2268 IEEE TRANSACTIONS ON MAGNETICS, VOL 32 NO 4 JULY 1996

Basic Design Considerations for a Medium-Size Superconducting Magnetic Energy Storage System (SMES)

W. Nick, K. Prescher Siemens AG, Mulheim, Germany

Abstract - In a feasibility study a 2 MWh magnet system suitable for utility application (primary control reserve) is analyzed. Basic choices for a realistic SMES design are consi- dered taking into account practical industrial feasibility. Com- pared to a large solenoid advantages are seen for a modular, toroidal configuration with circular coils. To achieve a robust coil pack internal cooling with supercritical helium for an im- pregnated winding is selected. System parameters are derived from a cost optimization approach. Possible conductor concepts, a thin-walled coil casing, and a corresponding refrigeration system based on a cryogenic loss estimate are presented.

I. INTRODUCTION

This paper reports on results of a feasibility study per- formed by Siemens AG in cooperation with German power utilities PreussenElektra AG and RWE Energie AG. The goal is to find a commercially attractive replacement for throttling of the main turbine inlet valve in thermal power stations to supply primary control power according to the requirements of the European grid (UCPTE). Various alter- native methods are considered for this purpose (not to be reported here), combined with a superconducting magnet system as a fast source capable of high power.

derived for a typical 10 GW system [l] , [Z]: maximum useable energy 1.5 MWh maximum power 50 MW approx. 20 "large" pulses per day with a typical profile according to Fig. 1, amplitude 20 to 50 MW approx. 100 "small" sinusoidal pulses per hour with a max. amplitude of 10 MW and a typical period of 12 s high reliability for a useful life of 30 years Comparable medium-size SMES are presently in develop-

ment by Babcock & Wilcox [ 3 ] and by the Bechtel ETM team [4].

From this application the following requirements were

11. BASIC MAGNET CHOICES

In our effort to define a cost effective magnet system design we are utilizing the existing knowledge developed mainly for other applications, e.g. for the magnet system of a fusion reactor. However we are free to choose parameters conveniently to optimize manufacturing cost and technical risk. A. Solenoid versus Toroid

The potential advantages of solenoids are most efficient use of conductor and structural material per stored energy. This optimization leads to high-aspect-ratio solenoids, with a

Manuscript received June 12,1995 This work was financially supported by the German government, Federal

Ministry of Education, Science, Research, and Technology (index 0329574). The authors are responsible for the contents of this publication.

I:h 20 Discharg

\ . Charg;

0

-20 H

_.

0 100 200 300 400 500 600

Time (s)

Fig. 1: Standard profile of required reserve power versus time.

short length compared to coil diameter, which result in signi- ficant stray fields. For the specified energy content such a magnet is not transportable, so it would have to be manufac- tured on-site. Also shielding of the stray field is not easy. The toroidal configuration requires more material (factor 2 to 3), but offers other advantages that may offset the reduced utilization of material. The toroid can be constructed from magnet modules that are transportable, so they can be manu- factured under quality conditions in the factory with potential for cost reductions due to series production. A modular system design allows to use basically the same modules for different customer requirements (stored energy). And, last but not least, the toroidal configuration has the inherent pro- perty to generate only negligible stray field levels already a few meters away from the magnets. This seems to be the most decisive argument when considering to introduce SMES into utility operation e.g. in Europe to ensure public acceptance. Thus we intend to use the toroidal configuration.

B. Shape of Coils For fusion magnet it is common practice to determine an

optimized D-shape to minimize bending forces on the win- ding, due to the stress levels caused by high magnetic fields required for plasma confinement. This however means a winding process with variable bending radius, with special precaution needed to achieve a compact coil package.

Instead we prefer to utilize circular coils (= solenoids) as modules for the toroid to reduce these difficulties in the win- ding process. Unlike for fusion magnets we have the freedom to scale the load levels by selecting magnetic field or coil dimensions in such a way that the critical stresses can be tolerated. Therefore our design contains a rather low field and a large number of magnets, so that each magnet is only at a small angle with its neighbours and thus "feels" almost the symmetric configuration in a solenoidal arrangement.

