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THE ELECTRIC WARSHIP VI by C G Hodge and D J Mattick OBE A paper for the Institute of Marine Engineers To be read at 1730 on Tuesday 12 December 2000 The views expressed are those of the authors alone. BIOGRAPHIES DAVID MATTICK David Mattick’s early career resulted from specialisation as a nuclear submarine weapon engineer. After service as the Assistant Weapon Electrical Officer of HMS WARSPITE and in the MoD, he rejoined WARSPITE as the Weapon Electrical Officer of HMS WARSPITE. In 1982 he was appointed as the Marine Engineer Officer of HMS SWIFTSURE. After promotion to Commander in 1984 he headed the electrical power systems specialist group within the MoD, subsequently serving with the VANGUARD Class submarine project and then as a Project Manager with Director General Ship Refitting. In 1994 he served as the Surface Ship Marine Engineering Desk Officer in Director Future Projects (Naval), tasked with concept design of future naval vessels and was appointed as the MoD Electric Ship Programme Manager in February 1996. Since retiring from the Royal Navy in 1999 he has been employed in the Integrated Propulsion Systems Division of Rolls-Royce Marine Power as the Electric Ship Manager. He was awarded the OBE in the 1999 New Year’s Honours List. CHRISTOPHER HODGE After initial training as a mechanical engineer, and service as an Assistant Marine Engineer in HMS WARSPITE, Christopher Hodge joined HMS ORPHEUS as the Marine Engineer Officer in 1982. He subsequently took an MSc in Electrical Marine Engineering and served in the MoD as the project officer for electrical ship propulsion. After promotion to Commander in 1989 he served as the Marine Engineer Officer of HMS CONQUEROR before returning to MoD as the head of the Nuclear Steam Raising Plant electrical design authority. He was the head of the electrical power system specialist group within the MoD until April 1998 after which time he joined Rolls Royce Marine Power initially in their Integrated Propulsion Systems Division and now in the Engineering and Technology Division . THE ELECTRIC WARSHIP VI C G Hodge BSc MSc CEng FIMarE D J Mattick OBE BSc CEng FIMarE MIEE MINucE ABSTRACT This paper which follows on from five earlier IMarE papers by the same authors, reviews the progress of the Electric Ship Programme over the last year and focuses on some of the technological developments in equipment and architectures associated with the Electric Warship INTRODUCTION This is the sixth paper in the Electric Warship series presented at the IMarE and as before it aims to re-state and update the Electric Warship vision and inform readers of progress towards realising the concept in a warship and to record significant developments in its associated technology. BACKGROUND The following is a very brief summary of the first five Electric Warship papers, References 1, 2, 3, 4 and 5 The benefits of employing a common power system for both propulsion and ship’s services can, in an optimised merchant ship installation, allow Running Cost savings of up to 25%. In a warship, where the constraint of reducing Unit Purchase Cost as well as Running Costs does not allow a system to be fully optimised for fuel economy, the savings may nevertheless be sufficient to allow purchase of one extra ship in a class of around thirty vessels. However as was explained in the previous papers, the size of motors and converters needs to be reduced if the equipment is to be installed in current frigate sized warships. The Electric Ship (ES) concept was developed from Integrated Full Electric Propulsion (IFEP) and it was aimed to reduce Unit Purchase Cost (UPC) and yet retain as much as is practical of the IFEP reduced Running Cost (RC). In the Electric Ship this is achieved with two main features above those to be found in a traditional IFEP vessel. Minimum Generator Operation: The UPC and space constraints require fewer but more highly rated prime movers than would be found in a merchant IFEP vessel. In order to restore the fuel savings conceded by fitting fewer prime movers the ES runs under a regime of minimum generator operation, often with only one prime mover operational. This brings significant gains, partly in fuel consumption but mainly from maintenance costs due to the minimised engine running hours. Electrification of Auxiliaries: Additional maintenance and manpower reductions can be achieved by using electric auxiliaries wherever possible. In addition there will be benefits to be gained from this policy in terms of overall weight of equipment fitted (central energy storage and high reliability ship wide electric power systems). The previous papers described our early vision of an Electric Ship Power System, shown at Figure 1. GEN GEN WR21 CONV PMPM 1-2 MW SPC En Store SPC GEN CONV PMPM 6-8 MW GEN WR21 750 V DC Figure 1: Electric Warship Power System REVIEW OF CURRENT ISSUES AC or DC As has been discussed in may forums, not least the earlier versions of this paper, one of the principal decisions to be made early in the concept work for an electric warship is whether AC or DC should be used as the medium for power distribution. The authors have consistently maintained that the advantages offered by DC in its cost of installation, efficiency of distribution and control require it to be treated as a serious option for both the propulsion and ships service duties. In Reference 1 the authors stated that DC should be used for ships service distribution and AC for the propulsion bus bar. This decision was based on the belief that technology would limit the power that could be transmitted safely by a DC system. However industry is now becoming increasingly confident that that the power for propulsion could be safely controlled using DC. Reference 6 a paper by PMES and the UK MoD presented at AES 2000, makes a strong case for using DC for the propulsion system of a frigate sized electric warship and many power electrical companies are now starting to believe that DC could provide the supply to the Propulsion Motor Converters. The main issue with DC as a traditional distribution medium is the expense and mass of DC air circuit breakers and this has been the reason why most assessments to date the advantages of DC have only been marginal. The authors now consider that convergence between power electronic converters and circuit breakers is feasible. Indeed, as is discussed later in this paper, the USA IPS Programme changed the functional definition of one of their modules from solid-state circuit breaker to DC-DC converter when the use of the equipment became predominantly one of DC voltage conversion, nevertheless the circuit protection and fault current interruption functionality of the module was retained. The authors believe that DC is the distribution medium of the future and can today be used for ships service distribution. Further, if the power electronic converters that will necessarily be required also provide the fault interruption duty then such a system will have clear advantages over its AC counterpart, including initial cost and mass. It is worth mentioning that one often quoted advantage of DC systems, namely that of simpler and more robust stability, needs some qualification. It is true that the same issues of transient stability do not arise. However if controlled in constant power mode, as is the case for motor drives, power electronic converters exhibit a