**Hydrodynamic Study of AR Coupling Effects on Solid Circulation and Gas Leakages in a High-Flux In Situ Gasification Chemical Looping Combustion System**

#### **Xiaojia Wang \*, Xianli Liu, Baosheng Jin and Decheng Wang**

Key Laboratory of Energy Thermal Conversion and Control of Ministry of Education, School of Energy and Environment, Southeast University, Nanjing 210096, China; liuxl\_seu@163.com (X.L.); bsjin@seu.edu.cn (B.J.); dechengw76@126.com (D.W.)

**\*** Correspondence: xiaojiawang@seu.edu.cn

Received: 27 September 2018; Accepted: 16 October 2018; Published: 18 October 2018

**Abstract:** In situ gasification chemical looping combustion (iG-CLC) is a novel and promising coal combustion technology with inherent separation of CO2. Our previous studies demonstrated the feasibility of performing iG-CLC with a high-flux circulating fluidized bed (HFCFB) riser as the fuel reactor (FR) and a counter-flow moving bed (CFMB) as the air reactor (AR). As an extension of that work, this study aims to further investigate the fundamental effects of the AR coupling on the oxygen carrier (OC) circulation and gas leakages with a cold-state experimental device of the proposed iG-CLC system. The system exhibited favorable pressure distribution characteristics and good adaptability of solid circulation flux, demonstrating the positive role of the direct coupling method of the AR in the stabilization and controllability of the whole system. The OC circulation and the gas leakages were mainly determined by the upper and lower pressure gradients of the AR. With the increase in the upper pressure gradient, the OC circulation flux increased initially and later decreased until the circulation collapsed. Besides, the upper pressure gradient exhibited a positive effect on the restraint of gas leakage from the FR to the AR, but a negative effect on the suppression of gas leakage from the AR to the FR. Moreover, the gas leakage of the J-valve to the AR, which is directly related to the solid circulation stability, was exacerbated with the increase of the lower pressure gradient of the AR. In real iG-CLC applications, the pressure gradients should be adjusted flexibly and optimally to guarantee a balanced OC circulation together with an ideal balance of all the gas leakages.

**Keywords:** in situ gasification chemical looping combustion; high-flux circulating fluidized bed; counter-flow moving bed; gas leakage; coupling mechanism

#### **1. Introduction**

Chemical looping combustion (CLC), which possesses an inherent feature of isolating CO2 during the combustion process, has been regarded as a promising novel combustion technology with a low energy penalty for carbon capture [1,2]. A conventional CLC system is mainly comprised of a fuel reactor (FR) and an air reactor (AR). The fuel introduced into the FR is oxidized to CO2 and H2O by a solid oxygen carrier (OC). The reduced OC particles are then transferred to the AR where they are re-oxidized upon contact with air. Compared to conventional combustion methods, the fuel will no longer mix with N2 in a CLC process, by means of the circulation of OC between the two reactors. Therefore, the flue gas leaving the FR will only contain CO2 and H2O with a complete conversion of the fuel, which enables efficient and energy-saving CO2 capture via the condensation of H2O [3,4].

On the basis of fuel types, CLC technology can be broadly categorized into gas-fueled CLC and solid-fueled CLC. So far, studies on gas-fueled CLC have been extensively conducted in different prototype reactors [5–14]. In recent years, considering the price advantage of solid fuels (coal as the main representative) in comparison with those of gaseous fuels, solid-fueled CLC began to attract increasing attention [15–21]. One of the coal-fueled CLC approaches is the so-called in situ gasification chemical looping combustion (iG-CLC) which involves three reaction steps between two reactors [20,21]. As shown in Figure 1, two of the reaction steps occur in the FR, i.e., the coal gasification reaction (Step 1) and the subsequent oxidation reactions of the gasification gases by the OC (Step 2). The other reaction step proceeds in the AR, i.e., the reoxidation reaction of the OC by air (Step 3).

**Figure 1.** Schematic diagram of an in situ gasification chemical looping combustion (iG-CLC) system.

Up to now, some experimental studies on iG-CLC systems have been carried out with various reactor designs [16,18,22–28]. An iG-CLC configuration with a bubbling fluidized bed (BFB) as the FR and a circulating fluidized bed (CFB) as the AR was first proposed by Berguerand and Lyngfelt [16] from Chalmers University of Technology (10 kWth), in view of the advantages of sufficient solid residence time in the BFB and favorable gas-solid contact in the CFB. Still, with CFBs as the ARs, some other FR designs have also been put forward. Shen et al. [18] from Southeast University adopted a spout-fluid bed (SFB) instead of the BFB as the FR (10 kWth), in which the strong solid mixing enhanced the gas-solid contact and reaction efficiencies. Thon et al. [22] from Hamburg University of Technology designed a two-stage BFB as the FR (25 kWth) for the purpose of enhancing the conversion of combustible gases. Bayham et al. [23] from Ohio State University utilized a counter-current moving bed as the FR (25 kWth) with OC particles flowing downwards and gases flowing upwards, showing high competitiveness in the combustion performance and CO2 purity. Adánez et al. [24] from CSIC proposed a CFB as the FR (20 kWth) in order to achieve intense gas-solid contact and reaction. This kind of dual circulating fluidized bed (DCFB) design (i.e., both the FR and AR are CFBs) can also be found in the units by Markström et al. [25] from Chalmers University of Technology (100 kWth), Ma et al. [26] from Huazhong University of Science and Technology (5 kWth), and Ströhle et al. [27] from Technische Universität Darmstadt (1 MWth). In addition, some other units with different design concepts have also been constructed and operated. Xiao et al. [28] from Southeast University developed a 50 kWth iG-CLC unit where the AR and FR were designed as a BFB and a CFB, respectively. Pressurized conditions were achieved in their work, demonstrating the positive role of high-pressure operation in carbon conversion, CO2 capture concentration, and combustion efficiency.

In our previous studies, we also proposed an iG-CLC system based on the high-flux circulating fluidized bed (HFCFB) technology [29–32]. Specifically, an HFCFB riser was designed as the FR because it can provide not only a favorable gas-solid contact but also sufficient solid holdups over the whole reactor height, which should be very beneficial to the interphase reaction efficiency. Specifically, compared to the common CFB FRs [24–27], the higher OC inventory in this HFCFB FR can potentially compensate for the low reactivity of OC to a certain degree, and hence this high-flux iG-CLC system will have a large advantage on the operation cost by virtue of the use of low-grade natural iron ores with a lower reactivity as the OCs. Certainly, the high-flux operation will inevitably lead to a higher pressure drop. Therefore, greater energy consumption is required to compensate for the pressure loss. However, compared to the cost reduction from the use of cheap OCs, the slight increase in energy consumption should be acceptable. A counter-flow moving bed (CFMB), in view of its advantages of steady solid flow, low pressure drop, and compact structure, was employed as the AR. Besides, it can be simply placed in the middle of the HFCFB downcomer to enhance the simplicity and stabilization of the whole system. Moreover, an inertial separator-based carbon stripper was specially designed to separate the reduced OC from the FR to the AR for reoxidation, and recirculate the unconverted char for further conversion.

Based on the design concept of this high-flux iG-CLC system, we successively built and tested a cold visualization experimental device operating at the ambient temperature [30,31] and a hot pilot-scale 20 kWth unit operating at the high-temperature conditions between 800–1000 ◦C [32]. After a series of tests, the steady operation with favorable gas-solid flow and reaction performance of the whole system could be realized under certain conditions, preliminarily verifying the feasibility of this design. With the deepening of the research, we have found that, compared to most iG-CLC systems with indirect serial structures of the two reactors [16,24–28], the direct coupling of the CFMB AR into the downcomer of the HFCFB FR in our system indeed contributes to the stabilization and controllability of the whole system. However, this coupling method of the AR also inevitably brings about greater challenges on the control of gas leakages between the two reactors, which will, in turn, affect the matching principle of reactors, such as the OC circulation characteristics.

The aim of this study is to investigate the fundamental effects of the AR coupling on the system operation stability and gas leakages. The pressure distribution characteristics of the whole system and the adaptability of solid circulation flux were first investigated with the cold-state visualization device for the proposed iG-CLC system. The effects of the upper and lower pressure gradients of the AR on the OC circulation and the gas leakages were further established. Moreover, by giving consideration to the AR coupling effects under various operational conditions, one optimal operating condition was recommended to demonstrate the adjustment feasibility for balanced solids circulation with low gas leakages during the iG-CLC process.

#### **2. Materials and Methods**

#### *2.1. Experimental Device*

Figure 2 gives a schematic representation of the experimental device for the proposed iG-CLC system under cold conditions. Here, only a brief description of the experimental setup is provided. A more detailed description can be found in our initial studies of this system [31].

#### 2.1.1. Main Assembly

The main assembly predominantly consists of a FR (5), a carbon stripper (6), a downcomer (7), an AR (8), a J-valve (11), and a bag filter (12).

The FR (5) is an HFCFB riser with a height of 5.8 m and an inner diameter of 60 mm. The AR (8) is a CFMB, mainly consisting of a gas inlet (10), a gas outlet (9), a cylindrical channel (0.418 m inner diameter × 0.5 m height), and a cone channel (0.2 m height). The carbon stripper (6) is an inertial separator. For the purpose of visualization, some sections of the FR, the AR, and the downcomer are made of plexiglas.

**Figure 2.** Schematic of the visualization experimental device of the iG-CLC system. 1—computer, 2—A/D converter, 3—differential pressure transducer, 4—riser gas chamber, 5—fuel reactor (FR), 6—carbon stripper, 7—downcomer, 8—air reactor (AR), 9—AR gas outlet, 10—AR gas inlet, 11—J-valve, 12—bag filter, 13—filter, 14—gas analyzer, 15—digital camera, 16—rotameter, 17—FR gas chamber, 18—AR gas chamber, 19—90 kW air compressor, 20—18 kW air compressor, 21—tracer gas. P—pressure gauge, Q—gas flow.

#### 2.1.2. Gas Supply System

The air stream from a 90 kW air compressor (Nanjing Compressor Co. Ltd., Nanjing, China) (19) was introduced into the FR to circulate the OC particles. Another air stream from an 18 kW air compressor (Guangzhou Panyu JOYO Air Compressor Factory, Guangzhou, China) (20) was fed into the AR through a tube distributor.

A high-purity carbon monoxide (99.99%) stream was used as the tracer gas (21) which would be fed into the FR, the AR, and the J-valve in turn during the testing stages.

#### 2.1.3. Data Acquisition System

Gas flow rates in the FR, AR, and J-valve were controlled and measured by calibrated rotameters (16).

The pressures of monitoring nodes were measured by pressure manometers and a multi-channel differential pressure transducer (3). During the experiments, the pressures of the two reactors could be adjusted by back-pressure regulators.

The tracer gas concentrations at the outlets of the two reactors were measured by a gas analyzer (MRU, Neckarsulm, Germany) (14).

A digital camera (15) was used to photograph the flow regimes and capture some special flow phenomena through the visualization sections during the experiments.

#### *2.2. Material*

The OC used in this study was a natural iron ore from Harbin, China. Prior to the experiments, the OC particles were crushed and sieved to a mean diameter of 0.43 mm. More details on this OC material can be found in Table 1.


**Table 1.** Main physical properties of the oxygen carrier.

#### *2.3. Experimental Procedures*

The particles of the OC were first packed in the downcomer. Then, the gas stream from the 90 kW air compressor was introduced into the FR from the J-valve and FR distributor to drive the OC particles for circulation. After the particle circulation became balanced, another gas stream from the 18 kW air compressor was introduced into the AR and coupled into the original circulation system. As the system rebounded to a steady state, the effects of operating parameters began to be tested with the real-time monitoring of the system flow state, by means of pressure tracking, time sampling, gas tracer, and so on. A detailed explanation of the data processing can be found in Section 2.4.

#### *2.4. Data Evaluation*

#### 2.4.1. Gas Flow Rates

The gas flow rates in this study were all normalized to the standard state with the subscript *sta*. According to the conservation of mass, the total gas flow rate of the system can be calculated as [31]

$$Q\_{in,sta} = Q\_{1,sta} + Q\_{2,sta} + Q\_{3,sta} + Q\_{4,sta} = Q\_{a,sta} + Q\_{b,sta} = Q\_{out,sta} \tag{1}$$

where *Qin,sta* and *Qout,sta* represent the total inlet air flow rate and the total outlet air flow rate of the system, respectively. *Q*1,*sta*, *Q*2,*sta*, *Q*3,*sta*, and *Q*4,*sta* are the inlet air flow rate of the FR distributor, the fluidizing air flow rate of the J-valve, the aeration air flow rate of the J-valve, and the inlet air flow rate of the AR, respectively. *Qa,sta* and *Qb,sta* are the outlet air flow rates of the FR and the AR, respectively.

All the inlet/outlet gas flow rates mentioned above were measured by calibrated rotameters except *Qa,sta* which can be deduced from Equation (1).

$$Q\_{a,sta} = Q\_{1,sta} + Q\_{2,sta} + Q\_{3,sta} + Q\_{4,sta} - Q\_{b,sta} \tag{2}$$

The FR superficial gas velocity can be calculated as:

$$
\mathcal{U}\_{f,sta} = \frac{\mathcal{Q}\_{f,sta}}{A\_f} \tag{3}
$$

where *Af* is the sectional area of the FR. *Qf,sta* represents the total inlet air flow rate of the FR, which can be estimated as:

$$Q\_{f,sta} = Q\_{1,sta} + Q\_{2,sta} + (1 - f\_3)Q\_{3,sta} \tag{4}$$

where *f* <sup>3</sup> represents the J-valve leakage ratio, and a detailed explanation of this parameter can be found in Section 2.4.3.

