# Optimum Design of High-Temperature Steam Generator for Hydrogen Production Enhancement

Authors: Sunhee Oh; Seong-Ryong Park; Yong Tae Kang
Journal: International Journal of Energy Research
 (2023)

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## Abstract

In this study, an optimal design was proposed for the pressure drop of a high-temperature steam generator to enhance hydrogen production in a solid oxide electrolyzer cell (SOEC). The high-temperature steam generator (HTSG) was analyzed under experimental conditions of a steam temperature of 773 K, steam inlet pressure of 300–600 kPa, and steam flow rate of 10–35 kg/h. The pressure drop characteristics of the HTSG were analyzed under various variable conditions via a computational fluid dynamics simulation. The pressure drop according to the variable change was ranged from a minimum of 54 kPa (inlet nozzle angle: 25°) to a maximum of 309 kPa (body ratio: 0.63), and the valid range was28,700<Re<306,600. The hydrogen production capacity that can be produced utilizing the optimally designed HTSG was predicted to be up to 35.7 tons annually. In addition, CO2 emissions during hydrogen production under optimal conditions are 9.92 ton-CO2/ton-hydrogen, which can be reduced by up to 86.8% compared to the conventional high-temperature and high-pressure (HTHP) steam, thereby enhancing hydrogen production and energy harvesting.


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## Body

## 1. Introduction

The utilization of fossil fuels increases CO2 emissions and is emerging as a cause of air pollution [1, 2]. It is stipulated that 30 TW of zero CO2 electricity production is required by 2050 to keep the CO2 level in the atmosphere at an acceptable level [3, 4], and renewable energy production has arrived at a point where it is not an inevitable option but a necessity [5]. Each country has different opinions about reaching zero CO2 emissions by 2050, but to slow warming and maintain the stabilization trend [6], a system that equates energy production and consumption should be introduced [7].Hydrogen is a renewable energy source that can contribute to energy conversion and is considered to quench the thirst for energy consumption. In the past, hydrogen was produced by reforming fossil fuel; however, CO2 emissions were significant during the reforming process, thereby causing air pollution [8]. Therefore, research on environmentally friendly hydrogen production has recently been conducted. Solid oxide electrolyzer cells (SOEC) produce hydrogen and electricity utilizing water electrolysis and are very suitable for environment-friendly methods [9]. Although the proportion of hydrogen produced employing this method is negligible, it is already being practically utilized in the industrial field, garnering attention as a zero CO2 technology [10].In particular, for this technology to be practically utilized in other industries apart from specific industrial fields, it is necessary to develop a technology suitable for the operating conditions of a specific section required by the supplier. Kazepoor and Braun [11] studied the effects of operating variables on the performance of SOEC. To date, the operation section of the SOEC is not specified, and operating conditions may alter according to the needs of the demand institute. Therefore, there is a need to develop a steam generator suitable for operating conditions and precise control technology using the device.According to the previous research, SOEC operates under specific conditions with high-temperature and low-pressure (HTLP) steam [12, 13]. A temperature above 773 K is required. As the steam temperature increases, an endothermic reaction occurs and enthalpy increases [14]. In this process, electric energy demand is reduced, and hydrogen conversion is possible with less electric energy. Therefore, the electrolysis process using high-temperature steam is relatively more efficient than low-temperature (LT) steam and is suitable for utilization in hydrogen production [15]. However, in the case of pressure, low-pressure (LP) steam supply is essential because stacks that can withstand high pressure have a large risk of commercialization [16]. Chen et al. [17] performed numerical modeling in the pressure range of 100–500 kPa adopted for SOEC and obtained the result that 300 kPa is the optimal operating pressure. Accordingly, supplying an appropriate level of pressure and temperature to the steam affects significantly the system performance. In particular, Henke et al. [18] have proposed a theoretical model to increase the commercial effectiveness of the technology applying HTLP steam to SOEC and emphasized that the reference condition of the appropriate model should be selected.Furthermore, a steam generator that generates HTLP steam is the first factor required to ensure stable hydrogen production. Nevertheless, research on deriving design factors and optimizing devices based on experiments has not been properly conducted. In this study, the development of an original device suitable for HTLP steam generation is conducted as a more important point, and its importance is discussed. A high-temperature steam generator (HTSG) is developed for the specific operating conditions required by the SOEC. A lab-scale system applied with HTSG was built, and the linkage with SOEC was strengthened by producing HTLP steam. In addition, optimization of the HTSG shape via computational fluid dynamics (CFD) simulation was conducted to ensure that the user could consider the possibility of altering the operating conditions at the design stage of the system. Using the proposed characteristics of the HTSG, optimum conditions for hydrogen production are predicted, and the effect on SOEC performance was analyzed.


## 2. Experimental Study

### 2.1. Experimental Setup

#### 2.1.1. Model Design

The conventional steam generator generally utilized in the field is a helical coiled heat exchanger (HCHE). The HCHE has a large heat transfer area and high thermal conductivity; however, owing to its thin thickness, it is vulnerable to corrosion, erosion, and cracking when operated at high temperatures for a long time, which deteriorates the durability of the system [19]. Above all, it is difficult to generate HTLP steam because the pressure drop is small, and it is not suitable for utilization for SOEC that requires LP steam [20]. In addition, the device has an unnecessarily large surface area, and when installed outside in winter, the device is highly susceptible to the seasonal changes. In this study, a suitable form of the HTSG is developed to generate the HTLP steam. A cyclone cylindrical type is selected for the HTSG, as illustrated in Figure 1. The thickness is over five times that of a general HCHE, and it can withstand the high temperature. The HTSG is designed to generate a required pressure drop based on the effect of each section.Figure 1
Geometric model of high-temperature steam generator for pressure drop analysis.Although the external structure of the HTSG is simple, a complex three-dimensional turbulence occurs inside, and a pressure drop occurs owing to steam flow. In this study, this phenomenon was adopted to produce LP steam [21]. The cylindrical shape is generally utilized as a separator and has an excellent function in strengthening the internal pressure drop [22]. In addition, the inlet nozzle section is configured in the tangential direction as much as possible to maximize the centrifugal force generated when steam enters the HTSG inlet [23]. As the shape alters, the number of geometric vortices has a strong influence on the internal flow characteristics [24], and in particular, the streamline length according to the HTSG body diameter is an essential factor to consider in terms of performance improvement for LP steam production [25, 26].In the developed HTSG, there are three sections in which the pressure drop occurs, and it is analyzed based on (1) sudden expansion (∆Pse) from the steam inlet, (2) swirl flow (∆Psf) on the wall of the HTSG, and (3) sudden contraction (∆Psf) at the steam outlet. In addition, the steam inlet has characteristics of direction and angle. The steam inlet direction is selected to flow tangentially, and as steam flows through the inside wall, it is effective for the heat transfer to and from the wall, as illustrated in Figure 1(b). The inlet angle (φ) is selected as 5°, and the vortex generated on the inside wall is expanded [27].The vortex flow increases the streamline length of the steam, resulting in a pressure drop. Therefore, it is more efficient for generating LP steam compared to the HCHE, where the pressure drop due to change in steam flow rate hardly occurs. In addition, to maximize the swirl flow, the steam inlet velocity is designed to be less than 100 m/s [28], and the steam inlet diameter is 8 mm using Equation (1), where ṁ is the steam flow rate, ρ is the density of the steam, and a is the inlet area of the HTSG.
(1)Vi=ṁρa.The magnitude of the centrifugal force (Fc) applied to the steam increases as the diameter of the HTSG decreases, which is proportional to the speed of the streamline. The HTSG body diameter (Db) and length (Lb) are 350 mm and 550 mm, respectively. The steam outlet diameter (Do) is 17.3 mm and is designed to maintain 60% of the inlet velocity.


#### 2.1.2. System Apparatus

In Section2.1.1, the HTSG was designed considering the characteristics of pressure drop, and it is intended to verify the device by applying it to generate HTLP steam. In this chapter, a lab-scale system to which HTSG is fabricated via design is constructed, as illustrated in Figure 2.Figure 2
Schematic diagram HTSG system for pressure drop analysis.The main equipment of the HTSG system comprises a previously designed HTSG, steam boiler, and combustion furnace, as illustrated in Figure3. First (low-temperature and low-pressure) steam is generated in a steam boiler in a superheated vapor state. The first steam was maintained at the value set in the steam boiler, without pressure control. The operating range of steam that can be produced in the steam boiler is 100–1,000 kPa, and the steam temperature is generated in a superheated steam state of pressure.Figure 3
Experimental system for pressure drop analysis.Variations in the primary steam cause fluctuations in the entire system and are not suitable because uniform steam must be supplied to the SOEC; therefore, a pressure-reducing valve (PRV) is necessary. The PRV maintains the stability of the primary steam, and the secondary steam from the PRV flows uniformly into the HTSG inlet. The combustion furnace is maintained by supplying liquified natural gas (LNG) heat, and an electronic volume compensator for gas flow measurements is installed. The low calorific value of LNG is 26 MJ/Nm3. It supplies gas heat to the HTSG installed inside, and the gas flow rate is kept constant at 5.5 m3/h. The LTLP steam inside the HTSG is heat exchanged with gas heat at a temperature of 773 K or higher to generate HTLP steam applicable to SOEC. The operating conditions are presented in Table 1. The HTLP steam generated in this system is discharged through the steam outlet and can be provided in linkage with the SOEC.Table 1
The operating conditions of the HTSG system.
Systematic characteristicsOperating conditionsSteam inlet condition (kPa)Saturated vapor state (300–600)Steam inlet flow rate (kg/h)10–35Steam outlet temperature (K)773Combustion gas temperature (K)973


