Abstract
This study evaluates the efficacy of the mode-K+ control logic for daily load-follow operation (DLFO) in a two-batch LEU+ fuel-loaded APR1400 reactor with an accident tolerant fuel (ATF) cladding. The objective is to assess the performance of the control strategy under various operational conditions including high burnup. The simulation results showcase a robust control for both reactor power and the ASI of the equilibrium core by utilizing a manganese absorber for the partial strength control element assembly (PSCEA) instead of the weak Inconel absorber in the standard design, particularly under high burnup conditions. The simulation employs a two-step method, where cross sections are generated using the SERPENT2 Monte Carlo code, and the whole core 3D simulation is conducted using the KANT in-house diffusion code. The results demonstrate successful power and ASI control within the desired limits throughout the whole operational period.
1. Introduction
As a result of the increase in the contribution of renewable energy resources to total energy generation, nuclear power plants that have been considered as baseload are forced to perform load-follow operation (LFO). The load-follow operation can be classified into the daily LFO (DLFO) and primary or secondary frequency control [1].
Utilities operating nuclear power plants have developed diverse control algorithms customized to specific reactor designs and grid operator requirements. In France, the mode-X control logic has recently been developed, enabling LFO to be performed for 90% of the cycle length with a ramp speed of 5% full-power per minute. This operation mode requires control over both the control element assembly (CEA) and the soluble boron concentration in the primary coolant system [2]. Similarly, the USA has devised control algorithms such as MSHIM and mode-A [3]. In Korea, the mode-K control logic was specifically designed for the Korea Next Generation Reactor (KNGR), serving as the basis for the current APR1400 reactor [4, 5]. However, none of the Korean reactors has received licenses for LFO operations thus far.
Meanwhile, previous studies have explored the feasibility of implementing DLFO in APR1400, showcasing promising results for the potential introduction of mode-K for LFO operations in APR1400 [6]. Other studies have investigated DLFO performance in the APR1400 using advanced burnable absorbers like the centrally shielded burnable absorber (CSBA) [7]. Furthermore, methods such as model predictive control (MPC) have also been explored for the APR1400 reactor [8].
Depending on the electricity grid regulations, the requirements and capacity of the LFO control logic shall be determined. In general, the typical LFO scenarios considered for the daily LFO include power variation from full power to 80%, 50%, and, in some extreme cases, 20% of the full power. The rate of power descending or ascending is another regulatory requirement, but in general, this is a slow process, except in some extreme cases where a fast return to full power is needed.
In addition to the importance of LFO to enhance the economy from the PWRs operation, recent studies have explored the potential of extending the nuclear fuel cycle to optimize nuclear power plant availability [9–14]. This extension would involve increasing the fuel BU while reducing the frequency of scheduled outages. However, with the current fuel assembly designs, it is impossible to increase the cycle length due to large excess reactivity issues. To address this, researchers introduced various types of burnable absorbers (BA) designs as a means of excess reactivity suppression from the early cycle depletion. These studies have investigated the potential BA materials that can be utilized including gadolinia, erbia, and hafnium diboride, as well as the form of BA material. These forms include using pyrex rods with various shapes, wet annular burnable absorber (WABA), integral fuel burnable absorber (IFBA), and even more uniformly distributed BAs such as gadolinia-bearing fuel rods [15]. Table 1 provides a description of different types of commercial BA materials along with their respective advantages and disadvantages.
Currently, the extension of the fuel cycle and improvements in BA materials are closely linked to advancements in nuclear fuel. Given the current nuclear fuel design limitations, the maximum allowable fuel BU is restricted. The introduction of accident tolerant fuel (ATF) can be achieved through enhancements in fuel performance, involving different fuel materials such as UN, UC, and U-Zr alloys, in contrast to the conventional UO2 fuel. Another approach is to use more accident-resistant fuel cladding materials, such as stainless steel (SS), FeCrAl alloy, and nickel-coated zircaloy [16].
This paper presents the introduction and implementation of a modified control logic called mode-K+ in an APR1400 reactor utilizing low-enriched uranium + (LEU+) fuel [17]. The reactor core design, including the ATF cladding and core configuration, will be discussed in Section 2. Section 3 provides a detailed description of the control logic, while section 4 presents the results and discussions.
2. Reactor Core Design
The APR1400 is a large-size PWR with a rated thermal power of 3983 MWth. The standard APR1400 is operated with an 18-month three-batch fuel management scheme [18]. In the modified APR1400, fuel assemblies with an enrichment higher than 5 wt.% are introduced to achieve a 24-month two-batch fuel management scheme. Table 2 provides the major important design parameters used in the LEU+ APR1400 core design.