#### 2.4.2. Solid Circulation Flux

The solid circulation flux can be estimated as:

$$\mathcal{G}\_s = \frac{\rho\_b u\_s A\_d}{A\_f} = \frac{\rho\_b A\_d}{A\_f} (\Delta H/\text{t}) \tag{5}$$

where *ρb*, *us,* and *Ad* represent the bulk density of the OC, the downward flow velocity of the OC particles in the upper dipleg, and the sectional area of the upper downcomer, respectively. Δ*H* is a scale height in the upper dipleg for the measurement of solid circulation flux, and *t* is the measured duration of the traced OC particles passing through the scale height.

#### 2.4.3. Gas Leakage Ratios

The distribution of the FR exhaust gas can be measured through the use of tracer gas 1. *f* 1, named as FR leakage ratio, represents the gas leakage ratio of the FR into the AR [31].

$$f\_1 = \frac{Q\_{b,sta} \mathbf{x}\_{b,CO}}{Q\_{a,sta} \mathbf{x}\_{a,CO} + Q\_{b,sta} \mathbf{x}\_{b,CO}} \tag{6}$$

where *xa,CO*, and *xb,CO* are the concentrations of tracer gas 1 measured at the outlets of the separator and the AR, respectively.

Similarly, the distribution of the exhaust gas from the AR inlet can be investigated by using tracer gas 2. *f* 2, named as the AR leakage ratio, represents the gas leakage ratio of the AR into the FR [31].

$$f\_2 = \frac{Q\_{a,sta} \mathbf{x}\_{a,CO}^{\prime}}{Q\_{a,sta} \mathbf{x}\_{a,CO}^{\prime} + Q\_{b,sta} \mathbf{x}\_{b,CO}^{\prime}}\tag{7}$$

where *x <sup>a</sup>*,*CO* and *x <sup>b</sup>*,*CO* are the concentrations of tracer gas 2 measured at the outlets of the separator and the AR, respectively.

Moreover, the distribution of the J-valve aeration air can be measured by the use of tracer gas 3. *f* 3, the so-called J-valve leakage ratio, represents the gas leakage ratio of the aeration air into the AR.

$$f\_3 = \frac{Q\_{b,sta} \mathbf{x}\_{b,CO}^{''}}{Q\_{a,sta} \mathbf{x}\_{a,CO}^{''} + Q\_{b,sta} \mathbf{x}\_{b,CO}^{''}} \tag{8}$$

where *x <sup>a</sup>*,*CO* and *x <sup>b</sup>*,*CO* are the concentrations of tracer gas 3 measured at the outlets of the separator and the AR, respectively.

#### 2.4.4. Pressure Gradients

The upper pressure gradient (Δ*P*1/*H*1) represents the pressure gradient between the AR and the separator.

$$
\Delta P\_1 / H\_1 = \left(\frac{P\_b + P\_c}{2} - P\_i\right) / H\_1 \tag{9}
$$

where *Pb* and *Pc* are the pressures of the AR outlet and inlet, respectively. *Pi* represents the pressure at the top position of the upper dipleg of the AR, i.e., the pressure at the interface of the dense phase and dilute phase of the upper downcomer. As the pressure loss is very small in the dilute phase region of the upper downcomer, the value of *Pi* can be approximated by the pressure at the underside of the separator (i.e., *P*11). *H*<sup>1</sup> is the solid-seal height of the upper dipleg of the AR.

The lower pressure gradient (Δ*P*2/*H*2) represents the pressure gradient between the J-valve and AR.

$$
\Delta P\_2 / H\_2 = \left(P\_d - P\_{12}\right) / H\_2 \tag{10}
$$

where *Pd* is the pressure at the bottom position of the lower dipleg of the AR, and *P*<sup>12</sup> the pressure at the top position of the lower dipleg. *H*2, named as the solid-seal height of the lower dipleg, is represented by the height difference between the pressure monitoring nodes *P*<sup>12</sup> and *Pd*.

#### **3. Results and Discussion**

As shown in Table 2, experiments were carried out with wide ranges of operating conditions (e.g., solid mass flux, superficial gas velocity, and pressure gradients) in order to present the primary flow behaviors of the proposed iG-CLC system. Thereinto, the operating condition of *Gs* = 310 kg/m2·s, *Uf,sta* = 10.7 m/s, *Q*4,*sta* = 48 m3/h is defined as the reference condition which realized a balanced operation but involved relatively obvious gas leakage features to facilitate the observation and analysis.

**Table 2.** Main operation conditions of the iG-CLC tests.


#### *3.1. Pressure Balance and Solid Circulation of the iG-CLC System*

An appropriate pressure balance is vital to the balanced gas flow and solid circulation in an iG-CLC system; therefore, 19 pressure measuring nodes were mounted on the main assembly, the labels can be seen in Figure 2. Figure 3 shows the pressure profile of the system under the reference condition. It can be found that the sum of the pressure drop within a solid circulation loop was zero. This indicates the whole system was self-stabilizing and hence a small perturbation would not break the balance of solid circulation. The pressure drop along the height of the FR was conspicuous with the total pressure difference between *P*<sup>1</sup> at the bottom and *P*<sup>10</sup> at the top reaching up to 12.8 kPa. In contrast, the pressure drop of the AR (i.e., *Pc*−*Pb*) was only about 1.5 kPa with a CFMB structure.

Due to the coupling of the AR in the middle of the downcomer, the dipleg of the downcomer was divided into two parts, the upper dipleg and the lower dipleg of the AR. Here, we adopted *H*1, see Figure 2, as the solid-seal height of the upper dipleg and *H*2, see Figure 2, as the solid-seal height of the lower dipleg. As shown in Figure 3, the pressures of the two diplegs both linearly increased from top to bottom, indicating that the solid flow structures in the two diplegs under the reference condition both belonged to the negative pressure differential flow. Moreover, the existence of the AR made the pressure distributions of the two diplegs relatively independent of each other. To be specific, the pressure distribution along the upper dipleg was mainly determined by the outlet pressures of

the AR and FR, and the solid-seal height of the upper dipleg (i.e., *H*1), while the pressure distribution along the lower dipleg was determined by the pressures of the AR and the J-valve together with the solid-seal height of the lower dipleg (i.e., *H*2).

**Figure 3.** Pressure profile of the iG-CLC system under the reference condition.

The solid circulation rate is significant to the performance of an iG-CLC system, which determines the carrying capacity of oxygen and heat transferred by the OC from the AR to the FR. Figure 4 shows the distributions of the solid circulation flux *Gs* corresponding to the FR superficial gas velocity *Uf,sta*. It can be seen that a wide range of *Gs* from 170 to 480 kg/m2·s had been achieved, indicating a good adaptability of this iG-CLC system on the OC circulation flux. This also demonstrates the positive role of the direct coupling method of the AR in the stabilization and controllability of the whole system. Thus, in the hot operation process, this iG-CLC system can achieve the feasible adjustment of oxygen and heat transfer according to actual situations. In particular, the capacity of high solid circulation flux (*Gs* ≥ 200 kg/m2·s) can greatly increase the OC inventory in the FR, and thus compensate for the possible low reactivity of the OC. Therefore, this iG-CLC system also provides the feasibility of the use of low-grade natural iron ores with lower reactivity as the OCs.

**Figure 4.** Distributions of the solid circulation flux corresponding to the FR superficial gas velocity.

#### *3.2. Effect of the AR Coupling on the Solid Circulation*

The direct coupling of the CFMB AR into the HFCFB system will inevitably affect the gas-solid flow behaviors of the system. In our previous work, we indeed notice that the OC circulation seemed to be affected during the adjustment process of the AR back pressure [31]. Therefore, it is significant to understand the fundamental effects of the AR coupling on the flux and further the stability of OC circulation.

**Figure 5.** Pressure profiles of the parallel AR tube distributors under the reference condition: (**a**) schematic of the AR tube distributors, and (**b**) pressure profiles.

Figure 5 shows the pressure profiles of the parallel AR tube distributors under the reference condition. It can be seen that the inlet pressures of the five distributors were basically the same, demonstrating the realizability of even flow and distribution of gas-solid phases in the AR under certain conditions. In this context, a series of tests were carried out to investigate the fundamental effects of the AR coupling on the OC circulation flux and the system stability. We found that the OC circulation was actually determined by more factors that had interactions with each other (e.g., the pressures of the two reactors and the J-valve, the solid-seal heights above and under the AR), not just the AR back pressure. Here, we proposed a combined influence factor, the so-called upper pressure gradient Δ*P*1/*H*1, which integrated the pressures of the two reactors and the solid-seal height of the upper dipleg, and hence should be able to more reasonably reflect the characteristics of OC circulation. Figure 6a shows the variations of the solid circulation flux with the upper pressure gradient Δ*P*1/*H*<sup>1</sup> while keeping all other parameters constant. It can be seen that the whole process could be divided into three sequential stages. In the first stage (−3.3 kPa/m ≤ Δ*P*1/*H*<sup>1</sup> ≤ 8.9 kPa/m), the solid circulation flux had a linear increase from 215 to 260 kg/m2·s with the increase of <sup>Δ</sup>*P*1/*H*1, which was defined as the stage of circulation strengthening. In the second stage (8.9 kPa/m < Δ*P*1/*H*<sup>1</sup> < 10.7 kPa/m), the so-called transition stage, the solid circulation flux began to decrease. In the third stage (Δ*P*1/*H*<sup>1</sup> ≥ 10.7 kPa/m), the solid circulation flux would drastically decline until the circulation collapsed completely, which was defined as the stage of circulation collapse.

The above results indicated that, with the increase in the upper pressure gradient Δ*P*1/*H*1, the solid circulation flux would increase initially and later decrease until the circulation collapsed. First, in the stage of circulation strengthening, the material seal of the upper dipleg could overcome most of the gas leakage from the AR to the separator in spite of the increase of the upper pressure gradient. Better still, the increasing upper pressure gradient due to the increase of the AR back pressure also gave rise to the decrease of the lower pressure gradient Δ*P*2/*H*2, and further, the decline of the aeration air leakage ratio from the J-valve to the AR (i.e., *f* 3), which indirectly enhanced the driving force of the J-valve aeration air for the OC circulation, and hence led to the increase of solid circulation flux. Then, starting from the transition stage, the material seal of the upper dipleg would be gradually destroyed by the increasing upper pressure gradient with the creation of large bubbles in the upper dipleg, as

shown in Figure 6b (the bubbles were circled in red). The retrograde motion of the bubbles in the upper dipleg greatly blocked the downward flow of the OC, which was the reason the solid circulation of the whole system was broken. Worse still, the appearance of bubbles also meant the massive leakage of N2 from the AR into the separator, and hence the great reduction in the CO2 concentration. Given the above, under the operational conditions shown in Figure 6a, the upper pressure gradient Δ*P*1/*H*<sup>1</sup> should be limited to 8.9 kPa/m (i.e., within the stage of circulation strengthening), where the OC circulation was balanced and adjustable.

**Figure 6.** Variations of the solid circulation patterns with the upper pressure gradient: (**a**) variations of the solid circulation flux; and (**b**) snapshot of the state of circulation collapse.

#### *3.3. Effect of the AR Coupling on the Gas Leakages*

The gas leakages between reactors are one of the key factors affecting the performance of an iG-CLC system, such as the CO2 capture efficiency, the CO2 capture concentration, and even the solid circulation flux. As shown in Figure 7, there are three possible routes of gas leakages in our iG-CLC system: From FR to AR (i.e., FR leakage ratio *f* 1), from AR to FR (i.e., AR leakage ratio *f* 2), and from J-valve to AR (i.e., J-valve leakage ratio *f* 3). It can be found that the AR is the critical component associated with each leakage route. Hence, it is very meaningful to understand the influence mechanism of the AR coupling on the gas leakages in order to discover feasible solutions.

After a series of tests, the gas leakages were also found to be determined by the pressure gradients of the AR, which integrated the effects of the pressures of the two reactors and the J-valve, and the solid-seal heights of the upper and lower diplegs. Figure 8 shows the variations of the FR leakage ratio *f* <sup>1</sup> and the AR leakage ratio *f* <sup>2</sup> with the upper pressure gradient Δ*P*1/*H*<sup>1</sup> under a high solid flux condition (*Gs* = 200 kg/m2·s). It can be seen that the FR leakage ratio *<sup>f</sup>* <sup>1</sup> could be effectively reduced with an increase in the upper pressure gradient Δ*P*1/*H*1. When Δ*P*1/*H*<sup>1</sup> exceeded 0.7 kPa/m, the value of *f* <sup>1</sup> would decline to about zero, indicating the gas leakage from the FR to the AR had almost disappeared. However, we also observed a contrary trend of the AR leakage ratio *f* <sup>2</sup> with the upper pressure gradient Δ*P*1/*H*1. First, *f* <sup>2</sup> was kept near zero when Δ*P*1/*H*<sup>1</sup> increased from −2.2 to 1.6 kPa/m, indicating the upper dipleg could seal the AR gas stream perfectly within this range of Δ*P*1/*H*1. When Δ*P*1/*H*<sup>1</sup> exceeded 1.6 kPa/m, *f* <sup>2</sup> would rise rapidly, indicating the material seal in the upper dipleg began to gradually lose its effectiveness.

