Mass reduction of superconducting power generators for large wind turbines
2019; Institution of Engineering and Technology; Volume: 2019; Issue: 17 Linguagem: Inglês
10.1049/joe.2018.8145
ISSN2051-3305
AutoresKevin Kails, Quan Li, Markus Mueller,
Tópico(s)HVDC Systems and Fault Protection
ResumoThe Journal of EngineeringVolume 2019, Issue 17 p. 3972-3975 The 9th International Conference on Power Electronics, Machines and Drives (PEMD 2018)Open Access Mass reduction of superconducting power generators for large wind turbines Kevin Kails, Kevin Kails orcid.org/0000-0002-3418-5977 School of Engineering, Institute for Energy Systems, University of Edinburgh, Edinburgh, EH9 3JL UKSearch for more papers by this authorQuan Li, Corresponding Author Quan Li quan.li@ed.ac.uk School of Engineering, Institute for Energy Systems, University of Edinburgh, Edinburgh, EH9 3JL UKSearch for more papers by this authorMarkus Mueller, Markus Mueller School of Engineering, Institute for Energy Systems, University of Edinburgh, Edinburgh, EH9 3JL UKSearch for more papers by this author Kevin Kails, Kevin Kails orcid.org/0000-0002-3418-5977 School of Engineering, Institute for Energy Systems, University of Edinburgh, Edinburgh, EH9 3JL UKSearch for more papers by this authorQuan Li, Corresponding Author Quan Li quan.li@ed.ac.uk School of Engineering, Institute for Energy Systems, University of Edinburgh, Edinburgh, EH9 3JL UKSearch for more papers by this authorMarkus Mueller, Markus Mueller School of Engineering, Institute for Energy Systems, University of Edinburgh, Edinburgh, EH9 3JL UKSearch for more papers by this author First published: 26 April 2019 https://doi.org/10.1049/joe.2018.8145Citations: 5AboutSectionsPDF ToolsRequest permissionExport citationAdd to favoritesTrack citation ShareShare Give accessShare full text accessShare full-text accessPlease review our Terms and Conditions of Use and check box below to share full-text version of article.I have read and accept the Wiley Online Library Terms and Conditions of UseShareable LinkUse the link below to share a full-text version of this article with your friends and colleagues. Learn more.Copy URL Share a linkShare onFacebookTwitterLinkedInRedditWechat Abstract High temperature superconductors (HTS) enable very compact electric machines with high power density. Here, an HTS power generator based on authors' previous double claw pole design is studied and improved. The original design of the double claw pole machine features a stationary field core to provide support and access to the superconducting field winding and its cryostat. It does not serve any other purpose. However, a homopolar field passes through it. The improvement was made to replace the field core by iron-cored copper coils, creating an inner stator within the generator to increase the electric loading. Through this step, the power output of the generator was increased from 10 to 11.6 MW, increasing the power density from 54.28 to 63.54 W/kg while maintaining the same outer diameter. 1 Objectives and context One of the biggest challenges the wind energy sector faces is to reduce the cost of energy. For several decades, now there has been a trend towards higher power-rated wind turbines [1], which help reduce the cost of energy through lower installation and maintenance costs per kilo-watt hour. However, one critical issue with large wind turbines is the tower head mass problem. The tower head becomes extremely heavy for large wind turbines; this leads to a need for more robust foundation towers for support which in turn dramatically increase the cost of the whole system [2]. To solve this issue, a novel lighter topology of power generators based on superconductor technology is required which would enable 10 MW and even higher rated wind turbines. With this objective, the double claw pole machine was designed. It is a 10 MW, 10 rpm superconducting generator. Referring to [3], the main advantages and the concept of the double claw pole machine can be described as follows. It is an axial flux machine. Claw poles are oriented around the superconducting field winding. These are oriented in an alternating fashion creating a North-South pole flux variation across the stator teeth. A stationary field core is required in between the small claw poles in order to support the superconducting winding and its cryostat. The stationary field winding eliminates the need for cryocouplers and brushes operating the superconducting winding with a DC current ensures that there are no AC losses present. The double claw pole machine only uses a small amount of superconducting tape due to the iron-cored design and the loop-shaped field winding. At 30 K, the design only requires 3.4 km of YBCO tape and for 65 K, it requires 13.5 km. In addition, the machine is highly modular, the two stator sides can operate independently from each other in case of a fault. The field winding can be wound into separate loops with separate cryostats, hence even if one of the field windings has a fault, the generator can continue to work under partial load. One disadvantage of the double claw pole machine is that it is heavier than other superconducting machine designs. This is due to the iron-cored structure. Table 1 summarises the original design of the generator. Table 1. Double claw pole machine mass Original generator design total mass 184.2 t outer diameter 6.37 m power output 10 MW rotational speed 10 rpm power density 54.28 W/kg efficiency 94.50% Here, the double claw pole machine is further improved by increasing its power density. The stationary field core in between the claw poles is replaced by copper coils, creating another stator within the generator. With the added stator, the electric loading of the generator is increased and hence a higher power output can be achieved. Fig. 1 shows the transition to the new design of the double