An experimental study on the early-age hydration kinetics of cemented paste backfill

Construction and Building Materials 212 (2019) 283–294

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An experimental study on the early-age hydration kinetics of cemented paste backfill Lang Liu a,b, Pan Yang a, Chongchong Qi c,⇑, Bo Zhang a, Lijie Guo d, KI-IL Song e a

Energy School, Xi’an University of Science and Technology, Xi’an 710054, China Key Laboratory of Western Mines and Hazards Prevention, Ministry of Education of China, Xi’an 710054, China c School of Civil, Environmental and Mining Engineering, University of Western Australia, Perth 6009, Australia d Beijing General Research Institute of Mining & Metallurgy, Beijing 100160, China e Department of Civil Engineering, Inha University, Incheon 402-751, Republic of Korea b

h i g h l i g h t s

g r a p h i c a l a b s t r a c t


The early-age heat evolution of CPB

was studied by the isothermal calorimeter.
The coupled effect of cement content and temperature was investigated.
The results were analysed from a chemical point of view.
The Krstulovic-Dabic model was utilised for hydration kinetics analysis.

a r t i c l e

i n f o

Article history: Received 18 October 2018 Received in revised form 18 February 2019 Accepted 29 March 2019

Keywords: Hydration kinetics Heat evolution Cemented paste backfill Cement content Temperature

a b s t r a c t Understanding the early-age hydration kinetics of cemented paste backfill (CPB) is necessary for its successful application. In this paper, the early-age hydration heat evolution of CPB was experimentally studied using an isothermal calorimeter considering the coupled effects of cement content and temperature. We employed the Krstulovic-Dabic kinetic model to investigate the hydration mechanism and kinetic parameters of CPB. The results showed that increasing the tailings-cement ratio (TCR) decreased the hydration heat flow (N) and cumulative heat (Q) of CPB due to the decreased amount of cement. Additionally, increasing the TCR elongated the induction stage while shortening the acceleration stage. Increasing the temperature, increased the peak value of N, shortened its appearance time, and shortened/diminished the induction stage due to the increased chemical kinetics of hydration. The exponent and rate constant values in the Krstulovic-Dabic model decreased with the increase of TCR and the decrease of temperature. The hydration mechanism was nucleation and crystal growth (NG) ? phase boundary reaction (I) ? diffusion (D) for most CPB slurries, except for CPB slurries at TCR = 10/12 and temperature = 40 °C, where the hydration mechanism was changed to NG ? D. This study indicates that the isothermal calorimeter, together with kinetic models, is a reliable tool for investigating the hydration kinetics of CPB, which can be incorporated into CPB design in the future. Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction

⇑ Corresponding author. E-mail address: [email protected] (C. Qi). https://doi.org/10.1016/j.conbuildmat.2019.03.322 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

The large-scale development of the global economy depends heavily on the mining industry to provide raw materials. In an era with increasing environmental concerns, a compromise

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Nomenclature Cu Cc

Coefficient of uniformity Coefficient of curvature a Hydration degree n Reaction series t Reaction time, h t0 The time when the induction stage ends, h r Reaction particle diameter, lm K 1 ðK 01 Þ; K 2 ðK 02 Þ; K 3 ðK 03 Þ Reaction rate constants of NG, I and D hydration process, respectively. Qmax Maximum hydration heat release, J/g dQ/dt,da=dt Hydration rate, mW/g Q Accumulated heat, J/g

between more sustainable production and less environmental impact is a continuous challenge [1,2]. This issue cannot be adequately addressed without the safe disposal of mine tailings, which are currently stored and maintained in tailings ponds in most cases. In the last two decades, cemented paste backfill (CPB) has become a promising alternative for the safe and environmentally responsible disposal of tailings [3,4]. In addition, CPB can enhance ore recovery by allowing the excavation of massive pillars, improve mining conditions by ensuring the stability of underground voids and minimising surface subsidence by providing ground support [5,6]. All of the above benefits promote the wide application of CPB in underground mines worldwide [7–15,63–68]. Typically, CPB is a kind of cementitious materials produced using dewatered mine tailings (75–85% wt.% of solids), binders (3–7 wt% by dry weight), and mixing water (15–25%) [16,17]. Understanding the hydration kinetics of CPB, as well as other types of cementitious materials, is of interest to scientists in both academic and industrial fields. From an academic viewpoint, cement hydration involves complex and interdependent mechanisms, making it difficult to investigate individual mechanisms and their reaction rates [18,19]. From an industrial viewpoint, complete knowledge of hydration mechanisms is needed for rational mix designs. For example, it is well accepted that the application of CPB considers its initial setting and early-age strength development, both of which are highly dependent on the early-age hydration of CPB. On the one hand, high hydration kinetics with a quick initial setting is undesirable due to the lack of workability, which is detrimental for pipe transport of fresh CPB [20]. On the other hand, high hydration kinetics leads to high early-age strength of CPB and helps achieve its required functions. Therefore, a thorough understanding of CPB hydration kinetics is an essential prerequisite for successful CPB design. Hydration heat evolution has been widely used to investigate the hydration kinetics of cementitious materials, including CPB [21–23]. It can provide continuous information during the curing of cementitious materials and the hydration heat data has been successfully correlated with kinetic parameters (i.e., the degree of hydration) over many years of investigations [21]. There are three main techniques available for hydration heat measurement, including the semi-adiabatic technique, the full-adiabatic technique and the isothermal calorimeter technique [24]. Although it is probably the most widely used technique, the semi-adiabatic technique has difficulty measuring the heat loss from the sample to the surrounding environment [25]. The full-adiabatic technique is limited by the equipment to achieve a full ‘heat isolation’ from the surrounding environment [26]. Compared with semiadiabatic and full-adiabatic techniques, the isothermal calorimeter technique is gaining popularity in hydration heat measurement due to its unique advantages, including direct measurement of