2.1. Accident Tolerant Fuel (ATF) Clad Design
With a 24-month fuel cycle length, the fuel BU would significantly increase, which in turn requires improving the clad performance to enhance its integrity. Additionally, during LFO, the temperature in the core is varying a lot. To secure the required clad improvements, an ATF clad shall be utilized [19].
The proposed solution to address this issue is by employing the swaging technique to enclose the conventional zircaloy clad with an additional SS316 tube. The process involves the use of a 10 μm thick inner tube and a 30 μm thick outer SS tube, which surround the standard zircaloy clad [20, 21]. These extra layers serve the purpose of protecting the fuel clad from the simultaneous interaction of coolant and fuel material. It is worth noting that the reactivity penalty associated with the added clad layers is estimated to be less than 700 pcm, and there are no noticeable effects on the thermal-hydraulic parameters or the fuel centerline temperature, assuming there is no gap between the cladding layers [17]. By implementing this proposed solution, the fuel-clad performance can be significantly improved without compromising the fuel safety parameters.
2.2. Reactor Core Configuration
In this specific core design, all noncore-related design parameters adhere to the standard APR1400, which includes the steam generator (SG) design, core thermal parameters, and CEA configuration. To illustrate the LEU+ APR1400 equilibrium core configuration, Figure 1 showcases the core loading pattern and shuffling scheme where the centerline is denoted by the red dashed line. In this design, two types of fuel assemblies are utilized, each with different BA loadings aimed at regulating the radial power profile. To achieve the desired radial power distribution, the loading pattern and shuffling scheme were optimized through trial and error, taking into account specific constraints such as shuffling most of the fuel assemblies to nearby positions, situating the fresh fuel assemblies at the core peripheral region, and placing the least BA-loaded fuel assemblies at the peripheral positions. This arrangement enhances neutron leakage and promotes a more uniform radial power distribution with maximum assembly-wise power peaking of 1.4, 1.34, and 1.25 at the beginning, middle, and end of cycle (BOC, MOC, and EOC) [17].
Erbia (Er2O3)-bearing fuel is used by blending 0.5 wt.% of erbia with UO2 fuel in all fuel pins. The inclusion of erbia is crucial to reduce excess reactivity in the core and ensure compatibility with existing front-end nuclear fuel fabrication facilities, possibly eliminating the need for license renewal by reducing the ex-core criticality values to the levels of LEU fuel. The addition of erbia is minimized to minimize the reactivity penalty at the EOC due to an unburned BA and to limit its impact on fuel thermal conductivity since a high fraction of BA admixed with the UO2 fuel is well known to significantly reduce the thermal conductivity of the fuel, which would lead to an increase in the fuel centerline temperature [22]. To control the radial power profile and excess reactivity, fuel pins containing gadolinia are employed. This core design utilizes two types of fuel assemblies, one with 12 gadolinia pins and the other with 16 gadolinia pins. The gadolinia composition consists of 8 wt.% gadolinia (Gd2O3) admixed with 2 wt.% enriched UO2 fuel, matching the composition used in the standard APR1400. The erbia is eliminated from the fuel pins that contain gadolinia since the enrichment is already lower than 5 wt.%. The axial configuration includes a 15 cm BA cutback region at the top and bottom of the core. With a two-batch fuel management scheme, a total of 121 fuel assemblies, including the center fuel assembly, will be discharged per cycle.
Figure 2 shows the CEA pattern of the APR1400 CEA. The configuration includes five full strength control element assembly (FSCEA) regulating banks and three partial strength control element assembly (PSCEA) banks. During operation, the PSCEAs can be grouped in different configurations, such as three or two independent groups [18]. Table 3 provides a summary of the important design parameters for the APR1400 CEA. B4C is used as the absorber material for regulating banks, while weak Inconel-625 absorbers are employed for the standard PSCEAs [23].
The APR1400 steam generator operation strategy is based on the sliding average core coolant temperature. In this strategy, the core inlet coolant temperature remains nearly constant within a specific deadband, while the exit and average coolant temperatures vary depending on the reactor core power. Figure 3 illustrates the core inlet, outlet, and average temperature as a function of the reactor power under the sliding average coolant temperature strategy utilized in the APR1400 reactor. This approach employs the arithmetic average of the core inlet and outlet coolant temperatures to determine the necessary movement of the CEA in order to restore the programmed power level. By adopting this approach, an ideal steam temperature and pressure at the main turbine inlet would be achieved at full power, resulting in a lower hot leg temperature compared to alternative control strategies like constant steam pressure or fixed average temperature.
