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Article

Cost-Effective Winding Strategy and Experimental Validation of a Real-Scale HTS Field Coil for 10 MW Class Wind Turbine Generators

Future Core Technology R&D Division, Research Institute of Medium and Small Shipbuilding, Busan 46757, Republic of Korea
Energies 2025, 18(18), 4892; https://doi.org/10.3390/en18184892
Submission received: 10 August 2025 / Revised: 1 September 2025 / Accepted: 10 September 2025 / Published: 15 September 2025
(This article belongs to the Section A3: Wind, Wave and Tidal Energy)

Abstract

In this study, real-scale high-temperature superconducting (HTS) field coils for a 10 MW class rotating machine were designed, fabricated, and experimentally evaluated. The aim was to propose a cost-effective winding strategy by combining two types of HTS wires with different angular dependencies of critical current. The 3D FEM simulations were performed to determine the coil layout by considering the magnetic field magnitude and incidence angle. Based on this design, two HTS field coils were fabricated, one wound with two different types of wire and the other with a single wire type. For application to an actual HTS generator, the coil was equipped with an iron core to evaluate its influence on critical current and magnetic field distribution. Experimental results at 77 K showed that the coil combined with two types of HTS wire achieved 112 A without the core and 105 A with the core, while the single-wire coil reached 101 A and 93 A, respectively. The measured results showed good agreement with the simulations, with deviations within 3.7% for the combined-wire coil and 1.9% for the coil equipped with the iron core. These findings indicate that the proposed winding method can maintain high performance while lowering material cost, providing useful guidelines for the design of large-scale HTS rotating machines.

1. Introduction

Global climate change and the increasing urgency of achieving net-zero carbon emissions have significantly accelerated the global transition toward renewable energy sources [1,2,3]. Among these, wind power has emerged as one of the most promising technologies due to its scalability, maturity, and declining cost per megawatt-hour. As countries implement aggressive decarbonization policies, the share of wind energy in global electricity generation is expected to grow steadily over the coming decades [4,5,6,7]. To improve power output and reduce the levelized cost of energy, wind turbine systems have been rapidly increasing in size. The rated capacity of a single wind turbine has grown from the multi-megawatt range to over 10 MW, and projections indicate that 15–20 MW turbines could become commercially viable within the next decade [8,9,10,11,12]. Larger turbines enable higher annual energy production and reduce the number of units required per wind farm, leading to savings in foundations, electrical infrastructure, and overall maintenance. The permanent magnet synchronous generators (PMSGs) face limitations in scaling to these power levels. The gearboxes or direct-drive systems with large-diameter PMSGs substantially increase the nacelle weight and size, which, in turn, escalates installation costs and complicates maintenance. These technical constraints have driven interest in alternative generator technologies capable of delivering high torque and magnetic field strength while minimizing weight [13,14,15,16,17].
High-temperature superconducting (HTS) generators have attracted growing interest as an alternative for large-scale wind power applications to overcome these limitations. HTS generators eliminate the need for permanent magnets by employing field coils capable of producing strong magnetic fields at cryogenic temperatures. HTS generators offer various advantages such as high torque density, reduced nacelle mass, and direct-drive operation without gearboxes [18,19,20,21,22]. When levelized cost of energy is considered, the HTS generators become competitive with or superior to PMSGs beyond certain capacities and offshore [23,24]. Accordingly, to enhance system-level cost effectiveness, recent work has focused on HTS field coil optimization, including winding methods tailored to local magnetic field distributions and mechanical constraints, insulation and impregnation techniques, and placement strategies for HTS wind power generators [25,26].
The authors proposed and designed a Performance Evaluation System (PES) to verify the characteristics of HTS field coils prior to integration into a full generator system [27,28,29]. The PES is designed to replicate the operational environment of one pole pair of the 10 MW HTS generator, including the magnetic field distribution and the torque applied to the HTS field coil. A full-scale HTS field coil was fabricated based on the coil design of the 10 MW generator, and the PES was constructed to test its behavior under equivalent electromagnetic loads. By supplying the same field current and simulated armature current as in the actual generator, the PES enables evaluation of Lorentz forces, magnetic flux density, and mechanical stress on the HTS coil and its support structure. The fabrication of the HTS field coil for the 10 MW class wind power generator is required to enable experimental testing using the PES.
In this paper, real-scale HTS field coils were designed and experimentally evaluated for application in a 10 MW class HTS wind power generator. An HTS coil winding strategy was proposed using two different types of HTS wire with distinct angular dependencies of critical current. The goal was to enhance the placement of HTS wires within the coil based on the local magnetic field distribution, thereby improving performance and reducing manufacturing costs. Three-dimensional FEM simulations and experiments were carried out to verify the effectiveness of the proposed design for HTS coils with and without an iron core.
The angular dependence of the critical current for each wire was first characterized through Ic–B(θ) analysis. Based on this, three HTS field coils were analyzed through simulation, including a coil wound with wire A, a coil wound with wire B, and a coil wound with wire A and wire B. Additionally, two HTS field coils were fabricated. One was wound with wire A and wire B and is referred to as HTS field coil 1, while the other was wound with only wire B and is referred to as HTS field coil 2. The performance of the HTS field coils was evaluated under liquid nitrogen, both with and without an iron core. The critical current, magnetic flux density, and inductance were measured and compared with the simulation results.
The HTS field coil 1 without an iron core demonstrated a critical current that was 10.7% higher than that of HTS field coil 2, while maintaining identical inductance and magnetic field characteristics. The HTS field coils with an iron core exhibited reduced critical current due to changes in magnetic field magnitude and incidence angle caused by core saturation. The experimental results showed good agreement with simulations, with errors under 4% for HTS field coil 1.
This paper is organized as follows. Section 2 introduces the specifications of the 10 MW class HTS wind power generator applied to the HTS coil in this study. The characteristics of the HTS wires and the iron core material are also investigated. Based on these parameters, Section 3 presents 3D FEM simulations of the HTS field coil with and without the iron core to examine the influence of magnetic field distribution on the critical current. Section 4 describes the fabrication of two real-scale HTS field coils, one wound with two types of HTS wires and the other with a single type, and their experimental validation at 77 K. Section 5 presents the discussion of iron core effects, validation of 35 K operation from 77 K data, cost implications, and scalability considerations. These results validate the effectiveness of the wire placement and confirm the reliability of the simulation results. This outcome provides useful guidance for the design of HTS coils in large rotating machines.

