Design, Implementation and Testing of a Ship-Board Medium-Voltage DC Generation System

Based on a Ultra-High Speed 12-Phase Alternator

G. Sulligoi, A. Tessarolo Dept. of Industrial Engineering and

Information Technology University of Trieste

Trieste, Italy gsulligoi@units.it

atessarolo@units.it

V. Benucci, A. Millerani Trapani 5th Division - NAVARM

Italian Navy Rome, Italy

varo.benucci@marina.difesa.it andrea.milleranitrapani@marina.difesa.it

M. Baret, F. Luise Test Room - Engineering Dept.

Ansaldo Sistemi Industriali S.p.A. Monfalcone, Italy

massimo.baret@asiansaldo.com fabio.luise@asiansaldo.com

Abstract—This paper reports on the design, development and validation of an advanced prototype 2MVA generation equipment (Naval Package) for a shipboard Medium-Voltage DC integrated power system. The generation equipment is based on a ultra-high-speed 22500-rpm 12-phase alternator which feeds an AC/DC power electronics converter composed of four diode rectifiers and four IGBT choppers. The prototype realization constitutes a follow-up of a previous Naval Package version employing a wound-field 6300-rpm alternator feeding non-controlled AC/DC converters. The major technical challenges faced in the design and development of the advanced Naval Package prototype are outlined in this paper taking the previous lower-speed version as a technology reference. The system performance highlighted by the testing campaign are finally addressed along with the lesson learnt to be capitalized for future developments.

I. INTRODUCTION The most important navies in the world are adopting the

electrical propulsion in many new projects and constructions. One of the most promising developments of this technology is the MVDC Integrated Power System (IPS), for which an IEEE Standard has been recently published [1]. Challenges related to the implementation of MVDC IPS have been investigated in [1] and involve subsystems technology, system integration, controls, fault tolerance, power quality, reconfiguration, power management, electric grounding and safety, etc. Power generation is certainly one of these key issues, as MVDC IPS are designed to employ different types of power sources such as electric machinery, accumulators, fuel cells. As regards electric machinery, generators can be of different type (wound-rotor, permanent-magnet, multi-phase), size, prime mover (diesel engines, gas turbines) and work at different or even variable speed and load.

In this scenario, in 2006 the Italian Navy, through the General Direction of Naval Armaments and Weapons (NAVARM), awarded a contract for the realization of the so called Naval Package (NP), a demonstrator of an innovative generation system for naval ships. The realization of the NP is part of an extended program of Italian Navy aimed at obtaining experimental results in the field of shipboard MVDC IPS for naval applications.

The NP project involved the realization of two 2-MW generation system prototypes suitable for supplying a 3000-V shipboard DC network using a 22500-rpm Gas Turbine (GT) as prime mover. The first and more conventional system prototype (referred to as “Naval Package 1”, NP1) has been described in a previous work [3]. This paper is devoted to the second, more technologically-advanced prototype, which will be called “Naval Package 2” (NP2) in the following. The main purpose in NP2 realization philosophy was to push ahead the electric generator power density enhancement target by further raising the machine speed up to 22500 rpm, making possible a direct coupling with the GT. From system conception, through its development, design and testing, remarkable technical challenges had to be faced, which will be outlined in this paper bearing in mind the previous NP1 experience as a benchmark and technology reference. The paper finally reports on system testing results highlighting the main lessons learnt from the entire project and their prospective impact on future developments.

II. OVERALL SYSTEM DESCRIPTION The overall structures of NP1 and NP2 systems are

sketched in Fig. 1. As illustrated in the figure and discussed in [3], NP1 is based on a dual-star alternator supplying two series-connected diode rectifiers. The cascade of the two rectifiers feeds the DC load constituted by the shipboard electric grid. The alternator is coupled to the gas turbine (GT)

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by means of an epicycloidal gear-box and is equipped with a wound-field rotor. Therefore the system voltage regulation is performed by acting on the generator excitation and the power electronics equipment performing the AC/DC conversion is passive (non-controlled).