N Nmax t 50 T C-S-H CH C3S (b-) C2S C3A C4AF TCR

Normalised heat flow, mW/g Peak value of N, mW/g The time when 50% of Q max is released, h Curing temperature, °C Calcium silicate hydrates Calcium hydroxide Tricalcium silicate (b-)dicalcium silicate Tricalcium aluminate Tetracalcium aluminoferrite Tailings-cement ratio

the heat generation rate, high reliability, real-time measurement and less calibration [27,28]. The hydration heat evolution, especially early-age hydration heat evolution, and the kinetics of cementitious materials have been widely investigated using the isothermal calorimeter technique. For example, Han et al., [29] investigated the hydration heat evolution and kinetics of blended cement containing steel slag at different temperatures. Tydlitát et al., [24] developed a largevolume isothermal calorimeter to explore the early-age hydration heat evolution in cement paste and mortar as a function of the water-cement ratio. The isothermal calorimeter was also utilised by Wang et al., [30] to investigate the hydration of early-strength low-heat Portland cementitious materials. Other noteworthy studies include those of Zhu et al., [31], Wang et al., [22], and Jin and Stephan [32]. Since containing a minor binder content, CPB might be more prone to the sulphate attack that deteriorates the strength development [33]. Therefore, the hydration heat evolution and kinetics of CPB could be different, which has not been thoroughly addressed in the literature. In this paper, the hydration heat evolution and kinetics of earlyage CPB under the coupled effects of cement content and temperature were studied. The cost of cement can contribute to up to approximate 75% of the total CPB cost, therefore, investigating the effect of the cement content on the CPB performance is significant for the mining industry [34]. The temperature can influence the CPB performance, from at least two aspects. On the one hand, the temperature of CPB structures can be significantly affected by various heat sources, such as the mining depth and human-induced temperature [35]. On the other hand, investigating the influence of the temperature on the hydration kinetics of CPB might reveal the hydration mechanisms from a thermodynamic perspective. Note that there are other influencing variables for the hydration heat evolution of CPB, such as the curing conditions [36,37], which should be investigated to gain a better understanding. The primary novelties of this study are as follows: (i) the coupled effects of the cement content and temperature on the hydration heat evolution of early-age CPB (up to 72 h) were investigated, (ii) the hydration heat evolution and kinetics of early-age CPB were discussed from a chemical (thermodynamics and kinetics) perspective, (iii) the Krstulovic-Dabic model was used to analyse the hydration kinetics of early-age CPB under the coupled effects of cement content and temperature.

2. Materials and methods 2.1. Materials The materials utilised in CPB consist of cement, mine tailings and mixing water, which are introduced as follows.

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L. Liu et al. / Construction and Building Materials 212 (2019) 283–294 Jidong No.425 common Portland cement was used according to China’s National Standard GB175-2007. The particle size distribution (PSD) of cement was measured by a laser diffraction particle size analyser (Malvern Mastersizer 2000, England). Fig. 1a shows the PSD of cement, as well as several widely-used indicators associated with PSD. The chemical composition of Jidong No. 425 cement was investigated by X-ray fluorescence (XRF) using an S8 Tiger spectrometer from Bruker, Germany. Fig. 1b shows the XRF pattern of cement and Table 1 summaries the chemical composition in the form of oxide content. As different clinker phases react with water at different rates, the phase composition was also investigated. The Bogue method, which is the most widely used mathematical procedure for determining the phase composition [38], was utilised in this study. Table 2 summarises the phase composition of the No. 425 cement. A detailed explanation of the PSD, XRF and Bogue methods has been well documented in the literature [38,39]. The mine tailings were sampled from a copper mine in China. First, fresh tailings slurry with a solids content of approximately 25 wt% was transported from the ore processing plant to plastic buckets for flocculation and sedimentation (FS). Second, the FS process was continued for 7 days before obtaining upperlayer clarified water, and the thickened tailings slurry was separated. The solids content for the thickened tailings slurry was approximately 60 wt%. Finally, the thickened tailings were dried and transported to Xi’an University of Science and Technology for analysis. The physical and chemical characteristics of the tailings were determined following the same procedure as that used for cement. Fig. 2 shows the PSD and XRF of the tailings, while Table 3 summaries its chemical composition. As suggested in [40], tailings can be regarded as well-graded when its coefficient of uniformity Cu  5. Also, a good gradation and a high compaction rate can be indicated by a coefficient of curvature (Cc) of 1–3. Therefore, the tailings used in this study were well-graded with a Cu of 8.55 and had a good gradation with a Cc of 1.71. According to the size classification from International Society of Soil Mechanics (ISSS) [41], the proportions of gravel (>2 mm), coarse sand (2–0.2 mm), fine sand (0.2–0.02 mm), silt (0.02–0.002 mm), and clay (<0.002 mm) in the copper tailings are 0.0%, 13.7%, 71.3%, 14.8% and 0.2%, respectively. Based on the tailings classification [42], the copper tailings can be categorised as fine tailings (the proportion of <20 lm is larger than 60%). SiO2 was the main component of the tailings, accounting for 58.7 wt%, followed by Al2O3 (14.8 wt%) and CaO (6.3 wt%). The authors also note that the tailings are free of sulphide minerals, as shown in Table 3. The tap water was used as the mixing water in this paper.