2. Description and Design of the HTS Field Coil for 10 MW HTS Wind Power Generator

2.1. Specification of the 10 MW HTS Wind Power Generator

Table 1 shows the specifications of the designed 10 MW HTS wind power generator. The generator is designed to deliver an effective output power of 10.5 MW, considering both mechanical and electrical efficiency losses. It operates at a rated line-to-line voltage of 6.6 kV and a rated armature current of 918 A.
Based on wind resource data from the Ulsan offshore region in Korea, where the wind turbine is intended to be deployed, the average wind speed is approximately 8.5 m/s. Accordingly, the rated wind speed of the turbine is set to 11.3 m/s, and the corresponding rotor blade length is calculated to be 89.1 m. In terms of electromagnetic configuration, the generator employs HTS field coils to produce a high magnetic flux, while conventional copper windings are used in the armature coil.
The HTS field coils are designed to operate at 35 K to ensure sufficient magnetic field generation. At the rated wind speed, the rotor operates at 9.48 rpm, and the corresponding mechanical torque is estimated to be 10.57 MN·m. The generator is designed with 40 poles, and the resulting Lorentz force acting on each HTS field coil is calculated to be approximately 70.6 kN. The 10 MW HTS wind power generator is designed to operate at 35 K using a neon–helium hybrid cooling system, as proposed in [30].

2.2. Specifications of the HTS Wires and HTS Field Coils

Table 2 shows the specifications of the HTS wires used for the HTS field coils. Figure 1 shows the Ic–B(θ) characteristics curve of wire A and B at 77 K. Both wire A and B are GdBCO wires with a thickness of 0.15 mm and a width of 12 mm. The critical currents at 77 K in self-field are 500 A for wire A and 600 A for wire B. Under the operating conditions of the wind generator, which correspond to 35 K and a magnetic field of 2 T, the critical currents are 652 A for wire A and 375 A for wire B. Wire A exhibits superior performance in high magnetic fields. However, its cost is approximately 40~50% higher than that of wire B. The HTS wires have different critical current characteristics for each manufacturer due to the introduction of an artificial pinning center for improving their magnetic field performance. The incidence angles of the applied magnetic field of the minimum critical current of wire A and B are 120° and 0°, respectively. The HTS field coils for a 10 MW HTS wind generator are designed with a 40% current margin. The HTS field coil wound with two different types of HTS wire is designed to have a 40% current margin with the same specifications as the actual wind power generator. Table 3 shows the specifications of the HTS field coil and the iron core. Each coil adopts a racetrack-type winding with four stacked layers, each containing 310 turns. The inner and outer radii of the coil are 125 mm and 202.5 mm, respectively. The effective length of the HTS field coil is 700 mm. An iron core assembly was placed at the center, top, and bottom of the HTS field coil to increase the magnetic flux density. The width of the center iron core is 200 mm, and the width of the iron core for the top and bottom is 445 mm. To improve mechanical stability during operation, stainless steel tape was co-wound with the HTS wire.
The total height of the HTS field coil equipped with and without the iron core is 125 mm and 70 mm, respectively. The operating current is set to 221 A at 35 K, determined from the electromagnetic design of the generator. The stycast 2850 FT was used as the impregnation material for both thermal and structural stability. The total length of HTS wire required per pole is approximately 3.0 km. Figure 2 shows the configuration and geometry of the HTS field coil and iron core. The materials of the iron core are S45C and 50PN470. 50PN470 has a thickness of 0.5 mm, which complicates the mechanical integration for configuring the rigid geometry of the top and bottom of the HTS field coil. Therefore, S45C was selected for the top and bottom of the iron core. The hysteresis curves of S45C and 50PN470 are shown in Figure 3. The material of the 50PN470 offers excellent magnetic properties and is frequently used in rotating machines. 50PN470 has higher initial permeability and greater saturation flux density, with saturation occurring near 1.7 T, while S45C saturates at approximately 1.5 T. 50PN470 is suitable for the center of the iron core and HTS field coils because of its magnetic characteristics, which allow efficient flux conduction.

