are discussed. A conclusion is drawn that an ISS for the ignition phase of an ETC tank gun is a feasible ... Indeed, mounting a several ton power supply seems less attractive than ... translates to about 100kg of additional weight. This scenario ...
2nd Int. Conf. on All-Electric Combat Vehicle (AECV), Michigan, 8-12 June 1997
A Pulsed Power System for an ETC Tank Gun A. Pokryvailo, M. Kanter and D. Melnik Propulsion Physics Laboratory, Soreq NRC Yavne 81800, Israel
Abstract A study of a pulsed power system for the ignition of an Electrothermal Chemical (ETC) gun for anti-armor mission is presented. It is assumed that 300-400kJ of electrical energy is sufficient to control the ignition of advanced consolidated propellants in order to improve the ballistic process. Systems based on capacitors, inductors and rotating machines are analyzed and compared. Taking into account the system operational performance volume, reliability, safety considerations and cost, the inductive based storage system (ISS) was selected as preferable. Among general issues in implementing ISS technology, that of optimizing the coil-battery volume ratio, transfer efficiency and matching to an ETC load are discussed. A review of opening switch technology is provided. Preference is given to a hybrid, or multistage, switch technology. Other subsystems required for the practical implementation of the ISS in a tank are discussed. A conclusion is drawn that an ISS for the ignition phase of an ETC tank gun is a feasible near-midterm solution. An 0.5MJ ISS, capable of delivering the first round of an eight-shot burst in 0.3s, will occupy 0.6m3 net volume, weighing 1000kg. The anticipated system energy density is 0.85kJ/liter. At higher energy levels, battery-based ISS will offer higher energy density owing to better utilization of the storage inductor.
Introduction Electrothermal Chemical (ETC) propulsion technology is now recognized as a probable near-midterm candidate for gunnery upgrading, while pure electromagnetic launchers are considered as a long-term technology in view of their enormous demand of electric pulsed power. ETC guns embrace two basic concepts, that of a boost, when a sizable amount of electric energy is conversed into mechanical energy during the projectile movement inside the barrel, and that of a plasma-assisted ignition of solid propellant[1]. The first approach requiring several MJ per pulse may be viable for stationary gunnery, but will hardly be feasible for armored vehicles for at least 10 years. Indeed, mounting a several ton power supply seems less attractive than slightly increasing the gun caliber. Plasma ignition, offering controlled burning at a wide temperature range, is much more power saving. An estimated energy demand is several hundred kJ per pulse, with 400kJ defined as a near-term goal. Techniques of implementing this approach are the subject of this work. With regard to electric guns, three energy storage methods are considered suitable for further power compression: the energy storage in electric field of capacitors, in magnetic field of inductors and as mechanical energy of rotating machines. The theoretical energy density ratio of these methods is commonly estimated as 1:10:100[2]. This clearly indicates the attractiveness of the inertial storage. Practical considerations dictates other priorities, such as the technology availability, cost, safety issues, compliance with demands of the firing scenario, etc. Capacitive storage is a mature technology, offering modularity of design and relative flexibility of pulse shaping using closing switches (CS). The best achieved energy density of 3MJ/m3 Aerovox type LM capacitors made it possible to Hammon et al[3] to construct a compact 250kJ PFN module having the overall energy density of 1.2MJ/m3. This appears to be fair enough for mounting onto an armored vehicle. One should be aware of the following requirements and constraints. The transfer efficiency of the LM capacitors is 50-75%, the lower figure relating to higher temperatures. This implies that the installed energy capacity of a PFN may be twice as high as the required energy, namely up to 800kJ for a 400kJ pulse. In addition, capacitors’ thermal management may be obligatory at high firing rate, both to stay within a specified temperature range and to guarantee an acceptable lifetime. Therefore, it