Table 2 Phase composition of cement obtained from the Bogue method. Formula

Component

Mass (%)

C3S C2S C3A C4AF CaSO4 Total

Tricalcium silicate Dicalcium silicate Tricalcium aluminate Tetracalcium aluminate ferrite Calcium sulphate

53.0 23.2 6.3 10.7 2.1 95.3

During the experiments, the sample chamber is used to house the fresh CPB sample while the reference chamber is used to house the reference sample. In this paper, water with the same thermal mass was used as the reference sample according to the suggestions in [44]. Because the influence of curing temperature was investigated in this study, considerable temperature variations may exist between the fresh CPB and the curing temperature. Thus, the CPB materials were stored and mixed at a temperature close to the curing temperature to minimise the influence of a temperature difference. The homogeneous CPB was immediately placed into the sample chamber, and continuous monitoring of N and Q was conducted up to 72 h. The solids content of the fresh CPB was kept at 76% based on the trial tests from the mine site. An orthogonal experimental design was performed with tailingscement ratios (TCRs) of 4, 6, 8, 10 and 12 and curing temperatures (T) of 20 °C, 30 °C and 40 °C. The selection of these values was conducted based on trial tests, mining operations (i.e., the mining depth) of the copper mine, and recommendations in the literature [20,32,45,46]. A detailed experimental procedure is shown in Fig. 3. The authors note here TCRs were used in the current study to represent the relative content of tailings and cement as the TCR is more widely used for CPB investigation. Together with solids content, the proportion of water, cement and tailings can be as easily calculated. In this study, the water/cement ratios for 4, 6, 8, 10 and 12 TCRs were 1.6, 2.2, 2.8, 3.5 and 4.1, respectively.

2.3. The Krstulovic-Dabic kinetic model 2.2. Experimental methods In this paper, the Krstulovic-Dabic model was employed to analyse N and Q measured with different TCR values and temperatures using the isothermal calorimeter [47]. The objective of utilising the Krstulovic-Dabic model is to reveal the hydration kinetics and mechanisms of CPB. The Krstulovic-Dabic model is one of the most frequently used kinetic models for different types of cementitious materials, such as cement containing slag [29], calcium aluminate cement containing MgAl2O4 spinel [31], and cement containing co-combustion ash [22].

The normalised heat flow (N) and cumulative heat (Q) of CPB were measured for 72 h using a TAM air isothermal calorimeter (Thermometric AB, Sweden) according to ASTM C 1702 [43]. Each channel in the isothermal calorimeter has two measuring chambers, namely the sample chamber and the reference chamber. This instrument has a wide temperature measurement range from 5 °C to 90 °C with a precise temperature control within ±0.02 °C.

Fig. 1. (a) PSD and (b) XRF patterns of cement.

Table 1 Chemical composition of cement. Oxide

CaO

SiO2

Al2O3

Fe2O3

MgO

SO3

K2O

Na2O

TiO2

Not identified

Mass [%]

64.13

19.19

4.50

3.33

1.82

1.06

1.04

0.41

0.24

2.28

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Fig. 2. (a) PSD and (b) XRF patterns of tailings.

above-mentioned basic processes than this paper are referred to [48–50]. There is one kinetic equation describing the relationship between the degree of hydration and the reaction time for each basic process, as follows.

Table 3 Chemical composition of tailings. Composition

SiO2

Al2O3

CaO

TFe

MgO

Other

wt .%

58.7

14.8

6.30

6.18

2.79

11.23

According to the Krstulovic-Dabic model, cement hydration can be divided into three basic processes, namely, nucleation and crystal growth (NG), a phase boundary reaction (I), and diffusion (D). During cement hydration, all three processes can occur separately or simultaneously. However, one process has the slowest reaction rate, namely, the rate-determining process, and it controls the overall development of cement hydration. Interested readers for a more detailed explanation of the

NG : ½lnð1  aÞ

1=n

¼ K 1 ðt  t0 Þ ¼ K 01 ðt  t0 Þ

ð1Þ

h i1 I : 1  ð1  aÞ1=3 ¼ K 2 r1 ðt  t0 Þ ¼ K 02 ðt  t0 Þ

ð2Þ

h i2 D : 1  ð1  aÞ1=3 ¼ K 3 r2 ðt  t0 Þ ¼ K 03 ðt  t0 Þ

ð3Þ

    where a represents the hydration degree; K 1 ðK 01 Þ; K 2 K 02 and K 3 K 03 represent the (apparent) rate constants of the NG, I and D processes, respectively; t 0 represents the time when the induction period ends; n is an exponent representing geometrical

Fig. 3. The experimental procedure in this study.

L. Liu et al. / Construction and Building Materials 212 (2019) 283–294

287

crystal growth (n = 1 represents needle growth, n = 2 represents sheet growth and n = 3 represents isotropic growth) [50,51]; and r represents the diameter of the participating particles in the reaction. By differentiating Eqs. (1)–(3), the kinetic Eqs. (4)–(6) can be obtained, which represent the hydration rates of the NG, I and D processes, respectively.

NG : da=dt ¼ F 1 ðaÞ ¼ K 01 nð1  aÞ½ ln ð1  aÞ

ðn1Þ=n

ð4Þ

I : da=dt ¼ F 2 ðaÞ ¼ 3  K 02 ð1  aÞ2=3

ð5Þ

h i D : da=dt ¼ F 3 ðaÞ ¼ 3  K 03 ð1  aÞ2=3 = 2  2ð1  aÞ1=3

ð6Þ

where F 1 ðaÞ, F 2 ðaÞ, and F 3 ðaÞ represent functions for the NG, I and D processes, respectively. The degree of hydration, a; can be calculated using the following equations from the hydration heat data [52]:

aðtÞ ¼ Q ðtÞ=Q max da=dt ¼ dQ =dt 

1 Q max

1 1 t 50 ðKnudsen equationÞ þ ¼ Q Q max Q max ðt  t0 Þ

ð7Þ ð8Þ Fig. 4. Typical hydration heat flow and cumulative heat patterns of early-age CPB.

