-conservative predictions of u(ti). As the range of applicability of codified expressions is limited to concrete strengths above 20 MPa, predictions of u (ti) below this range would be unreliable. An assessment of the u(ti) using fc = 20 MPa, together with the test geometries and conditions described in Section 2, shows that the predicted values overlap the test data for R60. The observations noted above indicate that Eurocode 2 offers relatively conservative estimates of u(ti) for r/fc within the linear creep range for high-strength CCM, yet within reasonable margins. This suggests that the codified provisions can be reliably used for creep assessments. In contrast, the limitations of the model do not permit reliable assessments of creep response for rubberised concrete with relatively high rubber proportions as compressive strengths of these materials would be outside of the specified ranges. Importantly, u(ti) for the FRP-confined rubberised concrete are reliably estimated, since the material may possibly be regarded as an unconfined normal-strength CCM. Nonetheless, further studies on rubberised concrete materials with lower replacement ratios and other confinement types and levels are necessary for wider practical application in structural elements. Additionally, further detailed experimental assessments incorporating a broader range of r/fc for unconfined RuC, as well as r/fc ratios within the transition and hardening zones of FRP-confined RuC materials, needs to be considered in order to support the development of full design expressions. 4. Concluding remarks This study investigated the creep response of unconfined and FRP-confined rubberised concrete materials provided with relatively high rubber content, as well as of high-strength conventional concrete from which the rubberised concrete was derived. A full account of creep tests carried out under sustained axial compression, and of material tests focusing on the complete constitutive
response of the materials investigated, was given. The key observations are outlined below. The compressive strength of the considered rubberised concrete material with 60% volumetric rubber replacement was about one-tenth of the reference conventional concrete from which it was derived, yet it exhibited enhanced energy dissipation and comparatively ductile response. External confinement measures by means of FRP recovered a significant proportion of the strength lost due to the presence of rubber and enhanced the deformation characteristics further. The compressive strength-time curves for both unconfined rubberised concrete and its reference concrete followed similar logarithmic trend lines, with 18% and 10% increase in strength over one year after casting, respectively. For confined elements, along with inherent experimental variations, the hardening of the epoxy resin may have increased the stiffness of the jacket with time, allowing less dilation under axial loading leading to a reduction in strength and deformation. For two layers of FRP-confinement, the confined-to-unconfined compressive strength ratio of the rubberised concrete with 60% rubber replacement was 4.9 at 28 days. After a year from casting, these ratios were 4.0 for non-loaded control specimens and 3.5 for specimens subjected to an axial load ratio of 20%. At stress-to-strength ratios within the linear creep range, about a year after loading, the creep strains of the high-strength conventional concrete were 35% lower than that of the rubberised concrete elements. For the same stress-to-strength ratio, creep strains of FRP-confined rubberised concrete were higher than those of unconfined rubberised concrete by a factor of 2.80. Although the stress-to-confined strength ratio was 0.20, for the FRP-confined concrete, the applied stress corresponded to a strain level within the crushing and non-linear creep ranges of the unconfined material, leading to an unstable crushing propagation throughout the concrete core, which increased the creep strains for the confined elements. Comparative assessment indicated that Eurocode 2 offers relatively conservative estimates of the creep coefficient for stress-to-strength ratios within the linear creep range for high-strength conventional concrete and confined rubberised concrete. However, it provides overly-conservative estimates for rubberised concrete, particularly since its concrete strength is outside of the codified model ranges. Conflict of interest None.
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