3. Simulation Analysis of the HTS Coil Using Two Types of the HTS Wire

3.1. Estimation of the Critical Current Distribution of HTS Coil with Wire A and B

Figure 4 shows the magnetic field distributions and incidence angle of the applied magnetic field for the HTS coil at 77 K. The FEM-based electromagnetic simulation was conducted to evaluate the operating characteristics of the HTS coil. The distribution of the magnetic field and the corresponding incidence angle with respect to the HTS wire surface were derived under the 77 K condition. These parameters are essential for estimating the local critical current based on the angular dependence of the HTS wires. In the curved section of the HTS field coil, a high magnitude of the magnetic field is formed within the coil volume due to magnetic flux linkage. The incidence angle of the applied magnetic field is defined as the angle between the field direction and the wire surface, and it varies rotationally from the center of the coil cross-section.
Figure 5a,b show the simulation results for the HTS field coil wound with wire A at 77 K and 35 K, respectively. In both cases, the lowest critical current was observed at the first inner turn in the curved section of the coil, where the magnitude of the magnetic field was highest. The estimated critical currents were 108 A at 77 K and 440 A at 35 K. Figure 5c,d present the corresponding results for the HTS field coil wound with wire B. Similar to the case of the HTS field coil wound with wire A, the minimum critical current occurred at the first inner turn of the curved section, where the magnetic field intensity peaked. However, due to the critical current characteristics of wire B, which exhibits lower values in high magnetic field regions, the estimated critical currents were lower compared to the HTS field coil wound with wire A, at 92 A at 77 K and 307 A at 35 K, respectively.
As shown in Figure 6, the HTS field coil wound with wire A and B exhibited the same critical current of 108 A at 77 K as the coil wound only with wire A. The distribution of allowable current within the coil volume differed from that of the HTS field coil wound with only wire A or wire B due to the varying angular performance of each wire type. The HTS field coil wound with wire A and B also exhibited the same critical current of 440 A as the coil wound only with wire A at 35 K. However, the critical current was found to occur at the 240th turn, which corresponds to the region wound with wire B. This combined winding configuration was designed to improve the critical current capacity by aligning the angular performance characteristics of each wire with the local magnetic field conditions. As shown in the critical current distribution in Figure 5, the HTS coil wound with wire A exhibits higher allowable current in the inner and outer regions of the coil volume, because wire A has relatively high allowable current in areas where the magnetic field incidence angle ranges from 0° to 45°. The HTS coil wound with wire B shows a higher allowable current near the central region, where the incidence angle approaches 90°, due to its favorable angular characteristics in that range. Based on these complementary properties, an HTS field coil combining wire A and B was designed, as illustrated in Figure 6 and Table 4. To optimize both performance and cost, a combined winding strategy was adopted. Wire A was assigned to inner and outer turns, where magnitude of the magnetic field and angular sensitivity were highest, while wire B was applied to middle turns. Specifically, the HTS coil was wound with 180 turns of the inner winding and 70 turns of the outer winding of wire A, corresponding to magnetic field incidence angles of approximately 0° to 40°, and 60 turns of the middle winding of wire B, spanning angles close to 90°.
Table 5 presents a comparative analysis of critical current performance and relative wire cost for different HTS field coil winding strategies. For the HTS field coil wound entirely with wire A, the critical currents are 108 A at 77 K and 440 A at 35 K, with the cost set as the baseline. For the HTS field coil wound with only wire B, the critical currents are reduced to 92 A at 77 K and 307 A at 35 K, but the cost is significantly lower at 60% of that of wire A. In the HTS field coil wound with a combination of wire A and B, the critical current is the same as that of the HTS field coil wound entirely with wire A, 108 A at 77 K and 440 A at 35 K, while the cost is reduced to 85%. These results show that the combined winding approach effectively balances performance and economic feasibility, offering a practical solution for large-scale HTS coil applications.

3.2. Estimation of the Critical Current Distribution of HTS Coil with Iron Core

Figure 7 shows the magnetic field distributions and incidence angle of the applied magnetic field for the HTS coil equipped with an iron core at 77 K. An iron core was applied to the HTS field coil to achieve the target output of a 10 MW class generator, which requires a strong magnetic field. The maximum magnetic flux density at the surface of the iron core was observed to be approximately 2.6 T, exceeding the saturation levels of both materials. Beyond the saturation point of the iron core, the core behaves nearly like air as its relative permeability approaches unity, and the magnetic flux density increases only gradually with additional ampere-turns. Additionally, due to the geometric structure of the iron core, the incidence angle of the magnetic field acting on the HTS coil was found to differ compared to the HTS field coil without iron core type.
Figure 8 shows the critical current distributions of HTS field coils equipped with an iron core. Figure 8a shows the coil wound with wire A, and Figure 8b shows the coil wound with wire B. Figure 8c illustrates the coil wound with a combination of wire A and B, as described in the previous section. The critical current of the HTS field coil equipped with an iron core and wound entirely with wire A was estimated to be 103 A, occurring at the 310th turn. In the case of the HTS field coil wound with wire B, the critical current was estimated to be 87 A, with the identified position at the first turn. For the HTS field coil wound with a combination of wire A and B, the critical current was also estimated to be 103 A, identical to that of the coil wound entirely with wire A, and it occurred at the 310th turn.
These results indicate that the presence of an iron core reduces the overall critical current of the coil due to variations in magnetic field strength and incidence angle. Additionally, the 3D FEM simulation results revealed that the critical current distribution of the HTS field coil equipped with iron core differs significantly from that of the HTS field coil without iron core configuration.