ð9Þ

where Q ðtÞ is the heat released at the hydration time, t, after the induction period; dQ =dt is the hydration rate; Qmax is the maximum hydration heat release when the cementitious materials terminate the hydration; t 0 is the time when the induction stage ends; and t50 is the time when 50% of the Qmax is released. During the application of the Krstulovic-Dabic model, the hydration heat data measured using the isothermal calorimeter are substituted into Eq. (9) for the determination of Q max and t50 . This can be achieved from a linear fitting between 1=Q ðtÞ and 1=ðt  t 0 Þ. Then, aand da=dtcan be calculated by substituting Q max into Eqs. (7) and (8). The n and K 01 can be calculated by substituting a into Eq. (1). A linear fitting between lnðt  t0 Þ and lnðlnð1  aÞÞ can be observed, and its corresponding slope and intercept are used to determine n and K 01 , respectively. K 02 and K 03 can be determined in a similar way using the same equations. Then, all kinetic parameters are substituted into Eqs. (4)–(6) to find the relationship between F n ðaÞ; n ¼ 1; 2; 3 and a. The hydration heat evolution is an overall performance by all reactions at a particular time, which are highly correlated and difficult to isolate. In other words, the heat evolution could be negligible when it is high, meaning that comparable exothermic and endothermic reactions take place simultaneously. For example, if the heat-releasing rate by the exothermic reactions and the heat-absorbing rate by the endothermic reactions are ±0.5 mW/g at one time, then the overall heat flow is zero without external disturbances. Thus, the hydration kinetics analysis using the above model is an overall apparent hydration process. This fact is a limitation for hydration heat analysis, which applies not only to CPB but also to other types of cementitious materials [21].

The acceleration stage is mainly characterized by the reaction of C3S and a small amount of b-C2S. In this stage, the dissolution of C3S accelerates, resulting in an increasing rate of C-S-H growth. Moreover, the concentrations of Ca2+ and OH– increase steadily until the calcium hydroxide (CH) begins to precipitate. During the acceleration stage, the smaller grains of C3S can react completely, and the larger grains are surrounded by the precipitated C-S-H and CH [21]. The hydration rate in the deceleration and slow reaction stages is considered to be determined by the diffusion process. At these two stages, the early-age strength of CPB is developed. Additionally, the calorimetry peak for C3A hydration is often observed in the deceleration stage for modern cement. In some cases, the influence of C3A calorimetry peaks is counteracted by the decreasing heat release from the calcium silicate phases, depending on multiple variables such as the hydration of calcium silicate and, the amounts of C3A and calcium sulphide. The deceleration and slow reaction stages are possibly caused by the consumption of small particles and the lack of water or space [18]. The reaction of bC2S occurs mainly at the slow reaction stage, as there is often a reaction delay for b-C2S until approximately two weeks after mixing [44].

3. Results and discussion 3.2. Effect of the TCR on the hydration heat characteristics of CPB 3.1. Hydration heat characteristics of CPB Fig. 4 shows a typical hydration heat flow (N) pattern and cumulative heat (Q) pattern of early-age CPB. As shown, the overall process of hydration is similar to that of most cementitious materials, which can be divided into five stages: (1) initial dissolution stage, (2) induction stage, (3) acceleration stage, (4) deceleration stage, and (5) slow reaction stage [18,27,30]. Notably, the precise determination of the division points in Fig. 4 is difficult, so the current division is somewhat arbitrary. The initial dissolution stage is characterized by the rapid reactions between C3S and water. Upon wetting, C3S undergoes rapid dissolution in a few seconds. The dissolution rate of C3S then decelerates very quickly, which might be explained by the metastable barrier hypothesis or the slow dissolution step hypothesis. An induction stage follows the initial dissolution stage, which enables the cementitious materials with high workability. This induction stage could last for approximately 3 h when a chemical retarder is added or materials are annealed. This critical point is mainly believed to be influenced by the nucleation and growth of the calcium silicate hydrates (C-S-H) [18].

In the current study, the effect of the TCR on the hydration heat characteristics of CPB was investigated by varying the TCR from 4 to 12, as shown in Fig. 5. The TCR is a frequently used indicator during the application of CPB. In the case of other types of cementitious materials, the water-cement ratio (WCR) is often used to represent the cement content. Considering the solids content of 76% in the CPB slurry, TCR = 4, 6, 8, 10, and 12 corresponded to WCR = 1.58, 2.20, 2.86, 3.43, and 4.14, respectively. It is widely considered that a WCR of approximately 0.4 is sufficient for the full hydration of cement [53]. Thus, full hydration of cement can be achieved in CPB given enough time, even considering the effect of retarding caused by the internal relative humidity [54]. From Fig. 5a, it is clear that the TCR has an important influence on the N of CPB. Regardless of the temperature, a higher TCR resulted in a lower N. For example, the maximum N within 72 h (Nmax) was 0.33 mW/g when the TCR was 4 at 20 °C, which was decreased to 0.27 mW/g, 0.21 mW/g, 0.17 mW/g, and 0.15 mW/g. When the TCR was increased to 6, 8, 10 and 12, respectively. A similar influence was also observed on the Q of CPB, as shown in Fig. 5b. Taking the maximum Q (Qmax) as an example, the Qmax

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Fig. 5. The influence of TCR on the N and Q of CPB: (a) N at 20 °C; (b) Q at 20 °C; (c) N at 30 °C; (d) Q at 30 °C; (e) N at 40 °C; and (f) Q at 40 °C.

was decreased from 40.38 J/g to 17.54 J/g when the TCR was increased from 4 to 12. Furthermore, the influence of the TCR on N and Q values was almost independent of temperature. For example, the ratio of the Qmax at TCR = 12 to the Qmax at TCR = 4 was 0.434 when the temperature was 20 °C, compared to the value of 0.441 when the temperature was 40 °C. Overall, the influence of the TCR on N and Q is straightforward because a higher TCR corresponds to a lower cement content, which reduces the hydration rate due to the decreased amount of reactants available in CPB.