3. Mode-K+ Control Logic
The control of reactor power in pressurized water reactors (PWRs) involves managing three parameters: soluble boron concentration, reactor coolant temperature, and the axial shape index (ASI). In mode-K+, the soluble boron scenario is assumed to be predetermined by the reactor operator, and boron dilution or boration is used to compensate for xenon concentration changes in the reactor core during power changes. Reactivity insertion or withdrawal during power ramp-up and ramp-down is solely handled by the CEA [24]. This constraint allows for efficient adjustment of the soluble boron concentration, considering only linear concentration variation with time. No boration or dilution occurs before the power ascension or descension is complete, minimizing the volume of liquid radiowaste and simplifying the operation of the chemical and volume control system (CVCS).
The control of reactor coolant temperature provides an indication of the criticality status in the reactor core. If the average coolant temperature is lower than the programmed value determined via the sliding average coolant temperature strategy, it indicates a decrease in reactor power, requiring positive reactivity insertion to increase the power. This is achieved by withdrawing the CEA. Similarly, if the reactor power is higher than the target value, CEA insertion is needed to introduce negative reactivity and decrease power. The degree of temperature deviation from the programmed one can also be utilized to judge how fast the CEA movement shall be to return to normal conditions. Figure 4 depicts the temperature control flag used in mode-K+. According to this control logic, if the temperature mismatch () exceeds the value, a CEA movement based on the sign of takes place at a minimum speed of 0.127 cm/sec. If the absolute value of continues to increase and exceeds 0.8°C, CEA movement with the maximum speed of 1.27 cm/sec. is initiated. However, if the absolute value of becomes less than , the temperature control flag is deactivated, and the normal temperature flag is activated. This way, the temperature deviation determines the direction and speed of the CEA movement [24]. Meanwhile, and values are optimized throughout different simulation scenarios to assure that the inlet coolant temperature variation is always within the allowable deadbands in the technical operation procedures [25].
The third parameter that requires careful control in large-size PWRs is the ASI, defined in the following equation:
where represents the reactor power, is the axial position, and is the active core height. represents the difference in power distribution between the upper and lower halves of the core. Deviations in ASI from a target value, typically determined by the ASI value for unrodded depletion of the equilibrium core, are used to select the CEA for movement, similar to the temperature mismatch. ASI control in large-size PWRs is crucial to guarantee the stability of the axial power profile and prevent xenon oscillation. Without controlling the power generation in each half of the core, the xenon oscillation will result in axial power fluctuation and then an increase of axial power peaking and imbalance of fuel BU, particularly in PWRs with a high BA loading. In general, a larger ASI deviation from the targeted value is allowed at low-power regions when there is a larger thermal margin, while stricter ASI control is usually required in the higher-power operation. Figure 5 illustrates the control flag. In this set point, a cutoff is determined at 90% of full power, and the deadband linearly changes between full power and 90% of full power to provide a smooth change in the deadband value [24]. For example, if the power is less than 90% of the full power and the ASI deviation exceeds 0.045, an ASI control flag is initiated. If it is reduced to less than 0.04, the ASI control flag is deactivated. In the APR1400, the allowable ASI deadband is ±0.27, while the target ASI used for mode-K+ is 0.01.
The CEA movement in mode-K+ is determined based on the temperature and ASI signals. Table 4 provides an explanation of the detailed CEA selection logic in mode-K, where is the core height, is the core top, is the bottom of the core, and indicates the CEA axial position from the top of the core. indicates that the ASI is more bottom-skewed compared to the target ASI while indicates a more top-skewed ASI. The selection criteria are based on physical phenomena. For example, if a CEA is inserted in the top half of the core, the power in the top half is reduced while the power in the bottom half increases, resulting in a more bottom-skewed power profile. Similarly, if a CEA is withdrawn from the top half, the power in the top half increases, causing a top shift in the axial power profile compared to the initial status. The same principles apply to CEA movements in the bottom half. ASI control is applied to determine the CEA selection during load-follow operation.
Another physical phenomenon considered is the change of ASI over the change of reactivity () depending on the location of CEA movement [26]. Figure 6 explains the effect of CEA movement on that ratio. It can be observed that when ASI control is needed without perturbing the reactor power, CEA movement at both ends of the reactor core is more favorable, while CEA movement at the core centerline produces the maximum reactivity effect with a minimum effect on the axial power profile.