negative impedance. That is if the system voltage falls the current drawn by the converter rises. This effect can cause voltage instability through interaction with the remaining voltage regulating control circuits in the complete DC system. Purdue University in the USA has conducted significant research (led by Dr Scott Sudhoff) into this phenomenon, References 7 and 8 report the latest results of their work, which offers design methods by which stability of power electronic DC systems can be assured. MOTORS AND CONVERTERS The quest for power dense motors suitable for warships continues at an ever-quickening pace with the Alstom Advanced Induction Motor at the forefront; a motor has been supplied to the USN Integrated Power System programme and selected for the Royal Navy Type 45 programme. In the commercial arena Jeumont Indutrie are providing radial flux permanent magnet machines to DCN for an export project and they also supply axial flux machines for wind generation duties. With their experience, they can readily scale and militarise these for warship propulsion. Siemens are offering radial flux permanent magnet machines in the Siemens-Schottel propulsors, a commercially available podded drive suitable for commercial ships. It is likely that development of other commercial permanent magnet motors is underway. Podded propulsors are yielding significant benefits to the commercial marine industry. In particular through reduced fuel consumption – good hydro-dynamic designed pod and ship back-end can realise propulsion efficiency gains in excess of 10% - and simplified ship installation - the complete pod can be supplied and fitted in days thus avoiding the lengthy installation of a traditional system. In the purely military arena, the Rolls-Royce development of the Transverse Flux Permanent Magnet Motor, discussed in the last paper of this series, continues with encouraging results emerging; ABB are offering a radial flux permanent magnet motor based on their development for the Italian Navy submarine programme; various companies and groupings in the US are developing permanent magnet motors targeted at the DD21 programme. As has been highlighted in many previous papers, the power electronics revolution continues and various new devices are being developed or are available, which enable different Converter topologies that can reduce volume, weight, cooling requirements and cost of the equipment; this thread weaves throughout this paper. ENERGY STORAGE There is a continuing debate about the need for and capacity of Energy Storage required to support Electric Ship propulsion and distribution. In practice, the need must be judged against the application and be tailored to the specific system requirements. There are a number of suitable technologies for this Energy Storage including traditional lead acid batteries widely used in submarines, advanced batteries such as the Zebra system which was reported at Reference 9, flywheels and, perhaps most recently, a Regenysis system based on that offered by Innogy Technology Ventures Ltd - previously a part of National Power. There is a web site dedicated to the Regenysis system and this may be found at Reference 10. The UK MoD has work underway to assess the Regenysis potential for a warship application; indeed it is one of the systems selected for the UK Electric Ship Technology Demonstrator. MOTORS TRANSVERSE FLUX The development of the TFM continues though not without its difficulties. The 2.5 MW Technology Demonstrator has been under test at DERA Pyestock with an Alstom Series IGBT PWM converter providing the electrical supply. Vibration at 33 Hz caused much early difficulty until a current instability caused by an interaction between the motor and converter was identified and rectified. The 2.5 MW machine has been built with external bearings that can be adjusted to investigate the effect of axial misalignment. It appears that the 33 Hz vibration was sufficient to slacken the bearings and allow the rotor to move and touch the stator. The damage to the stator C Cores was sufficient to necessitate an extensive repair. On completion of the repair full power was achieved and accepted by the UK MoD. After achieving full power the insulation on one of the eight coils failed and has again necessitated an extensive repair that is still in progress. It is likely - though not certain - that the insulation failure is associated with the design of the coil lead-outs and the difficulty of applying the same level of insulation during build as the rest of the coil. It is even for consideration that the lead-outs were insufficiently supported to withstand the 33 Hz vibration that occurred early in the testing programme. The design of the coil lead-outs is being re-assessed as part of the current repair programme. Because the 2.5 MW TFM, once built, exhibited a lower reactance than expected it proved impossible to achieve the design rating of the machine and although full power has been achieved and agreed at 2 MW this is less than originally anticipated. At that power the machine’s air gap shear stress has only reached 80 kN/m2 rather than the 100 kN/m2 necessary to design a 65 tonne 20 MW 180 RPM machine. Nevertheless the machine has demonstrated that the topology is not only viable but offers significant advantages in terms of torque density. As the specific torque of the TFM is increased the power factor of the machine reduces. For the 20 MW 180 RPM application being developed for the UK MoD the power factor is currently anticipated as being 0.6. This relates to an air gap shear stress of 100 kN/m2 however the machine could be designed for an air gap shear stress of 120 kN/m2 but with a consequent reduction in power factor to 0.4. A power factor of 0.4 implies a 50 MVA converter for a 20 MW TFM. With the converter technology currently available this requirement to provide 50 MVA reduces system efficiency unacceptably because of the increased switching and conduction losses. Nevertheless the TFM topology has a greater potential for torque density than is currently being exploited in the MoD development, all that is necessary to gain further size reductions – perhaps to as low as 40 tonnes for a 20 MW 180 RPM machine - is a more efficient form of power electronic conversion. The authors believe that this will arise in time as techniques such as resonant conversion and multi-level conversion are developed. It is of interest to note that the TFM topology is being actively considered for use in wave energy generation schemes where researchers anticipate utilising a linear TFM with an air gap shear stress of 120 kN/m2. AXIAL FLUX Jeumont Industrie is developing an Axial Flux PMPM under French Defence Ministry funding. The following tables show the principal parameters of their current machine development. The authors are grateful to Jeumont Industrie for permission to publish this information. Rated voltage V between ( tbd ) phases Phase Current A (8 disc 8 × ( tbd ) sides each fitted with 3 phase windings) Frequency Hz < 100 Number of discs 4 Power factor at full load 0.8 Efficiency @ 100% speed @ 80% 97 % 97.2 % @ 60% @ 40% @ 20 % 97.2 % 96.8 % 93 % Table 1: Jeumont Axial Flux PMPM Parameters As can be seen from the table the development is still at a relative early stage, nevertheless Jeumont remain confident of developing a suitable Axial Flux PMPM for naval applications. The next Table lists the principal