Conversely, NP2 employs a quadruple-star generator whose stator winding consists of four three-phase sets, each feeding a diode rectifier. Each diode rectifier in turn supplies a DC/DC power converter (chopper). The four chopper are finally series-connected and their cascade is closed on the DC load constituted by the shipboard grid. The alternator, in this case, is directly coupled to the GT and is then a ultra-high-speed permanent-magnet machine with a rated speed of 22500 rpm. Because the permanent-magnet rotor does not allow for any flux regulation, the output voltage control is demanded to the choppers. In order to cope with the high commutation frequency, IGBTs are used as chopper power switches.

III. ELECTRIC GENERATOR DESIGN The core of both NP1 and NP2 systems is constituted by

the electric generator. Its design, in fact, is expected to fully exploit the advantage of using a DC distribution network, which makes the choice of machine speed and pole number free and thus allows for wide design optimization margins in terms of machine size and weight.

NP1 and NP2 represent two possible ways of enhancing generator power density in shipboard DC power systems, the

Fig. 2. NP1 generator: (a) overall view from exciter side; (b) epicycloidal gear-box detail on the coupling side.

Fig. 3. NP2 generator: (a) overall view; (b) terminal box detail with 12 phase leads.

Fig. 1. Overall dimensions of (a) NP1 and (b) NP2 generators. The drawingshave equal scales.

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former based on a wound-field rotor solution, the latter on permanent magnet technology. The two generators are both innovative in this sense, but are based on strongly different design choices. A visual comparison between the machines is proposed in Fig. 2, Fig. 3 and Fig. 4. The latter, in particular, shows the overall external dimensions of the two machines drawn in the same scale.

The technical data of the two machines are compared in Table I, while a deeper insight into their technology differences is provided into Table II. Next the main choices made in developing the two machines are discussed in more detail.

A. Speed, number of poles and frequency The fundamental choice for dimensioning the two

generators consists of selecting their speed, while the GT speed (22500 rpm) is retained as a constant.

1) Case of NP1 generator

It is clear that for the NP1 solution, based on a wound-field generator, a direct coupling (implying a 22500 rpm speed) would not be feasible mainly because of the stresses that would act on the end windings of the rotor excitation circuit. Another major issue that limits the speed in wound field

generators is the lateral dynamics of the entire shaft line, which comprises not only the generator and turbine rotor, but also the exciter rotor [3]. The latter “hangs” out the generator main frame on the opposite side with respect to the coupled machines (see Fig. 2a and Fig. 4a) and may give rise to remarkable vibration issues [4].

Fig. 4. Overall dimensions of (a) NP1 and (b) NP2 generators (terminal boxes are omitted). The drawings have equal scales. The heat exchanged of NP2 is integrated in the generator frame and is thereby included in the overall dimensions.

TABLE II. TECHNOLOGY COMPARISON BETWEEN NP1 AND NP2

NP1 NP2

Winding type Embricated-coil, double layer, short-pitch winding Embricated-coil, double layer, short-pitch winding

Stator coil conductor type Flat strands Litz wire

Stator slot shape Rectangular Rectangular

Magnetic sheet type Mechanically cut and notched Mechanically cut, single punch

Stator wedge Non-magnetic Magnetic

Rotor type Wound-field solid-steel round rotor Surface Permanent Magnets (SPM)

Ventilation system Forced with electric blowers Forced with external fan (open air circuit)

Cooling system Top-mounted water-to-air heat exchanger Frame-integrated fluid heat exchanger

Coupling with GT Through epicycloidal gear-box Direct

TABLE I. TECHNOLOGY COMPARISON BETWEEN NP1 AND NP2

NP1 NP2 Number of phases 2 x 3 4 x 3 Number of poles 4 4

Rated voltage 1200 V 620 V Rated frequency 210 Hz 750 Hz

Rated power 2000 kVA 2000 kVA Rated speed 6300 rpm 22500 rpm

Rated power factor 0.95 0.95 Insulation class F H Thermal class B F

Generator weight 5200 kg 2300 kg Rotor weight 1035 kg 250 kg

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The rated speed of 6300 rpm for NP1 generator is chosen as the highest speed judged to be safely compatible with a machine construction including a hanging exciter, i.e. in the arrangement depicted in Fig. 4a. Such a choice implies the use of a gear-box with a gar ratio of 22500 rpm / 6300 rpm =3.6. An epicycloidal parallel-axis gear-box is chosen as shown in Fig. 2b, directly mounted on the machine end-shield and characterized by a highly compact design.