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seems that more conservative but more efficient capacitors having the energy density of 1MJ/m3 are better suited for building high rate PFNs for ETC applications. Capacitive systems need chargers capable of fully charging the storage between shots. At a firing rate of a shot per 6 seconds, a charger designated for a 800kJ PFN must have power of about 140-150kW, which translates to about 100kg of additional weight. This scenario is mild compared to the requirement of firing in one-two seconds after the alert, which demands either a powerful charger (400-800kW--the latter for a one second charge) or using energy storage capacitors capable of withstanding DC voltage for at least one hour. Such capacitors possess energy density that is only fractions of that indicated above, 0.5MJ/m3 being a typical value. Storing energy for prolonged time in capacitors is also undesirable because of safety issues and higher insulation demands. Similar reasoning led Podlesak et al[4] of the ARL to introduce an idea of “Silent Watch” requiring an intermediate energy storage (battery of flywheel) to provide capability for the first round. Thus, the requirement of readiness to the first round may lead to a very substantial increase of the size of a capacitive pulsed power supply. Inertial storage offers a clear-cut advantage of very high energy density, with demonstrated values above 500MJ/m[2,5]. The main problem associated with this method of storage is the safety issue. Indeed, the liberation of stored kinetic energy that is several times more than required for a single pulse, would be catastrophic for the vehicle, unless a massive protective housing is provided. For this reason alone Chrysler has frozen its research program on flywheel technology[6]. For a combat vehicle, implementation of inertial storage is aggravated by rough terrain conditions and possible damage resulting from hostilities. Similarly to capacitive systems, a requirement of readiness for the first round demands a storage that either is in standby or driven in a short time to full revolutions. In the first case, the lifetime and safety are at stake, while for the second one, a powerful motor and an intermediate storage are needed, since the alternator power would be at a premium, if sufficient at all. For several years a Soreq team has been pursuing the concept of inductive energy storage. The underlying philosophy of this determination relies on a potentially high energy density, low cost and a long life of inductive storage, combined with certain benefits regarding its mounting on a moving platform. Well known attractions of inductive storage systems (ISS) are their simplicity, static structure and low voltage prime power serving to improved safety compared with capacitive and inertial storage. A good coupling is achieved in the case of an ETC load[11]. Similar to capacitive systems, an ISS may be built modular, enabling flexible integration. Pulse shaping is possible in an XRAM configuration[11,12]. Inductive storage is especially attractive in view of the short charging time of the inductor that can be considered as a part of the firing sequence, thus creating no delay for the first round. Alternatively formulated, the silent watch capability is an inherent feature of ISS. This appears to be a major advantage over capacitive and inertial storage. A prime power supply of the ISS, i.e., a battery, may be a dual-use source, supplying electricity also to high-power actuators in the absence of firing. Such an approach supports, e.g., a concept of a “jumping tank” and smoothly integrates into the all-electric vehicle ideology. The ISS advance is hindered by the lack of a compact repetitive high current opening switch (OS) and the absence of commercially available high power density batteries. Considerable losses in the coil, or its low charge efficiency, if the coil is not cooled, are also a major disadvantage of ISS. We will continue with a detailed description of ISS and issues of their mounting on a tank. Special emphasis will be given to subjects experimentally verified in our laboratory.
General issues in implementing ISS technology A generic battery-based ISS shown in Fig.1 comprises of four main components, i.e., the battery, the inductor, the OS and the control system. The OS and the CS may be unified physically in one switch, e.g., a set of GTOs, or may exist as two independent units. The basic circuit of an ISS is illustrated in Fig. 2. Upon the CS/OS closure, the inductor L is charged to a desired current following an approximately exponential law1, with the time constant of τ ch = L R + R . b
i
When the OS breaks the current, the energy is transferred to the load. In case of the load malfunction, the coil energy is dumped via an additional CS (not shown) onto an emergency load.
1
Rigorous analysis shows that these parameters change during the both the charge and the discharge of the coil[7].