The conclusions about the effect of TCR on the hydration heat characteristics of cementitious materials are in good agreement with findings in the literature, such as in [55]. In addition to the local extremes of N and Q, the TCR was also found to influence the hydration stages of CPB. Fig. 6 illustrates the hydration stage when the temperature was 20 °C. As shown, an increasing TCR increases the induction stage and decreases the acceleration stage. For example, the acceleration stage lasted for 15.3 h when the TCR was 4, which was decreased to 13.4 h

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L. Liu et al. / Construction and Building Materials 212 (2019) 283–294 Table 4 Enthalpy values for main hydration reactions. Reaction

Equations

Enthalpy (kJ/mol)

C3S dissolution C2S dissolution C3A dissolution

C3S + (3-x + y)H ? Cx-S-Hy + (3-x)CH C2S + (2-x + y)H ? Cx-S-Hy + (2-x)CH C3A + CH + 12H ? C4AH13 C3A + 6H ? C3AH6 2C3A + C6A$3H32 + 4H ? 3C4A$H12(AFm) C3A + 3CSH2 + 26H ? C6A$3H32(AFt) C3A + 3C$H2 + 10H ? C4A$3H12 C4AF (+CH + H) ? C3(A,F)H6 C4AF + 3C$H2 + 30H ? C6A$3H32 + CH–FH3 C3S + (3-x + y)H ? Cx-S-Hy + (3-x)CH C2S + (2-x + y)H ? Cx-S-Hy + (2-x)CH C3S + (3-x + y)H ? Cx-S-Hy + (3-x)CH C2S + (2-x + y)H ? Cx-S-Hy + (2-x)CH

118 45 314 245 238 452 309 203 352 30

C4AF dissolution C-S-H precipitation CH precipitation

20

Note: In cement chemistry, the shorthand notation of C, S, H, A, $, F, represents CaO, SiO2, H2O, Al2O3, SO3, Fe2O3. Fig. 6. The influence of TCR on the hydration stage of CPB at 20 °C.

when the TCR was 12. This pattern was also observed at 30 °C and 40 °C. The duration of the deceleration stage was almost independent with of the TCR. Liu et al., [56] investigated the influence of basalt, which has a similar chemical composition to the tailings in this study, on the hydration heat of cementitious materials and found similar conclusions. As discussed in Section 3.1, the nucleation and growth process during cement hydration influences the termination of the induction stage. Gartner et al., [57] summarized four possible mechanisms for the onset of nucleation and growth, including nucleation and growth of C-S-H, growth of stable C-S-H, rupture of the initial barrier and nucleation of CH. In all of these mechanisms, the concentration of C-S-H/CH needs to be increased to a critical point. Since the cement content is decreased with a high TCR, more time might be required for the concentration of C-S-H/ CH to increase. Therefore, the induction stage is elongated. The authors also note that there are more tailings particles available

Fig. 8. Determination of Qmax and t50 from the linear fitting at TCR = 4 and temperature = 20 °C.

Fig. 7. The influence of temperature on the N and Q of CPB: (a) N at TCR = 4; (b) N at TCR = 6; (c) N at TCR = 8; (d) N at TCR = 10; (e) N at TCR = 12; and (f) Q at TCR = 4.

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with a higher TCR, thus providing more initial nucleation sites for C-S-H/CH and benefiting the nucleation. However, this increase in the initial nucleation sites may have a limited influence compared with the decreasing cement content. Moreover, the initial nucleation sites might be sufficient even at TCR = 4, in which case the tailings content is 4 times higher than the cement content. The shortening of the acceleration stage is mainly caused by the precipitation of C-S-H and CH around larger cement grains. With the increase in the TCR, the amount of larger cement grains is decreased. In these cases, a smaller amount of C-S-H/CH precipitation can surround the larger cement grains tightly, thus shortening the acceleration stage of hydration.

3.3. Effect of temperature on the hydration heat characteristics of CPB Fig. 7a–e shows the influence of temperature on the N of CPB. The most distinct influence was reflected in the Nmax and its time of appearance. An increase in the curing temperature from 20 °C to 40 °C increased the Nmax values and shortened its appearance time. Taking TCR = 4 as an example, the Nmax was 0.33 mW/g when the temperature was 20 °C, which was increased to 0.63 mW/g and 1.10 mW/g, when the temperature was increased to 30 °C and 40 °C, respectively. Therefore, the Nmax at 40 °C was approximately three times higher than that of the Nmax at 20 °C. In the case of the appearance time, the Nmax appeared at 20.6, 10.7 and 8.5 h when

Table 5 Knudsen equations for the determination of Qmax and t50 . TCR

Temperature/°C

Qmax (J/g)

t50(h)

Knudsen equation

4 6 8 10 12 4 6 8 10 12 4 6 8 10 12

20

60.3500 49.3340 38.7296 30.9215 28.2965 74.0741 61.9195 43.6681 38.4468 33.6360 77.8816 66.3130 46.8384 38.6100 33.1345