dimensions of their proposed machine and compares them directly to a traditional synchronous machine. Overall weight External diameter Overall length Axial field magnet 65 tonnes motor 2.7 m 3.0 m Wound synchronous motor 5.6 m 3.6 m 120 tonnes Table 2: Jeumont Axial Flux PMPM Dimensions It is of interest to note that the Jeumont Axial Flux PMPM and the Rolls-Royce Transverse Flux PMPM, at their current state of developments have identical predicted mass for a 20 MW 180 RPM machine: 65 tonnes. INDUCTION The ALSTOM Advanced Induction Motor has been selected for the Type 45 and the Electric Ship Technology Demonstrator and represents a remarkable achievement in terms of power and torque density. It has not been previously reported in this series of papers and the authors are grateful for ALSTOM’s permission to include the following details of the machine. Air-gap shear stress 100 kN/m2 Power factor at full-load > 0.8 Overall weight 70 tonnes External diameter 2.8m Overall length 3.0m Efficiency @ 100% speed @ 80% @ 60% @ 40% @ 20 % 97 % 97.1 % 95.5 % 93 % 80 % Table 3: ALSTOM Advanced Induction Motor Parameters ALSTOM Advanced Induction Motors have been designed and manufactured at a variety of ratings and phase numbers since their development for industrial applications in the early 90’s. The machines are fully compliant with Noise and Vibration, EMC/EMI and Shock requirements in accordance with the appropriate NES or DEFSTAN. Typical parameters for an H-bridge, PWM inverter fed, 15-phase motor for delivery in 2004 are shown in Table 3 and the construction of a typical advanced induction motor is shown in Figure 2. Figure 2: Alstom Advanced Induction Motor OPTIMISED CONTROL OF ELECTRICAL MACHINES Dr Chris French of Newcastle University has been working on a novel and much improved method of controlling motors with power electronic converters. This work was reported at the All Electric Ship Conference 2000 in Paris and is at Reference 11. The method employed by the researchers maximizes the full potential of an electric machine by optimizing the motors control as applied by the voltage waveform generated by its associated converter. The technique utilizes the magnetic characteristic of the machine that is determined during preliminary testing and ensures that the maximum torque is produced for any given input power. The method is generalized for any separately excited synchronous, permanent magnet or reluctance machine – including Axial and Transverse Flux toplogies. The approach is equally applicable to generators where the method then ensures the maximum electrical output for a given machine size; though, of course, the generator would need to provide its output via a converter (as is proposed by the authors for their electric ship power system). The technique ensures that the machine always receives the optimum voltage waveform and a second advantage of the technique is that torque oscillations – that may arise through saturation when a machine is driven close to its magnetic limits (as with the UK MoD PMPM) - can be entirely and predictably removed at all loads and speeds. DEVICES AND CONVERTERS DEVICES The development of power electronic devices has continued albeit not at as high a rate as formerly. The USA Office of Naval Research (ONR) funded Power Electronic Building Block (PEBB) programme previously reported at References 3,4 and 5 has now concluded with a replacement programme - Advanced Electrical Power Systems (AEPS) – continuing the overall development of power electronic equipment for the USA electric ship programme. The main development of power electronic devices has been performed G1 Cathode N Cathode G2 MOSFET G1 G2 Current Voltage P N P Anode Anode Figure 3: MTO Equivalent Circuits by SPCO who now, with the acquisition of the development division of Harris SemiConductors, develop both the monolithic devices (GTO, MTO Thyristor based) and the VLSI (IGBT, MCT, FTO). The main recent success has been the integration of the two technologies – monolithic and VLSI – into one device now termed by SPCO the Super- GTO. The authors are pleased that SPCO have chosen this paper to announce this device publicly for the first time and the previously unpublished short paper, Reference 12 produced by Dr Vic Temple is included in its entirety as an Annex. In order to place the development of the Super-GTO in context the MOS Turn Off Thyristor (MTO), which was described in detail in Reference 3, will be reviewed. The MTO has been developed by SPCO Inc in Pennsylvania USA; an equivalent circuit is given at Figure 3. The MTO follows the style of commutation found with a MOS Controlled Thyristor (MCT), where, referring to Figure 3, the upper of the device’s three junctions is short-circuited by a secondary MOSFET switch. This commutation process can be simplistically considered to be one of conversion of the four layer GTO into a three layer Transistor that then commutates by normal base voltage. Another view would be to imagine the stored charge in the upper middle layer, responsible for the GTO’s continuing state of conduction being drained away through the short-circuiting MOSFET. In either case the need for externally produced, stored and injected current – necessary for a GTO or force commutated Thyristor and part of the external circuits of an IGCT - is removed. The commutation process becomes entirely internal to the MTO. With the MCT these MOSFETs are integrated fully within the structure of the MCT itself. Conversely the MTO employs MOSFETs that are external to the GTO. Figure 4 shows the construction of a Super-GTO. An edited extract of Reference 12 follows to provide an indication of the devices advantages. The hybrid MOS Controlled GTO – or Super-GTO - is a planar processed, fine-line GTO EST lid top layout lid bottom Pebb insulating base or “sled”” MTO lid top with mounted off-FET’s Solder screened copper bonded to “sled” Figure 4: Super-GTO Construction combined with a thinPak lid that simultaneously packages the FET control element. This appears to provide both better GTO function and improved gate impedance. A list of the advantages of this approach are summarized as follows and include: • • • • • • • • • • • No costly single die special handling Many die per wafer is simple, leading to improved yields ThinPak is a lower cost package ThinPak is many time smaller and lighter No dry interfaces and only moderate mounting force needed Very low inductance and resistance with multiple gate and cathode contacts to reduce GTO current non-uniformity 3 times higher switching frequency 10 times (or more) higher cell turn-off current capability Much lower forward drop Very uniform on-state and transient current distribution. Flexible heat removal. CONVERTERS The advantages of multi level converters, as reported at Reference 5, are now being widely recognised this is mainly due to the fact that for the first time IGBTs are available at sufficiently high voltage levels so that series connection is avoided for a converter in the 10s of MW range. Both Ultra PMES in the UK and Power Paragon in the USA are actively developing these converters. In the case of the PMES converter the design is a Neutral Point Clamped topology which effectively uses a capacitor network to create an effective neutral voltage which can be used to provide a three level converter. When this is implemented with the 6.5 kV IGBTs now available series connection is not required. A