3
High-Power Battery
Charger
Closing& Opening Switch
Storage Inductor
Rb Emergency Dummy Load
Control
Ich
CS/OS
b a t t e r y
i n R I d u i c t L o r
ETC Load
Vb
Figure 1. Generic block diagram of battery-based ISS
D
RL
VL
Figure 2. Basic circuit of ISS
1. Assessment of the battery-coil volume One of the central optimization issues, that of the relationship between the battery and coil volumes, was treated earlier by Kanter et al[8,9], with an assumption of an ideal matching between the battery and coil internal resistances (Rb=Ri). The analysis for an aluminum room-temperature Brooks-shaped coil showed that the optimum was achieved at the ratio of VoB/VoI=3/2, where VoB is the battery volume, and VoI is the coil volume. Fig. 3a illustrates the total volume of a 0.5MJ ISS as a function of VoB/VoI for the battery volumetric power density of 20kW/liter. Three curves represent the battery volume for different coil charge times, namely, t/τch=1.5, 2, 3, where τch is the time constant of the charge circuit. The volumetric power density of 20kW/liter, with a correction for the packing factor, corresponds to values obtained in our characterization of Bolder thin-metal foil (TMF) batteries[10]. As follows, the battery-coil combined volume of an 0.5MJ ISS may be well within 0.25-0.3m3 for the present technology level.
0.4
100 100MJ
Stored energy 0.5MJ 10 Vo, m 3
Vo, m 3
0.35 0.3
2 t/τch=3
0.25
10MJ 1MJ
1
1.5 0.1
Stored energy 0.1MJ
0.01
0.2 0
2
4
(a)
6
VoB/V 8oI
1
10
sB, kW/liter 100
(b)
Figure 3. a--Total volume of 0.5MJ ISS as a function of VoB/VoI for the battery volumetric power density of 20kW/liter; b-- Total volume of ISS as a function of battery volumetric power density sB. The dotted curve shows the volume limited by the coil tensile stress (60MPa for this graph) The physical meaning of Fig. 3a is that if the coil is small and correspondingly has a high internal resistance per unit inductance, a large battery is required (right knee of the curve), while if the battery is small, a large low-resistance coil must be fitted to it to draw a sufficient current. The curve minimum is rather flat allowing for flexibility in the battery-coil matching. Note that at VoB/VoI> VB is made. At the time t = τ s , the load current iL equals the inductor current V i, since the OS current drops to zero: iL (τ s ) = i(τ s ) = 0 R (τ ) . With Eq. (2.4) this yields the peak load L s I0 RL ( τ s ) at the end of the switching, where the parameter β = τ τ s = L τ s RL ( τ s ) is the ratio of voltage V0 = 1 + 1 2β the discharge circuit time "constant" τ to the switching time. Integrating Eq. (2.2) with is , v s as defined by Eq. (2.3), we obtain the expression for the OS losses which substituted into Eq.(2.1) renders the transfer efficiency
η = ( β + 1 6) ( β + 1 2 )
(2.5)
Note that in the above analysis the only load parameter that matters is its resistance RL ( τ s ) at the end of the switching. 1 0.9 0.8
η
0.7 0.6 0.5 0.4 0
2
4
β
6
8
10
Thus, Eq. (2.5) is applicable for an ETC load, for it is a common practice to characterize such a load by its resistance at the peak load voltage that occurs upon complete current commutation. Defining the discharge time constant and allowing for a certain value of switching losses, one can assess the switching time required from the OS. For instance, if τ =1.5ms, and 10% of the stored energy can be tolerated to be dissipated in the OS, its opening time should be less than 0.5ms (Fig. 5, β=3).
Figure 5. Switching efficiency in an ISS (Eq.2.5) 2.2. Coil discharge efficiency The analysis of the coil losses was performed numerically, considering the intricacy of the problem, and verified experimentally. The results are partially reported in papers[7-9]. We present here only the summary of this work. Two coil designs of a Brooks-shaped coil were examined: a pancake coil (PCC) and a jellyroll coil (JRC) having the same size and the same number of turns. The analysis showed that a substantial part of the stored energy remains trapped inside the conductor, never reaching the load. Accordingly, the discharge efficiency, calculated by Eq.2.1, was found to be 76% for the PCC and 87% for the JRC. The JRC inductance is 10% higher compared to that of the PCC.