In the commercial application, if the FR leakage ratio *f* <sup>1</sup> is excessive, the CO2 capture efficiency will be greatly reduced. Moreover, more unreacted char will be carried by the leaked gas of the FR into the AR for combustion, which will increase the risk of OC sintering, and further influence the operation stability of the whole system. On the other hand, if the AR leakage ratio *f* <sup>2</sup> becomes too large, a large amount of N2 from the AR will bypass into the separator and mix with the FR exhaust gas, resulting in a substantial decrease in the CO2 concentration. Worse still, the gas leakage may also impede the downward flow of the solid in the upper dipleg and break the system circulation stability, as mentioned in Section 3.2. Therefore, both the FR leakage ratio *f* <sup>1</sup> and the AR leakage ratio *f* <sup>2</sup> must be limited to lower values. However, in view of the opposite effects of the upper pressure gradient Δ*P*1/*H*<sup>1</sup> on *f* <sup>1</sup> and *f* 2, a coordination control and optimized matching is inevitable. Under the involved operation conditions, the range of Δ*P*1/*H*<sup>1</sup> between −2.1 to 3.0 kPa/m should be an optimal region for adjustment, in which the values of *f* <sup>1</sup> and *f* <sup>2</sup> could be limited to 3%, together with a favorable solid circulation.

**Figure 7.** The possible routes of gas leakages in the proposed iG-CLC system.

**Figure 8.** Variations of FR leakage ratio and AR leakage ratio with the upper pressure gradient.

On the other hand, the lower pressure gradient Δ*P*2/*H*<sup>2</sup> could be used to reflect the gas leakage ratio of the J-valve to the AR. Figure 9 presents the variations of the J-valve leakage ratio *f* <sup>3</sup> with the lower pressure gradient Δ*P*2/*H*2. When the lower pressure gradient Δ*P*2/*H*<sup>2</sup> increased from 2.3 to 6.9 kPa/m, the J-valve leakage ratio *f* <sup>3</sup> had an obvious increase from 6.0% to 21.3% with a near-linear trend. This indicated a negative effect of the lower pressure gradient on the suppression of the gas leakage from the J-valve to the AR. In the iG-CLC application, a small amount of gas leakage from the J-valve to the AR will rarely affect the circulation stability of the system, and hence can be accepted. However, an excess gas leakage will cause the aeration gas stream of the J-valve to no longer contribute to the solid circulation, but impede the downward flow of the particles. Worse yet, the excess gas leakage will lower the air inflow to the AR, and hence reduce the thermal power of the proposed iG-CLC system. Therefore, under the involved operation conditions, the lower pressure gradient Δ*P*2/*H*<sup>2</sup> should be limited within 6.0 kPa/m, and thus the J-valve leakage ratio *f* <sup>3</sup> could also be controlled within a low value (20%) with a favorable solid circulation.

**Figure 9.** Variations of the J-valve leakage ratio with the lower pressure gradient.

#### *3.4. Performance Optimization of the AR Coupling*

Based on the above analysis, the coupling of the AR has important effects, by virtue of the upper pressure gradient Δ*P*1/*H*<sup>1</sup> and the lower pressure gradient Δ*P*2/*H*2, on the characteristics of gas leakages and even the solid circulation stability. In the practical operation process, we can adjust the relevant parameters (i.e., the pressures of the two reactors and the J-valve, and the solid-seal heights in the downcomer) flexibly and optimally to guarantee the pressure gradients within the optimal ranges for an ideal performance of operation and reaction.

In order to better exhibit the effect of the AR coupling, we carried out a comparison of gas leakages between the reference condition and an optimal condition, as shown in Table 3. On the basis of the coupling criteria proposed in Section 3.3, the upper pressure gradient Δ*P*1/*H*<sup>1</sup> and lower pressure gradient Δ*P*2/*H*<sup>2</sup> in the optimal test were selected to be 3.5 kPa/m and 5.0 kPa/m, respectively. It could be found that for the reference condition, although achieving a balanced solid circulation in the whole system, an unsatisfactory gas leakage of the AR (i.e., *f* 2) was observed, indicating considerable mixing of N2 from the AR into the FR exhaust gas stream, and hence an obvious reduction in the CO2 concentration. However, the good news is that, with an optimization of the pressure gradients, the iG-CLC unit, under the optimal condition, inhibited the gas leakages, together with a favorable solid circulation. This demonstrates the significance of the study of the AR coupling mechanism in the high-flux iG-CLC system for the achievement of high CO2 capture efficiency and

CO2 capture concentration under a balanced system operation, which could provide vital information and experience for the design of future large-scale coal-fired CLC power plants.


**Table 3.** Comparison of gas leakages between the reference condition and optimal condition.

#### **4. Conclusions**

On the basis of the previous feasibility studies of a high-flux iG-CLC system, this work further investigated the AR coupling effects on the system operation stability and gas leakages with a cold-state visualization device, enabling the design parameters and operating conditions. The following conclusions can be drawn from the present study:

(1) The iG-CLC system exhibited favorable pressure distribution characteristics and good adaptability of solid circulation flux, demonstrating the positive role of the direct coupling method of the AR in the stabilization and controllability of the whole system.

(2) With the increase of the upper pressure gradient of the AR, the OC circulation flux would increase initially and later decrease until the circulation collapsed, demonstrating the crucial effect of the AR coupling on the OC circulation flux and further the circulation stability. In the real iG-CLC applications, the upper pressure gradient of the AR should be limited within the stage of circulation strengthening in order to guarantee a balanced and adjustable OC circulation.

(3) The gas leakage ratios of the FR and the AR were determined by the upper pressure gradient of the AR, while the gas leakage ratio of the J-valve was determined by the lower pressure gradient. In the iG-CLC applications, we can adjust the pressures of the two reactors and the solid-seal heights in the downcomer flexibly and optimally to ensure the two pressure gradients within the optimal ranges for an ideal balance of all the gas leakages.

(4) By giving consideration to the AR coupling effects under various operation conditions comprehensively, one operating condition with 3.5 kPa/m for the upper pressure gradient and 5.0 kPa/m for the lower pressure gradient was recommended. Under this condition, the gas leakages between the two reactors could be limited to 3%, and the gas leakage of the J-valve could also be below 20% to guarantee the solid circulation. This demonstrates the significance of the study of the AR coupling mechanism in the high-flux iG-CLC system for the achievement of high CO2 capture efficiency and CO2 capture concentration under a balanced system operation.

**Author Contributions:** X.W. conceived and designed the experiments; X.L. performed the experiments; X.W. and X.L. analyzed the data; B.J. supervised the research; X.W., X.L. and D.W. wrote the manuscript.

**Funding:** This research was funded by the National Natural Science Foundation of China (grant numbers 51741603, 51806035, 51676038), the Natural Science Fund project in Jiangsu Province (grant number BK20170669), the Fundamental Research Funds for the Central Universities (grant number 2242018K40117), and the Guangdong Provincial Key Laboratory of New and Renewable Energy Research and Development (grant number Y707s41001).

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


© 2018 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

## *Article* **Experimental Study on Spray Breakup in Turbulent Atomization Using a Spiral Nozzle**

#### **Ondˇrej Krištof 1, Pavel Bulejko 1,\* and Tomáš Sv ˇerák 1,2**


Received: 15 September 2019; Accepted: 26 November 2019; Published: 3 December 2019

**Abstract:** Spiral nozzles are widely used in wet scrubbers to form an appropriate spray pattern to capture the polluting gas/particulate matterwith the highest possible efficiency. Despite this fact, and a fact that it is a nozzle with a very atypical spray pattern (a full cone consisting of three concentric hollow cones), very limited amount of studies have been done so far on characterization of this type of nozzle. This work reports preliminary results on the spray characteristics of a spiral nozzle used for gas absorption processes. First, we experimentally measured the pressure impact footprint of the spray generated. Then effective spray angles were evaluated from the photographs of the spray and using the pressure impact footprint records via Archimedean spiral equation. Using the classical photography, areas of primary and secondary atomization were determined together with the droplet size distribution, which were further approximated using selected distribution functions. Radial and tangential spray velocity of droplets were assessed using the laser Doppler anemometry. The results show atypical behavior compared to different types of nozzles. In the investigated measurement range, the droplet-size distribution showed higher droplet diameters (about 1 mm) compared to, for example, air assisted atomizers. It was similar for the radial velocity, which was conversely lower (max velocity of about 8 m/s) compared to, for example, effervescent atomizers, which can produce droplets with a velocity of tens to hundreds m/s. On the contrary, spray angle ranged from 58◦ and 111◦ for the inner small and large cone, respectively, to 152◦ for the upper cone, and in the measured range was independent of the inlet pressure of liquid at the nozzle orifice.

**Keywords:** spiral nozzle; gas absorption; spray atomization; droplet size; droplet velocity

#### **1. Introduction**

Gas–liquid absorption processes for gaseous contaminants removal are crucial in diverse industrial fields and are basic means for air pollution mitigation associated with large-scale industrial operations. To ensure the reduction of the gaseous pollutants released from such processes, many industrial facilities adopt a gas scrubbing system as a post-treatment of produced polluted air. Gas scrubbers are rather complicated devices, in which the polluted air is cleaned using a sprayed liquid, mostly various aqueous solutions depending on the gas to be removed. This includes, for example, CO2 and VOC removal [1–3], flue gas desulphurization [4–6], and ammonia separation [7]. Last but not least, scrubbing systems have widely been recognized in separation of particulate matter [8–12], the release of which is due mostly to various combustion processes [13–15] and, for example, comminution technologies [16,17]. This is very important due to associated health and environmental concerns [18,19]. Comparing the ability to remove particulate and gaseous pollutants, gas scrubbers have lower efficiency in one cleaning cycle but are able to treat a significantly higher amount of polluted air compared to, for example, air filtration or membrane contactors [20–23]. In these applications, filter/membrane causes additional

resistance to airflow, thus disabling processing of huge amounts of polluted air generated in large-scale industrial processes. On the other hand, gas scrubbers work with sprayed absorption liquid media, which must then be subjected to another processing step, such as regeneration (if possible) or disposal, which can be very expensive and energy demanding [24]. Due to these facts, there is a large field for development and optimization of these systems in terms of operational parameters, chemicals used or spraying arrangement including usage of extremely broad spectrum of nozzles. Nozzles are crucial in a plethora of industrial applications including dust control [25], spray cooling [26–30], hydraulic descaling [31], and play a vital role in gas scrubbers as they can provide an adequate spray pattern of the absorption liquid. This is necessary to create the largest possible contact surface with gas phase to ensure process intensification, appropriate mass transfer and thus separation efficiency. Such a type of nozzle, which has scarcely been studied in terms of the spray characteristics, is the spiral nozzle.

Spiral nozzles have been found in different applications including flue gas desulfurization and spray towers [4,5,32], spray drying [33], distillation [34], petrochemical industry [35], and fire suppression [36,37]. Several studies tried to focus on the basic spray characteristics including droplet-size distribution, the inlet liquid pressure–flowrate relationship, and mass spray density [38–41] at different hydrodynamic conditions. Some authors did a research on the spray surface geometry [42] and even air-assisted atomization using the spiral nozzle [43]. Important works related to the present study are further described in detail. Li et al. [4] investigated spray characteristics of spiral nozzles used in flue gas desulfurization. They observed the flowrate was linear with a square root of pressure and droplet diameter was a power function of pressure. The droplet diameter variations with pressure were similar for spiral nozzles with different orifice, while the spray angle varied slightly at pressures higher than 40 kPa. Zhang et al. [32] studied spray characteristics of spiral nozzles with different diameters using particle image velocimetry at different spray pressures. They observed both, the spray pressure and nozzle diameter, to have an influence on the spray angle, droplet diameter, droplet size uniformity, and sprayed area diameter. In a biomass pyrolysis experiments, they found a nozzle with a diameter of 5.6 mm and the liquid pressure of two bars to be ideal for the quenching of pyrolysis vapors. In another study, Zhou et al. [41] conducted experiments on the flow distribution characteristics of a low-pressure high-flux spiral nozzle using a flow distribution testing system. They analyzed the effect of nozzle size on the flow distribution by varying several parameters including nozzle dimensions (nozzle length to diameter ratio), radial flowrate, position of sprayed surface, and spraying angle. The results indicated that with increasing nozzle length to diameter ratio, the flowrate decreases and spraying angle increases. With increasing pressure, the relationships were the same as in the above mentioned works. Dong et al. [5] compared desulfurization performance of scrubbers with spiral and Dynawave nozzle. Li et al. [42] developed a spray surface geometry model of a spiral nozzle with involute atomization. They further simulated the model using MATLAB, which validated its effectiveness and provided a theoretical basis for designing and manufacturing the spiral nozzles. Wasik et al. [38] studied the influence of nozzle type (including a TF6 spiral nozzle) on a mass spray density. For other works on spiral nozzles, especially those with different application, refer to [33–35].

In this study, we focus on a spiral nozzle (Figure 1a), which generates an atypical cone-like spray consisting of three concentric hollow cones (Figure 1b), thus creating a full cone pattern (for further details refer to [44]). The pattern is formed via a complicated nozzle geometry without axis of symmetry and gradually narrowing orifice, in which the liquid is sprayed from several rebound surfaces at different rake angles. This nozzle is quite different compared to spiral channel nozzles [45] or spiral flow nozzles [46,47] with which it can be confused. Generally, a lot of works have been carried out on nozzles creating a full-cone spray in various application fields [48–58]. Conversely, the amount of works on nozzles with a spiral geometry is quite limited despite the fact that they possess several very interesting features [59]:


**Figure 1.** A 3D model of the spiral nozzle studied (**a**) and the concentric hollow cone pattern produced by the nozzle (**b**).