The possible number of poles that can be reasonably chosen with a rotational speed of 6300 rpm are two and four, respectively leading to a 105 Hz and 210 Hz stator frequencies. The choice of a four pole, although disadvantageous due to the higher frequency and to the consequent core losses is made preferred for two main reasons:

• Four-pole machines have a much better vibration performance than two-pole machines thanks to the absence of the twice-line-frequency excitation force component [4], [5].

• The four pole design usually leads to a shorter rotor with larger diameter, which means a reduced bearing span and a better placement of the critical (resonance) speeds [5].

• From the electromagnetic dimensioning viewpoint, the four pole design generally leads to a smaller volume of machine active parts and thereby to a higher utilization coefficient, mainly due to the thinner stator yoke [5], [6].

2) Case of NP2 generator

In the case of NP2 generator the direct coupling with the GT imposes a rated speed of 22500 rpm, which requires moving from a wound-field to a permanent-magnet rotor design. The four pole design is chosen for NP2 generator following the same reasons as discussed for NP1. In this case, however, the choice brings more challenging consequences as it determines a stator frequency of 750 Hz. Managing the core losses associated with such a fundamental frequency is one of the major issues in NP2 generator design (Sections III.B, III.C).

B. Number of stator phases The choice of a dual-star configuration for the NP1

machine has already been extensively analyzed in [3], where performance and fault tolerance issues have been discussed as the main reasons. Further studies [7], however, prove that the output performance advantages that could be gained by further increasing the number of phases above six are not so significant.

As concerns NP2 generator, a very critical aspect of its design is constituted by the losses induced in the rotor by revolving space harmonics which appear in the air-gap due to the phase current distortion. This phenomenon can be thoroughly studied by means of the time-domain (or time-stepping) Finite Element (FE) analysis method (Fig. 5).

It consists of preparing a 2D cross-sectional machine

model associating each of the phases to the external circuits (modeled with a lumped-parameter approach) which are connected to stator terminals. In this instance, as illustrated in Fig. 5, the twelve machine phases, arranged in four star-connected sets, are interfaced to diode rectifier models. The downstream choppers and the DC load are represented as four equivalent time-variant resistances connected to the bridge rectifiers. This modeling approach (based on interfacing the FE machine model with lumped-parameter external circuits) enables to predict how the real machine (modeled in its

Fig. 6. Generator voltage (a) and current (b) waveforms resulting from time-stepping FE analysis simulation of NP2 system.

Fig. 5. Scheme for time-stepping FE analysis simulation of NP2 system including both the generator and the coupled power electronics circuitry.

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detailed geometry and including saturation and eddy current effects) reacts when operating in combinations with the power electronics equipment to which it will be actually connected. Examples of time-stepping FE simulation outputs are provided in Fig. 6. In particular, one can observe (Fig. 6b) the noticeable distortion of phase current waveforms (subsequently confirmed by testing, see Sections C, D).

In a three-phase machine, such a distortion would give rise to important air-gap rotating harmonic fields, which would cause flux pulsations and consequent eddy-current losses in rotor conductive regions. As known from the literature [8], the adoption of multiple-star stator winding dramatically reduces the phenomenon thanks to the mutual cancellation effects among the harmonics produces by different stator stars. In particular, a dual-star winding configuration (as used in NP1) would lead to the cancellation of air-gap field harmonics of order 5 and 7 [9]; adopting a quadruple star arrangement causes also the 11th, 13th, 17th, 19th harmonic orders to disappear [9], thereby allowing for an almost sinusoidal armature reaction field distribution in the air-gap even in presence of the important phase current distortion shown in Fig. 6b and also observed in [5]. The effect of residual harmonic flux pulsations in rotor conductive regions can be investigated with the time-stepping FE method, as discussed later (see Section III.D). Of course the choice of twelve phases leads to some practical drawbacks, such as the larger complications in cabling and connections and the increase in the size of the terminal box, which needs to accommodate twelve phase leads (see Fig. 3b).