In this analysis, the exact core midplane at 190.5 is considered as the boundary between the two halves, which is a safe choice for standard PWRs with typical control rod worth. To perform a load-follow operation in mode-K+, the direction of CEA movement is selected based on the temperature mismatch value and sign. Then, the designated CEA movement is chosen depending on the value and sign. If is within the deadband and temperature control is still needed, the ASI flag is switched to AAS (acceptable ASI), and the CEA selection criteria are determined as shown in Table 5. Since power control takes priority over ASI, if no ASI-favorable movement is found, a CEA with a specific order will always be implemented to return to the target power. In mode-K+, the CEA insertion order is PSCEA 3, 2, and 1, followed by regulating banks R5, R4, and R3, respectively. Additionally, P3 is considered as the leading PSCEA bank, meaning that no PSCEA should be inserted in the core beyond P3.
If the temperature is within the deadband but ASI control is still required, the CEA movement is always chosen to yield a favorable ASI result. Figure 7 illustrates the CEA selection logic when the temperature mismatch flag is normal. In this logic, the direction of movement is determined based on the sign of . If an ASI-favorable movement is found, action is taken. Otherwise, the logic searches for CEA movement in the opposite direction. If that movement leads to a favorable ASI and is smaller than another temperature deadband (chosen to be smaller than the temperature mismatch deadband to ensure it does not affect the reactor power), which in this analysis is set to 0.1°C, then it is executed. However, if no CEA action can achieve the desired ASI condition, the logic is terminated without any action.
The mode-K+ version significantly improves upon the original mode-K by introducing flexibility to accommodate any CEA configuration. In contrast, the original mode-K was specifically designed for a core configuration consisting of two groups of PSCEA. Any changes made to the CEA grouping in the original version necessitated a complete rewrite of the control logic.
Moreover, the control logic in mode-K+ has undergone extensive simplification. For instance, instead of using a two-stage flag for ASI control as in the original version, a single flag is now utilized. Additionally, the CEA selection criteria have been streamlined to eliminate any possibility of incorrect selection or deselection during power control. These modifications enhance the overall functionality and ease of implementation of mode-K+.
4. Load-Follow Operation Simulation and Discussion
4.1. Computational Tool
This analysis was conducted using the conventional two-step method. The cross-sections were obtained from the SERPENT2 continuous energy Monte Carlo code, associated with the ENDF/B-VII.1 library [28]. The whole core calculation was performed using an in-house code called KAIST Advanced Nodal Tachygraphy (KANT) [29]. KANT is a multiphysics nodal code based on NEM-CMFD diffusion and capable of performing coupled T/H (thermal-hydraulic) depletion with a critical boron concentration (CBC) search. In terms of LFO, KANT includes two modules: one for LFO in small modular reactors (SMR) using the mode-Y approach and another for LFO in large-size PWRs that utilize soluble boron and ASI control, known as mode-K+. For the APR1400, users can choose between the original mode-K, developed for two groups of PSCEA, or the modified control logic mode-K+ with a flexible CEA configuration.
In the KANT code, the time-dependent solution of the CMFD-NEM is directly utilized to perform the transient LFO solution, with a time step of less than 10 seconds. Figure 8 illustrates a schematic representation of the calculation procedure in KANT during the transient LFO simulation. To streamline the SG model in the analysis, the difference between the target outlet coolant temperature and the current value is calculated. This difference is then applied to determine the inlet coolant temperature in the next time step. Equation (2) defines the formula for determining the inlet coolant temperature at each time step where represents the inlet coolant temperature at time step , denotes the outlet coolant temperature at time step , and signifies the targeted outlet coolant temperature at time step . This approach enables the simulation to realistically capture the variation of inlet coolant temperature over time. In this simplified SG model, it is assumed that the SG extracts a consistent amount of energy to meet the power demand in the slow DLFO scenarios. However, in order to account for variations, the deviation of the exit coolant temperature from the desired value at a particular time is taken into consideration. This information is then utilized to determine the necessary inlet coolant temperature for the subsequent step. To illustrate, if the actual outlet coolant temperature is lower than the target value, the inlet coolant temperature in the subsequent time step will be reduced by the same deviation value. Conversely, if exceeds , it indicates that more energy is being generated than needed. In such cases, will be increased by the same deviation value to match the power demand.