schematic for a single phase NPC converter is shown at Figure 5. DCPOS IGBT Gate Drive PEC IGBT Gate Drive PEC IGBT Gate Drive PEC IGBT Gate Drive PEC OUTPUT IGBT Gate Drive PEC IGBT Gate Drive PEC IGBT Gate Drive PEC IGBT Gate Drive PEC DCNEG Figure 5: Ultra-PMES NPC Converter Topology The NPC topology can also be applied to a poly-phase arrangement which will have space and weight advantages although, in the marine environment, control of earth circulating currents becomes more difficult. The creation of intermediate voltages to enable multi-level conversion is not limited to the neutral voltage; any other number of intermediate voltages can be developed through a capacitor network. The same technique is used, although on a larger scale, with diodes Figure 6: Power Paragon Diode Clamped Converter being used to clamp circuit points to desired voltages. The topology may then be referred to as a Diode Clamped Converter. The advantage is that with more intermediate voltage levels in play the harmonic distortion of the converter is reduced, as are the overall switching losses since each device switches at reduced voltage. Figure 6 shows a schematic for the Diode Clamped Converter being developed by Power Paragon in the USA. The Multi Port Converter, reported at References, 3, 4 and 5 continues to be developed by SAIC in the USA. The topology was recently recognised by ONR as having particular merit and they have funded a $2M development programme aimed at producing a working prototype of around 300 kW. The programme has only just started and the authors hope to be able to report more on this topology in the future. The UK MoD assessed the use of the converter for application in an Electric Warship power system and it was noted that its inherent flexibility allowed its use to be contemplated in all areas of the electric ship power system. Its inherent capability to integrate disparate electrical power systems offers much to the power system designer. GENERAL HARMONICS Power electronic devices are, by necessity, used as a switch and they have, by design, very fast switching times and provide complete electrical isolation. As such they interrupt current flow virtually instantaneously but, as they usually operate repetitively, they cause significant steady state disturbances to both the current and voltage waveforms of their power supply system; known as harmonic distortion. Harmonic distortion has several deleterious effects including causing insulation system degradation and heating of generators windings - thus a good knowledge of the harmonic burden imposed by all power electronics equipments is necessary when designing systems and specifying equipments. One way of visualising the Electro-Magnetic Interference (EMI) problem caused by harmonics is to undertake a Fourier Transform of the distorted waveform. Generally, there is a large number of odd harmonics based on the power electronic switching frequency, often 2kHz or higher. The lower order harmonics can be difficult to filter and can propagate around the galvanically connected system and can feed into sensitive equipment such as lighting, broadcast and telephony equipments. The higher order harmonics are in the radio frequency and can easily transmit into the ship and, in the case of a warship, can be received at the inputs to combat systems. As a typical combat system has a high gain amplifier at the front end, severe degradation or even failure of the combat system can result. In general, there are military specifications for the quality of the power supply to equipment, which controls this interface. However, the cost and size of suitable filtering to meet these specifications can be considerable and system design must undertake trade-offs to determine where and how best to achieve an acceptable harmonic distortion level. The design of the filtering systems is never simple and one method becoming available is that of system simulation. If accurate results are to be obtained the problem of simulating harmonics is complicated. Due to the need to accurately take account of multiple distorting loads interacting with each other it is necessary to work in the time domain and at an integration step interval consistent with the switching frequency of the converters. Even with today’s computing power, with increasing system size and numbers of power electronic converters, this rapidly becomes almost intractable and the problem is exacerbated in the marine field where high impedance power systems exhibit high susceptibility to voltage distortion. While it is now feasible – though time consuming - to conduct time domain simulations of limited power electronic esystems the overall scope of the system is still limited by computing power, in addition the wide disparity between the time constants of power electronics, the electro-mechanical generation and distribution system and of course the ship dynamics itself effectively prevents a unified simulation being developed. Indeed this may always prove impossible even with increased computing power dude to the conflict between rounding errors (which prevent small integration steps being used over large time intervals) and the short transient time constants (which force small integration intervals to track rapid changes). As a result it is likely that a dual approach will be required for the simulation and modelling of marine electrical power systems for the foreseeable future. The challenge is perhaps to form open simulation architectures that allow simple integration of the results from time domain simulation using packages such as Power System Blockset into the more traditional electrical power system frequency domain analysis packages such as Viper. In the particular case of harmonic distortion in a marine power system it should be possible to 'map' the harmonic distortion of the power system over a range of operating conditions generated through frequency domain analysis. These results could also be fed into models of the weapon systems' detectors for assessment of military effectiveness. The authors are grateful for the advice and assistance of The Mathworks (formerly Cambridge Control) in the development of this section. PROTECTION As the capability of the power electronics and the intelligence that can be embedded in the control increases, the option of using the power electronics for system protection becomes more practicable. The issues involved are seen as galvanic isolation and losses. Power electronics can fail short circuit and conducting maintenance where only silicon isolates the maintainer from a potentially lethal power supply is undesirable. The solution is relatively straight forward as off-load isolators can provide the galvanic isolation, however, these all add cost, volume and weight to a system. This impact can be minimised by allowing a larger proportion of the system to be de-energised when maintenance is underway and in many respects the traditional system using switchgear resolves this. The other facet of power electronics is that when a silicon junction is conducting physics demands that there is a volt drop across it. This in turn means that power is dissipated across the junction whenever the device is turned on. At the current ratings of devices with a full duty function, these losses are significant, as much as 2% of the power being handled in a particular application. These losses are allowed for when designing the