C. Total System Energy

The total energy of the magnet system must be larger than the required useable energy specified by the customer. To maintain a given output power while discharging the magnet needs a continuously rising voltage to be generated by the

0018-9464/96$05,00 0 1996 IEEE

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power conversion system. In order to keep this voltage below a reasonable threshold the system can only be discharged partially, details depending on the power profile. In addition this reduces operating cost of the refrigeration system, because magnetic hysteresis losses in the superconducting filaments become dominant in the lower field range. For our case 2 MWh were selected instead of 1.5 MWh.

D: Winding Concept

An outstanding requirement for the SMES is the number of discharge cycles the magnets must safely perform without degradation, more than 200,000 "large" pulses over the expected lifetime of 30 years. Thus long-term mechanical strength (fatigue) is essential for all components subject to pulsating loads. The configuration should provide maximum mechanical robustness to minimize quench initiating pro- cesses while allowing sufficient and flexible cooling:

conductor cooled internally by forced flow of

coil consists of double pancakes cooled in parallel, durable insulation system of glass-fibre tapes and G-11 parts, vacuum-impregnation of the winding block. This concept yields a conductor block with prospect for

good long-term voltage strength. Stabilization ratio (cross sections of copper to superconductor) for the conductor should be selected as 30 to 35, to limit the maximum hot- spot temperature for an emergency energy dump to 200 K. By taking into account space needed for cooling channel and insulation an effective overall current density of about 30 A/mm2 is derived.

supercritical He

111. SYSTEM OPTIMIZATION

Given these boundary conditions and choices we had to find a most cost-effective design for the magnet system as a whole. Especially we wanted to determine

the maximum magnetic induction at the conductor, the radius of module coils r and torus radius R,

0 the number and length of magnets, and required wall

A. Dependences between Parameters Each of these items forms one optimization parameter.

The radii r and R are interconnected because for a given average induction the space filled by magnetic field must be the same.

For a constant current density the radial thickness of the winding will mainly depend on the maximum field. Thus for a fixed quantity of ampere-turns (determined from energy) the length of the individual magnet and the number of coils required are directly related.

If the coils are chosen sufficiently short, the outer coil casing wall thickness is determined only from hoop stresses caused by the radial Lorentz pressure. If the coils grow longer, then combined s t r e s x s including bending have to be considered in the edges of the casing (see Fig 5), resulting in the need for either some structural reinforcement or equiva- lently an increased wall thickness. The centripetal force on

thickness of coil casing.

each magnet in the toroid is supported by a cold vault structure formed by the coil casings and wedges in between.

B. Layout Procedure

To perform the optimizations a PC code was developed covering the whole coil design process in an approximate manner. For a given magnetic field, coil radius, and wall thickness all other parameters can be determined quickly. Results include coil number and dimensions, conductor quantity, weight of cold mass, mechanical forces and stres- ses, up to an estimate of cryogenic losses and required power input of the refrigerator. An overview of possible configu- rations with coil radius varying from 2 to 4 m and maximum induction from 3.5 to 5 T is presented in the following figure: Fig. 2a can be used for given single coil dimensions to find the maximum field to store 2 MWh, while Fig. 2b allows to determine the required total number of module coils (all designs in these graphs for a fixed steel wall thick- ness of 70 mm).

C. Cost optimization

system can be calculated from the following factors: We assume that the design-dependent variable costs of the

Cost proportional to conductor length. This includes the conductor itself but also insulation and winding efforts. Cost proportional to total mass of coil casings. Cost proportional to total magnet number. This includes most of the assembly processes, such as electrical and cryogenic interconnections, heat exchangers and valves, support structure to transfer loads to the neighbouring magnets etc. Cost proportional to surface of the vacuum vessel. This contains vacuum vessel and its reinforcement, also super- insulation and shields.

2.5 T

0.5 4 I 45 f

c . x'.

25 I

1.5 2 2.5 3 3.5 4 Outer Coil Radius (m)

Fig. 2. Overview of possible layouts generated by approximate PC code for total energy of 2 MWh and casing wall thickness of 60-70 mm. a) magnet length as function of coil radius, b) magnet number as function of coil radius

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1.5 2 2.5 3 3.5 4 Coil Radius (m)

Fig. 3. Approximate magnet system total cost as function of coil radius, maximum magnetic induction from 3.5 T to 5 T, energy 2 MWh, other system parametem (magnet number, length) as given in Fig. 2.