C. Stator technology For both NP1 and NP2 generators, the a double-layer

short-pitch imbricated-coil stator winding is selected along with a rectangular slot shape.

1) Magnetic laminations

The laminated magnetic sheet used in NP1 generator is relatively traditional: mechanically cut and notched, it features 0.5 mm thickness and 1.3 W/kg rated specific losses at 1 T and 50 Hz. In NP2 generator, in order to cope with the expected large core losses due to the high stator frequency (750 Hz), special punch and die laminations are employed with 0.1 mm of thickness, less than 6% silicon content and reduced specific losses (Fig. 7).

2) Stator slot wedges

While in NP1 generator an ordinary non-magnetic wedge is used to retain stator coils in slots, in NP2 machine a sintered-power magnetic wedge is used. The material selected is Somaloy 500, whose B-H characteristic is shown in Fig. 8. The main reason for using a magnetic wedge is to reduce the slotting effects. In fact, using a rectangular slot cross section implies a relatively large ratio R between the slot opening width and the tooth width at the air-gap. High values of R are known to produce important slot harmonics in the air-gap magnetic field, with consequent flux pulsations and eddy-current losses in rotor conductive regions [10]. The use of magnetic wedges of suitable magnetic permeability strongly helps reducing slot harmonic amplitude, which is highly recommendable in this instance due to the high rotor speed and the consequent high eddy-current frequency. The main drawback of using magnetic wedges is the growth in slot leakage inductance. A higher leakage inductance, in fact, increases the internal voltage drop and requires higher permanent-magnet excitation (and volume as a consequence) to obtain the same output voltage at full load.

3) Stator winding technology

In NP1, stator coils are formed by multiple turns in series, each composed of some parallel-connected flat conductors (Fig. 9). The use of flat conductors instead of wire, typical of high power machines, is chosen as it allows for a better turn insulation while keeping a high slot fill factor. In NP2, the use of flat conductors is discouraged because the high operating frequency would produce too important skin and proximity effects in rectangular-section flat conductors, causing large copper stray load losses. The solution adopted is then to form each turn using litz wire, that is a pressed bundle of twisted parallel-connected round-section thin conductors. The very small cross section area of litz wire strongly limits skin effects, while the expedient of twisting conductors brings the benefit that each of them has nearly the same flux linkage, so that internal circulation currents between shut-connected wires are minimized.

Fig. 7. Rated specific core losses for the NP2 generator laminations (x-axisin logarithmic scale).

Fig. 8. B-H curve of Somaloy 500 magnetic wedge.

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D. Rotor technology In NP1 generator, the speed of 6300 rpm, combined with

the four-pole rotor diameter necessary to develop the required torque, suggests the use of a solid-steel round rotor whose design is derived from large turbo-alternator technology. Examples of such design in the case of a four-pole machine is provided in Fig. 10: a forged cylindrical shaft is first turned on a lathe and then the axial slots are machined by milling (Fig. 10a). Concentric excitation turns are then wound or mounted around each pole (Fig. 10b). The most mechanically critical aspect of such a design is the centrifugal force acting on excitation end-coils. To keep them in place a stainless steel retaining ring is mounted by shrink fit on both ends of the rotor body.

For such a speed of 22500 rpm, the mentioned design

would be inadequate and a permanent-magnet rotor solution appears mandatory. In the case of NP2 generator, in particular, a Surface-mounted Permanent-Magnet (SPM) rotor design is used as illustrated in Fig. 11. Because of the relatively low number of poles, it is found convenient to use multiple equally-oriented permanent magnets for each pole. With respect to the single-piece solution, this enables to use simple Sm-Co permanent-magnet blocks with parallel magnetization instead of an arc-shaped permanent-magnet with radial magnetization. Between the permanent-magnet arrays pertaining to adjacent poles, a spacer is placed for mechanical compactness. Due to the high centrifugal force acting on them, gluing permanent magnets to the rotor surfaced is far not sufficient for their retention. As a necessary provision, a retaining carbon fiber cylinder is wrapped around the permanent magnets. The dimensioning of the retaining cylinder is a crucial task both from a mechanical and electromagnetic viewpoint. As concerns the mechanical dimensioning, Fig. 12a, shows the forces that act on the PM and on the cylinder portion on it. Fm and Fc are centrifugal forces and T is the tensile force acting on carbon fibers. The mechanical stress distribution in PMs and in retaining cylinders is not uniform and needs to be locally determined by FE methods in order to check for possible critical points (Fig. 12b).