claw pole machine. Fig. 1Open in figure viewerPowerPoint Concept of the new design 2 Methods In order to study the double claw pole machine, a reluctance network model was developed in MATLAB [4]. The reluctance network is shown in Fig. 2. Each component of the machine is represented by several reluctances. Each reluctance is calculated according to its geometry and material properties using (1). (1) where l e is the equivalent length through the component, μ 0 is the permeability of air, μ r is the relative permeability of iron, and A is the cross-sectional area of the component. It was decided to use Vacoflux50 for the active mass due to its high saturation limit of 2.28 T at 16 kAm−1 [5]. In addition, all major leakage paths were considered such as leakage flux between the claw poles and zigzag leakage flux between stator teeth. A Newton–Raphson algorithm is used to solve the non-linear flux (Fig. 3). Fig. 2Open in figure viewerPowerPoint Reluctance network of the double claw pole machine Fig. 3Open in figure viewerPowerPoint Algorithm used to solve the non-linear flux equations From the reluctance network, the stator tooth flux is obtained and hence the power output can be calculated. In addition to the reluctance network model, magnetostatic analysis using the simulation software Infolytica MagNet was performed. The finite element analysis (FEA) is used to verify the results from the reluctance network model. Using the stator tooth flux, the power output of the stators is determined using (2). (2) where N COIL is number of coils per stator, E RMS is the induced voltage for one coil, I RMS is the stator current and R COIL is the resistance of a stator coil. E RMS is calculated according to Faraday's law using (3). (3) where N TURN is the number of turns per coil, ϕ PEAK is the peak stator tooth flux and f is the electrical frequency. I RMS is calculated using (4). (4) where J RMS is the current density of copper which is assumed to be 5 × 105 A/m2 and A COIL is the area of a stator coil. 3 Outcomes From the research, it became clear that a novel design approach can be introduced to reduce the mass per kW of the machine. The field core in between the small claw poles is required in order to give support and access to the stationary superconducting field winding, it does not serve any other purpose. However, a homopolar field crosses the field core similar to the homopolar machine described in [6]. Additional power can be extracted from the generator by replacing the field core with copper coils and hence increasing the electric loading of the machine. Fig. 4 shows how the field core was modified to accommodate the new inner stator coils. The forces acting on the coils are balanced hence there is no net force acting in either direction. Similar to the C-GEN concept, the stator coils can be immersed in epoxy resin. They are then held in place by an aluminium band [7]. Similar to the outer stators, the new inner stator also consists of 66 coils. The stator teeth of the three stators are aligned with each other. This ensures that the overlapping area between the small claw poles and the stator teeth is always equal for all four air gaps. Fig. 4Open in figure viewerPowerPoint Concept for the new design The magnetic fields of the inner stator coils will have an effect on the performance of the superconducting tape. However, the impact on the superconducting winding is expected to be relatively small. The critical current density of superconductors mainly depends on the perpendicular flux that penetrates the tape [8]. The field produced by the coils however mainly penetrates the superconducting coils in a parallel manner. Hence, the critical current density of the superconducting tape will not be affected very strongly. Further research into the effect of the armature reaction onto the superconducting field winding is, however, required. Fig. 5 shows the active mass of the new generator design without the structural mass. Fig. 5Open in figure viewerPowerPoint Active mass of the new generator design The reluctance network model was then applied in conjunction with the new design. From the model, the peak flux for the outer stator teeth was found to be 58.1 mWb and for the inner stator teeth 74.7 mWb. Fig. 6 shows the flux variation for the outer and inner stators with the assumption made that the flux variation is sinusoidal. It can be seen that the flux variation for the inner stator is homopolar since it only varies between 0 and ϕ MAX. In addition, the frequency is only half of that of the outer stators since the inner stator only sees half the number of poles. Fig. 6Open in figure viewerPowerPoint Stator tooth flux variation for the outer and inner stator In order to verify the results, a 3D magnetostatic analysis using the simulation software Infolytica MagNet was performed. It is essential to use 3D FEA since the flux path through the machine is three-dimensional. Fig. 7 shows the magnetic flux density distribution when the large claw poles are aligned with the outer stator teeth. Fig. 8 shows the flux density distribution when the small claw poles are aligned with the inner and outer stator teeth. From the FEA results, the outer stator tooth flux density was found to be equal to 1.26 T which is equivalent to a flux of 57.7 mWb and the inner stator tooth flux density was found to be equal to 1.40 T which is equivalent to a flux of 74.2 mWb. It can be seen that there is a very good correlation between the reluctance network model and the FEA results. Fig. 7Open in figure viewerPowerPoint Flux density distribution for large claw poles aligned with outer stator teeth Fig. 8Open in figure viewerPowerPoint Flux density distribution for small claw poles aligned with stator teeth Using (1)–(3), the power output was calculated to be 5 MW for each outer stator and 1.6 MW for the inner stator. Hence, an additional 