4.2. Results and Discussions
Figure 9 depicts the equilibrium cycle CBC curve calculated using KANT, showing an initial CBC of approximately 1470 ppm and a cycle BU of 26.3 GWD/MTU. Additionally, Figure 10 illustrates the radial and axial unrodded power profiles for the same equilibrium core [17]. The normalized power distribution in the equilibrium core exhibits a well-balanced pattern, with maximum peaking around the core center. This is a result of the initially high concentration of soluble boron in the core. However, as BU increases, the fuel in the core central region is consumed more than in the upper and lower regions. Consequently, a saddle-shaped axial power distribution appears, becoming more pronounced by the EOC. This leads to a shift in power peaking towards the upper and lower parts of the core. This unrodded axial power shape is of significant importance for the LFO since the unrodded ASI would affect the ASI control during the LFO.
(a)
(b)
In this core design, the observed maximum local power peaking in the single assembly lattice calculation is rather small as it was found in the previous study of around 1.14 in fresh fuel and gradually decreased with BU [17]. Equation (3) demonstrates the method for assessing the moderator temperature coefficient (MTC).
At point , and represent the reactivity and temperature, respectively. The reactivity was evaluated at 600, 575, and 550 K in this analysis, with 575 K as the reference normal temperature. The average MTC value was then considered for the upper and lower branches. The evaluation indicates highly negative values of -22.4, -33.6, and -64.4 pcm/K at the BOC, MOC, and EOC hot-full power (HFP) conditions, respectively. The significantly negative MTC at the EOC is a result of higher BU compared to the standard APR1400. Meanwhile, the strictly negative MTC at BOC is primarily due to the utilization of erbia as the BA material, which possesses a high epithermal absorption cross-section, leading to a more negative MTC. These conditions necessitate modifications to the absorber material of the PSCEA to enhance the control rod’s worth. Table 6 provides a description of the worth of the CEAs that will be employed in the load-follow operation at different BU conditions. The analysis revealed that employing manganese as a burnable absorber increases the worth of PSCEAs to 312 and 309 pcm at the BOC and EOC, respectively, in comparison to the values for Inconel-PSCEA represented in Table 6. This approach minimizes the insertion of strong regulating banks, thereby facilitating better ASI control.
In this analysis, DLFO was conducted in LEU+ APR1400 at the BOC, MOC, and 90% of the cycle length. The first LFO scenario was executed at the BOC condition. In this scenario, the reactor power was reduced from the initial full power to 50% of the full power over a period of 2 hours, followed by a 6-hour operation at the low-power region. Subsequently, a power ramp-up back to full power was carried out within another 2-hour period. The reactor core power was then maintained at full power for the remaining 24 hours. As depicted in Figures 11 and 12, utilizing the Inconel PSCEA resulted in a deep insertion of the regulating banks R5 and R4, leading to a sharp increase in the ASI value during the power ramp-down period. In this scenario, the boron concentration was controlled by initiating a dilution immediately after the power ramp-down completion, which lasted for four hours. The boration process was then performed for another 4 hours. In this analysis, the radial peaking is the peak assembly power in an axially integrated power profile, while the axial power peaking is the maximum axial power obtained from a radially integrated axial power shape. Similarly, the 3D power peaking is the KANT-calculated peak nodal power multiplied by a conservative pin power peaking factor of 1.14.
Figures 13 and 14 illustrate the LFO scenario at the BOC condition using Mn-PSCEA. It is evident that ASI control was effectively carried out with the stronger PSCEA groups, primarily due to the lesser insertion of the regulating bank R4. It can also be found that the power-peaking values are slightly smaller in this case compared to the Inconel-PSCEA case. Consequently, the power profile was more balanced, resulting in a more uniform axial burnup of the nuclear fuel and better fuel utilization. The LFO capability was assessed under MOC conditions using Mn-PSCEA, which involved simulating a 72-hour scenario with varying power levels each day. In this simulation, the power was gradually reduced from full power to 20% within one hour, followed by 6 hours of sustained low power before returning to full power within another hour. On the subsequent day, the power was reduced to 80% of full power within 2 hours. Finally, on the third day, the power was decreased to 50% of full power. Figures 15 and 16 depict the MOC scenario, demonstrating a close resemblance between the ASI, power control, deadbands, and targeted values. During the multiple days of LFO, the scenario was optimized to ensure that all CEAs are almost fully withdrawn at the end of each day. This is necessary to prepare the reactor core for operation on the following day.