Fig. 12. (a) Forces acting on a PM and the retaining cylinder portion above it. (b) Local stress distribution resulting from mechanical FE analysis.

Fig. 11. SPM rotor design with retaining cylinder.

Fig. 10. Sketch of a rotor design derived from turbo-alternatortechnology. (a) before winding; (b) after winding assembly.

Fig. 9. Slots structures for (a) NP1 and (b) NP2.

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The mechanical dimensioning of the retaining cylinder

has noticeable impacts from the electromagnetic point of view. In fact, as the cylinder thickness increases, the magnetic air-gap of the machine increases as well.

The increase in the air-gap has two main consequences:

• Permanent magnet thickness needs to grow due to the increased reluctance of the main flux path;

• Stator phase inductance decreases. The first consequence is certainly a drawback because it

pushes the design towards larger dimensions and higher costs due to the increased permanent magnet weight. The second consequence is a drawback under some respects (for example a small phase inductance determines large short-circuit currents), but may have also positive aspects. A low phase inductance, in fact, makes the machine more suitable to operate in combination with diode rectifiers at high frequencies as it speeds up commutation transients.

The retaining cylinder also plays a role in terms of eddy-current losses due to the fact that its material exhibits non negligible magnetic permeability and conductivity properties. Taking into account such properties in the time-stepping simulations illustrated in Fig. 5, it can be observed that the majority of rotor eddy-current losses due to air-gap space harmonics occurs just in the carbon fibers (Fig. 13). The eddy currents which arise in the retaining cylinder act as a sort of shield. The latter prevents flux high-order harmonics to penetrate deeper towards the permanent magnets, which are then expected not to suffer dangerous overheating problems.

As an alternative, the adoption of a suitable electromagnetic shielding (in the form of a copper layer) is evaluated both in terms of overall rotor eddy current losses amount and in terms of practical feasibility. The final construction does not adopt it, since the stator twelve-phase solution shows no convenience (it would be needed, instead, with a lower number of phases).

E. Cooling method The cooling method used in NP1 generator is quite usual

for medium-sized electric machines. It is based on a shell-and-tube air-to-water heat exchanger mounted on the top of machine frame (Fig. 2a). The internal generator air is forced to flow through the exchanger by means of electric blowers to

guarantee a sufficient ventilation also during low speed operation, which would not be possible with shaft-mounted fans. In NP2, a more effective cooling method is used, based on a heat exchanger integrated into the frame. In fact, refrigerating coils (carrying a pressurized water flow) are directly mounted inside the machine frame (see Fig. 3a) so that a direct heat transfer can take place by conduction from the stator to the cooled frame. The internal generator air is forced through an open ventilation circuit by means of external fans. In particular, the stator stack has a middle slit through which the cool air enters the air-gap, to flow from the stack middle section towards its end portions.

IV. SYSTEM TESTING The performance of the NP2 generator prototype (named:

Permanent Magnet Ultra High Speed Generator, PMUHS) is assessed with the test set-up shown in Fig. 14-15. The machine is driven by a DC motor through an auxiliary gear-box with a gear ratio equal to 22500 : 1800. The output of the rectifying power electronics is connected to a purely resistive load, constituted by forced-ventilated resistor banks. The ratings of the DC motor and of the DC load are reported in Fig. 14.

During the system test under load, the generator phase current distortion measured was completely in agreement with predictions (Fig. 15). A machine characterization at no load was also performed by driving it at various speeds with the DC machine (Fig. 14) used as a drive and previously characterized. A curve of the no-load losses versus the speed is reported in Fig. 16. Such no-load losses include the sum of friction and windage losses plus core losses (due to the fundamental no-load flux and its harmonics). Joule and stray-load losses, estimated from the short-circuit test, are around 25 kW. The loss distribution is visually illustrated in Fig. 17. The losses resulting from test essentially match design predictions. In particular, the core losses, which were a major reason for concern, proved to be well controlled by the design provisions adopted. The effect of the losses on machine performance is also assessed by performing a heat-run test in the complete system configurations shown in Fig. 14.