1.6 MW were added to the original 10 MW. Since the power generated in the inner stator is at a different frequency as compared to the outer stators, a separate power converter is required in order to extract the power. Next to the power density, another important aspect of generators is their efficiency. The efficiency calculation of the original design of the double claw pole machine was described in detail in [3]. To calculate the efficiency, the copper losses, iron losses, cooling power, and air blowers for the armature were considered. The overall efficiency is dominated by the copper losses in the stators. The copper losses are calculated according to the dimensions of the copper wire and the number of turns. The iron losses are calculated using the Vacoflux50 datasheet and using the operating frequency in the machine, which is 7.3 Hz at 10 rpm. The cooling power required depends on the design of the cryostat. Similarly to the original design, the cryostat and superconducting field winding can be split up into several independent sections to increase the modularity of the machine. This is shown in Fig. 9. Fig. 9Open in figure viewerPowerPoint Sectioned cryostat The initial radiation heat loss was calculated to be 545 W; however, using 30 layers of multi-layer insulation (MLI) material, it was reduced down to 17.4 W. The total cooling power required was calculated to be 154.2 W. Table 2 shows the thermal budget split up into its components. Since the superconducting winding was not changed for the new design, the thermal budget remains the same as for the original generator. The required cooling power can be supplied by four 50 W cryocoolers, which allow for an additional safety margin of 25%. The four cryocoolers have a total input power of 24 kW [9]. Table 2. Thermal budget for generator design [3] Thermal budget temperature 65 K gas conduction (at 10−3 Pa) 3.4 W suspension straps 50 W radiation 17.4 W current leads 48.8 W cold-head sleeve 15.6 W eddy current 4 W other 15 W total loss 154.2 W The copper losses were increased since an additional stator was created. The copper losses of the outer stators are equal to 510 kW. The copper losses of the inner stator can be calculated to be 296 kW. Since some of the iron in the field core was eliminated, the iron losses were reduced down to 6.67 kW. Table 3 shows all the losses for the new generator design. As can be seen, the overall losses are dominated by the copper losses in the stators. Table 3. Total losses of the 11.6 MW generator design Total losses copper losses 806 kW iron losses 6.67 kW cryocoolers 24 kW air blowers 50 kW total loss 886.67 kW Hence, with the power output of 11.6 MW, the efficiency of the generator can be calculated to be equal to 92.94%. This is slightly lower than the original generator design, which has an efficiency of 94.5%. The parameters of the new generator design are highlighted in Table 4. The total mass of the generator is lighter since some of the iron in the field core was eliminated. The generator still has the same outer diameter as for the original design. With the new power output and the new mass, the power density of the generator was increased from 54.28 to 63.54 W/kg. Table 4. New generator design summarised New generator design total mass 182 t outer diameter 6.37 m power output 11.6 MW rotational speed 10 rpm power density 63.54 W/kg efficiency 92.94% 4 Conclusion Higher power density generators are needed in order to lower the cost of energy of large direct-drive wind turbines. The double claw pole machine was introduced as a promising option for this objective. However, it was found to be still too heavy to satisfy the future needs of the wind energy industry. To further increase the power density of the double claw pole machine, a new design approach was introduced. The stationary field core was replaced by iron-cored copper coils to create another stator. A reluctance model was developed to verify the design in conjunction with a FEA to verify the results. It was shown that an additional 1.6 MW to the original 10 MW can be gained with the new design, increasing the power density from 54.28 to 63.54 W/kg, while maintaining the same machine diameter. 5 Acknowledgments This work was supported by the Engineering and Physical Sciences Research Council. [Grant number: EPSRC EP/N509644/1] 6 References 1 Wind Europe: 'The European offshore wind industry – key trends and statistics'. Annual report, 2016. Available at http://windeurope.org 2 INNWIND: ' Reference wind turbine report', Work package 1 – Deliverable 1.21, 2013. Available at http://www.innwind.eu 3Keysan O., Mueller M.: 'A modular and cost-effective superconducting generator design for offshore wind turbines', Supercond. Sci. Technol., 2015, 28, (3), p. 034004 4Keysan O., Mueller M.: 'A homopolar HTSG topology for large direct-drive wind turbines', IEEE Trans. Appl. Supercond., 2011, 21, (5) 5Bernholz J. J., Mueller M.: 'Analytical model of superconducting generators for wave energy systems'. 8th IET Int. Conf. on Power electronics, Machines and Drives, Glasgow, 2016 6Ubani O. G., Mueller M. A., Chick J. et al.: 'Analysis of an air-cored axial flux permanent magnet machine with Halbach array'. 8th IET Int. Conf. on Power Electronics, Machines and Drives (PEMD 2016), Glasgow, UK, 2016 7 Vacuumschmelze: ' Vacoflux datasheet', 2017. Available at http://www.vacuumschmelze.com 8Kiss T., Inoue M., Nishimura S. et al. 'Angular dependence of critical current properties in YBCO coated tape under high magnetic field up to 18 T', Physica C, 2001, 378–381, pp. 1113 – 1117 9 Cryomech: ' AL230 with CP950-cryorefrigerator specification sheet', 2013. Available at http://cryomech.com/specificationsheet/AL230_ss.pdf Citing Literature Volume2019, Issue17June 2019Pages 3972-3975 FiguresReferencesRelatedInformation
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