4. Experiments on HTS Field Coils for a 10 MW Class HTS Wind Generator

Simulation results showed that the HTS field coil wound with wire A and the HTS field coil wound with a combination of wire A and B had similar critical currents, whereas the HTS field coil wound with wire A and the HTS field coil wound with wire B exhibited different critical currents due to variations in magnetic field strength and incidence angle. In this study, two types of HTS field coils were fabricated for experimental evaluation. One of the HTS field coils was wound with a combination of wire A and B, referred to as HTS field coil 1, and the other was wound with only wire B, referred to as HTS field coil 2. To verify the improvement of the critical current for HTS field coils with the same geometry, two types of HTS field coils were manufactured and tested. The physical and electrical specifications of both coils were kept identical, and the winding configurations were implemented in accordance with the simulation models described in the Section 3.

4.1. Manufacturing of the HTS Field Coil

Figure 9a shows preparation for winding. On the left, incoming inspection of the HTS wire was performed, including checks of width, thickness, and available length. The required length for the coil was calculated, and the cassette separation and winding sequence were planned accordingly. On the right, the prearranged HTS wire cassettes and the coil bobbin were installed on the winding machine. The bobbin was fixed on a winding jig, and a stainless steel (SUS) tape was mounted on the opposite spindle to enable simultaneous co-winding for the metal insulation concept.
Figure 9b shows the HTS coil after completing 310 turns for the first layer. During winding, tension was controlled with tension controllers, set to about 5 kg for the HTS wire and about 4 kg for the SUS tape to ensure tight and uniform co-winding. The field coil for one pole consists of four layers in total.
Figure 9c shows the assembly of the wound coils. Two single-layer coils were connected to form a two-layer subassembly, and two such subassemblies were then combined to obtain the final four-layer HTS field coil.
Figure 9d shows the final assembly of the HTS field coil with the iron core. The core uses 50PN470 in the central section and S45C plates on the top and bottom. The assembled cross-section matches the 3D CAD model shown in Figure 3.

4.2. Test of the HTS Field Coils in LN2

Figure 10 presents the test results of HTS field coil 1 and HTS field coil 2 without an iron core under liquid nitrogen cooling at 77 K. The critical current was determined from the sum of the inductive voltage generated by the coil inductance and the resistive voltage corresponding to the 1 μV/cm criterion. Given the total conductor length of 3 km for a single coil and a current ramp rate of 0.2 A/s, the inductive voltage was calculated to be 309 mV and the resistive voltage 300 mV, resulting in a critical voltage threshold of 609 mV. Under these conditions, the measured critical currents were 111.9 A for HTS field coil 1 and 101.0 A for HTS field coil 2. The magnetic field profiles of the two coils were nearly identical, indicating that no significant current bypass occurred between turns. Both coils generated a magnetic flux density of 0.30 T at the coil center when the operating current reached 100 A, confirming uniform current distribution within each winding.

4.3. Test of the HTS Field Coils Equipped with the Iron Core in LN2

Figure 11 shows the test results of HTS field coil 1 and HTS field coil 2 equipped with an iron core under liquid nitrogen cooling at 77 K. The presence of the iron core caused the inductance to vary significantly depending on the saturation state of the core material. At low currents, before magnetic saturation, the peak inductance was measured to be 8.91 H, corresponding to a peak inductive voltage of 891 mV when the current ramp rate was set to 0.1 A/s. When the iron core reached magnetic saturation, the inductance of the HTS coil decreased to 2.24 H, and the saturated inductive voltage was measured to be 224 mV. The critical current was determined from the sum of the saturated inductive voltage, which was 224 mV, and the resistive voltage calculated using the 1 μV/cm criterion, which was 300 mV, resulting in a total critical voltage threshold of 524 mV. Under these conditions, the measured critical currents were 105 A for HTS field coil 1 and 93 A for HTS field coil 2. The reduction in critical current compared to the air–core configuration is attributed to changes in both the magnitude and the incidence angle of the magnetic field caused by the presence of the iron core.
Table 6 compares the simulated and measured critical currents of the HTS field coils, both with and without an iron core. For the configuration without an iron core, the measured critical currents were 112 A for HTS field coil 1 and 101 A for HTS field coil 2, corresponding to deviations of 3.7% and 8.9% from the simulated values, respectively. With the iron core, the measured critical currents were 105 A for HTS field coil 1 and 93 A for HTS field coil 2, yielding deviations of 1.9% and 6.5%. The manufacturer provided wire B for critical current between 600 and 650 A at 77 K in self-field, which is slightly higher than the nominal specification. In our study, critical current sample tests and visual inspections were performed to verify quality, and wire segments with higher performance were placed in more field-sensitive regions. These process discrepancies result in a greater measured critical current than that predicted by the simulation model.