At the EOC condition, a DLFO scenario was simulated, similar to the scenario at the BOC. Figures 17 and 18 illustrate the EOC DLFO scenario using Inconel-PSCEA, while Figures 19 and 20 show the scenario using Mn-PSCEA. Performing LFO at such a high BU poses challenges due to the lower boron concentration in the core, which slows down and complicates the dilution process. Furthermore, the highly negative MTC value at EOC requires a higher worth of the CEAs to manage the power deficit. Utilizing the Inconel-PSCEA in this scenario resulted in deeper insertion of the regulating banks and worsened ASI control. The ASI value exceeded the operation limit of -0.27 in the low-power range. However, using the Mn-PSCEA resolved this issue and successfully maintained ASI control as well as much lower power-peaking values.
In some extreme conditions where a larger power variation is required, the second scenario showcases a reactor power change from full power to 20% of the full power within a 2-hour ramp-down period, followed by an increase back to full power after 6 hours of low power operation within another 2-hour ramp-up speed. As expected, the results in Figures 21–24 reveal that with Mn-PSCEA, ASI control has been significantly improved, leading to much better axial power distribution and improved power peaking values. The simulation was conducted at the EOC, which is a more challenging situation compared to the BOC condition. Furthermore, it can be observed that in the low-power region, the ASI remained flat and close to zero.
To assess the applicability of the LFO using mode-K+ during multiple days of operation, simulations were performed at both the BOC and EOC for a repeated scenario. Figures 25 and 26 illustrate the 72-hour LFO scenario at the BOC. The simulation results indicate that the LFO on the third day closely resembles the second-day scenario, including CEA movement, boron scenario, and ASI control. A similar trend was observed at the EOC condition, as depicted in Figures 27 and 28, where the simulation was performed over a period of 4 consecutive days. It is observed that the maximum axial peaking occurred during the second day of power ramp-down, while for the subsequent days, the value remained below 1.4. Simultaneously, the variation in the assembly-wise radial power peaking remained minimal. Furthermore, Figure 29 shows the variation in transient xenon concentration during the same 96-hour DLFO scenario at the EOC. It is worth noting that the boron scenario on the second and third days differs slightly from that of the first day due to lower transient xenon concentration on consecutive days.
5. Conclusions and Future Works
In this study, daily load-follow operation (DLFO) was conducted in LEU+ fuel-loaded APR1400 using accident tolerant fuel (ATF) clad, employing the control logic mode-K+. This control logic allows for a flexible configuration of the partial strength control element assembly (PSCEA), enabling improved reactor performance. Mode-K+ utilizes temperature signals and the axial shape index (ASI) measurement to select control element assemblies (CEAs) that effectively control both reactor power and axial power distribution, while the soluble boron scenario is assumed to be predetermined by the operator to compensate for the xenon concentration change during the DLFO.
The analysis conducted at the beginning of cycle (BOC), middle of cycle (MOC), and at 90% of the cycle burnup (BU) for various DLFO scenarios revealed that using the standard Inconel PSCEA would result in poor ASI control due to the highly negative moderator temperature coefficient (MTC) observed in high BU cores. To address this issue, a slightly stronger absorber material was introduced, namely, the manganese-based PSCEA. The calculations demonstrated that the total worth of all PSCEAs increased from approximately 172 and 182 pcm at the BOC and EOC, respectively, to 312 and 309 pcm. This improvement in the worth of the PSCEAs led to the reduced insertion of strong regulating banks and improved ASI performance. With the utilization of Mn-PSCEA, successful LFO was achieved at the BOC, MOC, and EOC, considering the 100-50-100 power scenario, the 100-20-100 case, and a combination of them. It is worth mentioning that the maximum power peaking factor occurred at the low-power region where a more thermal margin is guaranteed. The soluble boron scenario was chosen to be simplified and linearly changing with time, which would simplify the operation of the reactor core chemical and volume control system (CVCS).
In future research, the applicability of mode-K+ during frequency control will be thoroughly investigated. Frequency control is a crucial operational aspect that aims to compensate for the uncertainty between power production and demand at any given moment. Exploring the effectiveness of mode-K+ in this context will provide valuable insights and contribute to enhancing the overall operational capabilities of the nuclear power plant.
Data Availability
The data used to support the findings of this study are available from the corresponding author upon reasonable request.
Conflicts of Interest
The authors declare that they have no conflicts of interest.
Acknowledgments
This work was supported by the National Research Foundation of Korea (NRF) Grant funded by the Korean Government (MSIT) (RS-2022-00144429 and 2022M2E9A304619011) and BK21 FOUR (Fostering Outstanding Universities for Research) Project No. 4120200313637.