Therefore, we investigate the pressure impact footprint of the falling liquid (cross-sectional pattern of the spray), spray radial, and tangential velocity distribution using the laser Doppler anemometry (LDA). The pressure impact footprint of the spray is generally important in spray cooling applications; in gas absorption processes this is not typically used. However, this can be easily used for validation of CFD data, as it is practically simpler to measure pressure impact compared to using the LDA method which is not always available. Further, we aimed at the spray morphology. This involved droplet-size distribution, spray angle, and determining the area of primary and secondary breakup of the spray using a classical photography modified with an Nd:YAG (neodymium-doped yttrium aluminum garnet) pulse laser, and assess the breakup areas using selected dimensionless numbers, which has not been carried out before for similar type of nozzle. We also compared the droplet diameter with selected theoretical relationships used for prediction of droplet and ligament diameter. Finally, we tried to outline flow conditions for ideal nozzle operation.

#### **2. Underlying Phenomena**

Liquid atomization (i.e., a breakup of liquid jet into dispersed fine droplets) is a complex process involving several physical/chemical phenomena that take place simultaneously. The spray characteristics (e.g., morphology, droplet size distribution, etc.) are strongly dependent on the atomizer used, especially on its size and geometry. Further, it is dependent on the physical properties of the fluids involved (i.e., the atomized liquid and the environment into which the liquid is sprayed (usually ambient air)). The spray atomization is strongly influenced by the liquid density, viscosity, and surface tension. The effect of density is rather lower, as indicated by the experimental data [60]. Conversely, the influence of the surface tension is quite essential. Surface tension is a force acting against the formation of a new surface area. Atomization involves two main phases (i.e., primary and secondary atomization). In the primary atomization, the disruptive forces act against the consolidating forces and cause oscillations of the liquid. Once the disruptive forces are stronger than the consolidating ones (surface tension), the bulk liquid disintegrates into smaller formations (ligaments, larger drops). Then the secondary atomization occurs (i.e., larger droplets or ligaments split into smaller droplets in a gas caused by either greater relative velocity or turbulence) [61]. A governing parameter relating disruptive inertial and restorative surface tension forces is the non-dimensional Weber number:

$$\mathcal{W}\varepsilon = \frac{\rho \nu^2 d}{\sigma},\tag{1}$$

where ρ, *v*, *d,* and σ are the fluid density, absolute radial velocity of liquid, characteristic dimension, and surface tension, respectively. The larger the Weber number, the higher the tendency toward the liquid breakup [62]. Viscosity is another very important parameter affecting the droplet size distribution and mainly the flow mode inside the atomizer, thus influencing the spray pattern/morphology. The influence of viscosity on the flow in the nozzle is quite complex and depends on the type of atomizer. Generally, drop size increases with increasing viscosity and delays the liquid jet breakup [60]. Both surface tension and viscosity decrease the tendency of the jet/sheet to disintegrate, which is accounted for by the Ohnesorge number (i.e., the ratio of viscous to surface tension forces):

$$Oh = \frac{\mu\_1}{\sqrt{\sigma \rho \mathbf{u} \mathbf{d}}} \,\tag{2}$$

where ρ<sup>l</sup> and μ<sup>l</sup> are the liquid density and viscosity, respectively. Many spray nozzles form a liquid sheet from bulk liquid prior to the atomization itself. The liquid sheet exits the nozzle orifice and may oscillate, which results in the formation of liquid ligaments, which are then broken into droplets. The droplet size is mostly in the same order as the liquid sheet thickness [63]. The primary atomization is strongly dependent on the liquid jet Weber number:

$$\mathcal{W}c\_{\text{jet}} = \frac{\rho\_1 \nu\_{\text{jet}}^2 d\_0}{\sigma},\tag{3}$$

where *d*<sup>o</sup> is the nozzle orifice diameter and νjet is the jet velocity (i.e., liquid velocity inside the nozzle prior to exiting the orifice). If the jet Weber number is lower, the surface tension forces impede the formation of a new surface area, thus preventing the liquid sheet breakup. Conversely at a larger jet Weber number, the breakup occurs due to the inertial forces to completely dominate over the surface tension forces causing the liquid sheet to tear into ligaments and droplets [64]. A critical value of the jet Weber number describes the onset of a decrease of the radial breakup distance. This value is typically around 1000 depending on the nozzle type [65]. After the first breakup phase, the secondary atomization may occur which is indicated by the gas Weber number:

$$\mathcal{W}e\_{\mathfrak{E}} = \frac{\rho\_{\mathfrak{E}} \nu\_{\text{AR}}^2 d\_{\text{D}}}{\sigma},\tag{4}$$

where ρ<sup>g</sup> is the ambient gas density. Droplet viscous forces are significant for *Oh* > 0.1. For *Oh* below this value, the breakup was observed to be independent of *Oh*. For *Oh* < 0.1, the transition *We* for individual breakup modes were practically constant as reported previously [66–68]. The individual modes are droplet deformation/vibrational breakup for 0 < *We* < 11, bag breakup for 11 < *We* < 35, multimode breakup for 35 < *We* < 80, sheet thinning for 80 < *We* < 350, and catastrophic breakup for *We* > 350. These are; however, not valid for conditions with higher *Oh*. The individual modes of secondary atomization are often depicted in *We*-*Oh* space, refer, for example, to [68–71]. To assess objectively which regime factually took place in the atomization process, several correlations for critical Weber number (*We*c) were proposed for *Oh* → 0. One of such correlation was suggested for Oh < 4 by Gelfand [72] as follows:

$$\mathcal{W}e\_{\mathbb{C}} = \mathcal{W}e\_{\mathbb{C}\text{Oh}\to\mathbb{0}} \Big(1 + 1.5\mathcal{O}h^{1.74}\Big),\tag{5}$$

where *We*c*Oh*→<sup>0</sup> is the critical *We* at low *Oh*, as listed above for individual modes.

The liquid sheet is formed via impinging the edge layer of the liquid jet on the surface of the helix as indicated in the CFD model (Figure 2). Therefore, we can expect an analogy with the radial spread of a liquid jet over a horizontal plane. This can be used to estimate the liquid sheet thickness (*t*sh) based on the free-surface similarity boundary layer concept as developed by Watson [73] and used

by Ren et al. [74] and Zhou and Yu [75]. Assuming the sheet flow on the helix surface be turbulent, the sheet thickness at the edge of the helix can be estimated as follows:

$$t\_{\rm sh} = \frac{d\_{\rm o}^2}{8r\_{\rm h}} + \frac{0.0245 d\_{\rm o}^{1/5} r\_{\rm h}^{4/5}}{Re\_{\rm o}^{1/5}},\tag{6}$$

where *d*<sup>o</sup> and *r*<sup>h</sup> are the nozzle orifice diameter (11.25 mm) and width of the helix (6.45 mm), respectively, and *Re* is the Reynolds number at the nozzle orifice calculated as follows:

$$\mathcal{R}\mathfrak{c}\_{\mathfrak{o}} = \frac{d\_{\mathfrak{o}}\nu\_{\mathfrak{j}\mathfrak{e}}\rho\_{\mathfrak{l}}}{\mu\_{\mathfrak{l}}}\tag{7}$$

**Figure 2.** A CFD model illustration of the formation of the liquid sheet by impinging the peripheral jet layer onto the surface of the helix (**a**) and a photograph of the same with the applied screen filter (**b**).

To estimate the spray velocity in a given radial location (*r*), it is necessary to consider the effect of viscous interaction with the helix surface. Therefore, a non-dimensional sheet thickness is defined as the ratio of the actual thickness (*t*sh) to an inviscid sheet thickness (*t*sh0):

$$\delta = \frac{t\_{\rm sh}}{t\_{\rm sh0}} = 1 + \frac{0.196}{Re\_{\rm o}^{1/5}} \left(\frac{r}{d\_{\rm o}}\right)^{9/5}.\tag{8}$$

Note that *t*sh0 is the first term in Equation (6). The average sheet velocity at the edge of the helix can be calculated as [74]:

$$\nu\_{\rm sh} = \frac{K \sqrt{\Delta p}}{2 \pi r\_{\rm h} t\_{\rm sh0} \delta} \,, \tag{9}$$

where *K* is the flow factor, which is a characteristic constant of the nozzle and Δ*p* is the pressure drop at the nozzle (inlet pressure of the liquid). There is a relationship between the volumetric flowrate (*Q*V) and the inlet water pressure (Δ*p*) as follows:

$$Q\_V = K \sqrt{\Delta p}.\tag{10}$$

For the spiral nozzle used in this study, the value is about 75. Further breakup of the sheet into ligaments and droplets is due to inherent instabilities caused by the wave growth. The wavelength at the sheet breakup governs the size of the ligaments and ultimately the droplet diameter. Dombrowski and Johns [76] developed a theory to predict the wave instability of liquid sheets in an inviscid gas. The model assumes sinusoidal waves to be on the liquid sheet, and the force balance is performed considering the inertial, pressure, and surface tension forces be associated with the wave displacement. After simplification, the force balance can be expressed as follows:

$$
\left(\frac{\partial f}{\partial \tau}\right)^2 + \frac{\mu\_\mathrm{L} \bar{\omega}^2}{\rho\_\mathrm{l}} \left(\frac{\partial f}{\partial \tau}\right) - \frac{2\left(\rho\_\mathrm{g} \bar{\omega} \nu\_\mathrm{sh}^2 - \sigma \bar{\omega}^2\right)}{\rho\_\mathrm{l} t\_\mathrm{sh}} = 0,\tag{11}
$$

where τ is time and *f* and ω˜ are the breakup parameter and wavenumber, respectively. The breakup parameter, also called dimensionless wave amplitude, was first investigated by Weber [77] who obtained a value of 12 and further confirmed by Dombrowski and Hooper [78] who stated that this value is constant regardless of the experimental conditions. However, in this study we better used the following correlation as it can be a function of the nozzle geometry [63]:

$$f = R e^{0.07} W e^{0.37},\tag{12}$$

where the Weber and Reynolds numbers are calculated for jet (i.e., liquid properties), liquid jet velocity prior to entering the nozzle orifice, and nozzle orifice diameter as the characteristic length. Another attempt was to estimate theoretically the ligament and droplet sizes based on the wave instabilities, which are the main cause of sheet disintegration. With an assumption of attenuating sheet and the formed ligaments be of cylindrical shape, the diameter estimate can be expressed as follows [76]:

$$d\_{\rm L} = 2 \left(\frac{4}{3f}\right)^{1/3} \left(\frac{k^2 \sigma^2}{\rho\_{\rm g} \rho\_l \nu\_{\rm sh}^2}\right)^{1/6} \left(1 + 2.6 \mu\_1 \sqrt[3]{\frac{k \rho\_{\rm g}^4 \nu\_{\rm sh}^8}{6f \rho\_l^2 \sigma^5}}\right)^{1/5},\tag{13}$$

where *k* for a uniform velocity radiating sheet can be expressed as follows:

$$k = \frac{r t\_{\rm sh}}{\nu\_{\rm sh}}.\tag{14}$$

Accordingly, the estimate of the ligament breakup time τL,bu can be written as follows [74,77,79]:

$$
\pi\_{\rm L,bu} = 24 \sqrt{\frac{2\rho\_l}{\sigma}} \left(\frac{d\_{\rm L}}{2}\right)^{3/2}. \tag{15}
$$

Based on the ligament size *d*L, the droplet diameter *d*<sup>D</sup> can be calculated [76,77]:

$$d\_{\rm D} = 1.882 d\_{\rm L} \left( 1 + 3 \text{Oh}\_{\rm sh} \right)^{1/6} \text{ } \tag{16}$$

where *Oh*sh is the Ohnesorge number based on the sheet thickness.

#### **3. Experimental**

#### *3.1. Spiral Nozzle*

A TF-28 150 asymmetric spiral nozzle (BETE, USA) was used in this study. Figure 3 shows a 3D scan of the nozzle obtained using an ATOS Tripple Scan 8M camera and its profile. The nozzle is made of Teflon and has a narrowing spiral-like orifice. The spiral is divided into three sections based on the rake angle of the rebound surfaces related to the nozzle *Z*-axis. Liquid falling on the individual rebound surfaces of the helix forms a full cone. The full cone can be considered for a combination of three water curtains in the form of hollow concentric cones (Figures 1 and 4). Based on their position to each other, they can be called as the outer (upper) cone, inner large, cone and inner small cone, as indicated in Figure 4a.

**Figure 3.** The BETE TF-28 150 spiral nozzle with designation of the longitudinal axis (*Z*-axis), views from different planes, and longitudinal cross sections (plane cuts).

**Figure 4.** An illustration of the main streams of the spray generated (**a**) and a depiction of the sprayed water beams numbered (blue); the distances of the beginning of the water beams from the nozzle orifice and the distances of individual impact planes from the nozzle orifice (**b**).

#### *3.2. Pressure Impact Footprint*

Measurement of the pressure impact footprint was done to visualize the pressure patterns generated by the sprayed water. During spraying the water curtain fell down on surface connected to an electronic tensometric pressure sensor. This surface was placed on a position system, which was operated using a computer software. The real impact pressure values were recorded depending the sensor position related to the nozzle. The impact measurement was performed twice, each with a different arrangement of the impact surface. In each experiment, different hydrodynamic conditions and different distances of the nozzle from the impact surface (planes, Figure 4b) were adopted.