equipment and are acceptable for such applications as Converters where the functionality they offer is essential. However, to introduce another device in series such as solid states switchgear, can double the losses adding significant inefficiency to the system. In principal then, traditional switchgear, which has very little volt drop when conducting and good galvanic isolation when open, remains the preferred choice for purely protection functions. However, hybrid switchgear based on solid-state devices is being developed and will become available at some stage. These are likely to use power electronics to make and break the supply followed up by a mechanical conductor to minimise conduction losses. Once available one of the disadvantages of electrical DC distribution will vanish as the difficulty of extinguishing the arc drawn when DC contacts open will have been resolved. SECURITY OF SUPPLY In today’s world, even outside of essential military applications, losses of electrical supply are becoming ever more unacceptable. Indeed, the situation is exacerbated by the fact that even a very short interruption or disruption of the supply waveform can cause computer based systems to crash. Even if single generator operation is not adopted with a particular Electric Ship system there is a need to thoroughly address security of supply. The UK MoD Electric Ship Technology Demonstrator is assessing a Zonal Power Supply Unit concept, which has been outlined in previous papers and is based on energy storage. The US Navy IPS programme has successfully demonstrated a similar capability to ensure the continued provision of high quality supplies despite severe disruption from failures or action damage – known as ‘fight through’ and this will be discussed later in the paper. This concept is based on immediate switching between two separate distributed supplies. AES on the World Stage GENERAL Electric warship concepts are now being developed by several nations. The USA has a fully funded development programme aimed at proving the technology for their 1100 VDC Port 4160 VAC Zone A PMM-1 PCM-4 900 VDC PCM-1 Swbd PCM-2 860 VDC PMM-1 Stbd 4160 VAC Figure 7: USA IPS Schematic PCM-4 Zone B PCM-1 860 VDC PCM-2 900 VDC PCM-1 1100 VDC PCM-1 implementation of an electric warship – this is discussed below – and have stated that their new surface combatant, DD21, will have an integrated electric propulsion system. As reported at Reference 13 the UK continues to further develop the electric warship concept through their Electric Ship Technology Demonstrator and the Type 45 class of destroyer will have an integrated full electric propulsion system. In addition France, The Netherlands and Germany are developing equipment for an electric warship with Italy, Spain and Japan, at least, undertaking concept studies. THE USA As has already been stated in this paper the concept of the Electric Warship is now receiving significant attention in many countries. The USA is one of the leaders and the authors are pleased to be able to report on their recent activities in this paper. The Electric Warship Programme in the United States of America has two distinct facets. The acquisition programme office for the replacement surface combatant, DD-21, has announced that Integrated Electric Propulsion will be used in the new warship. In addition the Naval Sea Systems Command has an active development programme working towards implementing not only integrated propulsion and power systems (IPS) but also of a topology and control regime that allows, in their terms, “Zonal Fight Through”. A system exhibiting Zonal Fight Through will be resilient to damage and selfhealing to the extent that zones on either side of the damaged section will suffer neither interruption of degradation of their power supply. It is important to note that although the DD-21 acquisition programme office and the Naval Sea Systems Command IPS team both refer to Integrated Propulsion Systems the system actually referred to in each case is different with the acquisition programme office’s definition of IPS being much more loosely defined. Indeed it is possible that the DD-21 final IPS system will not be that being developed by the Naval Sea Systems Team. PGM Power Generation Module PDM Power Distribution Module PCM Power Conversion Module PMM Propulsion Motor Module ESM Energy Storage Module PCON Power Control Module Notes: • Where two differing equipments with the same function are used they are discriminated by their numbering: hence PCM-1 and PCM-4. • The PMM includes the Propulsion Motor and its converter. • Not all the modules referred to in this table appear in the current IPS scheme – ESM for example. Table 4: USA IPS Component Designations A diagram of the Naval Sea Systems Command IPS concept is at Figure 7 and shows two zones though in practice there would be more, perhaps as many as six. The design and operation of the IPS system is significantly different to the IFEP system originally proposed by the authors at Reference 1 and so some explanation is worthwhile. The system is conceived as having three 20 MW Gas Turbine Alternators and this increase in propulsion power over the original UK concept is simply due to the different and much larger ship displacement: perhaps 8 or 9 thousand tonnes for the USA ship against 3 of 4 thousand tonnes for the UK. In a similar fashion to the UK concept the propulsion and ships service systems are segregated and AC and DC respectively with power electronics performing the integration. However in the USA system all power generation is performed on the (in the USA case) 4.16 kV AC 60 Hz system and therefore the interconnecting power electronics need not be bi-directional. The diagram at Figure 1 uses the Naval Sea Systems Command nomenclature and abbreviations. These are explained in Table 4. At first inspection the system seems to be inherently inefficient with two Power Conversion Modules in simple series where one would suffice. The answer lies in the desire to implement robust Zonal Fight Through. By using both PCM-4s and PCM-1s it is possible to choose which of the Starboard or Port main 1100 V busbars provide power to individual zones. Thus in Figure 7 the Zone A would receive its power from the Port DC busbar because its PCM-1 output voltage is set at 900 V – higher that the Starboard PCM-1 (860 V). Conversely Zone B is receiving its power from the Starboard busbar. In this way it is possible to control the loads seen by the Port and Starboard busbars whilst at the same time guaranteeing uninterrupted supply should one of the main 1100 V DC busbars fail. It is of interest that the PCM-1 module Full Zone PCM-4 1/2 Zone PCM-1 PCM-1 PMM-1 PCM-2 Load Bank PCM-2 PGM-1 AC Swbd PDM-1 PGM-3 Load Bank Load Bank PCM-4 PCM-1 Prototypical Functionally Equivalent LBES Use only Figure 8: Philadelphia IPS Test System Schematic started as a solid state DC circuit breaker, but was re-used when the principal use became voltage conversion to achieve zonal fight through. This concept is being extensively tested at the USA shore test facility at Philadelphia. This facility is part of the Naval Surface Warfare Centre – Carderock Division (NSWCCD) and is termed the Advanced Propulsion and Power Generation Test Site (APPGTS). The APPGTS has been under development since January 1993 and although originally planned for testing of the US Navy’s Intercooled Recuperated (ICR) gas turbine engine development program, the facility has proven its versatility, having provided a venue for performance testing of the Advanced Turbine System (ATS) compressor for Westinghouse Electric Corporation prior to being converted for testing of the IPS. In order to limit costs the IPS system is reduced in scope and in addition some components are not fully representative of equipment that would be used in a warship. 