#### 2.1.3. Data Reduction and Uncertainty Analysis

The system was operated for 11 h during the experiment and maintained in a steady state for approximately half a day after the initial operation of the system. The pressure drop occurring in the HTSG was evaluated using Equation (2). In particular, in pipe friction caused by sudden contraction, the pressure drop is approximately negligible, equivalent to 1.19% of the swirl flow [27].
(2)∆Pt=∆Pse+∆Psf+∆Psc.The error of the pressure drop derived from the correlation and experimental data was evaluated using Equation (3), where the subscript “pre” and “exp” indicate the prediction and experiment, respectively.
(3)Error=ΔPpreΔPexp.The system error is derived based on the method of Abernethy et al. [29] and is expressed in Equations (4)–(7). Uc denotes the experimental error, A is the bias error, B is the random standard error, and σ is the standard deviation of the experimental values (x), including the temperature, pressure, and mass flow rate.
(4)Uc=4∙∑n=1N∂X∂xn∙BPn2+∑n=1N∂X∂xn∙APn2,(5)BP=σN,(6)X=Xx1,x2,x3,⋯,xn,(7)σ=1N−1∑n=1NXn−X¯2.The specifications and accuracies of the measuring devices in the system are presented in Table2, and the errors for each parameter are summarized in Table 3. The experimental uncertainty in the pressure drop was estimated to be 9.49%.Table 2
The specifications of the measuring devices.
ParameterTypeOperating rangeAccuracySteam temperature (K)K-type-473–1,523±0.02Steam pressure (kPa)Pressure transducer0–1,000±4.71Steam flow rate (kg/h)Volumetric flowmeter10.0–47.6±0.35Gas flow rate (m3/h)Electronic volume corrector95–300±5.17Table 3
Experimental uncertainty for HTLP steam generation.
ParametersA (%)B (%)Steam temperature0.01480.61Gas temperature0.01522.62Pressure1.560.056Steam flow rate1.751.26Gas flow rate1.721.48


### 2.2. Experimental Results

The performance of the developed HTSG was evaluated under the influence of geometric conditions, and the pressure drop effect occurred, as illustrated in Figure4. When the steam flow rate of the experiment was converted to speed, the pressure drop increased proportionally according to the local velocity under all pressure conditions. This is an obvious result that can be interpreted as local velocity having a dominant effect on the HTSG. Figure 5 illustrates the friction factor based on the pressure drop data obtained from the experiment. ∆Pse and fse for sudden expansion were calculated using Equations (8) and (9), respectively, and Vin was the steam velocity in the inlet nozzle section. For the steam properties, the average of the inlet and outlet values of the HTSG was adopted, and B was the ratio of the cross-sectional area.
(8)ΔPse=fse·ρ·Vi22,(9)fse=1−1B2.Figure 4
Pressure drop of HTSG system according to operating conditions.Figure 5
Friction factor for swirl flow in variation of inlet steam pressure.∆Psc and fsc for sudden contraction were calculated using Equations (10) and (11), where Vout represents the steam velocity in the inlet nozzle section.
(10)ΔPsc=fSC·ρ·Vo22,(11)fsc=0.45·1−1B.ΔPsf and fsf for the swirl flow are calculated using Equations (12) and (13), where Vh represents the local velocity in the HTSG. The main friction coefficient (f) was derived from the Blasius correlation in Equation (14). It was adopted when the Reynolds number (Re) in the experiment was 104 scale. The length of the steam stream inside the HTSG was estimated based on the inlet nozzle angle (φ) and steam flow in the experiment. k is calculated by Equation (15) from the pressure drop data obtained from the experiment, and the maximum error of the model prediction was 11.5%.
(12)ΔPsf=f∙LD∙ρ∙Vh22+fsf∙ρ∙Vin22,(13)fsf=πLD∙tanφ∙k,(14)f=0.3164Re0.25,(15)k=2.428Vi−0.00248.Similar results were obtained, regardless of the variation in inlet steam pressure. However, as the local velocity of the inlet steam increased, the friction factor decreased because the steam density and steam viscosity under each operating condition were different. In particular, when the local velocity was 150 m/s or higher in the low-pressure condition, the friction factor increased considerably. This phenomenon is interpreted as the system becoming unstable because the energy loss increased as the steam velocity approached the compressibility. Definitely, there are advantages in terms of hydrogen production and system stability at low pressure, but the energy loss in the system due to the increase in pressure drop also increased. Therefore, for the inlet condition of the HTSG, 600 kPa, which can supply steam constantly without affecting system stability, was selected as an appropriate value. A strategy is required to strengthen the pressure drop inside the HTSG to generate steam at 300 kPa. The optimization of the HTSG shape is an important task to predict the steam pressure, and specific design factors affecting the pressure drop can be derived via simulations based on experimental values.


## 2.1. Experimental Setup

### 2.1.1. Model Design

The conventional steam generator generally utilized in the field is a helical coiled heat exchanger (HCHE). The HCHE has a large heat transfer area and high thermal conductivity; however, owing to its thin thickness, it is vulnerable to corrosion, erosion, and cracking when operated at high temperatures for a long time, which deteriorates the durability of the system [19]. Above all, it is difficult to generate HTLP steam because the pressure drop is small, and it is not suitable for utilization for SOEC that requires LP steam [20]. In addition, the device has an unnecessarily large surface area, and when installed outside in winter, the device is highly susceptible to the seasonal changes. In this study, a suitable form of the HTSG is developed to generate the HTLP steam. A cyclone cylindrical type is selected for the HTSG, as illustrated in Figure 1. The thickness is over five times that of a general HCHE, and it can withstand the high temperature. The HTSG is designed to generate a required pressure drop based on the effect of each section.Figure 1
Geometric model of high-temperature steam generator for pressure drop analysis.Although the external structure of the HTSG is simple, a complex three-dimensional turbulence occurs inside, and a pressure drop occurs owing to steam flow. In this study, this phenomenon was adopted to produce LP steam [21]. The cylindrical shape is generally utilized as a separator and has an excellent function in strengthening the internal pressure drop [22]. In addition, the inlet nozzle section is configured in the tangential direction as much as possible to maximize the centrifugal force generated when steam enters the HTSG inlet [23]. As the shape alters, the number of geometric vortices has a strong influence on the internal flow characteristics [24], and in particular, the streamline length according to the HTSG body diameter is an essential factor to consider in terms of performance improvement for LP steam production [25, 26].In the developed HTSG, there are three sections in which the pressure drop occurs, and it is analyzed based on (1) sudden expansion (∆Pse) from the steam inlet, (2) swirl flow (∆Psf) on the wall of the HTSG, and (3) sudden contraction (∆Psf) at the steam outlet. In addition, the steam inlet has characteristics of direction and angle. The steam inlet direction is selected to flow tangentially, and as steam flows through the inside wall, it is effective for the heat transfer to and from the wall, as illustrated in Figure 1(b). The inlet angle (φ) is selected as 5°, and the vortex generated on the inside wall is expanded [27].The vortex flow increases the streamline length of the steam, resulting in a pressure drop. Therefore, it is more efficient for generating LP steam compared to the HCHE, where the pressure drop due to change in steam flow rate hardly occurs. In addition, to maximize the swirl flow, the steam inlet velocity is designed to be less than 100 m/s [28], and the steam inlet diameter is 8 mm using Equation (1), where ṁ is the steam flow rate, ρ is the density of the steam, and a is the inlet area of the HTSG.
(1)Vi=ṁρa.The magnitude of the centrifugal force (Fc) applied to the steam increases as the diameter of the HTSG decreases, which is proportional to the speed of the streamline. The HTSG body diameter (Db) and length (Lb) are 350 mm and 550 mm, respectively. The steam outlet diameter (Do) is 17.3 mm and is designed to maintain 60% of the inlet velocity.