#### 3.2.1. Measuring the Impact Pressure in Four Planes

In the first set, pressure impact footprint of the water curtain was measured. This was done in five planes (Figure 4b) (i.e., in the distances of 60, 80, 100, and 140 mm from the nozzle orifice). In each experiment, an inlet water pressure was 2 bars corresponding to a flowrate of 1.76 L/s. These are optimal hydrodynamic conditions for the operation of the nozzle tested. The experimental setup is shown in Figure 5. It was a metal plate (1) placed on an upper moving arm (5), in the middle of which was a dismountable metal casing (2) with a bevel-like top. At the top of the casing, an impact surface (3) was attached and connected to a tensometric pressure sensor placed inside the casing. The impact surface was circular with a diameter of 12 mm. The water falling down on the impact surface generates an impact pressure, which is taken by the sensor and the signal is recorded by the data acquisition system. The upper arm (5) was moved using a step motor in the direction of *X*-axis (Figure 5). This system was placed on the top of a bottom moving arm (6), which was moving in the direction of *Y*-axis (Figure 5). The area under the nozzle (4) was scanned in a length of 300 mm in the *X*-direction and 600 mm in the *Y*-direction, which was the maximum possible range of the experimental device used. The shift of both arms was set to 5 mm.

**Figure 5.** Experimental setup for the impact measurement: Bed plate (1), metal casing (2), impact surface (3), spiral nozzle (4), upper movable arm (5), lower movable arm (6).

#### 3.2.2. Measuring the Impact Pressure in One Plane at Different Water Inlet Pressures

These experiments (Figure 6) were performed to clarify the effect of different inlet water pressures (the pressure of water measured prior to entering the nozzle orifice) on the impact footprint. The impact surface (2) was in the form of a circular plate with a diameter of 200 mm, in the center of which was an opening (3) with a diameter of 1 mm. The opening was connected to a closed space under the plate where the tensometric pressure sensor was placed (4). The space was completely filled up with water prior to the measurement to ensure the impact pressure be transferred to the sensor through the water fluctuations. The remaining parts of the experimental setup including the moving arms in the *X* and *Y* axes were identical with the previous measurements (Figure 5). The only difference was the extent of the area scanned, which was 240 mm in both directions with a step of 3 mm. All the measurements were performed for plane 2 (i.e., in a distance of 70 mm from the nozzle orifice; Figure 4b). The hydrodynamic conditions are shown in Table 1.

**Figure 6.** Experimental setup for the impact measurement with different water inlet pressures: Spiral nozzle (1), impact surface (2), pressure sensor opening (3), pressure sensor covering (4), bed plate of the impact surface (5).


**Table 1.** Experimental conditions at pressure impact measurement.

#### *3.3. Spray Morphology*

Spray morphology was observed inside a transparent laboratory-scale sprinkle chamber made of PVC glass (Figure 7a). The chamber with a water reservoir were fixed on a frame (Figure 7b) together with a centrifugal pump providing water circulation through the nozzle. The experiments were performed at 20 ◦C and an inlet pressure of 0.9 bar corresponding to a flowrate of 1.19 L/s. This flowrate was selected due to ideal properties of the spray for measuring in the experimental device due mainly to the size of the transparent chamber. Lower flowrate caused improper spray properties including very narrow span and practically no breakup of the liquid stream exiting the nozzle into ligaments/droplets. Higher flowrate caused significant rebound of the liquid stream from the chamber wall, which had a negative effect on the snapshots taking.

**Figure 7.** A 3D model of the laboratory sprinkle chamber (**a**) and a real view of the same (**b**).

For a detailed evaluation of the spray, an apparatus as shown in Figure 8 was used (for the real view refer to supplementary Figure S1). A classical photography with a modified type of lightning (i.e., an Nd:YAG pulse laser lightning (2) with a pulse length of 5 ns) was used (a spray snapshot using classical photography and modified with laser is shown in Figure S2). The images were taken using a Canon D70 camera (1) with a Canon EF 10 mm f/2.8 USM Macro objective. From the images, the areas of primary and secondary atomization of the spray were observed. The images were further used to measure droplet size using the Stream Motion software (Olympus Corporation, Shinjuku, Japan).

**Figure 8.** A scheme of the experimental setup for the assessment of the spray morphology/kinetics with the Canon D70 camera (1), Nd:YAG pulse laser (2), spiral nozzle (3), centrifugal pump (4), flowmeter (5), manometer (6), and valve (7).

*Processes* **2019**, *7*, 911

From the measured droplet sizes, distribution curves were plotted and fitted with suitable distribution functions. The appropriateness of the fitting was then tested using the Kolmogorov–Smirnov test at a significance level of 0.05. Two density distribution functions, which are frequently used to describe droplet size in sprays were chosen (i.e., log-normal and Rosin-Rammler distribution) [80]. Log-normal density distribution can be expressed as follows:

$$f(D) = \frac{1}{\overline{\sigma}\_D \sqrt{2\pi}} \exp\left[-\frac{\left(\ln D\_i - \overline{\mu}\_D\right)^2}{2\overline{\sigma}\_D^2}\right] \tag{17}$$

where *Di*, μ*D*, and σ*<sup>D</sup>* are the droplet diameter, mean, and standard deviation of the distribution, respectively. Rosin-Rammler density distribution is described by the following equation:

$$f(D) = \frac{\alpha}{\beta} \left(\frac{D\_i}{\beta}\right)^{\alpha - 1} \exp\left[-\left(\frac{D\_i}{\beta}\right)^{\alpha}\right].\tag{18}$$

where α and β are empirical constants of the distribution. Further representative means of the droplet size were calculated (i.e., surface-weighted mean), which is used in the area of absorption processes where interphase area between two phases is important:

$$D\_{20} = \left[\frac{\sum N\_i D\_i^2}{\sum N\_i}\right]^{\frac{1}{2}}.\tag{19}$$

In addition, it is widely used Sauter mean diameter, which characterizes the spray fineness and is often used to calculate the efficiency and rate of mass transport in chemical reactions. It is a ratio of droplet volume to its surface area:

$$D\_{32} = \frac{\sum N\_i D\_i^3}{\sum N\_i D\_i^2}.\tag{20}$$

Finally, volume weighted mean was calculated as follows:

$$D\_{43} = \frac{\sum N\_i D\_i^4}{\sum N\_i D\_i^3}.\tag{21}$$

Another parameter studied was the spray angle, which was assessed using two methods. The first method was based on the pressure impact footprint curves, which were approximated using an Archimedean spiral. The other method was based on a visualization of the liquid spray in an open space to determine the spraying angle of individual cones using the Stream Motion software. Both methods are further explained in detail in the Results and Discussion section.

#### *3.4. Spray Kinetics*

Another method used was the laser Doppler anemometry (LDA), which provided us with information about the spray velocity distribution. The LDA setup was arranged on the same apparatus as for the experiments for the spray morphology evaluation (Figure 8). Thus, we obtained radial and tangential velocities from the upper and inner large cone. The LDA was performed for eight horizontal planes (Figure 9). In each plane, the spray velocity was measured in two perpendicular axes (the *X*-axis in yellow, and the *Y*-axis in red).

The beginning of the spray velocity measurement in the *X*-axis was at a distance of 110 mm from the nozzle center, as shown in Figure 9a (first position). Then, the spray velocity was measured after each 5 mm up to a distance of 165 mm from the nozzle center (12 positions in total). The spray velocity in the *Y*-axis was measured in a distance of 145 mm from the nozzle center and ranged from −80 to 80 mm (Figure 9a, detail shown in Figure 9b) perpendicular to the *X*-axis (33 positions in total). In each

position, the measurement took 30 s or 30 thousand samples was taken. In the *Z*-axis, the point 0 mm corresponds to the position above the spray. The radial velocity was then measured up to 70 mm from this point in a step of 10 mm (8 positions in total; Figure 9b).

**Figure 9.** A representation of the measured position in relation to the spray (**a**) with a detailed demonstration/description of the measured positions (**b**).

#### **4. Results and Discussion**

#### *4.1. Pressure Impact Footprint*

Data from measurements in individual planes are presented in Figure 10. In the 3D graphs of Figure 10, the axes are not equidistant and the 2D graphs show the Cartesian coordinate system from the view above, turned according to the nozzle position in the experiments. Therefore, the 2D graphs can be considered for the spray cones' baselines and serve mainly as a visualization of the spray impact footprint of the two inner cones (the upper cone was out of the scanned area). The values of the impact pressure are very low (0.61 kPa at the highest). The reason for this is the size of the impact area, which is quite large for such a fine spray. The pressure impact intensity decreases with increasing distance from the nozzle orifice from about 0.61 kPa in the plane 1 to less than half (0.28 kPa) in the plane 6, which corresponds to the distances of individual planes from the orifice (60 and 140 mm of the plane 1 and 6, respectively). The effective spraying angle of the inner cones was further evaluated. This was done only for the pressure of 2 bars because with increasing inlet pressure, the effective spraying angle varied very slightly. Figure 10d shows the maximum impact pressure in relation to set inlet pressure of the liquid at the nozzle. We can see a linear increase up to an inlet pressure of 1.5 bar, then the impact pressure increase slowed down. The same is shown in supplementary Figure S3, which compares the varying intensity of the impact pressure at a distance of 70 mm (plane 2) from the nozzle in relation to changing increasing inlet pressure.

**Figure 10.** Pressure distribution pattern of the sprayed liquid measured in the plane 1 (**a**), 3 (**b**), 6 (**c**), and a relationship between maximum impact pressure and inlet pressure of liquid at nozzle (**d**).

#### *4.2. Spray Angle*

The base of the inner cones can be expressed using the parametric equation of the Archimedean spiral as follows:

$$\mathbf{x} = \mathfrak{a}\theta\cos\theta + \mathcal{S}\_{\mathbf{x}\prime} \tag{22}$$

$$y = a\theta\sin\theta + S\_{y\prime} \tag{23}$$

where *x* and *y* are the points of the curve in the coordinate system, *a* is the parameter, *Sx* and *Sy* are the shifted centers of the spirals, and θ is the angle between half line describing the spiral trajectory and polar axis of the system. The values of the parameter *a* for the inner cones are shown in (Table S1). An example of the expression for plane 3 (Figure 11a) is for the small cone as follows:

$$
\lambda = 2.69 \cos \theta + 145,\tag{24}
$$

$$y = 2.69\sin\theta + 250.\tag{25}$$

In addition, for the large cone:

$$\mathbf{x} = \theta.0 \theta \cos \theta + 145,\tag{26}$$

$$y = 9.0 \theta \sin \theta + 250,\tag{27}$$

where for each plane θ ∈ (2.95π; 4.95π). The upper cone could not be evaluated using this method due to its large extent and was obtained using another method as explained further. The spiral curves delineate the trajectory of the impact footprint and well corresponds to experimental data (Figure 11a). An exception can be found in the fourth quadrant (beam number 5, Figure 11a). This is due to the nozzle geometry and transition between individual cones resulting in a local change of liquid flow.

**Figure 11.** Pressure impact footprint with numbered water jets (individual numbers relates to individual water beams, refer to Figure 4b) and a graphical depiction of the spirals tracing the cones' baselines used for spraying angle assessment (in plane 3) (**a**), a photograph of the spray with an image filter applied to determine the main liquid streams (**b**), a visualization of the spray with assessment of inner

angles formed by the liquid beam and nozzle longitudinal axis (**c**), and comparison of the spray formed at an inlet liquid pressure of 1 bar (**d**) and 2 bars (**e**).

The extent of the small cone angle was calculated as a sum of the inner angles of the beams no. 1 and 2 (refer to Figure 11a). The extent of the large cone was a sum of the beam no. 4 and an arithmetic average of beams no. 3 and 5. The inner angles of individual water beams were determined in relation to the nozzle longitudinal axis (the *Z*-axis) in the same position as in experiments (i.e., in the *XZ* plane (refer to Figure 3)). Based on the extent of water beams of both cones, the effective spraying angle γef was calculated. The average values of effective spraying angle were 61.7◦ and 115.5◦ for small and large cone, respectively, (for detailed results, refer to Table S2).

Another method to determine the spraying angle was using the photographs of the spray with applied screen filters to assess the main liquid beams (Figure 11b) and comparing with the photographs of the nozzle without spray (Figure 11c). Evaluation of the angles was done in the Stream Motion software (refer to Table S3). The small, large, and upper cone angles were 58.0◦, 111.4◦, and 152.3◦, respectively. Comparing the results of the angles for the small and large cone with those obtained with the previous method (Table S2), we get values, which are smaller by 6% and 3.5%, respectively. This difference is mainly caused by a subjective assessment of the 2D photograph of a 3D spray, which may bring an error into the determination of the spraying angle. Both discrepancies are; however, very small and can be neglected.

Comparing the results of the spray angle with other studies, one significant difference can be observed (i.e., an independence of the spray angle on the inlet pressure (flowrate) of the liquid; Figure 11d,e), refer also to (Figure S4). This is in agreement with a previous study on characterization of a spiral nozzle, in which authors observed minimal variation of the spray angle at pressures above 0.4 bar [4]. On the contrary, such behavior is quite different compared, for example, to an air–water impinging jet atomizer [81], in which the spray angle increased with increasing pressure, or pressure swirl atomizer, in which the spray angle increased up to 7 bar and then decreased [82]. Some researchers also observed a decrease in spray angle with increasing inlet pressure in the whole measured range [48]. Another typical feature of the nozzle studied is its large spray angle, thus large spray coverage. This can be observed in nozzles mostly used for fire suppression applications [36]. However, this property is also required in the applications related to gas cleaning via absorption processes, especially for gas scrubbers with larger spray tower diameter.

#### *4.3. Spray Breakup*

Here we discuss the breakup of the outer upper cone only (Figure 4a). The other two inner cones are not evaluated as it was not allowed by the experimental setup. Moreover, the results are compared with mathematical models assuming the analogy with radial spread of a liquid jet over a horizontal plane. The condition of the horizontal plane is fulfilled for the upper cone only (Figure 2), which is formed by impinging the liquid jet on the first twist of the helix, which is horizontal (refer to CFD model at cross-section, Figure 2). The other helix surfaces (second and third twist) have different rake angles and are not horizontal, so the model could not be applied to these.