3000 VOLTAGE 2800 2600 PGM-1 VOLTAGE 5000 3500 4500 2400 3000 4000 2200 2000 LM2500 POWER TURBINE SPEED 3500 2500 1800 3000 1600 1400 1200 2000 2500 2000 1000 800 600 1500 PGM-1 CURRENT 1500 1000 1000 PCM-2 VOLTAGE 400 200 0 SPEED 500 500 0 0 200 205 210 215 Time Elapsed (sec.) 220 225 230 CURRENT Figure 9: Philadelphia IPS Results: Load Transients However there is in general at least one fully specified example of each of the equipment that would be required to implement IPS in a warship. It is of note that the Propulsion Motor Module is an Alstom 15 phase Advanced Induction Motor rated at 20 MW and 150 RPM together with its associated Alstom 20 MW PWM series connected IGBT converter. The initial testing results from Philadelphia have been reported at Reference 14 and it is clear that the system is indeed extremely robust to damage and failure. Testing continues but early results are extremely encouraging. The total harmonic distortion levels are higher than the original target but nevertheless the harmonic performance is reasonable and the original target for total harmonic distortion was extremely demanding. A schematic of the Philadelphia IPS test system is shown at Figure 8. The testing to date has included efficiency measurements and system transient responses. Perhaps the most remarkable result is the stability of the in zone 450 V 60 Hz three phase supply during loss of one 1100 V DC busbar – no transient appears on the converted supply. Figures 9, 10 and 11 show some of the results obtained for the Philadelphia test programme. TOTAL HARMONIC DISTORTION (%) 45 ORIGINAL PREDICTION REVISED PREDICTION MEASURED DATA 40 35 30 25 20 15 10 5 0 0 20 40 60 80 PERCENT OF RATED POWER Figure 10: Philadelphia IPS Results : Harmonic Distortion As can be seen in Figure 9 the voltages remain stable during load shedding but most notably the Gas Turbine speed is controlled satisfactorily during the load reduction. Figure 10 illustrates the level of harmonic distortion present in the 4.16 kV system and 100 PERCENT EFFICIENCY 90 80 70 60 50 40 30 20 10 0 0 20 40 60 80 100 PERCENT POWER Figure 11: Philadelphia IPS Results : System Efficiency Figure 11 the variation of efficiency with system load and this is particularly noteworthy as it demonstrates the way that the overall system efficiency for an Integrated Electrical Propulsion System remains high throughout the majority of the load range. The following tables give a more analytical assessment of the success of the IPS system. 100 CRITERIA Frequency Level Frequency Droop Voltage Level Voltage Droop Current Harmonics (PGM-1 design limit) Voltage Harmonics (PCM-4 design limit) NOMINAL VALUE 60 HZ 3.3 % @ rated power 4160 V 3.0 % @ rated power NA NA TARGET RESULTS +/- 5 % +/- 1 % + 0 % / - .1 % 3.3 % @ rated power +/- 10 % +/- 5 % 291 A or 8 % of 5th IHD @ rated power +2%/-0% Within – 1.5 % Over power range 185 A or 7.2 % @ 5th IHD Any IHD < 8 % THD < 10 % 11th IHD @ 8.9 % THD @ 18.5 % TABLE 5: Philadelphia IPS Results : Steady State Main Power System Interface Goals and Results CRITERIA Frequency Level Frequency Response Voltage Level Voltage Response NOMINAL VALUE 60 HZ NA TARGET RESULTS +/- 10 % 5 seconds +5 % / - 0 % 1 second 4160 V NA +20 % / -15 % 1.5 seconds +5 % / - 0 % .6 second TABLE 6: Philadelphia IPS Results : Transient Main Power System Interface Goals and Results CONCLUSION Whilst conclusions should be drawn solely from the paper to which they refer, the authors feel that in this particular case they can be a little wider as it is worth taking a brief review of the electric warship changes over the period spanned by this sequence of papers. At the time of the first Electric Warship paper in 1996 the UK expectation was that the anti-air warfare destroyer, now known as Type 45, would be combined diesel and gas propulsion, both the AO and LPD Replacement would be diesel mechanical, the future surface combatant would be combined diesel electric and gas propulsion and that the brand new engine under development for the US and UK, the WR21, would be supplied as a direct drive geared propulsion engine. Abroad, although the US IPS development was underway it was not related to a ship programme and all future USN surface warships and auxiliaries were to be mechanically propelled. France Holland and Germany were in similar positions with some electric propulsion activity underway but little expectation of shifting away from mechanical propulsion. The UK Marine Engineering Development Strategy was being formed with one of the main legs being electric propulsion and another being adoption of advanced cycle gas turbines. The only signposts presaging the change that has occurred were the commercial marine arena and, perhaps, the hugely successful Type 23 combined diesel electric and gas propulsion system. There has been a remarkable change over the period and this past year has possibly been the most dramatic. The WR21, now a US, UK and France tripartite programme, has completed development and the first orders are for application as generator engine. The Type 45 has selected electric propulsion, DD21 has been committed to electric propulsion, the IPS testing at Philadelphia has reported some stunning results and the UK Electric Ship Shore Technology Demonstrator programme has become a UK/French contract and will be producing results within 18 months to support the second generation of warships with electric propulsion. However, there is another strand to the UK Marine Engineering Development Strategy; the electrification of auxiliary systems and services where cost effective. Similarly, the US IPS programme had other goals besides electric propulsion, one of which has matured into the ‘fight through’ capability outlined earlier. Furthermore, the authors have always believed that widespread electrification of auxiliaries will prove cost effective and that such an approach can and must be harnessed to provide significantly improved war fighting capabilities. Although, inevitably, changes in submarine propulsion are not well publicised, there is sufficient information in the public domain to be certain that the electric propulsion option is being considered by several nuclear powered submarine owners. The first conclusion is thus that electric propulsion is likely to become the norm for warships of the future. Surveying the situation now, the authors remain enthused by the achievements around the world in electric propulsion. However, they remain doubtful that the totality of the Electric Warship is yet within reach. The increased war fighting capability and the reduced costs that should come with wider electrification has yet to be grasped. It will not be easy, as the gains need a bigger step into the unknown than electric propulsion and there are few commercial applications to indicate the way. That said, there are activities underway and the remarkable achievement of the security of supplies offered by the IPS fight through shows what can be done. All in all, it has been a remarkable year and the authors look forward to the next one. ACKNOWLDGEMENTS The authors are most grateful for the support they have received from many companies during the preparation of this paper. The following list records our gratitude: Alstom Power Conversion BMT FKI Innogy Technology Ventures Ltd Jeumont Industrie Lloyds Register L3 Communications - Power Paragon Rolls-Royce Kamewa SAIC SPCO The Math Works (formerly Cambridge Control) Ultra Electronics PMES USA DoD Naval Sea Systems Command USA DoD Office of Naval Research In addition the authors wish to record their gratitude to the many individuals without whose encouragement and help this paper would not have been completed. The list ranges form understanding Rolls-Royce seniors to their many academic and industrial friends – too many to mention individually but none have been forgotten. REFERENCES 1. Cdr C G Hodge and Cdr D J Mattick, ’The Electric Warship’ Trans IMarE, Vol 108, Part 2, The Institute of Marine Engineers (1996). 