38.9514 41.2158 41.2773 41.6921 48.7450 29.7852 30.2326 32.9043 33.4646 32.4732 25.7826 23.9032 22.5864 25.0816 22.4163

1/Q = 0.01657 + 0.56145/(t-t0) 1/Q = 0.02027 + 0.71714/(t-t0) 1/Q = 0.02582 + 0.90607/(t-t0) 1/Q = 0.03234 + 1.12581/(t-t0) 1/Q = 0.03534 + 1.44890/(t-t0) 1/Q = 0.01350 + 0.38348/(t-t0) 1/Q = 0.01615 + 0.46325/(t-t0) 1/Q = 0.02290 + 0.71738/(t-t0) 1/Q = 0.02601 + 0.80806/(t-t0) 1/Q = 0.02973 + 0.88534/(t-t0) 1/Q = 0.01283 + 0.30703/(t-t0) 1/Q = 0.01508 + 0.33653/(t-t0) 1/Q = 0.02135 + 0.44880/(t-t0) 1/Q = 0.02590 + 0.59509/(t-t0) 1/Q = 0.03018 + 0.61823/(t-t0)

30

40

Fig. 9. Determination of kinetics parameters from the linear fitting at TCR = 4 and temperature = 20 °C: (a) NG process, (b) I process and (c) D process.

Table 6 Kinetics parameters of hydration process of different CPB. TCR

Ta/°C

n

K 01

K 02

K 03

Hydration mechanism

a1

a2

a2  a1

4 6 8 10 12 4 6 8 10 12 4 6 8 10 12

20

1.3397 1.3227 1.2769 1.2487 1.2202 1.5352 1.5030 1.3806 1.3071 1.2342 1.5893 1.5182 1.3991 1.3513 1. 2737

0.0265 0.0259 0.0262 0.0254 0.0240 0.0437 0.0425 0.0384 0.0384 0.0391 0.0655 0.0668 0.0661 0.0631 0.0630

0.00712 0.00697 0.00695 0.00679 0.00645 0.01133 0.01020 0.00885 0.00883 0.00857 0.01555 0.01510 0.01451 – –

0.00135 0.00124 0.00120 0.00112 0.00080 0.00178 0.00153 0.00127 0.00120 0.00112 0.00180 0.00173 0.00163 0.00151 0.00148

NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-I-D NG-D NG-D

0.1548 0.1518 0.1300 0.1269 0.1179 0.1245 0.1181 0.0874 0.0696 0.0381 0.1254 0.1028 0.0748 0.1326 0.1267

0.2582 0.2438 0.2374 0.2273 0.1748 0.2315 0.2084 0.2004 0.1907 0.1840 0.1640 0.1621 0.1599 0.1326 0.1267

0.1034 0.092 0.1074 0.1004 0.0569 0.107 0.0903 0.113 0.1211 0.1459 0.0386 0.0593 0.0851 0 0

Note:

30

40

a

represents temperature.

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the temperature was 20 °C, 30 °C and 40 °C, respectively. Thus, the appearance time was shortened by half to two-thirds when the temperature was increased from 20 °C to 40 °C. The conclusions about the influence of temperature on the Nmax and its appearance time are in good agreement with findings in the literature [58]. Moreover, such an influence of temperature was almost independent of the TCR. In other words, the Nmax at 40 °C was approximately three times higher than the Nmax at 20 °C when the TCR was 12. The foregoing results indicate that the acceleration stage was shortened with increasing temperature. According to the Maxwell–Boltzmann distribution, a higher temperature results in a larger fraction in the system with sufficient energy to overcome the energy barrier for dissolution [59]. As a result, high temperatures accelerate the dissolution of the cement phases. This is also applicable to the nucleation and growth of C-S-H/CH, which also has energy barriers that must be overcome for these processes to happen. Thus, the accelerating rate is increased with increasing temperature. This acceleration process advances the time when smaller cement grains are exhausted and the deceleration stage begins. The above discussions were mainly in terms of the reaction kinetics, instead of the thermodynamics where chemical equilibrium and the equilibrium constant are involved. It can also be observed that an increasing temperature shortened or diminished the induction stage. As previously discussed, the termination point of the induction stage is influenced by the nucleation and growth processes, which are in turn affected by the concentrations of C-S-H and CH in water. As high temperatures accelerate the dissolution of the cement phases, there are more reactants in water available for producing C-S-H and CH. Additionally, high temperatures provide essential energy for reactions that require some energy barriers to overcome. Therefore, the concentrations of C-S-H and CH are increased quickly, and the induction stage is short or non-existent with high temperatures. The shortening and disappearance of the induction stage can also be analysed by employing the metastable barrier hypothesis. High temperatures promote a quick dissolution of the metastable barrier by allowing more metastable C-S-H to overcome its dissolution energy barrier. Thus, the acceleration stage appears quickly once the metastable barrier is dissolved to a critical point. Fig. 7f presents Q curves at TCR = 4 and different curing temperatures. As shown, increasing the curing temperature increased the Q value throughout the measuring period. For example, Q at 72 h was increased by 30.9% from 40.4 J/g to 52.9 J/g when the temperature was increased from 20 °C to 30 °C. When increasing to 40 °C, Q was further increased to 57.8 J/g, corresponding to a 9.3% increase in Q. The influence of temperature on Q, remains unclear, although it has been widely observed without being clearly explained. Cement hydration is an overall exothermic process. Based on Le Chatelier’s principle, it seems correct to predict that the overall cement hydration will decrease with increasing temperature. Table 4 shows the enthalpy values of the main reactions during cement hydration [21,60]. As shown, the dissolutions of C3S, C2S, C3A, and C4AF are exothermic reactions (enthalpy < 0), while the precipitations reactions of C-S-H and CH are endothermic (enthalpy > 0). According to Le Chatelier’s principle, increasing the temperature will ‘retard’ the dissolution of cement phases, while it will ‘promote’ the precipitation of C-S-H and CH. However, the influence of temperature on Q within 72 h is also a chemical kinetics issue, instead of a thermodynamics issue. Thus, it can only be concluded that increasing the temperature will shift the dissolution of cement phases to reactants, while shifting the precipitation of C-S-H and CH to products at chemical equilibrium. In other words, the equilibrium constant of cement dissolution is decreased while the equilibrium constant of hydrate precipitation is increased with increasing temperatures.