### 2.1.2. System Apparatus

In Section2.1.1, the HTSG was designed considering the characteristics of pressure drop, and it is intended to verify the device by applying it to generate HTLP steam. In this chapter, a lab-scale system to which HTSG is fabricated via design is constructed, as illustrated in Figure 2.Figure 2
Schematic diagram HTSG system for pressure drop analysis.The main equipment of the HTSG system comprises a previously designed HTSG, steam boiler, and combustion furnace, as illustrated in Figure3. First (low-temperature and low-pressure) steam is generated in a steam boiler in a superheated vapor state. The first steam was maintained at the value set in the steam boiler, without pressure control. The operating range of steam that can be produced in the steam boiler is 100–1,000 kPa, and the steam temperature is generated in a superheated steam state of pressure.Figure 3
Experimental system for pressure drop analysis.Variations in the primary steam cause fluctuations in the entire system and are not suitable because uniform steam must be supplied to the SOEC; therefore, a pressure-reducing valve (PRV) is necessary. The PRV maintains the stability of the primary steam, and the secondary steam from the PRV flows uniformly into the HTSG inlet. The combustion furnace is maintained by supplying liquified natural gas (LNG) heat, and an electronic volume compensator for gas flow measurements is installed. The low calorific value of LNG is 26 MJ/Nm3. It supplies gas heat to the HTSG installed inside, and the gas flow rate is kept constant at 5.5 m3/h. The LTLP steam inside the HTSG is heat exchanged with gas heat at a temperature of 773 K or higher to generate HTLP steam applicable to SOEC. The operating conditions are presented in Table 1. The HTLP steam generated in this system is discharged through the steam outlet and can be provided in linkage with the SOEC.Table 1
The operating conditions of the HTSG system.
Systematic characteristicsOperating conditionsSteam inlet condition (kPa)Saturated vapor state (300–600)Steam inlet flow rate (kg/h)10–35Steam outlet temperature (K)773Combustion gas temperature (K)973


### 2.1.3. Data Reduction and Uncertainty Analysis

The system was operated for 11 h during the experiment and maintained in a steady state for approximately half a day after the initial operation of the system. The pressure drop occurring in the HTSG was evaluated using Equation (2). In particular, in pipe friction caused by sudden contraction, the pressure drop is approximately negligible, equivalent to 1.19% of the swirl flow [27].
(2)∆Pt=∆Pse+∆Psf+∆Psc.The error of the pressure drop derived from the correlation and experimental data was evaluated using Equation (3), where the subscript “pre” and “exp” indicate the prediction and experiment, respectively.
(3)Error=ΔPpreΔPexp.The system error is derived based on the method of Abernethy et al. [29] and is expressed in Equations (4)–(7). Uc denotes the experimental error, A is the bias error, B is the random standard error, and σ is the standard deviation of the experimental values (x), including the temperature, pressure, and mass flow rate.
(4)Uc=4∙∑n=1N∂X∂xn∙BPn2+∑n=1N∂X∂xn∙APn2,(5)BP=σN,(6)X=Xx1,x2,x3,⋯,xn,(7)σ=1N−1∑n=1NXn−X¯2.The specifications and accuracies of the measuring devices in the system are presented in Table2, and the errors for each parameter are summarized in Table 3. The experimental uncertainty in the pressure drop was estimated to be 9.49%.Table 2
The specifications of the measuring devices.
ParameterTypeOperating rangeAccuracySteam temperature (K)K-type-473–1,523±0.02Steam pressure (kPa)Pressure transducer0–1,000±4.71Steam flow rate (kg/h)Volumetric flowmeter10.0–47.6±0.35Gas flow rate (m3/h)Electronic volume corrector95–300±5.17Table 3
Experimental uncertainty for HTLP steam generation.
ParametersA (%)B (%)Steam temperature0.01480.61Gas temperature0.01522.62Pressure1.560.056Steam flow rate1.751.26Gas flow rate1.721.48


## 2.1.1. Model Design

The conventional steam generator generally utilized in the field is a helical coiled heat exchanger (HCHE). The HCHE has a large heat transfer area and high thermal conductivity; however, owing to its thin thickness, it is vulnerable to corrosion, erosion, and cracking when operated at high temperatures for a long time, which deteriorates the durability of the system [19]. Above all, it is difficult to generate HTLP steam because the pressure drop is small, and it is not suitable for utilization for SOEC that requires LP steam [20]. In addition, the device has an unnecessarily large surface area, and when installed outside in winter, the device is highly susceptible to the seasonal changes. In this study, a suitable form of the HTSG is developed to generate the HTLP steam. A cyclone cylindrical type is selected for the HTSG, as illustrated in Figure 1. The thickness is over five times that of a general HCHE, and it can withstand the high temperature. The HTSG is designed to generate a required pressure drop based on the effect of each section.Figure 1
Geometric model of high-temperature steam generator for pressure drop analysis.Although the external structure of the HTSG is simple, a complex three-dimensional turbulence occurs inside, and a pressure drop occurs owing to steam flow. In this study, this phenomenon was adopted to produce LP steam [21]. The cylindrical shape is generally utilized as a separator and has an excellent function in strengthening the internal pressure drop [22]. In addition, the inlet nozzle section is configured in the tangential direction as much as possible to maximize the centrifugal force generated when steam enters the HTSG inlet [23]. As the shape alters, the number of geometric vortices has a strong influence on the internal flow characteristics [24], and in particular, the streamline length according to the HTSG body diameter is an essential factor to consider in terms of performance improvement for LP steam production [25, 26].In the developed HTSG, there are three sections in which the pressure drop occurs, and it is analyzed based on (1) sudden expansion (∆Pse) from the steam inlet, (2) swirl flow (∆Psf) on the wall of the HTSG, and (3) sudden contraction (∆Psf) at the steam outlet. In addition, the steam inlet has characteristics of direction and angle. The steam inlet direction is selected to flow tangentially, and as steam flows through the inside wall, it is effective for the heat transfer to and from the wall, as illustrated in Figure 1(b). The inlet angle (φ) is selected as 5°, and the vortex generated on the inside wall is expanded [27].The vortex flow increases the streamline length of the steam, resulting in a pressure drop. Therefore, it is more efficient for generating LP steam compared to the HCHE, where the pressure drop due to change in steam flow rate hardly occurs. In addition, to maximize the swirl flow, the steam inlet velocity is designed to be less than 100 m/s [28], and the steam inlet diameter is 8 mm using Equation (1), where ṁ is the steam flow rate, ρ is the density of the steam, and a is the inlet area of the HTSG.
(1)Vi=ṁρa.The magnitude of the centrifugal force (Fc) applied to the steam increases as the diameter of the HTSG decreases, which is proportional to the speed of the streamline. The HTSG body diameter (Db) and length (Lb) are 350 mm and 550 mm, respectively. The steam outlet diameter (Do) is 17.3 mm and is designed to maintain 60% of the inlet velocity.


## 2.1.2. System Apparatus

In Section2.1.1, the HTSG was designed considering the characteristics of pressure drop, and it is intended to verify the device by applying it to generate HTLP steam. In this chapter, a lab-scale system to which HTSG is fabricated via design is constructed, as illustrated in Figure 2.Figure 2
Schematic diagram HTSG system for pressure drop analysis.The main equipment of the HTSG system comprises a previously designed HTSG, steam boiler, and combustion furnace, as illustrated in Figure3. First (low-temperature and low-pressure) steam is generated in a steam boiler in a superheated vapor state. The first steam was maintained at the value set in the steam boiler, without pressure control. The operating range of steam that can be produced in the steam boiler is 100–1,000 kPa, and the steam temperature is generated in a superheated steam state of pressure.Figure 3
Experimental system for pressure drop analysis.Variations in the primary steam cause fluctuations in the entire system and are not suitable because uniform steam must be supplied to the SOEC; therefore, a pressure-reducing valve (PRV) is necessary. The PRV maintains the stability of the primary steam, and the secondary steam from the PRV flows uniformly into the HTSG inlet. The combustion furnace is maintained by supplying liquified natural gas (LNG) heat, and an electronic volume compensator for gas flow measurements is installed. The low calorific value of LNG is 26 MJ/Nm3. It supplies gas heat to the HTSG installed inside, and the gas flow rate is kept constant at 5.5 m3/h. The LTLP steam inside the HTSG is heat exchanged with gas heat at a temperature of 773 K or higher to generate HTLP steam applicable to SOEC. The operating conditions are presented in Table 1. The HTLP steam generated in this system is discharged through the steam outlet and can be provided in linkage with the SOEC.Table 1
The operating conditions of the HTSG system.
Systematic characteristicsOperating conditionsSteam inlet condition (kPa)Saturated vapor state (300–600)Steam inlet flow rate (kg/h)10–35Steam outlet temperature (K)773Combustion gas temperature (K)973


## 2.1.3. Data Reduction and Uncertainty Analysis

The system was operated for 11 h during the experiment and maintained in a steady state for approximately half a day after the initial operation of the system. The pressure drop occurring in the HTSG was evaluated using Equation (2). In particular, in pipe friction caused by sudden contraction, the pressure drop is approximately negligible, equivalent to 1.19% of the swirl flow [27].
(2)∆Pt=∆Pse+∆Psf+∆Psc.The error of the pressure drop derived from the correlation and experimental data was evaluated using Equation (3), where the subscript “pre” and “exp” indicate the prediction and experiment, respectively.
(3)Error=ΔPpreΔPexp.The system error is derived based on the method of Abernethy et al. [29] and is expressed in Equations (4)–(7). Uc denotes the experimental error, A is the bias error, B is the random standard error, and σ is the standard deviation of the experimental values (x), including the temperature, pressure, and mass flow rate.
(4)Uc=4∙∑n=1N∂X∂xn∙BPn2+∑n=1N∂X∂xn∙APn2,(5)BP=σN,(6)X=Xx1,x2,x3,⋯,xn,(7)σ=1N−1∑n=1NXn−X¯2.The specifications and accuracies of the measuring devices in the system are presented in Table2, and the errors for each parameter are summarized in Table 3. The experimental uncertainty in the pressure drop was estimated to be 9.49%.Table 2
The specifications of the measuring devices.
ParameterTypeOperating rangeAccuracySteam temperature (K)K-type-473–1,523±0.02Steam pressure (kPa)Pressure transducer0–1,000±4.71Steam flow rate (kg/h)Volumetric flowmeter10.0–47.6±0.35Gas flow rate (m3/h)Electronic volume corrector95–300±5.17Table 3
Experimental uncertainty for HTLP steam generation.
ParametersA (%)B (%)Steam temperature0.01480.61Gas temperature0.01522.62Pressure1.560.056Steam flow rate1.751.26Gas flow rate1.721.48