Our study is not primarily aimed at the cost for a first-of-a- kind system, but rather at the cost of future commercial installations. Thus it is necessary to make realistic estimates for the magnitude of these cost factors in the foreseeable future.

When scaling factors are inserted into the range of magnet designs as presented in Fig. 2 a cost optimum is found (see Fig. 3) for a coil radius of 2.5 m and a maximum induction of 4.5 T. Some further parameters of this specific design are listed in table I. For this design the magnet system cost can be broken up into the following parts: 48% proportional to conductor length, 33% proportional to coil casing mass, 11% proportional to magnet number and 8% proportional to vessel surface.

TABLE I PARAMETERS OF SELECTED DESIGN

torus radius number of magnets outer coil radius coil length operating current max. induction total turns conductor radial Loreniz pressure. average strain force to torus center

13.6 m 34

2.5 m 1.6 m 15 kA 4.5 T

15000 230 km

70 .. 80 bar 0.08 % 1500 t

While the cost level is strongly dependent on the assump- tions used, the position of the optimum is only weakly influenced. The existence of the minimum is due to: * decreasing conductor cost with increasing coil radius. This

is the well-known effect of high torus aspect ratio r/R [5] . increasing coil casing cost with increasing coil radius. This is not only the effect of growing surface area, but also a consequence of the larger asymmetry of the bending forces on the winding which requires increased structural efforts to deal with.

A conclusion to be drawn from these analyses is that un- less there are better (= more cost effective) manufacturing technologies it is not recommended, not necessary, for an energy of 2 MWh to design the magnet system with a larger coil radius (or equivalently bigger r/R-ratio) than the one

presented here. Thus we arrive at dimensions that are still transportable e.g. by road. The size is well within the limits of many existing manufacturing machines (e. g. turning lathes). This, however, was not our guiding criterion.

Iv. MAIN COMPONENTS

A: Conductor

This activity was subcontracted to Vacuumschmelze (Hanau, Germany) to develop a scalable conductor concept compatible with the above choices that allows cost effective production in large quantities. The essential requirements were large stabilization cross section and an internal cooling channel for forced-flow cooling. Operating current level was set at 15 kA to allow the specified discharge cycle with a maximum voltage of less than 5 kV. Alternative designs (Fig. 4) were developed and checked according to the criteria aic losses, mechanical performance and cost.

The "cheapest" version (a) uses a flat cable from standard NbTi multifilamentary wires coextruded into a high-purity Al matrix. Cooling channels can be introduced during the same process. This offers the potential for least price, how- ever aic losses are large and mechanical properties are very bad. Thus the winding package would not be able to carry a significant fraction of the Lorentz loads. Together with the high thermal shrinkage of Al it is impossible to prevent formation of gaps between coil and casing during cooldown or electrical excitation of the coil. Hard Al alloys do exist, but they cannot be used for electrical stabilization. Preferably something between these extremes would be needed, but that would require a material development and qualification effort which is beyond our present scope.

The massive copper version (b) uses the same flat super- conductor cable, soldered into the slot of a Cu profile. This could be extruded as a whole or be composed of brazed parts, provided the connection is able to survive the cyclic mecha- nical loads due to SMES operation. This conductor can carry about half of the operational Lorentz hoop load. The thermal contraction behaviour of the winding block would be very close to that of the steel casing. A/c losses are smaller than for the high-purity Al, they consist mainly of coupling through the copper matrix and hysteresis in NbTi filaments. The latter can easily be reduced by selecting superconductor wires with more filaments, i.e. smaller filament diameter, without impact on the overall concept. This is our presently preferred solution.

The third version (c) should provide an alternative with

b C a Fig. 4. Conductor concepts: a) flat cable coextmded with high purity AI b) flat cable soldered into groove of copper profile c) "bundle" conductor in external pipe for reduced aic losses

227 I

small alc losses. To achieve the stabilization, there are many more Cu strands than superconductor wires (with small fila- ments in this case). In order to keep the number of cabling stages to the minimum the outer Cu wires of the subcable are arranged in concentric rings, twisted in opposite directions. The central cooling channel allows cooling with very little pressure drop. The cable made up of 6 subcables has to be inserted into a steel jacket, which is then drawn to reduced diameter (round or squared) to get an acceptable void frac- tion in the strand region. Due to the steel tube this conductor is mechanically strong but also more difficult to wind. A/c losses in NbTi filaments and Cu wires are low. Inserting the cable into the jacket may be similar to what is being deve- loped for fusion magnets, however at present this process is not performed in a large scale and still contains technical risk. We are not sure the improved a/c performance (i.e. refrigerator operating cost) are worth the significantly higher price of this alternative.