5. Discussion

This work investigated a cost-effective winding strategy for a real-scale HTS field coil by combining two types of wires that exhibit different angular dependencies of critical current. The FEM analysis mapped the local magnetic field magnitude and incidence angle along the racetrack geometry and guided the placement of the two wires. Wire A shows higher critical current near an incidence angle of about 30 degrees and wire B performs better near 90 degrees. Alternating the two wires turn by turn in the appropriate regions equalized the local current-carrying capability and produced a coil in which both wires operated close to their favorable angular ranges. The good agreement between simulation and measurements indicates that the angle-aware placement is an effective design lever rather than an ad hoc choice.
The iron core reduced the measured critical current compared with the air–core configuration. This behavior is consistent with changes in both field magnitude and incidence angle near the conductor when the core is introduced. The reduction is therefore interpreted as a magnetic design effect that follows the critical current as a function of field and angle characteristics rather than as a sign of degraded manufacturing quality.
Once the 77 K critical-current measurements are found to agree with simulations based on the Ic–B(θ) curve, the critical current at 35 K likewise matches the simulated values with good fidelity [31]. In HTS rotating machinery, the operating current at 35 K is typically set to about 60% of the critical current, providing a conservative margin for reliable operation [32]. This validation and margin method are widely adopted in HTS coil research and manufacturing. Therefore, measurements at 77 K are sufficient to predict stable operation under the target 35 K condition.
From a manufacturability standpoint, the proposed winding strategy that uses two types of HTS wire maintains the required coil performance while lowering material cost at the coil level. Because HTS wire accounts for roughly 30% of the total generator cost [13], savings achieved in the coil propagate to the system. For the design with 40 poles considered here, and including jointing and soldering between coils, the two-wire allocation is expected to reduce the overall generator cost by about 5%. These results indicate that the cost advantage demonstrated in the prototype coils scales to full generator production. The methodology appears transferable beyond the 10 MW class. As the number of poles increases, the Lorentz force per coil generally decreases, which relaxes mechanical demands on each coil and its support. Nevertheless, scaling to larger ratings requires careful attention to support thickness, thermal conduction through the support path, and potential changes in magnetic loading. The angle-aware mapping and turn-by-turn allocation remain applicable, but the optimal partition between the two wires can shift with geometry and operating temperature.
Overall, the results support the use of angle-aware winding with two wire types as a practical route to maintain performance while improving cost efficiency.

6. Conclusions

In this study, real-scale HTS field coils for a 10 MW class HTS rotating machine were designed, fabricated, and experimentally evaluated. Two types of HTS wires, designated as wire A and wire B and each having different angular dependencies of critical current, were analyzed considering the magnitude and incidence angles of the applied magnetic field through 3D FEM. The simulation results indicated that, by considering the angular dependence of critical current for each manufacturer in the coil design, it was possible to reduce material cost while designing an HTS coil that maintains the same critical current.
Based on these results, two HTS field coils were manufactured for experimental testing. A comparison between simulation and experimental results confirmed good agreement, with errors within 3.7% for HTS field coil 1 without an iron core and 1.9% with an iron core. The relatively larger discrepancies for HTS field coil 2, 8.9% without an iron core and 6.5% with an iron core, were attributed to the processed HTS wire exhibiting slightly better performance than the specifications provided by the manufacturer.
In future work, we will use the PES to conduct electrical, mechanical, and thermal assessments under conditions representative of a wind generator environment, thereby verifying the applicability of the developed HTS field coil to wind turbine generators.
The results of this study validate the proposed conductor arrangement strategy and confirm the reliability of the simulation approach for predicting coil performance under various configurations. These findings provide practical design guidelines for optimizing both performance and cost in large-scale HTS field coils, contributing to the development of next-generation MW class superconducting rotating machines.

Funding

This work was supported by the Technology Innovation Program (Development of a marine electric propulsion system based on a 250 kW water-cooled axial flux motor, 00456966) funded by the Ministry of Trade, Industry & Energy (MOTIE, Korea).

Institutional Review Board Statement

Not applicable.

Informed Consent Statement

Not applicable.

Data Availability Statement

Data is contained within the article.

Conflicts of Interest

The authors declare no conflicts of interest.