## 2.2. Experimental Results

The performance of the developed HTSG was evaluated under the influence of geometric conditions, and the pressure drop effect occurred, as illustrated in Figure4. When the steam flow rate of the experiment was converted to speed, the pressure drop increased proportionally according to the local velocity under all pressure conditions. This is an obvious result that can be interpreted as local velocity having a dominant effect on the HTSG. Figure 5 illustrates the friction factor based on the pressure drop data obtained from the experiment. ∆Pse and fse for sudden expansion were calculated using Equations (8) and (9), respectively, and Vin was the steam velocity in the inlet nozzle section. For the steam properties, the average of the inlet and outlet values of the HTSG was adopted, and B was the ratio of the cross-sectional area.
(8)ΔPse=fse·ρ·Vi22,(9)fse=1−1B2.Figure 4
Pressure drop of HTSG system according to operating conditions.Figure 5
Friction factor for swirl flow in variation of inlet steam pressure.∆Psc and fsc for sudden contraction were calculated using Equations (10) and (11), where Vout represents the steam velocity in the inlet nozzle section.
(10)ΔPsc=fSC·ρ·Vo22,(11)fsc=0.45·1−1B.ΔPsf and fsf for the swirl flow are calculated using Equations (12) and (13), where Vh represents the local velocity in the HTSG. The main friction coefficient (f) was derived from the Blasius correlation in Equation (14). It was adopted when the Reynolds number (Re) in the experiment was 104 scale. The length of the steam stream inside the HTSG was estimated based on the inlet nozzle angle (φ) and steam flow in the experiment. k is calculated by Equation (15) from the pressure drop data obtained from the experiment, and the maximum error of the model prediction was 11.5%.
(12)ΔPsf=f∙LD∙ρ∙Vh22+fsf∙ρ∙Vin22,(13)fsf=πLD∙tanφ∙k,(14)f=0.3164Re0.25,(15)k=2.428Vi−0.00248.Similar results were obtained, regardless of the variation in inlet steam pressure. However, as the local velocity of the inlet steam increased, the friction factor decreased because the steam density and steam viscosity under each operating condition were different. In particular, when the local velocity was 150 m/s or higher in the low-pressure condition, the friction factor increased considerably. This phenomenon is interpreted as the system becoming unstable because the energy loss increased as the steam velocity approached the compressibility. Definitely, there are advantages in terms of hydrogen production and system stability at low pressure, but the energy loss in the system due to the increase in pressure drop also increased. Therefore, for the inlet condition of the HTSG, 600 kPa, which can supply steam constantly without affecting system stability, was selected as an appropriate value. A strategy is required to strengthen the pressure drop inside the HTSG to generate steam at 300 kPa. The optimization of the HTSG shape is an important task to predict the steam pressure, and specific design factors affecting the pressure drop can be derived via simulations based on experimental values.


## 3. Simulation Methods

### 3.1. Numerical Analysis

Because the pressure drop inside the HTSG significantly depends on the steam flow, the flow should be checked via a CFD simulation. Fluent 19.0 from ANSYS was adopted for the simulation. The physical model of the HTSG and combustion furnace was the same as the experimental design, and for comparison with the experimental values, it was constructed based on the numerical values illustrated in Figure1. It represents the physical model of the HTSG, and the design module was utilized in Fluent 19.0. To increase the precision of the pressure drop occurring inside the HTSG, a k–ω SST model with high accuracy for the free shear flow according to the steam flow was utilized. The model is more suitable for a fully developed turbulence analysis than the k–ε model which has limited accuracy and is advantageous for analyzing phenomena in the boundary layer of the wall. In particular, owing to the large Reynolds number of the 104 scale, it can be accurately predicted considering the semilaminar flow, buffer zone, and fully developed zone. If the characteristic length of the current is negligible, the transport effect can be neglected. The law of mass conservation for vapor gases involving density, time, and local velocity is considered. Polynomials were adopted for properties such as density, thermal conductivity, and specific heat capacity, and regression analysis was employed to increase the accuracy of the simulation analysis [30–32]. For the eddy motion of a turbulent flow, each term can be adopted, as expressed in Equation (16), by applying the Reynolds average numerical simulation (RANS) method. X represents the velocity, enthalpy, and pressure, and it can be expressed as the sum of the time average value and fluctuation term; the density fluctuation term is ignored.
(16)X=X¯+X′.The fluctuation term can be estimated based on the concept of molecular motion. This term is modeled as the eddy viscosity based on Boussinesq’s hypothesis:μt, i, j, and δ denote turbulent viscosity, mean free index, repeated index, and Kronecker delta, respectively.


#### 3.1.1. Grid Independence Analysis

To ensure similarity between the experiments and simulations, it is necessary to validate the predicted results and experimental data. Mesh quality is an important indicator for ensuring the reliability of geometric models. Therefore, to select a suitable mesh quality for the simulation, a grid independence analysis was performed. The initial mesh size of the model was set to 5 mm, average skewness was 0.22, average orthogonal quality was 0.86, and mesh geometry was confirmed, as illustrated in Figure6. Skewness refers to the degree to which the mesh deviates from the equilateral polygon; the closer it is to zero, the higher the mesh quality. Orthogonal quality is an index calculated using a normal vector and one from the mesh center to the center of the face. The closer the value is to one, the higher the mesh quality. Therefore, this model satisfied the acceptable criteria.Figure 6
Geometric model for numerical simulation.For the grid independence analysis in the simulation, the test was conducted by dividing the mesh size into four, as presented in Table4. Figure 7 illustrates the results according to the mesh size and pressure drop of the experimental data, and the pressure drop (steam inlet pressure: 600 kPa, steam flow rate: 35 kg/h) of the experimental data used for comparison was 80.5 kPa. As the number of cells increased, the difference between the simulation results and experimental data decreased remarkably, and as the number of cells increased, the pressure drop increased. However, in the simulation, it was more important to reduce the error with the experimental value than to check the mesh size, which increased the pressure drop. The pressure drop of grids 3 and 4 and values of the experimental data differ by only 1.23% and 1.5%, respectively. In particular, the difference between the predicted pressure drops in grids 3 and 4 was less than 2%, indicating a similar value. Therefore, it can be confirmed that grid 3 provides grid independence results that increase the accuracy more closely, with respect to the pressure drop for the flow, suggesting that it is suitable for the simulation. Figure 8 illustrates a comparison of the experimental and simulation data. The maximum error of the comparison value was 6.2%, indicating that the results were within 10% of the simulation error standard [33].Table 4
Grid independence test.
CaseGrid 1Grid 2Grid 3Grid 4Mesh size (mm)52.510.1Number of cells1,725,9173,525,8196,173,50013,371,645Figure 7
Grid independence with number of cells in steam domain mesh.Figure 8
Validation of pressure drop analysis.The pressure drop inside the HTSG can be clearly observed in Figure9. Sudden expansion in the steam inlet, swirl flow rotating in the axial direction, and sudden contraction occurring when exiting the outlet were visually confirmed. It was confirmed that each pressure level is different, and pressure drop occurring in the term can be calculated. In addition, the effective length inside the HTSG can be estimated based on the angle of the inlet nozzle and path of the vapor movement through the swirl flow.Figure 9
Flow analysis of pressure drop in the HTSG system.


### 3.2. Dynamic Analysis

The pressure drop has a significant effect on SOEC requiring low-pressure steam and affects the stable and efficient production of hydrogen. Therefore, in Section3.2, dynamic simulation modeling linking the HTSG and SOEC is developed to enhance hydrogen production. TRNSYS 18 has excellent functions for system performance modeling, parameter variation, and long-term operation based on physical characteristics. Figure 10(a) is a model for high-temperature and low-pressure steam generation using HTSG, configured in a similar way as the experimental device, and each component element is inputted similar to the parameter values of the numerical simulation. In Figure 10(b), it is possible to predict the annual hydrogen production according to pressure drop using the SOEC model selected within the TRNSYS environment. The results can be compared with the numerical analysis, and it can be determined whether the optimal design of the HTSG is appropriate.Figure 10
Linked HTSG-SOEC modeling for hydrogen production prediction: (a) steam generation system using HTSG; (b) hydrogen production process using SOEC.


## 3.1. Numerical Analysis

Because the pressure drop inside the HTSG significantly depends on the steam flow, the flow should be checked via a CFD simulation. Fluent 19.0 from ANSYS was adopted for the simulation. The physical model of the HTSG and combustion furnace was the same as the experimental design, and for comparison with the experimental values, it was constructed based on the numerical values illustrated in Figure1. It represents the physical model of the HTSG, and the design module was utilized in Fluent 19.0. To increase the precision of the pressure drop occurring inside the HTSG, a k–ω SST model with high accuracy for the free shear flow according to the steam flow was utilized. The model is more suitable for a fully developed turbulence analysis than the k–ε model which has limited accuracy and is advantageous for analyzing phenomena in the boundary layer of the wall. In particular, owing to the large Reynolds number of the 104 scale, it can be accurately predicted considering the semilaminar flow, buffer zone, and fully developed zone. If the characteristic length of the current is negligible, the transport effect can be neglected. The law of mass conservation for vapor gases involving density, time, and local velocity is considered. Polynomials were adopted for properties such as density, thermal conductivity, and specific heat capacity, and regression analysis was employed to increase the accuracy of the simulation analysis [30–32]. For the eddy motion of a turbulent flow, each term can be adopted, as expressed in Equation (16), by applying the Reynolds average numerical simulation (RANS) method. X represents the velocity, enthalpy, and pressure, and it can be expressed as the sum of the time average value and fluctuation term; the density fluctuation term is ignored.
(16)X=X¯+X′.The fluctuation term can be estimated based on the concept of molecular motion. This term is modeled as the eddy viscosity based on Boussinesq’s hypothesis:μt, i, j, and δ denote turbulent viscosity, mean free index, repeated index, and Kronecker delta, respectively.