B: Structure of Coil Casing

The winding creates a radial Lorentz pressure of about 80MPa. The complete force pattern with radial and axial components, including forces that are not rotationally symmetric due to the toroidal geometry, must be carried by strain in the coil composite and casing. Most critical are combined stresses including bending in the case corners and shear within the winding pack. All these have to be assessed by FEM analysis, using reduced material property data in order to take into account the large number of load cycles. The tolerable limits for the coil composite are improved by properly prestressing in axial and radial direction.

With manufacturability and cost in mind we decided for a rather thin steel casing with parts that are cast or welded (sketch see Fig. 5). Two halves of the impregnated winding block will be placed on an inner cylinder, two outer cylinders with end caps are mounted on top and fixed axially by bolts. Radial precompression is done by steel cushions inflated with resin.

C: Refrigeration System

Each coil module has only two helium ports (internally all pancake cooling paths are in parallel) which also serve for the electrical interconnection of coils or to the current leads.

We suggest a dual-circuit cooling system to combine a standard refrigerator supplying a LHe bath in the primary

I &-. I hoop stresses. outer cylinder with end cap

-0 # double combined stresses pancake

and local peaks including bending

-\ - Fig. 5. Sketch of coil casing components indicating mechanical interaction of winding and regions with limiting stresses.

Fig. 6. Schematic of dual-circuit cooling system.

loop with a secondary circuit of supercritical helium cooling the magnets (see Fig. 6). In normal operation the coolant is driven by a cold pump and passes a heat exchanger before entering each magnet, Due to the large cooling channel diameter in the magnets there is so little pressure drop that all magnets can be driven hydraulically in series. Mass flow can be adjusted to the momentary SMES operational need.

The helium liquifier for the primary circuit fills the LHe bath which serves as a buffer, thus its capacity is matched to average cooling requirements. Presently estimated require- ments are about 500 W refrigerator power at 4 K, plus 50 l/h liquifaction, and 2 kW refrigeration at 80 K (magnets only). The total power is estimated for the specified operating pro- file of the SMES and can be broken down into 32% due to "large" pulses, 16% due to "small" pulses, 24% caused by conventional current leads, 9% generated in soldered coil connections, 9% from thermal radiation to 4 K, 5% from coil suspensions (G-11 straps), and 5% due to thermal radiation to the shield at 80 K.

The modular design of the distribution system (heat ex- changers, valves) allows for parallel operation of the magnets as needed during system cooldown and for hydraulic isola- tion of a magnet after a quench and separate recooling there- after, while the unaffected modules are kept cold.

V. CONCLUSION

Basic magnet design concepts are presented for a cost- effective magnet system design based on draft specifications for a SMES to be used in electrical utility primary control. For a usable energy of 1.5 MWh there is no advantage to design coils with a radius of more than 2.5 m radius.

REFERENCES [1] R. Voelzke, M. Kleimaier, K.P rescher, U. Radtke: "Network Simulations

for the Sizing of SMES Systems for Providing Primary Reserve", 2nd Int. SMES Symposium, Karlsruhe, Germany, 1994, pp. 295-308

[2] M. Kleimaier, K. Prescher, U. Radtke, U. Voelzke: "Supraleitender magne- tischer Energiespeicher", ETZ, vol. 116 (1995) 9, pp.38-45 (in German)

[3] C.M. Rey, X. Huang, S.F. Kral, Y. Lvovsky, M.F. Xu: "Key Design Issues of a 30 MW Babcock and Wilcox SMES", 2nd Int. SMES Symposium, Karlsruhe, Germany, 1994, pp. 182-191

141 C. Luongo: "Review of the Bechtel Team's SMES Design and Future Plans for a Technology Demonstration Unit", proc. ASC 1994, IEEE Trans. Appl. Superconductivity, vol. 5, (1994) 2, pp. 422-7

Requirements", LA--10320-MS, Appendix B, Los Alamos, 1984 [5] W.V. Hassemhl: "Comparison of Coil Geometries Based on the Conductor