2. Cdr C G Hodge and Cdr D J Mattick, ’The Electric Warship II’ Trans IMarE, Vol 109, Part 2, The Institute of Marine Engineers (1997). 3. Cdr C G Hodge and Cdr D J Mattick, ’The Electric Warship III’ Trans IMarE, Vol 110, Part 2, The Institute of Marine Engineers (1998). 4. Cdr J M Newell, Cdr D J Mattick Royal Navy and C G Hodge, ’The Electric Warship IV’ Trans IMarE, Vol 111, Part 2, The Institute of Marine Engineers (1999). 5. C G Hodge and D J Mattick, ’The Electric Warship V’ Trans IMarE, Vol 112, Part 2, The Institute of Marine Engineers (2000). 6. R G Blakey and S C Mason, ‘A developed scheme for high power electric propulsion’, Proceedings AES 2000, SEE France (2000). 7. SD Sudhoff, S.F. Glover, P.T. Lamm, D.H. Schmucker and D.E. Delisle, Stability Analysis of DC Power Electronics Based Distribution System Using Admittance Space Constraints, IEEE Trans (2000) 8. SD Sudhoff and S.lover’Three-Dimensional Stability Analysis of DC Power Electronics Based Systems, accepted for publication by IEEE (2000) 9. A J Donaldson and R C Galloway, ‘Zebra Batteries for Marine Applications’, Proceedings AES 2000, SEE France (2000). 10. http://www.regenesys.com/brochure_FSET.htm 11. Chris French, Tony Roskilly and Paul Acarnley, ‘Maximising the Power:Weight Ratio of Marine Electric Propulsion and Generation Systems’, Proceedings AES 2000, SEE France (2000). 12. Dr. Vic Temple, “Advanced GTO/MOSFET Hybrids”, Unpublished Short Paper, Annex to Electric Warship 6 13. Cdr J M Newell and Cdr S S Young, ‘Beyond Electric Ship’, Trans IMarE Vol 112, Part The Institute of Marine Engineers (2000). 14. Thomas Dalton, Matthew Stauffer, CDR(sel) Timothy J. McCoy, Edward Harvey, “Initial Testing Results Of The Integrated Power System At The Full Scale Land Based Engineering Site (Rev-F)”, Association of Scientists and Engineers 37th Annual Symposium, 10 May 2000. ANNEX to Electric Warship 6 Dr. Vic Temple, Senior VP R&D, Silicon Power Corporation, e-mail [email protected] Advanced GTO/MOSFET Hybrids Several years ago SPCO described a GTO whose turn-off was effected by discrete FET chips packaged inside a PressPak packaged GTO which could be gated to effectively short the GTO gate and cathode. Both 53 and 75 mm devices have been produced by a team headed by Dante Piccone at SPCO, Malvern, PA which successfully demonstrate this principle at voltages as high as 9000 volts. The value of such an approach is that it greatly reduces the cost and size of the turn-off gate drive. Compared to an MCT, in which the FET is built into every cell, the advantage is simply an ability to yield larger area die than a MOS gated device is yet capable. In principle, one can think of this as producing a high voltage MOS gated device without taking all of the silicon through both the FET and the HV GTO processes. SPCO used, for its initial products, commercial GTO’s with very minor changes and paralleled inexpensive commercial FET chips or surface mount packaged devices, each of about a 5 milli-ohm rating. In principle sufficient FET’s can be paralleled to turn off any current. In practice, the amount of current that can be turned off is dominated by the worst GTO finger to gate FET circuit inductance, which, for a practical device must be in the sub-nanohenry region. If the inductance can be made to approach zero then the turn-off is dominated by 1) GTO finger current uniformity and 2) the finger width and upper base sheet resistance under the GTO fingers. In order to increase the theoretical maximum current that can be turned off various researchers, including one of the authors of this paper, have experimented with finger sizes of the order of those found in planar transistors, ie, instead of 20 to 30 mil finger repeat distances 2 to 3 mils was proposed. The difficulty here is that this small geometry was not compatible with the traditional closed tube approach in which the fingers were defined by a deep etch and the cathode pressure contacted. SPCO has found solutions to these problems and is in the process of building >4.5KV GTO’s on a 6 inch float zone wafer running an IC foundry whose highest voltage prior device was of the order of 100 volts. Of course, the immediate advantages of small size and fine geometries at extremely large process yields was one of the advantages achieved. In fact we chose to a 2 mil cell repeat distance. The next problem is how to interdigitate high current gate and cathode contacts with very small (microns) separation. The solution here was in part the use of a second metal and in part SPCO’s thinPak package technology which is described in some detail below. The final problem was how to achieve stable, high voltage breakdown voltage without the normal deep, closed tube diffusion. For this SPCO turned to its JTE approach.. Given that the combination of fine geometry, thinPak package and JTE termination worked we would expect a rather remarkable turn-off capability of well over 1000A/cm2. Figure 1 shows how such a device could be incorporated, using a thinPak lid approach to minimize inductance, to make various MOS controlled or gated alternative high performance GTO-like or based devices. Note that a small MCT can be used as the MOS gated turn-on amplifying gate with the unusual advantage that it can remain gated on and will re-gate the GTO as necessary if the current in the main circuit oscillates through zero, for example. Standard MTO MTO Block w/ high performance turn-on Standard IGCT Advanced MTO Block w/ high performance turn-on, ultra-high current turn-off and current sense Emitter Switched Thyristor Block w/ high performance turn-on, ultra-high current turn-off, current sense and current limiting IGCT Block w/ high performance turn-on Figure 1. Advanced hybrid GTO-like devices with MOS control. Figure 2 shows how one might make an MTO power module using one of these “superGTO’s”. The base is aluminum nitride upon which are mounted several S-GTO’s each lidded with a thinPak lid. The thinPak lid top sides are patterned to mount the gate-cathode FET die or surface mount pre-packaged FET’s as the case may be. Also shown is the MCT turn-on device with a series diode that is intended to allow the pilot MCT to run at low current in the on-state, ie on-state MCR/diode drop is to be considerably larger than the S-GTO.. insulating substrate metal matrix heat spreader low R-on FET device metal planar super-GTO power electrode diode (or resistor) thinPak lid high voltage MCT Figure 2. Hybrid MTO example. Figure 3 shows the edge regions, including the JTE termination of our high voltage MCT and S-GTO devices. In the figure the JTE zone is the varying doping P- region at the edge of the device whose goal is to allow us to rate at a very high percent of theoretical one-D breakdown voltage and to simultaneously reduce surface field to less than one half the peak bulk electric field in order to enhance breakdown voltage reliability. Figure 4 shows experimental data from an 18 mil thick set of float zone devices with a 4KV one-D theoretical BV. 75% of our tested devices with the optimum JTE implant dose made our 90% design target. The inset table provides an insight into another of the advantages of JTE, namely the much smaller wasted device area which is typically reduced factors of 3 or higher. For example a 5 mil deep one degree bevel uses an edge width of just under 400 mils and achieves 5340 volts (see table) while about 35 mils, 10 times less, are needed for an equivalent BV using JTE. SGTO termination and cell MCT termination and cells Figure 3. SPCO JTE terminated high voltage devices. 