### 3.1.1. Grid Independence Analysis

To ensure similarity between the experiments and simulations, it is necessary to validate the predicted results and experimental data. Mesh quality is an important indicator for ensuring the reliability of geometric models. Therefore, to select a suitable mesh quality for the simulation, a grid independence analysis was performed. The initial mesh size of the model was set to 5 mm, average skewness was 0.22, average orthogonal quality was 0.86, and mesh geometry was confirmed, as illustrated in Figure6. Skewness refers to the degree to which the mesh deviates from the equilateral polygon; the closer it is to zero, the higher the mesh quality. Orthogonal quality is an index calculated using a normal vector and one from the mesh center to the center of the face. The closer the value is to one, the higher the mesh quality. Therefore, this model satisfied the acceptable criteria.Figure 6
Geometric model for numerical simulation.For the grid independence analysis in the simulation, the test was conducted by dividing the mesh size into four, as presented in Table4. Figure 7 illustrates the results according to the mesh size and pressure drop of the experimental data, and the pressure drop (steam inlet pressure: 600 kPa, steam flow rate: 35 kg/h) of the experimental data used for comparison was 80.5 kPa. As the number of cells increased, the difference between the simulation results and experimental data decreased remarkably, and as the number of cells increased, the pressure drop increased. However, in the simulation, it was more important to reduce the error with the experimental value than to check the mesh size, which increased the pressure drop. The pressure drop of grids 3 and 4 and values of the experimental data differ by only 1.23% and 1.5%, respectively. In particular, the difference between the predicted pressure drops in grids 3 and 4 was less than 2%, indicating a similar value. Therefore, it can be confirmed that grid 3 provides grid independence results that increase the accuracy more closely, with respect to the pressure drop for the flow, suggesting that it is suitable for the simulation. Figure 8 illustrates a comparison of the experimental and simulation data. The maximum error of the comparison value was 6.2%, indicating that the results were within 10% of the simulation error standard [33].Table 4
Grid independence test.
CaseGrid 1Grid 2Grid 3Grid 4Mesh size (mm)52.510.1Number of cells1,725,9173,525,8196,173,50013,371,645Figure 7
Grid independence with number of cells in steam domain mesh.Figure 8
Validation of pressure drop analysis.The pressure drop inside the HTSG can be clearly observed in Figure9. Sudden expansion in the steam inlet, swirl flow rotating in the axial direction, and sudden contraction occurring when exiting the outlet were visually confirmed. It was confirmed that each pressure level is different, and pressure drop occurring in the term can be calculated. In addition, the effective length inside the HTSG can be estimated based on the angle of the inlet nozzle and path of the vapor movement through the swirl flow.Figure 9
Flow analysis of pressure drop in the HTSG system.


## 3.1.1. Grid Independence Analysis

To ensure similarity between the experiments and simulations, it is necessary to validate the predicted results and experimental data. Mesh quality is an important indicator for ensuring the reliability of geometric models. Therefore, to select a suitable mesh quality for the simulation, a grid independence analysis was performed. The initial mesh size of the model was set to 5 mm, average skewness was 0.22, average orthogonal quality was 0.86, and mesh geometry was confirmed, as illustrated in Figure6. Skewness refers to the degree to which the mesh deviates from the equilateral polygon; the closer it is to zero, the higher the mesh quality. Orthogonal quality is an index calculated using a normal vector and one from the mesh center to the center of the face. The closer the value is to one, the higher the mesh quality. Therefore, this model satisfied the acceptable criteria.Figure 6
Geometric model for numerical simulation.For the grid independence analysis in the simulation, the test was conducted by dividing the mesh size into four, as presented in Table4. Figure 7 illustrates the results according to the mesh size and pressure drop of the experimental data, and the pressure drop (steam inlet pressure: 600 kPa, steam flow rate: 35 kg/h) of the experimental data used for comparison was 80.5 kPa. As the number of cells increased, the difference between the simulation results and experimental data decreased remarkably, and as the number of cells increased, the pressure drop increased. However, in the simulation, it was more important to reduce the error with the experimental value than to check the mesh size, which increased the pressure drop. The pressure drop of grids 3 and 4 and values of the experimental data differ by only 1.23% and 1.5%, respectively. In particular, the difference between the predicted pressure drops in grids 3 and 4 was less than 2%, indicating a similar value. Therefore, it can be confirmed that grid 3 provides grid independence results that increase the accuracy more closely, with respect to the pressure drop for the flow, suggesting that it is suitable for the simulation. Figure 8 illustrates a comparison of the experimental and simulation data. The maximum error of the comparison value was 6.2%, indicating that the results were within 10% of the simulation error standard [33].Table 4
Grid independence test.
CaseGrid 1Grid 2Grid 3Grid 4Mesh size (mm)52.510.1Number of cells1,725,9173,525,8196,173,50013,371,645Figure 7
Grid independence with number of cells in steam domain mesh.Figure 8
Validation of pressure drop analysis.The pressure drop inside the HTSG can be clearly observed in Figure9. Sudden expansion in the steam inlet, swirl flow rotating in the axial direction, and sudden contraction occurring when exiting the outlet were visually confirmed. It was confirmed that each pressure level is different, and pressure drop occurring in the term can be calculated. In addition, the effective length inside the HTSG can be estimated based on the angle of the inlet nozzle and path of the vapor movement through the swirl flow.Figure 9
Flow analysis of pressure drop in the HTSG system.


## 3.2. Dynamic Analysis

The pressure drop has a significant effect on SOEC requiring low-pressure steam and affects the stable and efficient production of hydrogen. Therefore, in Section3.2, dynamic simulation modeling linking the HTSG and SOEC is developed to enhance hydrogen production. TRNSYS 18 has excellent functions for system performance modeling, parameter variation, and long-term operation based on physical characteristics. Figure 10(a) is a model for high-temperature and low-pressure steam generation using HTSG, configured in a similar way as the experimental device, and each component element is inputted similar to the parameter values of the numerical simulation. In Figure 10(b), it is possible to predict the annual hydrogen production according to pressure drop using the SOEC model selected within the TRNSYS environment. The results can be compared with the numerical analysis, and it can be determined whether the optimal design of the HTSG is appropriate.Figure 10
Linked HTSG-SOEC modeling for hydrogen production prediction: (a) steam generation system using HTSG; (b) hydrogen production process using SOEC.


## 4. Results and Discussions

### 4.1. Effect of HTSG Geometry

To maintain a steam condition suitable for the operating conditions, the design factor that has the greatest influence on pressure drop inside the HTSG should be derived. In this section, shape optimization is performed by varying the body ratio (Db/Lb), outlet diameter (Do), and inlet angle (φ) of the HTSG. A desired pressure input value can be selected when designing an HTSG based on the predicted and derived pressure-drop indices.


#### 4.1.1. Variation of Inlet Nozzle Angle

Selection of the inlet angle is an important factor that can enhance the swirl flow occurring on the inner wall of the HTSG. The centrifugal force generated by the swirl causes steam to move toward the outer wall. Wang et al. [34] conducted simulations for negative and positive entrance angles and analyzed the flow trajectories of particles that vary depending on each inlet angle. In particular, as suggested by Mahmoud et al. [35], varying the internal pressure drop using design factors according to the operating conditions is an important parameter that has a significant effect on performance. Therefore, in this section, simulations were conducted by altering the inlet angle to obtain the pressure drop and axial steam velocity distribution in the HTSG. The inlet nozzle angles are presented in Table 5. A geometric model based on the inlet nozzle angle is illustrated in Figure 11.Table 5
Variation of inlet nozzle angle.
Inlet angle,φ (°)Standard5Case 110Case 215Case 320Case 425Figure 11
Geometric model according to inlet nozzle angle.Figure12(a) illustrates the local velocity and friction factor in the pressure drop analysis for a change in the inlet angle. The pressure drop is lower when φ is 25° than when φ is 5°, and the change in the inlet angle indicates that the length of the steam stream inside the HTSG is affected. The Reynolds number of the steam in the HTSG inlet is approximately 70,000, which causes rapid turbulent flow; however, as φ increases, the local outlet velocity decreases and is interpreted as affecting the pressure drop. The effect of the pressure drop on the annual hydrogen production is illustrated in Figure 12(b). As the inlet angle increases, the pressure drop decreases by up to 32%, and annual hydrogen production decreases by approximately 4%.Figure 12
Parameter value prediction according to inlet nozzle angle: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