Fig. 14. Set-up for system testing of NP2 generator (PMUHS).

Fig. 13. Current density distribution from time-stepping FE analysis.

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The stator temperature was monitored during the test through Pt100 thermal detectors mounted on the end coils and embedded in the slots. The hottest Pt100 temperature trajectory during the test up to steady-state conditions is shown in Fig. 18. It confirms the full generator capability of operating within thermal class F temperature limits.

V. CONCLUSION In this paper the conception, design and testing of a 2

MVA prototype generation system, based on a 22500-rpm 12-phase SPM alternator, have been described. The alternator is connected to an active AC/DC converter to supply a MVDC integrated power system for shipboard applications. The technical challenges faced during system development have been reported and the main design choices have been illustrated. Based on testing results, it can be certainly concluded that the prototype has matched the expected requirements showing the practical feasibility of achieving outstanding power density targets by raising generator speed towards supersonic values. In view of future developments, the experience gained throughout the project makes it possible to say that further performance and compactness enhancements could be achieved by two design modifications: the use of magnetic bearings (which are definitely more suitable for similar speeds and also allow an active vibration control) and choosing a closed ventilation system based on a shaft-mounted suitably-designed fan (this would reduce the room needed for external ventilation circuit elements).

REFERENCES [1] Recommended Practice for 1 to 35 kV Medium Voltage DC Power

Systems on Ships, IEEE Std. 1709, DC Power Systems on Ships Working Group of the IEEE Industry Applications Society Petroleum & Chemical Industry (IAS/PCI) Committee.

[2] N.Doerry, J.Amy, “Functional decomposition of a medium voltage DC integrated power system, NGIPS Technology Development Roadmap”, NAVSEA 2007”.

[3] G. Sulligoi, A. Tessarolo, V. Benucci, M. Baret, A. Rebora, A. Taffone, “Modeling, simulation and experimental validation of a generation system for Medium-Voltage DC Integrated Power Systems”, IEEE Transactions on Industry Applications, vol. 46, no. 4, July/Aug. 2010, pp. 1304-1310.

[4] W. R. Finley, M. M. Hodowanec, W. G. Holter, “An analytical approach to solving motor vibration problems”, IEEE Trans. on Industry Applications, vol. 36, no. 5, Sept./Oct. 2000, pp. 1467-1480.

[5] A. Tessarolo, G. Zocco, C. Tonello, "Design and Testing of a 45-MW 100-Hz Quadruple-Star Synchronous Motor for a Liquefied Natural Gas Turbo-Compressor Drive", IEEE Transactions on Industry Applications, in press.

[6] S. Huang, J. Luo, F. Leonardi, T. A. Lipo, “A general approach to sizing and power density equations for comparison of electrical machines”, IEEE Trans. on Industry Applications, vol. 34, Jan./Feb. ’98, pp. 92-97.

[7] A. Tessarolo, “Experimental performance assessment of multiphase alternators supplying multiple AC/DC power converters”, Journal of Energy and Power Engineering, vol. 4, no. 12, Dec. 2010, pp. 43-50.

[8] E.A. Klingshirn, “High phase order induction motors−Part I−Description and theoretical considerations”, IEEE Trans. on Power Apparatus and Systems, Jan. 1983, vol. PAS-102, pp. 47-53.

[9] A. Tessarolo, “Analysis of split-phase electric machines with unequally-loaded stator windings and distorted phase currents”, International Conference on Electric Machines, ICEM 2010, 6-9 Sept. 2010, Rome, Italy, CD-ROM paper RF-013331.

[10] A. Still, C.S. Siskind, Elements of Electrical Machine Design, McGraw-Hill, 1954.

Fig. 15. Phase current waveform at 1140 kW DC output and 18000 rpm.

Fig. 16. No-load losses as a function of the speed (measurement points andinterpolating curve).

Fig. 17. Estimated loss distribution in rated operation.

Fig. 18. Maximum stator winding temperature during the heat-run test.

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