Individual atomization phases can be observed in the pictures (Figure 12) (i.e., primary/secondary atomization in various distances from the nozzle). The water sheet remains continuous up to a distance of about 80 mm (Figure 12a), even though some local perforations can be observed. Some perforations can be observed throughout the whole length of the liquid sheet (Figure 12b 2). The jet Weber number was higher than 22,000, which is far higher than the critical value of 1000 proposed by other researchers [65,83]. At this Weber number, we can expect very short radial breakup distance and the sheet to thin rapidly. The sheet thickness was estimated using the Equation (6) to be 2.47 mm at the edge of the helix and gradually decreased with increasing radial distance from the nozzle. At a distance of 78 mm, corresponding to the first sheet breakup into ligaments, the sheet thickness was estimated to 0.33 mm. The estimated sheet thickness also decreased with increasing liquid flowrate. At a radial distance of about 80 mm the sheet integrity is markedly disturbed (i.e., a primary atomization occurs

up to 125 mm forming larger ligament structures; Figure 12b 1). The first ligaments observed (at the distance of 78 mm) were as large as (1.86 ± 0.57) mm with a breakup time estimate of 10.9 ms (according to Equation (15)) and attenuating down to (0.75 ± 0.19) mm with a breakup time of 2.8 ms (at a distance of approximately 120 mm). The estimate of the ligament diameter was 0.70 mm (based on Equation (13) using the correlation in Equation (12) for the dimensionless wave amplitude, at a radial distance of 120 mm). However, conducting experiments at different flowrates is necessary to assess further this comparison.

**Figure 12.** Individual phases of spray atomization (**a**), primary spray atomization to ligaments, 1–forming of ligaments, 2–liquid sheet rupture (**b**) and secondary atomization to individual droplets (**c**), and a detail of the boundary of both phases (**d**).

Subsequently a breakup to individual droplets (Figure 12c) occurred in a very narrow area. From this area the droplet-size distribution and structure has changed very slightly indicating area of secondary (or better quasi-secondary) atomization of the spray. A detail of the boundary between primary (sheet breakup into ligaments) and the quasi-secondary atomization of the spray (ligaments breaking into droplets) is shown in a modified photograph (Figure 12d). Generally, similar qualitative spray characteristics were very scarcely found in the literature. Qualitatively the most similar

sprays were observed in studies focusing on characterization of fire sprinklers, refer, for example, to [64,74,75,84,85]. These sprinklers consist of a convergent nozzle with a cone-disc deflector placed under in a defined distance to create a circular liquid sheet. This is very similar to the spray generated by the spiral nozzle as the principle of forming the liquid sheet is practically the same (i.e., a rebound from a surface into the ambient space).

The gas Weber number calculated for the measured droplet sizes and radial velocities was always lower than 11. This may indicate, according to some studies, no secondary breakup [86–88], droplet deformation [67], or vibrational breakup [66]. The latter two may be true as indicated in obtained photographs (Figure 12d, also refer to an example shown in Figure S5). This is; however, a mere qualitative observation of the spray. As the liquid is sprayed into a quiescent air, it is more probable that the secondary atomization did not occur at the adopted experimental conditions. It is also important to say that the Weber number ranges defining individual secondary breakup modes were mostly derived for liquid jets. In liquid sheets or even conical sheets, the situation can be different and is rather a suggestion for future research. In general, vibrational breakup is not always observed. Oscillations at a natural frequency of the drop are typical and a formation of only a few fragments with comparable sizes as the original droplet are observable at this mode [68].

#### *4.4. Droplet Size Distribution*

Droplet size distribution was obtained from the modified classical photography of the upper cone in various distances normal to the nozzle axis. From detailed photographs of the secondary atomization area, droplet size was measured using the Stream Motion software. Prior to the measurement, a filter was applied to the photographs to better recognize the droplet edges. Spherical droplets' diameter was measured once, whereas the diameter of droplets of non-spherical shape was measured twice (two diameters normal to each other, refer to Figure S6), and then an arithmetic average was calculated. Using this method, 1773 droplets' diameters were evaluated. Droplet-size distribution curves were then plotted using the obtained data (Figure 13). The width of one size class was 0.07 mm. This was obtained by dividing the whole distribution width by square root of the number of measured droplet sizes. The agreement between experimental and theoretical distribution was assessed using the Kolmogorov–Smirnov test, which revealed the best fit only for the log-normal distribution for which the *P*-values were higher than the selected significance level of 0.05 (refer to Table S4).

**Figure 13.** Number (**a**), surface (**b**), and volume weighted (**c**) droplet-size distribution fitted with selected theoretical distribution functions.

Droplet-size distribution of a spray can be considered monodisperse if the standard deviation is approximately less than 10% of the mean particle diameter [80]:

$$\frac{\overline{\sigma}\_{D}}{\overline{\mu}\_{D}} < 0.1.\tag{28}$$

In our case, according to the number droplet-size distribution of the spray, the ratio is as high as 0.48, confirming a polydisperse spray. The average droplet size was 0.815 mm, while the mode was around 0.587 mm with a portion of 8.7%. The surface-weighted average (*D*20) was 0.997 mm and the Sauter mean diameter (*D*32) was 1.201 mm. The largest amount of droplet surface (6.6%) was represented by a droplet diameter of 0.911 mm. For the volume-weighted distribution, the mean value is shifted to larger droplet diameters as large as 1.423 mm and a mode of 1.108 mm representing 5.8% of the liquid volume. The theoretical predictions of the droplet diameter were based on the theory of the liquid sheet instability, assuming an analogy with a rebound of liquid jet from a horizontal plane. A droplet diameter of 1.413 mm was calculated based on Equation (16), which is quite different from the average value obtained experimentally (0.815 mm), despite the expected analogy with the theory adopted for the diameter estimate. It is; therefore, necessary to conduct further research in this area as the description of conical sheets breakup mechanisms is rather scarce in the literature [89]. The measured droplet diameters are mostly much larger compared to different types of atomizers, which produce droplets in the range up to 100–150 μm in diameter, refer to, for example, [58,90–93], but in a similar order as compared to fire sprinklers [64,79,94]. Comparing the Sauter mean diameter empirical formulas developed for different atomizers we can also see rather significant differences. For example, the Sauter mean diameter for pressure-atomized sprays is a function of several parameters, as proposed in the correlation by Elkotb [95]:

$$D\_{32} = 3.08 \mu\_1^{0.385} (\sigma \rho\_1)^{0.737} \rho\_\text{g}^{0.06} \Delta p^{-0.54},\tag{29}$$

where Δ*p* is the pressure drop at the nozzle (i.e., the pressure difference between pressure inside the nozzle and the ambient pressure of the space, into which the liquid is sprayed). This formula gives the Sauter mean diameter as large as 1.844 mm, which is more than 0.6 mm larger than that obtained experimentally.

#### *4.5. Liquid Velocity Distribution*

Using the LDA method, radial and tangential velocities in the outer upper cone (Figure 4a) were obtained. 3D surface graphs of absolute radial velocity in relation to position were created. Data was processed into velocity distribution histograms of radial and tangential velocity component at individual points (Figure 14). The absolute radial velocity (*v*AR) was obtained through a vector sum of radial and tangential velocity at a given point. An example of the data processing into histograms for the position (140,0,40) mm (*x*,*y*,*z*) is shown in Figure 14. The velocities minus signs are due to the LDA experimental setup, which measured the radial velocities as negative. An average radial and tangential velocity at the point (140,0,40) mm was (7.78 <sup>±</sup> 2.03) m s−<sup>1</sup> and (0.17 <sup>±</sup> 1.28) m s−1, respectively. The vector sum of the both velocities (i.e., the representative absolute radial velocity of the liquid) is 7.78 m s−1. This means that the contribution of the tangential velocity was very small with a very narrow distribution. This is in contrast with a spiral flow nozzle with an annular slit [47] using which a tangential velocity up to 8 m/s was observed (in relation to slit width). Looking at the geometry of the spiral nozzle studied, more noticeable tangential velocity contribution was expected. This is due to the centripetally-oriented impact surfaces of the spiral causing torque of the liquid at the outlet. This would probably be more noticeable at lower liquid velocities, at which the liquid sprayed copies the trajectory demarcated by the spiral. However, this effect is rather limited at high velocities due to high impact force causing a strong rebound of the liquid from the surface in the radial direction, thus eliminating the formation of the torqued flow.

**Figure 14.** Droplet radial (**a**) and tangential (**b**) velocity distribution at a position (140,0,40) mm (*x,y,z*).

Using the treated data, 3D surface graphs of absolute radial velocity in relation to radial *X*and *Y*-position at different axial distances (*Z*) were plotted and approximated using the distance weighted least squares algorithm in the TIBCO Statistica (USA) software. The 3D graphs are shown from two different orientations, and blue marks represent individual points approximated by the plane. Figure 15a shows a decrease in velocity with increasing distance. At the *Z*-axis 0 and 10 mm, the velocity was measured right above the liquid sheet; therefore, the measured velocities are low. The highest velocity is at the *Z*-axis between 40 and 50 mm, where the main liquid stream occurs and then slightly decrease (the same in 2D is shown in supplementary Figure S7). The highest radial velocity of 8.26 m s−<sup>1</sup> was measured at a radial and axial distance of 110 and 40 mm, respectively. At the same axial distance was also the lowest measured velocity of 7.22 m s−1, but at the position corresponding to a radial distance of 165 mm (the most distant position from the nozzle axis). This is due to the kinetic energy loss in the space with increasing distance from the nozzle orifice. However, at the 0 and 10 mm Z-distance, there is an obvious decrease of the velocity profile, which is higher than could be expected from the loss of kinetic energy. This is probably caused by the large spray angle (the liquid in the upper cone is sprayed from the nozzle almost horizontally) and a descent of the liquid sheet due to gravity. In a distance between 110 and 165 mm, the descent by 15 mm can be expected. The kinetic energy loss (liquid flow deceleration) is also due to the friction caused by the complicated nozzle orifice geometry. This can be expressed using the Euler number (*Eu*) (i.e., the ratio of pressure loss due to flow restriction and kinetic energy per volume of the flow):

$$Eu = \frac{2\Delta p}{\rho \eta v^2},\tag{30}$$

where ν is the liquid velocity inside the nozzle. Figure 16 shows the relationship between Euler number and Reynolds number at nozzle orifice (*Re*o) of the flow inside the nozzle (Equation (7)).

**Figure 15.** Absolute radial velocity in relation to *X*-distance (**a**) and *Y*-distance (**b**).

**Figure 16.** Euler number in relation to Reynolds number at nozzle.

We can see an atypical behavior of the nozzle studied. The friction losses expressed in terms of *Eu* are quite low (ideal frictionless flow corresponds to *Eu* = 1). However, *Eu* first decreases with increasing *Re* from approx. 1.4 <sup>×</sup> 105 to 1.6 <sup>×</sup> 105. Such a high *Re* corresponds to fully developed turbulent flow and any further increase is expected to rather increase pressure losses due to friction. Nonetheless, the extraordinary nozzle orifice geometry probably causes this slight anomaly and up to the point corresponding to *Re* of about 1.6 <sup>×</sup> 105 the friction losses slightly decreased. This can outline ideal operation conditions for the use of this nozzle in gas cleaning applications. At lower *Re*

the friction is higher but the turbulence is lower at the same time causing non-ideal conditions for the removal of the contaminative gases from air. With further increase of *Re*, *Eu* starts to also increase up to 1.28 corresponding to an inlet pressure of 2 bars (i.e., a flowrate of 1.76 L/s; refer to Table 1). It is further suggested for future work to measure a wider range of inlet pressures to obtain more complex idea about the nozzle behavior. However, we were limited to a pressure of 3 bars at the highest due to the nozzle made of Teflon (for higher pressures, metal nozzles are necessary). The nozzle behavior described in Figure 16 is quite different compared to a hollow-cone pressure swirl nozzle as studied by Nonnenmacher and Piesche [96]. The authors observed higher values of *Eu* increasing approx. from 50 to 80 with *Re* increasing approx. from 300 to 50,000.

#### **5. Conclusions**

This work tried to provide basic characteristics of a spray produced using a spiral nozzle and compare the spray properties with other atomizers. The spray produced using the spiral nozzle consists of three concentric liquid cones creating a full-cone pattern. Typical spray angle is quite large compared to other nozzles, the upper large cone having a spray angle as large as 150◦. At the adopted conditions, the droplet-size distribution was larger. Sauter mean diameter was 1.201 mm, the same is true for velocity which was, conversely, lower compared to, for example, effervescent atomizers. Ideal operational conditions were found to be at a liquid flowrate of 1.41 L/s corresponding to an inlet pressure of 1.25 bar. Based on the initial comparison with the theory, we can see a gap in the mathematical description of the conical sheets. Therefore, future research is necessary to further understand the breakup behavior of the liquid jets on plates horizontal as well as at different angles. Future work will also focus on the description of the conical sheet breakup using detailed observations with optical methods, and basic RANS simulations (Reynolds-averaged Navier-Stokes equations) and large eddy simulations (LES) with adaptable computational mesh to catch as many of the smallest structures and droplets formed during the sheet atomization as possible. The presented results will also serve as the experimental verification of the developed CFD models.