#### 4.1.2. Variation of Outlet Diameter

Currently, the inlet velocity of the HTSG is designed as 60–100 m/s. An industrial standard is utilized in the initial design, and the swirl flow inside the body was strengthened. If the velocity is high, the heat transfer to and from the gas inside the body is insufficient and is sucked out at a high speed. However, if the velocity is low, the steam stays inside the body and the discharge time may be delayed; hence, it may not be possible to obtain the HP steam required for SOEC. Therefore, it is necessary to predict∆Psc by altering the appropriate outlet diameter and determining the ideal outlet diameter. Table 6 and Figure 13 present the variation in the body diameter and length and the outlet diameter when the steam flow rate is fixed. A commercial standard is adopted to determine the outlet diameter.Table 6
Variation of outlet diameter.
Di (mm)Do (mm)StandardID 8OD 17.3 (10A)Case 1OD 21.7 (15A)Case 2OD 27.2 (20A)Case 3OD 34 (25A)Case 4OD 42.7 (32A)Figure 13
Geometric model according to outlet diameter.The pressure drop with respect to the outlet diameter is illustrated in Figure14(a). Regarding the Reynolds number, the pressure drop increases proportionally with the steam inlet flow rate. The friction factor for the sudden expansion and swirl flow does not change significantly because the inlet section is in a fixed state. However, the friction factor for sudden contraction and different aspects are indicated, which is clearly confirmed in the results. It is determined that the diameter of the HTSG body and the outlet diameter have a significant effect on the constant B of the friction factor for sudden contraction, and the steam flow at this time is analyzed as abruptly changing owing to the outlet diameter. At an outlet diameter of 42.7 mm (32A) in this system, the pressure drop is greatest and is approximately 151 kPa. Figure 14(b) illustrates that hydrogen that can be produced owing to the corresponding parameter change is up to 32 tons annually. When the Reynolds number is above approximately 43,000, hydrogen production tends to increase linearly. Because it is not solely affected by sudden contraction but also changes in the inlet local velocity, it is considered to exhibit a stable rate of change in low-pressure steam with a large pressure drop.Figure 14
Parameter value prediction for sudden contraction: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


#### 4.1.3. Variation of HTSG Body Ratio (Diameter and Length)

In a previous study [27], the optimal inlet nozzle angle and outlet diameter were selected via simulation. However, the outlet pressure was approximately 449 kPa, which did not reach the steam pressure of less than 300 kPa for supplying the SOEC. Therefore, a variation in additional parameters is required to enhance the pressure drop. The pressure drop for the swirl flow exhibited a significant difference depending on the body ratio. The length of the steam stream can depend significantly on the body diameter and length, and it is necessary to determine which factor has the greatest influence on the diameter and length. In Section 4.2, simulations are performed by altering the diameter and length. The local inlet velocity is 100–170 m/s, and the simulation variations are presented in Table 7. Case 3 was the same as the body ratio of the HTSG. The friction factor with respect to the body ratio is illustrated in Figure 15(a). It is confirmed that the pressure drop increases rapidly when the body volume is 0.4 m or more. In particular, in Cases 3 and 4 with the same body diameter, despite the 1.5 times difference in body length (0.55 : 0.825), there was no significant difference in the pressure drop. However, Cases 5 and 6 with the same body diameter exhibited a different trend, and the pressure drop increased according to the body volume under turbulent flow conditions. The outlet pressure in Case 6 was 309 kPa, which was determined to be the optimal condition for supplying the SOEC. Figure 15(b) illustrates the hydrogen production according to the pressure drop. The maximum amount of hydrogen that can be produced in Case 1 is 15.3 tons annually, and the amount of hydrogen that can be produced in Case 6, which has the lowest outlet pressure, is predicted to be approximately 35.7 tons annually. When steam is supplied under optimum conditions, the internal cell resistance is increased because mass transport is reduced in the SOEC and diffusion to dislocations is increased [36]. Because the diffusion resistance affects hydrogen production, hydrogen can be produced stably at low pressure.Table 7
Variation of HTSG body ratio.
Db (m)Lb (m)Db/LbVolume (m3)Case 10.1750.2750.630.04Case 20.2160.550.390.10Case 30.350.550.630.35Case 40.350.8250.420.40Case 50.5250.550.950.78Case 60.5250.8250.631.02Figure 15
Parameter value prediction according to HTSG body ratio: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


#### 4.1.4. Optimization of HTSG

The pressure drop of each term according to the variable factors was confirmed via numerical simulation. The corrected correlation equation based on the data is expressed in Equations (17)–(19). The swirl number (S) was added to Equation (12), which was estimated based on existing experimental data. The swirl number was highly dependent on the Reynolds number, inlet nozzle angle, and body ratio. In addition, the correlation with the number of geometric vortices in the HTSG geometry included the ratio of the steam profile to maximum velocity and is calculated using Equation (20) [37]. The swirl number in the HTSG shape was estimated to be less than 0.38. Different values of k were derived from the pressure drop data for the various changes. The range of the Reynolds number varied according to the change in the shape variable. kE was obtained from the results according to the inlet angle and outlet diameter in Equation (21). kH was obtained from the result according to the HTSG body ratio in Equation (22).
(17)ΔPsf=f∙LD∙ρ∙Vh22+fsf∙S∙ρ∙Vi22,(18)f=0.3164Re0.25,(19)fsf=πLD∙tanφ∙k,(20)S=2GγD∙Gx,(21)kΕ=2417.4Re−1.09,(22)kΗ=1×1010Re−2.16.kE and kH for the Reynolds number are confirmed in Figures 15(a) and 15(b), respectively, and the maximum error of the predicted data is 10.3%. A numerical value that is reduced by approximately 12% compared to the error of the experimental data proves that it is a suitable correlation for the shape characteristics of HTSG. Unlike coiled tube types with the same inlet and outlet diameters, HTSG is designed to produce steam suitable for operating conditions of the SOEC, and it is necessary to predict the pressure drop within a sophisticated range. It can be predicted in the Reynolds number range of 28,700–306,600. In the HTSG, the effect on the outlet diameter and body ratio is greater than that of the inlet nozzle angle and is a factor that increases the Reynolds number. Users can utilize it to characterize the swirl flow, which is useful for HTSG design based on the prediction results. In addition, because it is designed considering the feasibility of the variation in the operating pressure of the SOEC, unnecessary economic loss can be reduced.


### 4.2. Effectiveness of CO2 Reduction

The HTSG can be adopted as an energy harvesting method that satisfies uniform hydrogen production in the industrial field. To enhance the production of hydrogen in SOEC, stability and durability must be guaranteed, and the steam condition must be less than 0.3 MPa. Therefore, in this study, the HTSG for generating HTLP steam was optimized based on a numerical simulation. If the optimal HTSG geometry derived from the simulation is adopted, HTLP steam can be generated uniformly, thereby enabling a stable supply to the SOEC. In addition, the linkage between the HTSG and SOEC was improved by predicting the hydrogen production according to the parametric analysis via the dynamic simulation.Currently, there are various energy sources that can produce steam, such as solid refuse fuel (SRF), LNG, and coal, and they can be produced in a large capacity. However, to reduce fossil fuels and CO2 emissions, it should be replaced with an eco-friendly heat source by visualizing and analyzing the values of CO2 emissions for each heat source system. The CO2 emission from the heat source according to the hydrogen production method is evaluated by Equations (23)–(25) [38] and illustrated in Figure 16 [39, 40].
(23)FuelheatingvalueMJ=fuelconsumptionNm3×heatgenerationMJNm3,(24)CarbonemissionstC=fuelheatingvalueMJ×CO2conversionfactortC/TJ106,(25)CO2emissionstCO2=carbonemissionstC×4412.Figure 16
Comparison of CO2 emissions during hydrogen production according to steam heat source.When the HTLP steam produced by the SRF presented in this study is utilized for hydrogen production, the CO2 emission is estimated as 9.92 ton-CO2/ton-hydrogen. This is approximately 11.1% greater than high-temperature and high-pressure (HTHP) steam by LNG and 86.8% greater than HTHP steam by coal, confirming that steam conditions affect the hydrogen production significantly. Moreover, it is very encouraging that the lowest CO2 emission was calculated when the energy source for supply was SRF. It is therefore confirmed that this is an efficient method compared to other energy sources, and when hydrogen is produced utilizing SRF, it can be considered as an eco-friendly solution that can harvest cost and resources.Since the system considers the effect on the pressure drop, an optimal design is required in terms of thermal energy. In the future, it will be intended to derive a method for improving heat transfer according to the operation conditions of the HTSG. It will be an important guideline for industrial application to SOEC, where the higher the temperature, the higher the hydrogen production efficiency.


## 4.1. Effect of HTSG Geometry

To maintain a steam condition suitable for the operating conditions, the design factor that has the greatest influence on pressure drop inside the HTSG should be derived. In this section, shape optimization is performed by varying the body ratio (Db/Lb), outlet diameter (Do), and inlet angle (φ) of the HTSG. A desired pressure input value can be selected when designing an HTSG based on the predicted and derived pressure-drop indices.