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Since a JRC is able to store 10% more energy, it supplies to the load 26% more energy compared to a PCC carrying the same current. Another relation is obtained if the coils are loaded up to the strength limit. According to the stress analysis, the hoop stress, which is dominant, is greater in the JRC for the same current. It follows that with the same mechanical stress, a PCC is able to deliver to the load approximately 10% more energy than a JRC, at the expense of higher losses and higher power drawn from the primary source. Therefore, a conclusion can be drawn that with a powerful primary source, a PCC may be preferable, whereas a JRC is superior to its counterpart when the primary source is not able to charge the coil up to the limit of the mechanical strength. It seems that the latter will be the case for a decade or two. From the above analysis, it follows that the OS losses for β ≥ 3 are relatively small compared to the Brooksshaped coil losses. Thus, for ETC launchers an OS having an opening time of about 0.5ms will be satisfactory. Benefits of a soft switching will be a lower EMI and lower overvoltages impressed on the system components. We conducted extensive experimental work with an ISS described elsewhere[7,13] at the energy level of up to 500kJ that enabled us to verify the above analysis. The switching time varied from 15µs for GTOs to several ms for fuses, with 100µs for an explosively driven OS and 0.25-1ms for a two-stage OS developed by us[14], in between. Both PCC and JRC were used. In several experiments, the PCC and JRC were charged to the same current and discharged onto the same resistive dummy load. For one of the experiments, Fig. 6 shows the experimental load voltage and derived from them load energy curves. It follows that the experimental data agree fairly well with theoretical results.
Figure 6. Load voltage, Vload, and load energy, E, for PCC and JR1 We conducted the high-current testing with a larger 0.26mH jellyroll coil capable of storing up to 2MJ. Low coil resistance of 0.6mΩ allowed the coil current reach 64kA (Fig. 7) which translates to 515kJ of the stored energy. Consistently, the transfer efficiency was 75% owing to smaller losses in the jellyroll coil compared to the pancake one.
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Figure 7. Jellyroll coil charge current and voltage across resistive load. We intentionally skip here the aspects of the coil screening, since the requirements for magnetic field attenuation will be specified later. 2.3. Matching to ETC load It was shown earlier by Kanter et al[11], both theoretically and experimentally, the latter at a low energy level of 10-15kJ, that an inductor discharges onto a capillary in a triangle-shaped current pulse. It can be expressed by the following equation: 11
6 t 6 I L = I p 1 − 11 τ p
(3.1)
where Ip is the peak current at the beginning of the discharge, τ p ≡ L R and RLp is the capillary resistance Lp 6 at peak current. In contrast to a capacitor discharge, the inductor has a finite decay time t d = τ p , which 11 means that within a limited time more energy will be transferred to the load. A more general analysis of capillary operation is provided by Loeb and Kaplan[16]. Limited by acoustical noise, we operated a capillary load with our ISS up to an energy of 250kJ. The system performance closely follows the theoretical simulations based on the works[11,16]. Fig. 8 shows typical experimental load current and voltage traces obtained with a 22cm long, 0.63cm diameter polyethylene capillary installed in a reinforced fiberglass tubing. In this experiment, the energy of 320kJ was stored in the 0.26mH jellyroll coil, of which 240kJ were transferred to the load with an efficiency of 75%. Once more, it was corroborated that ISS are well suited to driving an ETC load.
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Figure 8. Load current and voltage traces of an inductor discharge onto ET load
Opening switch The problem of OS is a major one in pulsed power technology. Some of the following features are required from an OS for a battery-based ISS: •High current conduction for long times •Low conduction and switching losses •Short opening time •High hold-off voltage •Repetitive operation •Precise timing •Low mass and volume •High reliability •Low cost •Easy operation and service For ETC tank gunnery the following ratings have been estimated: •conduction time
0.4 s
•peak (breaking) current
60 kA
•hold-off voltage
6 kV
•opening time (see section 2.1)