**Supplementary Materials:** The following are available online at http://www.mdpi.com/2227-9717/7/12/911/ s1, Figure S1: Representation of the measured position in relation to the spray (a) and a detailed demonstration/description of the same (b); Figure S2: A snapshot of classical (a) and modified (b) photography, Figure S3: Impact pressures of individual measurements in plane 2; Table S1: The parameters of the Archimedean spiral equation; Table S2: Effective spraying angles γef of small and large cones for planes 1, 3, and 5; Table S3: Effective spraying angles γef of small and large cones for planes 1, 3, and 5; Figure S4: Comparison of the spray at two different flowrates of 1.25 l/s corresponding to a pressure at the nozzle of 1 bar (A) and 1.76 l/s corresponding the a pressure at the nozzle of 2 bars (B); Figure S5: Deformation (yellow-circled) and vibrational breakup (red-circled) of droplets; Table S4: Results of the Kolmogorov–Smirnov test; Figure S6: Detail of individual droplets (a) and evaluation of their size (b); Figure S7: Velocity profile of the liquid sheet in relation to radial distance *X* at individual axial positions *Z*

**Author Contributions:** O.K. conceived, designed, and performed the experiments and evaluated data; P.B. evaluated the data and wrote the paper; T.S. conceived the work, provided the laboratory with the equipment and supervision.

**Funding:** This paper has been supported by the project "Computer Simulations for Effective Low-Emission Energy Engineering" funded as project no. CZ.02.1.01/0.0/0.0/16\_026/0008392 by Operational Program Research, Development and Education, Priority axis 1: Strengthening capacity for high-quality research.

**Conflicts of Interest:** The authors declare no conflicts of interest.

#### **References**


© 2019 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

## *Review* **Carbon Mineralization by Reaction with Steel-Making Waste: A Review**

**Mohamed H. Ibrahim 1, Muftah H. El-Naas 1,\*, Abdelbaki Benamor 1, Saad S. Al-Sobhi <sup>2</sup> and Zhien Zhang 3,\***


Received: 9 February 2019; Accepted: 20 February 2019; Published: 24 February 2019

**Abstract:** Carbon capture and sequestration (CCS) is taking the lead as a means for mitigating climate change. It is considered a crucial bridging technology, enabling carbon dioxide (CO2) emissions from fossil fuels to be reduced while the energy transition to renewable sources is taking place. CCS includes a portfolio of technologies that can possibly capture vast amounts of CO2 per year. Mineral carbonation is evolving as a possible candidate to sequester CO2 from medium-sized emissions point sources. It is the only recognized form of permanent CO2 storage with no concerns regarding CO2 leakage. It is based on the principles of natural rock weathering, where the CO2 dissolved in rainwater reacts with alkaline rocks to form carbonate minerals. The active alkaline elements (Ca/Mg) are the fundamental reactants for mineral carbonation reaction. Although the reaction is thermodynamically favored, it takes place over a large time scale. The challenge of mineral carbonation is to offset this limitation by accelerating the carbonation reaction with minimal energy and feedstock consumption. Calcium and magnesium silicates are generally selected for carbonation due to their abundance in nature. Industrial waste residues emerge as an alternative source of carbonation minerals that have higher reactivity than natural minerals; they are also inexpensive and readily available in proximity to CO2 emitters. In addition, the environmental stability of the industrial waste is often enhanced as they undergo carbonation. Recently, direct mineral carbonation has been investigated significantly due to its applicability to CO2 capture and storage. This review outlines the main research work carried out over the last few years on direct mineral carbonation process utilizing steel-making waste, with emphasis on recent research achievements and potentials for future research.

**Keywords:** carbon capture; CO2 sequestration; steel-making waste; steel slag

#### **1. Introduction**

Fossil fuels are used as the main source of energy globally, and now they supply over 80% of the world energy demand [1]. Fossil fuels are expected to remain the most used energy source for years to come. This is due to the ever-increasing demand for energy created by the thriving economies around the globe. International Energy Agency reported a total energy demand of 574 exajoules globally in 2014 [2]. Although there are multiple sources for atmospheric CO2, human activities, such as transportation and electricity generation, which directly burn several kinds of fossil fuels (including coal, oil, and natural gas), release more CO2 into the atmosphere. This leads to increases in

the earth temperature and in turn causes global warming. Hence, mitigating CO2 emissions is a key to decrease global warming and sustain a better future for humanity [3]. Carbon capture and storage serves as the main technology for mitigating carbon emissions. Numerous conventional CO2 capture technologies based on a post-combustion approach are being used in the industry. Separation based methods, such as absorption, adsorption, and membrane separation, are the most utilized separation technologies available [4–7]. Mineral carbonation is one of few technologies that work as both capture and storage technologies [8]. It is based on the principles of natural rock weathering, where the CO2 dissolved in rainwater reacts with alkaline rocks to form carbonate minerals. The active alkaline elements (Ca/Mg) are the fundamental reactants for mineral carbonation reaction. Although the reaction is thermodynamically favored, it takes place over a large time scale. The challenge of mineral carbonation is to offset this limitation by accelerating the carbonation reaction with minimal energy and feedstock consumption. Calcium and magnesium silicates are generally selected for carbonation due to their abundance in nature. Industrial waste residues emerge as an alternative source of carbonation minerals that have higher reactivity than natural minerals; they are also inexpensive and readily available in proximity to CO2 emitters. In addition, the environmental stability of the industrial waste is often enhanced as they undergo carbonation. Recently, mineral carbonation has been investigated significantly, due to its applicability to CO2 capture and storage. Despite the growing interest in mineral carbonation research, there have not been any focused reviews that assess the status of CO2 sequestration using steel-making waste. In this review, mineral carbonation using steel-making waste is reviewed in the light of different process parameters and their effect on CO2 uptake. Potentials for future research in the area are highlighted.

#### *1.1. CO2 Storage*

Several storage techniques are used to store CO2, and the most feasible option to do so is geological sequestration [1]. Literature work investigating geological CO2 storage has seen a substantial increase in the last decade [2]. Practically, over 1 million ton CO2 is being sequestered in 14 individual different locations around the globe [3]. Estimates of CO2 storage capacity varies depending on the region at which the study has been conducted. Nonetheless, the capacity is in the range of 100–20,000 giga ton CO2 worldwide [4]. One of the mature CO2 storage techniques is to inject it into depleted gas or oil reservoirs. Carbon dioxide is used to increase reservoirs pressure to produce enough driving force to push the gas/oil out of it. In other words, it enhances oil recovery in active wells by extracting the residual oil left. Additionally, CO2 can be used to recover natural gas (methane, CH4) trapped in coal beds. The main premise behind the idea is that CH4 can be quickly displaced from coal by carbon dioxide injection, allowing CO2 to be stored in the porous structure of the coal bed [5]. Injecting CO2 into saline aquifers is also a viable option that commercially exists with an acceptable capacity [6,7]. Carbon dioxide is usually injected in its supercritical conditions [8]. At these conditions, CO2 is buoyant relative to porous rocks and saline aquifers. Thus, there is always a possibility that buoyant CO2 could leak to the surface and cause catastrophic environmental impacts. Most critically, monitoring programs for post-injection are limited and do not provide long-term detectability of the gas that can potentially escape from the storage medium [9]. Hence, these approaches cannot be taken for granted and considered as permanent and safe CO2 storage solutions. Mineral carbonation is one approach that can provide long-term storage solution in addition to being CO2 leak-free. This is due to the fact that carbonates are in a lower energy state than CO2 [10]. More importantly, it possesses extremely large sequestration capacity compared to other geological storage options, as indicated in Table 1.


**Table 1.** Storage capacities for several geological storage options [11].

<sup>a</sup> No specific number can be given, however, massive potential to sequester CO2 exists. <sup>b</sup> Including fields that are not economically viable to inject carbon dioxide into.

#### *1.2. Mineral Carbon Sequestration*

Mineral carbon sequestration is based on the principles of the natural carbonation process of natural rocks, where the CO2 dissolved in rainwater forms a weak carbonic acid. Consequently, alkali and alkaline earth metals (i.e., Ca and Mg) neutralize the acid to from insoluble carbonate minerals [12,13]. Sequestration happens in several alkaline minerals, such as calcite (CaCO3), dolomite (Ca/Mg(CO3)2), magnesite (MgCO3), siderite (FeCO3), and serpentine (Mg3Si2O5(OH)4) [14]. For mineral carbonation, having a sufficient amount of a certain natural mineral is an essential factor. Hence, magnesium-based silicates are utilized since they are available in considerable amounts globally [11]. However, increasing CO2 levels led to more CO2 absorption by the oceans, hence increasing its acidity by 30% since the industrial era started [15]. Hence, this limits the natural carbonation process. The formed carbonates are in solid form resulting from exothermic reaction (Equation (1)) and a certain amount of heat is released, depending on the type of metal oxide reacting.

$$\rm Mg\_3Si\_2O\_5(OH)\_4 + 3CO\_2 \rightarrow 3MgCO\_3 + 2SiO\_2 + 2H\_2O \ + 64 \ kJ/mol \tag{1}$$

Carbonates require a high amount of energy to decompose back into CO2. Hence, carbonates can be considered as thermodynamically stable CO2 sink [16]. CO2 will be fixed permanently without further monitoring to check its stability [17]. Table 2 shows the composition of different minerals rocks and the mass of CO2 that can be sequestered by a unit mineral mass (mass CO2/mass mineral). This ratio is based on the theoretical basis and considered as the maximum potential carbonation capacity for the specific mineral. Different minerals have different ratios according to their alkali metal content. Whether or not the maximum capacity can be reached, it is subject to the carbonation process and different operating parameters.

**Table 2.** Carbonation potential for different naturally occurring minerals [18].


Mineral sequestration technique was first proposed by Seiftriz [19]. The proposed idea suggested introducing high purity CO2 to accelerate the carbonation process. This ensured that carbonation time can be shortened from geological time scale to hours or minutes. Since then, the literature work has expanded greatly. However, it is clear that the research progress is facing challenges in enhancing the carbonation process to be viable to deploy on a large scale, as shall be demonstrated in the following sections. Nonetheless, the technique possesses several advantages over other sequestration techniques such as ocean and geologic sequestration, due to concerns over long term carbon leakage, as described previously [20]. Mineral carbonation produces more stable products that have the potential to be profitable and usually produced in fewer steps than other techniques. Additionally, the heat of the reaction can be further utilized as a source of energy.

Mineral sequestration techniques are often divided into in situ or ex situ manner. In situ sequestration requires injecting CO2 into underground reservoirs to start a reaction with the existing underground minerals to form carbonates. Ex situ sequestration is related to carbonation process above the ground, where the raw natural mineral needs to be mined and treated before it undergoes the carbonation process. The scope of this work focuses on ex situ mineral carbonation and its related mechanisms and applications.

Although naturally occurring minerals have the potential to sequester huge amounts of CO2 due to their abundance, it is not practically feasible due to the cost of extracting and pretreatment of the minerals and the impacts associated with it [21]. In addition to numerous process challenges in terms of carbonation efficiency and energy intensity (temperature and pressure). Alkaline industrial waste rich with Mg2+ and Ca2+ is an attractive alternative for CO2 sequestration. It can be used to imitate mineral carbonation without the additional mining cost associated with natural rocks. Even so, alkaline waste is available in less amounts than natural minerals. It is available at lower cost, higher reactivity, and uptake capacity, and less pretreatment is required. Table 3 summarizes the most studied industrial alkaline waste and their alkali earth metal composition (Mg and Ca), in addition to their production rate per year and CO2 emissions associated with their production. Examples of the industrial wastes include fly ash, such as coal and shale oil ashes, cement industry waste dust, and steel slag.


**Table 3.** Alkaline solids studied in the literature, their composition and global production per year and the CO2 emission.

<sup>a</sup> Retrieved from [10]; <sup>b</sup> Retrieved from [38].

Globally, cement industries account for 5% of the total CO2 emissions [39,40]. Furthermore, the steel industry accounts for 7% of CO2 emissions globally [41]. There are four main types of steelmaking slags, including blast furnace (BF), basic oxygen furnace (BOF), electric arc furnace (EAF), and ladle furnace (LF) slags. The slags consist of several oxides, primarily calcium, iron, and magnesium oxides that are present in different phases. On average, manufacturing 1 ton of steel produces approximately 420 kg of BOF and 180 kg of EAF [42]. There are two main approaches for ex situ alkaline waste carbonation: direct and indirect; each one has several sub-classifications based on the carbonation technique.

Direct carbonation technique implies that the carbonation process happens in one single step. On the other hand, indirect carbonation has more than one step (two or more) that usually involves pre-treatment of used minerals. Typically, mineral ores undergo a pre-treatment process where the reactive chemical components (i.e., alkali earth metals) in the rocks are separated for the mineral core. Pre-treatment usually involves mining, grinding, and activation of the rock minerals (Table 4). The end product of the pre-treatment process is almost pure carbonate form of the mineral. Then, the mineral carbonate is reacted with carbon dioxide in a separate step [12]. Table 5 shows the CO2 uptake capacity using natural rocks. It lists numerous studies that investigated the carbonation efficiency of minerals using natural minerals (i.e., olivine and serpentine) by dry and aqueous carbonation routes. The table includes details about the process parameters and carbonation conversion.