References

  1. Nguyen, V.G.; Sirohi, R.; Tran, M.H.; Truong, T.H.; Duong, M.T.; Pham, M.T.; Cao, D.N. Renewable energy role in low-carbon economy and net-zero goal: Perspectives and prospects. Energy Environ. 2025, 36, 2248–2287. [Google Scholar] [CrossRef]
  2. Gul, E.; Baldinelli, G.; Bartocci, P.; Shamim, T.; Domenighini, P.; Cotana, F.; Wang, J.; Fantozzi, F.; Bianchi, F. Transition toward net zero emissions-Integration and optimization of renewable energy sources: Solar, hydro, and biomass with the local grid station in central Italy. Renew. Energy 2023, 207, 672–686. [Google Scholar] [CrossRef]
  3. Lau, H.C.; Tsai, S.C. Global Decarbonization: Current Status and What It Will Take to Achieve Net Zero by 2050. Energies 2023, 16, 7800. [Google Scholar] [CrossRef]
  4. Summerfield-Ryan, O.; Park, S. The power of wind: The global wind energy industry’s successes and failures. Ecol. Econ. 2023, 210, 107841. [Google Scholar] [CrossRef]
  5. Global Wind Energy Council (GWEC). Global Wind Report 2021; Global Wind Energy Council: Brussels, Belgium, 2021; Volume 80. [Google Scholar]
  6. Sayed, E.T.; Olabi, A.G.; Alami, A.H.; Radwan, A.; Mdallal, A.; Rezk, A.; Abdelkareem, M.A. Renewable Energy and Energy Storage Systems. Energies 2023, 16, 1415. [Google Scholar] [CrossRef]
  7. Dorrell, J.; Lee, K. The Cost of Wind: Negative Economic Effects of Global Wind Energy Development. Energies 2020, 13, 3667. [Google Scholar] [CrossRef]
  8. Bilgili, M.; Alphan, H. Global growth in offshore wind turbine technology. Clean Technol. Environ. Policy 2022, 24, 2215–2227. [Google Scholar] [CrossRef]
  9. Bošnjaković, M.; Katinić, M.; Santa, R.; Marić, D. Wind Turbine Technology Trends. Appl. Sci. 2022, 12, 8653. [Google Scholar] [CrossRef]
  10. IRENA. Future of Wind Deployment, Investment, Technology, Grid Integration and Socio-Economic Aspects. Available online: https://www.irena.org (accessed on 20 July 2022).
  11. Sharma, V.; Sharma, S.; Sharma, G. Recent development in the field of wind turbine. Mater. Today Proc. 2022, 64, 1512–1520. [Google Scholar] [CrossRef]
  12. Ali, S.; Park, H.; Lee, D. Investigating the Structural and Power Performance of a 15 MW Class Wind Energy Generation System under Experimental Wind and Marine Loading. J. Mar. Sci. Eng. 2024, 12, 1485. [Google Scholar] [CrossRef]
  13. Jung, G.E.; Sung, H.J.; Dinh, M.C.; Park, M.; Shin, H. A Comparative analysis of economics of PMSG and SCSG floating offshore wind farms. Energies 2021, 14, 1386. [Google Scholar] [CrossRef]
  14. Dias da Silva, M. Design and Optimization of Spoke Type Permanent Magnet Synchronous Machines: A Rare-Earth Element Free Solution For Electromobility. Ph.D. Thesis, Acta Universitatis Upsaliensis, Uppsala, Sweden, 2025. [Google Scholar]
  15. Amato, A.; Becci, A.; Bollero, A.; Cerrillo-Gonzalez, M.D.M.; Cuesta-Lopez, S.; Ener, S.; Dirba, I.; Gutfleisch, O.; Innocenzi, V.; Montes, M.; et al. Life cycle assessment of rare earth elements-free permanent magnet alternatives: Sintered ferrite and Mn–Al–C. ACS Sustain. Chem. Eng. 2023, 11, 13374–13386. [Google Scholar] [CrossRef]
  16. Dayo, S.A.; Memon, A.; Memon, Z.A.; Jumani, T.A.; Abbas, G.; Othmen, S.; Yousef, A.; Wijaya, A.A. A new approach for improving dynamic fault ride through capability of gridctied based wind turbines. Sci. Rep. 2025, 15, 6144. [Google Scholar] [CrossRef]
  17. Wagh, C.; Boukhenfouf, J.; Colas, F.; Rouco, L.; Guillaud, X. Non-linear interaction between synchronous generator and GFM controlled wind turbines–Inertial effect enhancement and oscillations mitigation. Wind. Energy Sci. Discuss. 2025, 2025, 1–23. [Google Scholar]
  18. Liu, D.; Salmi, T.; Deng, F.; Ye, C. Characteristics of the superconducting field winding of an HTS wind turbine generator during a short circuit fault. IEEE Trans. Appl. Supercond. 2022, 32, 5200606. [Google Scholar] [CrossRef]
  19. Köster, R.; Binder, A. Multi-objective optimization of a direct-drive wind turbine generator with HTS excitation winding. IEEE Trans. Appl. Supercond. 2022, 32, 5200508. [Google Scholar] [CrossRef]
  20. Prajzendanc, P.; Kreischer, C. A Review of New Technologies in the Design and Application of Wind Turbine Generators. Energies 2025, 18, 4082. [Google Scholar] [CrossRef]
  21. Kolchanova, I.; Poltavets, V. Superconducting generators for wind turbines. In Proceedings of the 2021 International Conference on Electrotechnical Complexes and Systems (ICOECS), Ufa, Russia, 16–18 November 2021; IEEE: Piscataway, NJ, USA, 2021; pp. 529–533. [Google Scholar]
  22. Zhang, C.; Shen, L.; Wu, X.; Shan, F.; Chen, L.; Liu, S.; Zheng, Z.; Zhu, L.; Wang, J.; Zhai, Y. Review of Offshore Superconducting Wind Power Generation for Hydrogen Production. Energies 2025, 18, 1889. [Google Scholar] [CrossRef]
  23. Hoang, T.-K.; Queval, L.; Vido, L.; Nguyen, D.-Q. Levelized Cost of energy comparison between permanent magnet and superconducting wind generators for various nominal power. IEEE Trans. Appl. Supercond. 2022, 32, 5202606. [Google Scholar] [CrossRef]