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It is likely that a higher temperature will induce a smaller concentration of dissolved cement phases at chemical equilibrium. If there is no precipitation, then there would be less cement dissolution at a higher temperature. However, these dissolved cement phases are continuously consumed in terms of precipitated C-S-H/CH, leading to a continuous dissolution of cement phases. Therefore, the equilibrium constant, which determines the amount of reactants and products at chemical equilibrium, is less important until the point at which dissolution or precipitation cannot reach its chemical equilibrium. The influence of temperature is mainly concerned with chemical kinetics, and among the kinetic factors, increasing the temperature is

Fig. 10. Hydration rate curves of CPB at 20 °C: (a) TCR = 4, (b) TCR = 8 and (c) TCR = 12.

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likely to increase the rate of both dissolution and precipitation, as previously discussed. 3.4. Hydration kinetics of CPB The Krstulovic-Dabic model was utilised in this paper for hydration kinetics analysis. First, Qmax and t50 were determined using the Knudsen equation, where a linear fitting between 1=Q ðtÞ and 1=ðt  t 0 Þ was constructed. Fig. 8 illustrates the determination of Qmax and t 50 for CPB at TCR = 4 and temperature = 20 °C, and Table 5 summaries Qmax and t50 values, as well as their corresponding Knudsen equations, for all CPB slurries. As shown, the Qmax ranged from 28.30 J/g (TCR = 12, temperature = 20 °C) to 77.88 J/g

Fig. 11. Hydration rate curves of CPB at 30 °C: (a) TCR = 4, (b) TCR = 8 and (c) TCR = 12.

(TCR = 4, temperature = 40 °C), which was consistent with the findings in Sections 3.2 and 3.3. The t 50 value ranged from 22.42 h (TCR = 12, temperature = 40 °C) to 48.75 h (TCR = 12, temperature = 20 °C). The remaining kinetic parameters can be calculated through linear fitting based on the given equations. Fig. 9 illustrates the determination of the remaining kinetic parameters for the CPB slurry at TCR = 4 and temperature = 20 °C, and all results are summarized in Table 6. It can be seen that n decreased with the increase of TCR and the decrease of temperature. As n describes geometrical crystal growth, it is found that changing the TCR and temperature significantly influences the crystal growth geometry

Fig. 12. Hydration rate curves of CPB at 40 °C: (a) TCR = 4, (b) TCR = 8 and (c) TCR = 12.

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[61]. The reaction rate constant of the NG process (K1) was higher than that of the I process (K2) and the D process (K3), indicating that the reaction rate of NG was much faster than that of I and D. Furthermore, the K values decreased with an increasing TCR and a decreasing temperature because the cement content and hydration kinetics are influenced by temperature. The influence of tailings content, which is represented by TCR, on the n and K values found in this paper is consistent with findings in the literature [61,62]. The changing positions from NG to I and from I to D are also shown in Table 6, which are represented by a1 and a2, respectively. The value of a1 was decreased by increasing the TCR and temperature, indicating that the hydration reaction transformed from NG to I at a relatively low degree of hydration with an increased TCR and temperature. This result is consistent with the findings in Sections 3.2 and 3.3, in which the acceleration stage (controlled by the NG process) was shortened with an increasing TCR and temperature. Although the value of a2 was decreased by an increasing TCR and temperature, it cannot be concluded that the duration of the I process was also decreased because of the decreasing a1. The value of (a2  a2) did not show a clear pattern with the TCR and temperature. Additionally, the transformation from the I to D process cannot be correlated with the hydration stages in Section 3.1, as the role of the I process has not been well investigated in the hydration stages mentioned above. The hydration mechanism was found to be NG ! I ! D for most CPB slurries, except at TCR = 10/12 and temperature = 40 °C where the hydration mechanism was changed to NG ! D. By substituting the kinetics parameters listed in Table 6 into Eqs. (4)–(6), the hydration rate curves can be determined, as shown in Figs. 10–12. The Krstulovic-Dabic model was successful in representing the hydration heat evolution of CPB. These curves also confirmed that the cement hydration in CPB could be divided into three processes of NG, I and D and there was no I process at TCR = 12 and temperature = 40 °C. The CPB slurry at TCR = 10 and temperature = 40 °C, which also had no I process, is not shown in Fig. 12 for conciseness.