#### **Table 4.** Description of pretreatment process.

#### *1.3. Indirect Carbonation*

The term indirect mineral carbonation refers to the carbonation process that happens in two or more stages. Extraction of the reactive elements (Mg and Ca) form the mineral solid matrix is an essential step in indirect carbonation. Typically, strong acids are used as extracting agents. The extracted mineral then undergoes the carbonation process by reacting with CO2. One of the main advantages of using extraction is that it allows the production of almost pure mineral, as other impurities available in the natural mineral core can be removed after the reactive metal is extracted [50]. Numerous methods exist that can achieve mineral extraction, such as using acids, molten salts, caustic soda, and bioleaching. Table 6 lists all extraction methods studied in the literature with their associated extraction reactions. Every extraction technique possesses its intrinsic advantage and disadvantages. For example, using HCl produces pure alkali earth metal; however, it is significantly energy intensive if the recovery of HCl is required. Inversely, molten salts are less energy intensive in terms of regeneration. Nonetheless, molten salts are more corrosive compared to HCl [12]. Another approach of indirect carbonation is pH swing. pH swing refers to the extraction of mineral carbonate from the solid matrix at low pH condition in the first stage. In a second stage, the pH of the extraction solution is raised to improve carbonate formation [51]. Although acids are able to extract significant amounts of calcium and magnesium ions from the feedstock, as explained previously, pH plays a great

#### *Processes* **2019**, *7*, 115

role in the precipitation of calcium and magnesium carbonates. Therefore, increasing the solution pH to approximately 10 in the second stage of mineral extracting helps to increase the rate of carbonate precipitation [10]. Hence, due to the added cost of implementing a second process step and the extensive use of extraction agents, direct carbonation is a more practical mineral carbonation option. In this review, direct aqueous carbonation using steel-making waste is being reviewed.





#### **2. Direct Carbonation**

Direct carbonation includes the reaction of CO2 with a suitable feedstock or Calcium/Magnesium rich solid residue in a single step. It is relatively easier to implement compared to indirect carbonation. Hence, it has the potential to be used in industrial scale. The following sections explain the working principles of direct carbonation, discussing the operational parameters that have the most impact on the carbonation capacity.

#### *2.1. Gas-Solid Carbonation*

The reaction of gaseous CO2 with solid minerals is the most basic and straightforward approach of direct carbonation, first studied by Lackner et al [52]. In the case of olivine carbonation:

$$\text{Mg}\_2\text{SiO}\_{4(s)} + 2\text{ CO}\_{2(g)} \to \text{MgCO}\_{3(s)} + \text{SiO}\_{2(s)}\tag{2}$$

The reaction suffers from very slow reaction rates. Hence, it is usually carried over an elevated temperature and pressure. However, due to thermodynamics limitations, the temperature is restricted between 170–400 ◦C for most natural minerals, as equilibrium shifts to the reactant side with increasing temperature [52]. Hence, the maximum carbonation temperature is a function of CO2 partial pressure and the specific mineral used (Table 7). The process high temperature requirement can be further utilized to generate steam that will be used to produce electricity [12]. Nevertheless, the process slow kinetics is still the main obstacle hindering further progress even at high temperature and pressure.

**Table 7.** Maximum carbonation temperature for several minerals at CO2 partial pressure of 1 bar [59].


Thus, the experimentally obtained carbonation rate of direct dry rock minerals carbonation is insignificant, even at elevated temperature and pressure [12]. Kwon et al. [60] reported that introducing moisture to the flue gas in dry carbonation process can increase the carbonation rate significantly. However, due to low CO2 sequestering efficiency, 8 tons of olivine would be required to capture 1 ton of CO2. This renders the process practically not viable and reduces its wide scale applicability. Hence, focus shifted on dry carbonation of pure magnesium and calcium oxides [52]. Nevertheless, the limited availability of the used minerals hindered the research progress. However, alkali earth metals can be extracted from the mineral rocks and industrial wastes. This adds another process step (extraction) to the overall process scheme. Hence, the process becomes an indirect carbonation process, which is discussed previously.

#### *2.2. Direct Aqueous Carbonation*

As explained in the introduction section, natural carbonation occurs when CO2 is dissolved in rain water according the following equation:

$$\text{HCO}\_2\text{ (g)} + \text{H}\_2\text{O}\_{\text{(l)}} \leftrightarrow \text{HCO}\_{3\text{(aq)}}^- + \text{H}\_{\text{(aq)}}^+ \leftrightarrow \text{CO}\_{3\text{(aq)}}^{2-} + 2\text{H}\_{\text{(aq)}}^+\tag{3}$$

The aqueous solution becomes more acidic due to the presence of protons (H <sup>+</sup> (aq) ) resulting from CO2 solubility in water. Hence, by imitating natural carbonation, carbonation of natural minerals could take place in aqueous media in a single stage process. When the rock mineral is placed in aqueous solution, calcium or magnesium element in the solid matrix leaches from mineral ore according to Equation (4):

$$\rm{Ca(Ca/Mg)SiO\_4} + 2H^+ \rightarrow \rm{(Ca^{2+}/Mg^{2+})} + \rm{SiO\_2} + \rm{H\_2O} \tag{4}$$

Eventually, mineral carbonate is formed when the mineral reacts with dissolved CO2 (bicarbonate proton)

$$\left(\mathrm{Ca^{2+}/Mg^{2+}}\right) + \mathrm{HCO\_{3(aq)}^{-}} \rightarrow \left(\mathrm{Ca/Mg}\right)\mathrm{CO\_{3}} + \mathrm{H\_{2}O} \tag{5}$$

When studying direct aqueous carbonation, a good practice is to study the process parameters, such as temperature, pressure, and solution medium, that can maximize CO2 uptake capacity. For aqueous carbonation specifically, increasing process pressure enhances CO2 solubility in the aqueous medium. Therefore, according to Equation (5), the reaction will shift towards the products side, which is highly desirable. On the other hand, increasing the reaction rates can be enhanced by elevating temperature. However, this applies to a certain extent due to the decline in CO2 solubility in the solution with increasing temperature. In other words, CO2 solubility in a certain solution dictates the upper limit at which the process temperature can be elevated. Carbonation conversion is a way to measure the carbonation efficiency and is defined according to the following equation:

$$
\eta\_{(\text{Carbonation})}\%=\frac{\text{Quantity of Mg or Ca converted to carbonate}}{\text{Quantity of Mg or Ca available in mineral}}\times 100\tag{6}
$$

Hence, the efficiency is reported on the basis of magnesium and calcium content of the mineral, not the total quantity of the used mineral.

Studies investigating direct aqueous carbonation reaction mechanism revealed that aqueous carbonation proceeds in two distinct steps, as opposed to 1 step in dry carbonation [61]:


Typically, leaching of alkali metal into aqueous solution is the rate limiting step. Nonetheless, altering the process parameters, such as temperature and pressure, can make the carbonation step the rate limiting step [62]. This is explained by the formation of the carbonation products on the surface of the minerals that will increase the mass transfer resistance between the dissolved CO2 and mineral core. Hence, controlling the dissolution rate and finding the best process parameters is a must to ensure sufficient carbonation efficiency. Different minerals have different dissolution rates [63]. The rate depends on the morphology of the mineral (surface area and structure) [64]. Thus, pre-treatment techniques stated in Table 4 can be used to enhance the dissolution rate, hence, the carbonation efficiency.

#### **3. Steelmaking Waste Mineral Carbonation**

Solid industrial wastes are generally alkaline and rich in Ca/Mg and can therefore be applied as an additional feedstock for mineral CO2 sequestration. The main advantages of industrial waste are that they are available at low to no costs in proximity to industrial emitters, almost no pre-treatment is needed, and they are more reactive in less energy intensive conditions. In addition, the end product of the sequestration can be used in several applications, i.e., as a construction material and in fertilizers. The fundamental working principles for mineral CO2 sequestration apply for industrial waste in the same way. In fact, the major elements of e.g., steel slag (Mg, Ca, Si, and Fe) are present in a comparable concentration as in natural rocks. However, trace metals and soluble salt concentrations are available in more quantities compared to the average composition of natural rocks. Thus, steel industry waste can undergo the same direct and indirect carbonation techniques previously explained. Presently, the research is going towards optimizing the uptake capacity of CO2 by modifying the

operating parameters including pressure, temperature, liquid-to-solid ratio, CO2 gas flowrate, solid particle size, and pretreatment. Table 8 presents a summary of steel making waste which have been tested as mineral carbonation in terms of feedstocks; feed composition, the experimental CO2 capture capacity, and the different process conditions were investigated. The mineral carbonation uptake is a function of process temperature, CO2 partial pressure, and steel waste surface area, which affect the carbon dioxide dissolution rate, the diffusion rate of ions through the reaction with steel slag. The pH value is an additional essential parameter in mineral carbonation process. Optimum pH for aqueous carbonation is achieved at pH of 10 [65]. pH of the process influences the carbonation reaction, as the reaction is more favorable in alkaline mediums. In addition, the pH decreases continuously as carbonation, due to CO2 being dissociated into the solution. Eventually, the pH value remains unchanged at around 7 after the carbonation process ends. This signifies that the mineral carbonation process will not proceed in acidic mediums. Figure 1 summarizes the different aspects that affect the ex-situ mineral carbonation process, such as different reactor types and process parameters.

#### *3.1. Temperature and Particle Size*

Huijgen et al. [21] were among the first to utilized steel slag as feed stock for mineral carbonation. The authors studied parameters that could affect the carbonation rate, which include reaction temperature and steel slag particle size. An autoclave reactor was used to carry out the reactor, and a 450 mL of the slurry was used with a liquid to solid ratio of 20 kg/kg. A maximum conversion of 70% of the calcium in the feed stock was carbonated at a pressure of 19 bar and temperature of 100 ◦C was achieved. The authors reported that at higher temperatures leaching of calcium from steel slag components will proceed faster, hence increasing the reaction rate, but the solubility of CO2 in the solution decreases. This was also observed by Han et al. [65]. To achieve this carbonation percentage, the particle size was reduced from <2 mm to <38 μm. Reducing the particle size will produce more surface area for the carbonation reaction to occur, hence increasing the conversion. Particle size and specific surface area are among the most important factors affecting the dissolution kinetics of any kind of material. Mineral particle size determines its reactive surface area in addition to its leaching mechanisms. Typically, grinding is used to achieve a specific particle size. However, it is an energy intensive process. Hence, determining the optimal particles size will help in reducing the process cost in addition to increasing its efficiency. Baciocchi et al. [43] reported that the parameter that most affected the CO2 uptake of the slag was particle size, especially the specific surface of the particles. An increase in temperature also had a positive effect, achieving a maximum uptake of 130 g CO2/kg slag. The authors reported that an average particle size of less than 150 micrometers is considered as optimum.

#### *3.2. Liquid to Solid Ratio*

Liquid to solid ratio is defined by the amount of steel slag that is being utilized in a certain amount of aqueous medium (mass/mass). Revathy et al. reported that the carbonation efficiency increased when the S/L ratio decreased. The results indicate that when L/S is increased from 5 to 10 g/g, the carbonation degree of steel slag also increases. A further L/S increase causes a decrease in the carbonation degree of steel slag. This is caused due to the presence of extra liquid that leads to dilution of calcium ion concentration in the aqueous medium [66]. In a similar manner, the sequestration capacity of slag water slurry increased with the L/S ratio from 2 to 10, after which it decreased. This is due to the fact that a high amount of water inside the reactor causes blocking in the diffusion of gas molecules in the slurry [67].









 summary.

#### *3.3. Pressure*

At constant temperature, CO2 gas solubility increases along with pressure according to Henry's law. Hence, CO2 molecules that are involved in the carbonation process will be more as the pressure is elevated. The effect of pressure on CO2 uptake was tested at 10, 50, 100, and 150 bar under the same condition (50 ◦C, L/S = 1) by Han et al. [65]. The carbonation conversion was found to be 21% and 50.2% for 10 and 150 bar, respectively. Similarly, Ghacham et al. [68] reported that a higher CO2 partial pressure caused more CO2 to be soluble in aqueous medium, forming carbonic acid and consequently increasing the bicarbonate ions formation. Therefore, more bicarbonates will react with calcium ions. Hence, higher CO2 uptake. Fagerlund et al. [55] reported that that the carbonation rate and degree might increase exponentially with time, as long as a high enough CO2 pressure could be maintained. Additionally, high pressure will cause the reaction time to be shorter, hence, having lower carbonation time. Similarly, Eloneva et al. [69] reported shorter reaction times as the partial pressure of CO2 is increased.

#### **4. Summary and Future Prospective**

Carbon capture and sequestration can be achieved through different techniques that have the potential to capture substantial amounts of CO2 and help reduce its emissions. Mineral carbonation is evolving as a possible candidate to sequester CO2 from medium-sized emissions point sources. The process of natural carbonation forms the basis of mineral carbonation process. Active alkaline elements (Ca and Mg) are the fundamental reactants for mineral carbonation reaction. Industrial alkaline wastes, such as steel-making waste, are rich with these alkaline compounds, especially calcium and magnesium oxides. Hence, they are studied in the literature as a possible mineral carbonation process feedstock. Several parameters govern the carbonation process, including process temperature, pressure, and liquid to solid ratio. There is still a room for improvement by targeting higher CO2 uptake value. This can be achieved by using a different aqueous medium to carry out the carbonation process, i.e., reject brine and the development of reactor systems that minimize mass transfer limitations. Optimizing the interactions between process parameters, such as the interplay between temperature, pressure, and the degree of mixing, will contribute to the carbonation process. In addition, studying the adsorption behavior of CO2 on other elements, such as iron oxide, will give more insights into increasing CO2 uptake.

**Author Contributions:** Conceptualization, M.H.E.-N. and M.H.I.; Formal Analysis, M.H.I. and A.B.; Investigation, A.B., M.H.E.-N. and M.H.I.; Resources, S.S.A. and Z.Z.; Writing-Original Draft Preparation, M.H.E.-N. and M.H.I.

**Funding:** This research received no external funding.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


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