### 4.1.1. Variation of Inlet Nozzle Angle

Selection of the inlet angle is an important factor that can enhance the swirl flow occurring on the inner wall of the HTSG. The centrifugal force generated by the swirl causes steam to move toward the outer wall. Wang et al. [34] conducted simulations for negative and positive entrance angles and analyzed the flow trajectories of particles that vary depending on each inlet angle. In particular, as suggested by Mahmoud et al. [35], varying the internal pressure drop using design factors according to the operating conditions is an important parameter that has a significant effect on performance. Therefore, in this section, simulations were conducted by altering the inlet angle to obtain the pressure drop and axial steam velocity distribution in the HTSG. The inlet nozzle angles are presented in Table 5. A geometric model based on the inlet nozzle angle is illustrated in Figure 11.Table 5
Variation of inlet nozzle angle.
Inlet angle,φ (°)Standard5Case 110Case 215Case 320Case 425Figure 11
Geometric model according to inlet nozzle angle.Figure12(a) illustrates the local velocity and friction factor in the pressure drop analysis for a change in the inlet angle. The pressure drop is lower when φ is 25° than when φ is 5°, and the change in the inlet angle indicates that the length of the steam stream inside the HTSG is affected. The Reynolds number of the steam in the HTSG inlet is approximately 70,000, which causes rapid turbulent flow; however, as φ increases, the local outlet velocity decreases and is interpreted as affecting the pressure drop. The effect of the pressure drop on the annual hydrogen production is illustrated in Figure 12(b). As the inlet angle increases, the pressure drop decreases by up to 32%, and annual hydrogen production decreases by approximately 4%.Figure 12
Parameter value prediction according to inlet nozzle angle: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


### 4.1.2. Variation of Outlet Diameter

Currently, the inlet velocity of the HTSG is designed as 60–100 m/s. An industrial standard is utilized in the initial design, and the swirl flow inside the body was strengthened. If the velocity is high, the heat transfer to and from the gas inside the body is insufficient and is sucked out at a high speed. However, if the velocity is low, the steam stays inside the body and the discharge time may be delayed; hence, it may not be possible to obtain the HP steam required for SOEC. Therefore, it is necessary to predict∆Psc by altering the appropriate outlet diameter and determining the ideal outlet diameter. Table 6 and Figure 13 present the variation in the body diameter and length and the outlet diameter when the steam flow rate is fixed. A commercial standard is adopted to determine the outlet diameter.Table 6
Variation of outlet diameter.
Di (mm)Do (mm)StandardID 8OD 17.3 (10A)Case 1OD 21.7 (15A)Case 2OD 27.2 (20A)Case 3OD 34 (25A)Case 4OD 42.7 (32A)Figure 13
Geometric model according to outlet diameter.The pressure drop with respect to the outlet diameter is illustrated in Figure14(a). Regarding the Reynolds number, the pressure drop increases proportionally with the steam inlet flow rate. The friction factor for the sudden expansion and swirl flow does not change significantly because the inlet section is in a fixed state. However, the friction factor for sudden contraction and different aspects are indicated, which is clearly confirmed in the results. It is determined that the diameter of the HTSG body and the outlet diameter have a significant effect on the constant B of the friction factor for sudden contraction, and the steam flow at this time is analyzed as abruptly changing owing to the outlet diameter. At an outlet diameter of 42.7 mm (32A) in this system, the pressure drop is greatest and is approximately 151 kPa. Figure 14(b) illustrates that hydrogen that can be produced owing to the corresponding parameter change is up to 32 tons annually. When the Reynolds number is above approximately 43,000, hydrogen production tends to increase linearly. Because it is not solely affected by sudden contraction but also changes in the inlet local velocity, it is considered to exhibit a stable rate of change in low-pressure steam with a large pressure drop.Figure 14
Parameter value prediction for sudden contraction: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


### 4.1.3. Variation of HTSG Body Ratio (Diameter and Length)

In a previous study [27], the optimal inlet nozzle angle and outlet diameter were selected via simulation. However, the outlet pressure was approximately 449 kPa, which did not reach the steam pressure of less than 300 kPa for supplying the SOEC. Therefore, a variation in additional parameters is required to enhance the pressure drop. The pressure drop for the swirl flow exhibited a significant difference depending on the body ratio. The length of the steam stream can depend significantly on the body diameter and length, and it is necessary to determine which factor has the greatest influence on the diameter and length. In Section 4.2, simulations are performed by altering the diameter and length. The local inlet velocity is 100–170 m/s, and the simulation variations are presented in Table 7. Case 3 was the same as the body ratio of the HTSG. The friction factor with respect to the body ratio is illustrated in Figure 15(a). It is confirmed that the pressure drop increases rapidly when the body volume is 0.4 m or more. In particular, in Cases 3 and 4 with the same body diameter, despite the 1.5 times difference in body length (0.55 : 0.825), there was no significant difference in the pressure drop. However, Cases 5 and 6 with the same body diameter exhibited a different trend, and the pressure drop increased according to the body volume under turbulent flow conditions. The outlet pressure in Case 6 was 309 kPa, which was determined to be the optimal condition for supplying the SOEC. Figure 15(b) illustrates the hydrogen production according to the pressure drop. The maximum amount of hydrogen that can be produced in Case 1 is 15.3 tons annually, and the amount of hydrogen that can be produced in Case 6, which has the lowest outlet pressure, is predicted to be approximately 35.7 tons annually. When steam is supplied under optimum conditions, the internal cell resistance is increased because mass transport is reduced in the SOEC and diffusion to dislocations is increased [36]. Because the diffusion resistance affects hydrogen production, hydrogen can be produced stably at low pressure.Table 7
Variation of HTSG body ratio.
Db (m)Lb (m)Db/LbVolume (m3)Case 10.1750.2750.630.04Case 20.2160.550.390.10Case 30.350.550.630.35Case 40.350.8250.420.40Case 50.5250.550.950.78Case 60.5250.8250.631.02Figure 15
Parameter value prediction according to HTSG body ratio: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


### 4.1.4. Optimization of HTSG

The pressure drop of each term according to the variable factors was confirmed via numerical simulation. The corrected correlation equation based on the data is expressed in Equations (17)–(19). The swirl number (S) was added to Equation (12), which was estimated based on existing experimental data. The swirl number was highly dependent on the Reynolds number, inlet nozzle angle, and body ratio. In addition, the correlation with the number of geometric vortices in the HTSG geometry included the ratio of the steam profile to maximum velocity and is calculated using Equation (20) [37]. The swirl number in the HTSG shape was estimated to be less than 0.38. Different values of k were derived from the pressure drop data for the various changes. The range of the Reynolds number varied according to the change in the shape variable. kE was obtained from the results according to the inlet angle and outlet diameter in Equation (21). kH was obtained from the result according to the HTSG body ratio in Equation (22).
(17)ΔPsf=f∙LD∙ρ∙Vh22+fsf∙S∙ρ∙Vi22,(18)f=0.3164Re0.25,(19)fsf=πLD∙tanφ∙k,(20)S=2GγD∙Gx,(21)kΕ=2417.4Re−1.09,(22)kΗ=1×1010Re−2.16.kE and kH for the Reynolds number are confirmed in Figures 15(a) and 15(b), respectively, and the maximum error of the predicted data is 10.3%. A numerical value that is reduced by approximately 12% compared to the error of the experimental data proves that it is a suitable correlation for the shape characteristics of HTSG. Unlike coiled tube types with the same inlet and outlet diameters, HTSG is designed to produce steam suitable for operating conditions of the SOEC, and it is necessary to predict the pressure drop within a sophisticated range. It can be predicted in the Reynolds number range of 28,700–306,600. In the HTSG, the effect on the outlet diameter and body ratio is greater than that of the inlet nozzle angle and is a factor that increases the Reynolds number. Users can utilize it to characterize the swirl flow, which is useful for HTSG design based on the prediction results. In addition, because it is designed considering the feasibility of the variation in the operating pressure of the SOEC, unnecessary economic loss can be reduced.


## 4.1.1. Variation of Inlet Nozzle Angle

Selection of the inlet angle is an important factor that can enhance the swirl flow occurring on the inner wall of the HTSG. The centrifugal force generated by the swirl causes steam to move toward the outer wall. Wang et al. [34] conducted simulations for negative and positive entrance angles and analyzed the flow trajectories of particles that vary depending on each inlet angle. In particular, as suggested by Mahmoud et al. [35], varying the internal pressure drop using design factors according to the operating conditions is an important parameter that has a significant effect on performance. Therefore, in this section, simulations were conducted by altering the inlet angle to obtain the pressure drop and axial steam velocity distribution in the HTSG. The inlet nozzle angles are presented in Table 5. A geometric model based on the inlet nozzle angle is illustrated in Figure 11.Table 5
Variation of inlet nozzle angle.
Inlet angle,φ (°)Standard5Case 110Case 215Case 320Case 425Figure 11
Geometric model according to inlet nozzle angle.Figure12(a) illustrates the local velocity and friction factor in the pressure drop analysis for a change in the inlet angle. The pressure drop is lower when φ is 25° than when φ is 5°, and the change in the inlet angle indicates that the length of the steam stream inside the HTSG is affected. The Reynolds number of the steam in the HTSG inlet is approximately 70,000, which causes rapid turbulent flow; however, as φ increases, the local outlet velocity decreases and is interpreted as affecting the pressure drop. The effect of the pressure drop on the annual hydrogen production is illustrated in Figure 12(b). As the inlet angle increases, the pressure drop decreases by up to 32%, and annual hydrogen production decreases by approximately 4%.Figure 12
Parameter value prediction according to inlet nozzle angle: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