  24. Liu, D.; Polinder, H.; Abrahamsen, A.B.; Wang, X.; Ferreira, J.A. Comparison of superconducting generators and permanent magnet generators for 10-MW direct-drive wind turbines. In Proceedings of the 2016 19th International Conference on Electrical Machines and Systems (ICEMS), Chiba, Japan, 13–16 November 2016; IEEE: Piscataway, NJ, USA, 2016. [Google Scholar]
  25. Xu, Y.; An, L.-T.; Jia, X.-P.; Jia, B.-Z.; Maki, N. Optimization study of the main parameters of different types of wind turbine generators. Supercond. Sci. Technol. 2022, 35, 035007. [Google Scholar] [CrossRef]
  26. Köster, R.; Binder, A. Optimum 7 MW HTS direct-drive wind turbine synchronous generator designs with different rotor and stator iron topologies. Elektrotech. Informationstech. 2023, 140, 324–337. [Google Scholar] [CrossRef]
  27. Kim, C.; Sung, H.-J.; Go, B.-S.; Sim, K.; Nam, G.D.; Kim, S.; Park, M. Design, fabrication, and testing of a full-scale HTS coil for a 10 MW HTS wind power generator. IEEE Trans. Appl. Supercond. 2021, 31, 4900705. [Google Scholar] [CrossRef]
  28. Kim, C.; Sung, H.J.; Go, B.S.; Nam, G.D.; Kim, S.; Park, M. Design and property analysis of a performance evaluation system for HTS wind power generators. IEEE Trans. Appl. Supercond. 2020, 30, 5202805. [Google Scholar] [CrossRef]
  29. Sung, H.J.; Go, B.S.; Park, M. A performance evaluation system of an HTS pole for large-scale HTS wind power generators. IEEE Trans. Appl. Supercond. 2019, 29, 5203905. [Google Scholar] [CrossRef]
  30. Seo, G.; Mun, J.; Kim, D.; Park, M.; Kim, S. Neon-helium hybrid cooling system for a 10 MW class superconducting wind power generator. IEEE Trans. Appl. Supercond. 2021, 31, 5202205. [Google Scholar] [CrossRef]
  31. Senatore, C.; Barth, C.; Bonura, M.; Kulich, M.; Mondonico, G. Field and temperature scaling of the critical current density in commercial REBCO coated conductors. Supercond. Sci. Technol. 2015, 29, 014002. [Google Scholar] [CrossRef]
  32. Bo, K.; Chen, J.; Jiang, Y.; Wang, D. Numerical analysis of 40 MW HTS motor electromagnetic characteristics for ship electric propulsion. Sci. Rep. 2023, 13, 20261. [Google Scholar] [CrossRef]
Figure 1. Ic–B(θ) curve of (a) wire A and (b) wire B at 77 K.
Figure 1. Ic–B(θ) curve of (a) wire A and (b) wire B at 77 K.
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Figure 2. Composition of the HTS field coil: (a) real-scale HTS field coil with bobbin, (b) cross-section view of the HTS field coil, (c) iron core for the HTS field coil, and (d) cross-section view of the iron core.
Figure 2. Composition of the HTS field coil: (a) real-scale HTS field coil with bobbin, (b) cross-section view of the HTS field coil, (c) iron core for the HTS field coil, and (d) cross-section view of the iron core.
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Figure 3. Core material specifications including hysteresis curve of S45C and 50PN470.
Figure 3. Core material specifications including hysteresis curve of S45C and 50PN470.
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Figure 4. (a) Magnetic field distributions and (b) incidence angle of the HTS coil at 77 K.
Figure 4. (a) Magnetic field distributions and (b) incidence angle of the HTS coil at 77 K.
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Figure 5. Critical current distribution of the HTS field coils: (a) HTS coil wound with wire A at an operating temperature of 77 K, (b) HTS coil wound with wire A at an operating temperature of 35 K, (c) HTS coil wound with wire B at an operating temperature of 77 K, and (d) HTS coil wound with wire B at an operating temperature of 35 K.
Figure 5. Critical current distribution of the HTS field coils: (a) HTS coil wound with wire A at an operating temperature of 77 K, (b) HTS coil wound with wire A at an operating temperature of 35 K, (c) HTS coil wound with wire B at an operating temperature of 77 K, and (d) HTS coil wound with wire B at an operating temperature of 35 K.
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Figure 6. (a) Configuration and critical current distribution of the HTS field coil wound with wire A and B at (b) an operating temperature of 77 K and (c) an operating temperature of 35 K.
Figure 6. (a) Configuration and critical current distribution of the HTS field coil wound with wire A and B at (b) an operating temperature of 77 K and (c) an operating temperature of 35 K.
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Figure 7. (a) Magnetic field distributions and (b) incidence angle of the HTS field coil with iron core at 77 K.
Figure 7. (a) Magnetic field distributions and (b) incidence angle of the HTS field coil with iron core at 77 K.
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Figure 8. Critical current distribution of the HTS field coils equipped with an iron core: (a) HTS coil wound with wire A, (b) HTS coil wound with wire B, and (c) HTS coil wound with wire A and B at an operating temperature of 77 K.