5 KV standard edge bevel 5 KV JTE terminated device Ideal BV 6000 V Ideal BV 6000 V 1 degree bevel 5340 V 30 mils 5180 V 2 degree bevel 5040 V 45 mils 5690 V 3 degree bevel 4740 V 60 mils 5890 V Pebb device BV vs JTE implant step dose Breakdown voltage 4,000 75% 3,500 25% 3,000 2,500 2,000 1.5 2 2.5 3 TRP phase 1: thinPak packaged JTE terminated device breakdown voltage results for an 18 mil thick substrate 3.5 JTE dose (e12/cm2) Figure 4. Top: comparison (modeled) between bevel and JTE terminations. Bottom: experimental BV results for 18 mil thick devices as a function of JTE implant dose. One of the key advantages of JTE is that the upper p-base junction need no longer be as deep. In the case of the devices shown in figure 4 we chose a depth of the order of 10 microns, about an order of magnitude thinner than in standard GTO’s. With the cathode finger width being so narrow we were also able to increase upperbase sheet resistance and still expect to turn off very high current densities. The net result is that the upper transistor gain in our S-GTO is very high. This allows us to make the gain of the lower transistor much smaller than usual. The net result is a vastly improved tradeoff between forward drop and turn-off energy. In our detailed modeling turn-off energy is greatly reduced and forward drop decreased. Part of the advantage comes from the very density of the fingers which results in uniform current density through all but the upper 10 or 20 microns of the S-GTO. Forward drops of under 2 volts are easily achieveable at current densities of 100A/cm2, a current density that is very much higher than the rated RMS current densities of standard GTO’s. Figure 5 shows the sub 2 volt forward drops of the 3600 volt devices of figure 4 at 25C and at 125C for un-irradiated devices (slow) and irradiated devices of about half the turn-off loss. Thick, substrate based M’top Pebb diode with sub-1.5V Vf at 100A/cm2 Pebb device forward drop at 25C and 125C Current 400 Dev A 300 25 C 200 Dev B 25 C 100 0 Dev A .75 1 1.25 1.5 1.75 2 2.25 2.5 2.75 125 C Forward drop Device A: no radiation Dev B Device B: 0.2 MR radiation 125 C Active area: 1 cm2 Experimental Phase 1 TRP results for 18 mil thick, ~4KV device. Figure 5. Demonstration of low forward drop at 100A/cm2. The device shown has a 1 cm2 active device area. Status What has been presented above are good arguments for expecting a JTE terminated, thinPak packaged S-GTO to provide superior hybrid GTO-based devices and modules. Presently, SPCO has generated masks and run several lots of S-GTO’s at its California foundry. Before agreeing to contribute to this paper we met the key milestone of achieving high breakdown voltage, ie >4500 volts, on >50% of the tested die on our most recent lot. Figure 6 below shows a completed die. SGTO devices on a 6” Figure 6 34 S-GTO’s on a completed wafer. The metalization pattern that is visible is the top layer of a two metal layer contact scheme. The stripes are alternatively gate and cathode with each stripe making tens of contacts to each of the fingers which run at right angles. The ceramic thinPak lid mates exactly, on its underside, to the device metal. On its top the gates and cathodes are combined into two metal contacts as seen in figure 7 where the lid for a top mounted series FET for an emitted switched GTO is shown on the left and the lid for top mounted MTO FET’s is shown on the right. For the MTO lid, the metal contacts from left-to right are 1) GTO cathode, 2) GTO gate, 3) FET gate, 4) FET gate return, 5) FET gate, 6) GTO gate and 7) GTO cathode. With the 8 FET’s installed, each rated at 4 milliohms and < 2 nH inductance, we would expect to have < 0.5 uH and about 0.5 mOhms of gate-cathode turn-off path impedance for about 2 cm2 of active GTO silicon. If one tracks the various gate turn-off current paths one finds that all parts of all fingers are similar and very short. Therefore, we have the expectation of very uniform current in both on-state and transient condition. This is very different from the mechanically contacted fingers, all of which are contacted at different local pressures at random high points and with varying contact pressures. The lid provides another function in that it provides additional thermal mass that enhances 60 hertz surge current rating. This is because the lid is so well thermally connected to the die. Although the cooling of the figure 2 module is only one side, the module form allows the heat to be spread over a much larger effective area than does, say, the center square centimeter of the pressPak packaged GTO with a corresponding advantage in thermal resistance which more than off-sets (there also are no dry interfaces.) the 2-side cooling advantage of the pressPak. EST lid top layout lid bottom MTO lid top with mounted off-FET’s Pebb insulating base or “sled”” Solder screened copper bonded to “sled” Figure 7. ThinPak lids for MTO and EST hybrids. The designed surface mount FET’s for our initial experiments are shown on the bottom right. Summary This paper has described the advantages of a hybrid MOS Controlled GTO. Its planar processed, fineline GTO combined with a thinPak lid that simultaneously packages the FET control elements appears to provide both better GTO function and improved gate impedance. A list of advantages of this approach are summarized as follows and include: • • • • • • • No costly single die special handling Many die per wafer is simple, leading to improved yields ThinPak is a lower cost package ThinPak is many time smaller and lighter No dry interfaces and only moderate mounting force needed Very low inductance and resistance with multiple gate and cathode contacts to reduce GTO current non-uniformity 3 times higher switching frequency • • • • 10 times (or more) higher cell turn-off current capability Much lower forward drop Very uniform on-state and transient current distribution. Flexible heat removal. Acknowledements: To Steve Arthur for key contributions to high voltage diodes and MCT’s and to Dante Piccone for SPCO’s first hybrid FET/GTO devices. To Clarence Severt (Wright Patterson AFB) and to Terry Ericsen (ONR) for their support of advanced devices and packaging through the TRP and Pebb contract support. To Sabih Al-Marayati and Forrest Holroyd for their key roles in device design and fabrication.