## 4.1.2. Variation of Outlet Diameter

Currently, the inlet velocity of the HTSG is designed as 60–100 m/s. An industrial standard is utilized in the initial design, and the swirl flow inside the body was strengthened. If the velocity is high, the heat transfer to and from the gas inside the body is insufficient and is sucked out at a high speed. However, if the velocity is low, the steam stays inside the body and the discharge time may be delayed; hence, it may not be possible to obtain the HP steam required for SOEC. Therefore, it is necessary to predict∆Psc by altering the appropriate outlet diameter and determining the ideal outlet diameter. Table 6 and Figure 13 present the variation in the body diameter and length and the outlet diameter when the steam flow rate is fixed. A commercial standard is adopted to determine the outlet diameter.Table 6
Variation of outlet diameter.
Di (mm)Do (mm)StandardID 8OD 17.3 (10A)Case 1OD 21.7 (15A)Case 2OD 27.2 (20A)Case 3OD 34 (25A)Case 4OD 42.7 (32A)Figure 13
Geometric model according to outlet diameter.The pressure drop with respect to the outlet diameter is illustrated in Figure14(a). Regarding the Reynolds number, the pressure drop increases proportionally with the steam inlet flow rate. The friction factor for the sudden expansion and swirl flow does not change significantly because the inlet section is in a fixed state. However, the friction factor for sudden contraction and different aspects are indicated, which is clearly confirmed in the results. It is determined that the diameter of the HTSG body and the outlet diameter have a significant effect on the constant B of the friction factor for sudden contraction, and the steam flow at this time is analyzed as abruptly changing owing to the outlet diameter. At an outlet diameter of 42.7 mm (32A) in this system, the pressure drop is greatest and is approximately 151 kPa. Figure 14(b) illustrates that hydrogen that can be produced owing to the corresponding parameter change is up to 32 tons annually. When the Reynolds number is above approximately 43,000, hydrogen production tends to increase linearly. Because it is not solely affected by sudden contraction but also changes in the inlet local velocity, it is considered to exhibit a stable rate of change in low-pressure steam with a large pressure drop.Figure 14
Parameter value prediction for sudden contraction: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


## 4.1.3. Variation of HTSG Body Ratio (Diameter and Length)

In a previous study [27], the optimal inlet nozzle angle and outlet diameter were selected via simulation. However, the outlet pressure was approximately 449 kPa, which did not reach the steam pressure of less than 300 kPa for supplying the SOEC. Therefore, a variation in additional parameters is required to enhance the pressure drop. The pressure drop for the swirl flow exhibited a significant difference depending on the body ratio. The length of the steam stream can depend significantly on the body diameter and length, and it is necessary to determine which factor has the greatest influence on the diameter and length. In Section 4.2, simulations are performed by altering the diameter and length. The local inlet velocity is 100–170 m/s, and the simulation variations are presented in Table 7. Case 3 was the same as the body ratio of the HTSG. The friction factor with respect to the body ratio is illustrated in Figure 15(a). It is confirmed that the pressure drop increases rapidly when the body volume is 0.4 m or more. In particular, in Cases 3 and 4 with the same body diameter, despite the 1.5 times difference in body length (0.55 : 0.825), there was no significant difference in the pressure drop. However, Cases 5 and 6 with the same body diameter exhibited a different trend, and the pressure drop increased according to the body volume under turbulent flow conditions. The outlet pressure in Case 6 was 309 kPa, which was determined to be the optimal condition for supplying the SOEC. Figure 15(b) illustrates the hydrogen production according to the pressure drop. The maximum amount of hydrogen that can be produced in Case 1 is 15.3 tons annually, and the amount of hydrogen that can be produced in Case 6, which has the lowest outlet pressure, is predicted to be approximately 35.7 tons annually. When steam is supplied under optimum conditions, the internal cell resistance is increased because mass transport is reduced in the SOEC and diffusion to dislocations is increased [36]. Because the diffusion resistance affects hydrogen production, hydrogen can be produced stably at low pressure.Table 7
Variation of HTSG body ratio.
Db (m)Lb (m)Db/LbVolume (m3)Case 10.1750.2750.630.04Case 20.2160.550.390.10Case 30.350.550.630.35Case 40.350.8250.420.40Case 50.5250.550.950.78Case 60.5250.8250.631.02Figure 15
Parameter value prediction according to HTSG body ratio: (a) Reynolds number and friction factor in pressure drop analysis; (b) annual hydrogen production according to pressure drop.
(a)(b)


## 4.1.4. Optimization of HTSG

The pressure drop of each term according to the variable factors was confirmed via numerical simulation. The corrected correlation equation based on the data is expressed in Equations (17)–(19). The swirl number (S) was added to Equation (12), which was estimated based on existing experimental data. The swirl number was highly dependent on the Reynolds number, inlet nozzle angle, and body ratio. In addition, the correlation with the number of geometric vortices in the HTSG geometry included the ratio of the steam profile to maximum velocity and is calculated using Equation (20) [37]. The swirl number in the HTSG shape was estimated to be less than 0.38. Different values of k were derived from the pressure drop data for the various changes. The range of the Reynolds number varied according to the change in the shape variable. kE was obtained from the results according to the inlet angle and outlet diameter in Equation (21). kH was obtained from the result according to the HTSG body ratio in Equation (22).
(17)ΔPsf=f∙LD∙ρ∙Vh22+fsf∙S∙ρ∙Vi22,(18)f=0.3164Re0.25,(19)fsf=πLD∙tanφ∙k,(20)S=2GγD∙Gx,(21)kΕ=2417.4Re−1.09,(22)kΗ=1×1010Re−2.16.kE and kH for the Reynolds number are confirmed in Figures 15(a) and 15(b), respectively, and the maximum error of the predicted data is 10.3%. A numerical value that is reduced by approximately 12% compared to the error of the experimental data proves that it is a suitable correlation for the shape characteristics of HTSG. Unlike coiled tube types with the same inlet and outlet diameters, HTSG is designed to produce steam suitable for operating conditions of the SOEC, and it is necessary to predict the pressure drop within a sophisticated range. It can be predicted in the Reynolds number range of 28,700–306,600. In the HTSG, the effect on the outlet diameter and body ratio is greater than that of the inlet nozzle angle and is a factor that increases the Reynolds number. Users can utilize it to characterize the swirl flow, which is useful for HTSG design based on the prediction results. In addition, because it is designed considering the feasibility of the variation in the operating pressure of the SOEC, unnecessary economic loss can be reduced.


## 4.2. Effectiveness of CO2 Reduction

The HTSG can be adopted as an energy harvesting method that satisfies uniform hydrogen production in the industrial field. To enhance the production of hydrogen in SOEC, stability and durability must be guaranteed, and the steam condition must be less than 0.3 MPa. Therefore, in this study, the HTSG for generating HTLP steam was optimized based on a numerical simulation. If the optimal HTSG geometry derived from the simulation is adopted, HTLP steam can be generated uniformly, thereby enabling a stable supply to the SOEC. In addition, the linkage between the HTSG and SOEC was improved by predicting the hydrogen production according to the parametric analysis via the dynamic simulation.Currently, there are various energy sources that can produce steam, such as solid refuse fuel (SRF), LNG, and coal, and they can be produced in a large capacity. However, to reduce fossil fuels and CO2 emissions, it should be replaced with an eco-friendly heat source by visualizing and analyzing the values of CO2 emissions for each heat source system. The CO2 emission from the heat source according to the hydrogen production method is evaluated by Equations (23)–(25) [38] and illustrated in Figure 16 [39, 40].
(23)FuelheatingvalueMJ=fuelconsumptionNm3×heatgenerationMJNm3,(24)CarbonemissionstC=fuelheatingvalueMJ×CO2conversionfactortC/TJ106,(25)CO2emissionstCO2=carbonemissionstC×4412.Figure 16
Comparison of CO2 emissions during hydrogen production according to steam heat source.When the HTLP steam produced by the SRF presented in this study is utilized for hydrogen production, the CO2 emission is estimated as 9.92 ton-CO2/ton-hydrogen. This is approximately 11.1% greater than high-temperature and high-pressure (HTHP) steam by LNG and 86.8% greater than HTHP steam by coal, confirming that steam conditions affect the hydrogen production significantly. Moreover, it is very encouraging that the lowest CO2 emission was calculated when the energy source for supply was SRF. It is therefore confirmed that this is an efficient method compared to other energy sources, and when hydrogen is produced utilizing SRF, it can be considered as an eco-friendly solution that can harvest cost and resources.Since the system considers the effect on the pressure drop, an optimal design is required in terms of thermal energy. In the future, it will be intended to derive a method for improving heat transfer according to the operation conditions of the HTSG. It will be an important guideline for industrial application to SOEC, where the higher the temperature, the higher the hydrogen production efficiency.


## 5. Conclusions

In this study, an HTSG capable of generating HTLP steam for SOEC supply was developed. The experiments were performed under various operating conditions to consider the pressure drop in the HTSG. The derivation of the variable factors that enhance the pressure drop was achieved via numerical simulation and can be predicted using the developed correlation equation. The following conclusions were drawn from this study.(1)
In the experiment, an experimental correlation for predicting the pressure drop of the HTSG was developed with a maximum error of 11.5%. When the steam flow rate was 35 kg/h and inlet pressure was 300 kPa, the pressure drop was 170 kPa, which was estimated to be the optimal operating condition(2)
The pressure drop according to variable parameters was estimated to be a minimum of 54 kPa (inlet nozzle angle: 25°) and a maximum of 309 kPa (body ratio: 0.63). The amount of hydrogen that can be produced from the optimal HTSG was predicted to be a maximum of 35.7 tons annually(3)
In the HTSG, the simulation results were compared with the experimental ones, and it was concluded that it is a suitable correlation to determine the shape characteristics of the HTSG with a maximum error of 10.3%.(4)
The CO2 emission was estimated as 9.92 ton-CO2/ton-hydrogen under optimal conditions, which can be reduced by up to 86.8% compared to the conventional HTHP steam



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*Source: 9022385-2023-02-09.xml*