Figure 8. Critical current distribution of the HTS field coils equipped with an iron core: (a) HTS coil wound with wire A, (b) HTS coil wound with wire B, and (c) HTS coil wound with wire A and B at an operating temperature of 77 K.
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Figure 9. Fabrication, winding, and assembly of the HTS field coil: (a) preparation for winding of the HTS field coil, (b) HTS field coil winding setup and coil after 310 turns, (c) fabricated and assembled HTS field coil after winding, and (d) integration of HTS field coil and iron-core.
Figure 9. Fabrication, winding, and assembly of the HTS field coil: (a) preparation for winding of the HTS field coil, (b) HTS field coil winding setup and coil after 310 turns, (c) fabricated and assembled HTS field coil after winding, and (d) integration of HTS field coil and iron-core.
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Figure 10. Test results of the HTS field coils without iron core under LN2 cooling.
Figure 10. Test results of the HTS field coils without iron core under LN2 cooling.
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Figure 11. Test results of the HTS field coils equipped with iron core under the LN2 cooling.
Figure 11. Test results of the HTS field coils equipped with iron core under the LN2 cooling.
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Table 1. Specifications of the 10 MW HTS wind power generator.
Table 1. Specifications of the 10 MW HTS wind power generator.
ItemsValue
Rated output power10.5 MW
Rated L-L voltage6.6 kV
Rated armature current 918 A
Rotating speed9.48 rpm
Rated torque10.57 MN·m
Number of poles40
Type of the rotorHTS field coil
Type of the statorCopper winding coil
Operating temperature of the HTS field coil35 K
Cut-in wind speed5 m/s
Cut-out wind speed25 m/s
Rated wind speed11.3 m/s
Length of rotor blades89.1 m
Average wind speed of the site of installation8.5 m/s
Table 2. Specification of the HTS wires.
Table 2. Specification of the HTS wires.
ItemsWire AWire B
Type of HTS wireGdBaCuO (GdBCO)GdBCO
Thickness of HTS wire0.15 mm0.15 mm
Width of HTS wire12 mm12 mm
Critical current @77 K, self-field500 A600 A
Critical current @35 K, 2 T652 A375 A
Critical bend radius60 mm35 mm
Critical tensile strength>300 MPa>500 MPa
Table 3. Specifications of the HTS field coil and iron core.
Table 3. Specifications of the HTS field coil and iron core.
ItemsValue
Number of layers4
Turns of field coil310
Inner radius of field coil 125 mm
Outer radius of field coil 202.5 mm
Effective length of field coil700 mm
Width of center iron-core200 mm
Width of top and bottom iron-core445 mm
Co-winding materialStainless steel tape
Thickness of stainless-steel tape0.1 mm
Height of field coil70 mm
Height of field coil with iron-core125 mm
Operating current at 35 K221 A
ImpregnationStycast 2850 FT
Total length of a 1 pole HTS wire3.0 km
Materials of the salient poleS45C and 50PN470
Table 4. Number of turns for each winding section of the HTS field coil.
Table 4. Number of turns for each winding section of the HTS field coil.
Num.ItemsValue
1Inner winding180 turns
2Middle winding60 turns
3Outer winding70 turns
Table 5. Comparison of critical current and relative cost for each type of winding.
Table 5. Comparison of critical current and relative cost for each type of winding.
ItemsCritical Current (77 K)Critical Current (35 K)Cost Ratio
Wire A108 A440 A1
Wire B92 A307 A0.6
Combined two wires108 A440 A0.85
Table 6. Comparison of critical current of the simulation and experiment according to the iron core.
Table 6. Comparison of critical current of the simulation and experiment according to the iron core.
ItemsCritical Current
(Simulation)
Critical Current
(Experiment)
Error
(%)
HTS field coil 1108 A112 A3.7%
HTS field coil 292 A101 A8.9%
HTS field coil 1 with iron core103 A105 A1.9%
HTS field coil 2 with iron core87 A93 A6.5%
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Kim, C. Cost-Effective Winding Strategy and Experimental Validation of a Real-Scale HTS Field Coil for 10 MW Class Wind Turbine Generators. Energies 2025, 18, 4892. https://doi.org/10.3390/en18184892

AMA Style

Kim C. Cost-Effective Winding Strategy and Experimental Validation of a Real-Scale HTS Field Coil for 10 MW Class Wind Turbine Generators. Energies. 2025; 18(18):4892. https://doi.org/10.3390/en18184892

Chicago/Turabian Style

Kim, Changhyun. 2025. "Cost-Effective Winding Strategy and Experimental Validation of a Real-Scale HTS Field Coil for 10 MW Class Wind Turbine Generators" Energies 18, no. 18: 4892. https://doi.org/10.3390/en18184892

APA Style

Kim, C. (2025). Cost-Effective Winding Strategy and Experimental Validation of a Real-Scale HTS Field Coil for 10 MW Class Wind Turbine Generators. Energies, 18(18), 4892. https://doi.org/10.3390/en18184892

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