4. Conclusions An experimental investigation of the early-age hydration heat evolution of CPB was conducted and the coupled effects the TCR and temperature were analysed from a chemical viewpoint. The Krstulovic-Dabic model was utilised for hydration kinetics analysis. This study provided a better understanding of the hydration mechanisms of CPB, as well as its kinetics, which can improve CPB design and promote its application. Based on the results of this study, it was found that increasing the TCR decreased the values of the early-age N and Q due to the decreased cement content. Moreover, increasing the TCR elongated the induction stage and shortened the acceleration stage. Due to the influence of temperature on the kinetics of hydration reactions, increasing the temperature increased the Nmax values, shortened the appearance time of the Nmax, and shortened/diminished the induction stage. The n and K values from the Krstulovic-Dabic model decreased with the increase of TCR and the decrease of temperature. The hydration mechanism was found to be NG ! I ! D for most CPB slurries, except at TCR = 10/12 and temperature = 40 °C where the hydration mechanism was changed to NG ! D. In future studies, long-term hydration heat will be measured and other influencing variables, such as the consolidation effect, will be investigated. Additionally, the hydration mechanism can be better revealed using multiple techniques, such as thermogravimetry. Finally, the evolution of mineralogical composition

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during the curing of CPB can be investigated and correlated with the hydration heat results. Acknowledgements This research was supported by the National Natural Science Foundation of China (No. 51674188, 51874229, 51504182, 51404191), Shaanxi Innovative Talents Cultivate Program-Newstar Plan of Science and Technology (2018KJXX-083), Natural Science Basic Research Plan of Shaanxi Province of China (No. 2015JQ5187), Scientific Research Program funded by the Shaanxi Provincial Education Department (No. 15JK1466), China Postdoctoral Science Foundation (No. 2015M582685), and Outstanding Youth Science Fund of Xi’an University of Science and Technology (No. 2018YQ2-01). This research was also supported by the National Research Foundation of Korea (NRF) funded by the Minstry of Science, ICT & Future Planning (No. 2017R1E1A1A01075118). The corresponding author also thanks China Scholarship for the financial support during Ph.D. (No. 201606420046). Conflict of interest The authors declare no conflict of interest concerning the publication of this paper. References [1] E. Tranvik, M. Becker, B.I. Pålsson, J.P. Franzidis, D. Bradshaw, Towards cleaner production – Using flotation to recover monazite from a heavy mineral sands zircon waste stream, Miner. Eng. 101 (2017) 30–39. [2] Y. Zhao, R. Li, C. Ji, C. Huan, B. Zhang, L. Liu, Parametric study and design of an earth-air heat exchanger using model experiment for memorial heating and cooling, Appl. Therm. Eng. 148 (2019) 838–845. [3] M.B.C. Mangane, R. Argane, R. Trauchessec, A. Lecomte, M. Benzaazoua, Influence of superplasticizers on mechanical properties and workability of cemented paste backfill, Miner. Eng. 116 (2018) 3–14. [4] C. Qi, A. Fourie, Q. Chen, X. Tang, Q. Zhang, R. Gao, Data-driven modelling of the flocculation process on mineral processing tailings treatment, J. Cleaner Prod. 196 (2018) 505–516. [5] M. Fall, M. Benzaazoua, E.G. Saa, Mix proportioning of underground cemented tailings backfill, Tunn. Undergr. Space Technol. 23 (1) (2008) 80–90. [6] M. Fall, S.S. Samb, Effect of high temperature on strength and microstructural properties of cemented paste backfill, Fire Saf. J. 44 (4) (2009) 642–651. [7] Q. Chen, Q. Zhang, C. Qi, A. Fourie, C. Xiao, Recycling phosphogypsum and construction demolition waste for cemented paste backfill and its environmental impact, J. Cleaner Prod. 186 (2018) 418–429. [8] F. Cihangir, B. Ercikdi, A. Kesimal, H. Deveci, F. Erdemir, Paste backfill of highsulphide mill tailings using alkali-activated blast furnace slag: Effect of activator nature, concentration and slag properties, Miner. Eng. 83 (2015) 117–127. [9] A. Kesimal, E. Yilmaz, B. Ercikdi, I. Alp, H. Deveci, Effect of properties of tailings and binder on the short-and long-term strength and stability of cemented paste backfill, Mater. Lett. 59 (28) (2005) 3703–3709. [10] H. Lu, C. Qi, Q. Chen, D. Gan, Z. Xue, Y. Hu, A new procedure for recycling waste tailings as cemented paste backfill to underground stopes and open pits, J. Cleaner Prod. 188 (2018) 601–612. [11] E. Yilmaz, M. Fall, Paste Tailings Management, Springer, 2017. [12] L. Liu, Z. Fang, C. Qi, B. Zhang, L. Guo, K.-I. Song, Experimental investigation on the relationship between pore characteristics and unconfined compressive strength of cemented paste backfill, Constr. Build. Mater. 179 (2018) 254–264. [13] C. Qi, Q. Chen, A. Fourie, X. Tang, Q. Zhang, X. Dong, Y. Feng, Constitutive modelling of cemented paste backfill: A data-mining approach, Constr. Build. Mater. 197 (2019) 262–270. [14] L. Liu, Z. Fang, C. Qi, B. Zhang, L. Guo, K.I.I.L. Song, Numerical study on the pipe flow characteristics of the cemented paste backfill slurry considering hydration effects, Powder Technol. 343 (2019) 454–464. [15] C. Qi, X. Tang, X. Dong, Q. Chen, A. Fourie, E. Liu, Towards Intelligent Mining for Backfill: A genetic programming-based method for strength forecasting of cemented paste backfill, Miner. Eng. 133 (2019) 69–79. [16] B. Koohestani, A. Khodadadi Darban, P. Mokhtari, A comparison between the influence of superplasticizer and organosilanes on different properties of cemented paste backfill, Constr. Build. Mater. 173 (2018) 180–188. [17] C. Qi, Q. Chen, A. Fourie, Q. Zhang, An intelligent modelling framework for mechanical properties of cemented paste backfill, Miner. Eng. 123 (2018) 16– 27. [18] J.W. Bullard, H.M. Jennings, R.A. Livingston, A. Nonat, G.W. Scherer, J.S. Schweitzer, K.L. Scrivener, J.J. Thomas, Mechanisms of cement hydration, Cem. Concr. Res. 41 (12) (2011) 1208–1223.

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