[American Institute of Aeronautics and Astronautics 46th AIAA/ASME/SAE/ASEE Joint Propulsion...

American Institute of Aeronautics and Astronautics

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Evaluation and Optimization of an 8-GHz Microwave Electrothermal Thruster

Daniel E. Clemens1, Michael M. Micci2, Sven G. Bilén3, Jeffrey R. Hopkins4, Jacob H. Blum5, and Christopher A. DeForce4

The Pennsylvania State University, University Park, PA, 16802

Silvio G. Chianese6

Northrop Grumman Aerospace Systems, Redondo Beach, CA, 90278

A modular microwave electrothermal thruster (MET), operating at a microwave frequency near 8 GHz (X-band), was investigated for use with ammonia or simulated hydrazine decomposition products as propellants, under sponsorship of Northrop Grumman. The objective of this effort was to identify the thruster configurations and operating conditions that maximize specific impulse and thruster efficiencies. The effort involved both computational electromagnetic modeling using the COMSOL finite element multiphysics software and experimental characterization with plasma chamber pressure measurements. Configuration parameters varied during testing included antenna depth and shape, injector size, and nozzle throat size. Operating conditions varied included input microwave power and frequency and propellant flow rate. The configurations that achieved the highest specific impulse, thermal efficiency and thruster efficiency are discussed and considered to be the optimal configurations. Using ammonia propellant, thermal efficiencies as high as 75% and specific impulses as high as 540 s were achieved based on chamber pressure rise measurements. While not as high performance as ammonia, compared to previous testing with simulated hydrazine propellant, significant performance improvements have been achieved. A lightweight thruster design and associated vertical thrust stand have been fabricated for future testing.

Nomenclature a = cavity radius At = nozzle throat area c = cold Cd = nozzle discharge coefficient fres = resonant frequency h = cavity height

= mass flow rate MW = molecular weight P = power p0 = stagnation pressure Re = Reynolds number T0 = stagnation temperature TM = transverse magnetic 1Graduate Assistant, now a Propulsion Engineer at Johns Hopkins Applied Physics Laboratory, Aerospace & Materials Science Group, 11100 Johns Hopkins Rd, 21-S186, Senior Member.

2Professor, Department of Aerospace Engineering, 233E Hammond Building, Associate Fellow. 3Associate Professor, Department of Aerospace Engineering, 213N Hammond Building, Associate Fellow. 4Graduate Assistant, Department of Aerospace Engineering, 229 Hammond Building , Student Member. 5Graduate Assistant, now a Propulsion Engineer at Lockheed Martin, Newtown, PA, Member. 6Spacecraft Propulsion Engineer, Propulsion & Combustion Products Department, One Space Park, M2/2500AK, Senior Member.

46th AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit25 - 28 July 2010, Nashville, TN

AIAA 2010-6520

Copyright © 2010 by the American Institute of Aeronautics and Astronautics, Inc. All rights reserved.


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= conical nozzle half-angle = specific heat ratio ε = permittivity = nozzle divergence loss factor μ = permeability χ01 = first zero of the J0 Bessel function

I. Introduction he microwave electrothermal thruster (MET) uses microwave frequency energy to create and sustain a resonant cavity plasma to heat propellant which is expanded through a nozzle. Many MET propellants may be used,

including liquids such as hydrazine and self-pressurizing fluids such as ammonia. Compared to an arcjet, the MET is projected to have similar thrust with >25% higher Isp for a given power processing unit (PPU) input power and propellant flow. The MET has no electrodes in the flow path, resulting in lower thermal losses and less life-limiting erosion. The MET has less complex PPU requirements than Hall Effect or ion thrusters, with already space-qualified traveling wave tube amplifiers (TWTA) applicable for low-power operation and high-efficiency magnetrons appropriate for higher power applications. The MET also has low electromagnetic interference (EMI) and an exhaust divergence comparable to chemical thrusters, rather than the larger half-angles typical of electrostatic thrusters. The MET cavity can be constructed with non-exotic materials with a very low parts count. The MET may also be utilized as a dual-mode thruster, integrated into a hydrazine monopropellant or bipropellant propulsion system with common tanks, ducts, and valves for significant cost and mass savings, while enabling missions that non-integrated electric propulsion and conventional thrusters could not complete. The MET concept, shown in Fig. 1, consists of a circular cross-section metal-walled cavity operating in the

electromagnetic resonant mode. This mode, with field lines shown in Fig. 2 for an empty cavity, is characterized by high electric energy density on the cavity axis at the endplates and in an annular region

circumscribing the cavity mid-plane. One cavity endplate contains a nozzle and the other an antenna. The cavity is partitioned into a plasma chamber and an antenna chamber by a dielectric plate that allows the two halves of the cavity to be maintained at different pressures. A MET firing begins when microwave energy is input into the cavity with propellant in the plasma chamber at low pressures (<~0.25 psia). At low pressures, accelerated free electrons collide with propellant molecules, heating them and stripping additional electrons to create a plasma. Once initiated, the plasma can be sustained at higher pressures with propellant flowing through the choked nozzle. Since plasma initiation occurs at low pressures, the antenna chamber is maintained at relatively higher pressures, preventing plasma formation near the antenna. Propellant is tangentially injected from multiple ports and the free-floating propellant plasma sits in the region of high electric energy density at the nozzle entrance. The propellant is injected tangentially to form a vortex flow inside the chamber, cooling the chamber walls. This

also aids plasma stabilization due to the generation of a radial pressure gradient that helps maintain the plasma centered on the cavity axis. The propellant flowing around and through the plasma is heated and exhausted through the choked converging-diverging nozzle. The mode resonant frequency for a cavity of height h and radius a is given by the following equation, where μ is permeability, ε is permittivity, and χ01 is the first zero of the J0 Bessel function.1

zTM 2 2

res 01011

1

2f a h

(1) As the height or radius is increased, the resonant frequency decreases. Similarly, geometric perturbations in cavity affect the resonant frequency. Inclusion of dielectric plates, introduction of propellant, and temperature increase due to plasma initiation all serve to change cavity permittivity, and

T

Figure 1. MET schematic.

Nozzle

Plasma

Propellant Vortex

Separation Plate

Antenna

Figure 2. mode fields.

Electric Field

Magnetic Field


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thus resonant frequency. These effects require analytical predictions of MET characteristics to be multi-disciplinary, including coupled electromagnetic, plasma-dynamic, and fluid-dynamic models. For a given resonant frequency, there exists infinite combinations of cavity height and radius that satisfy Eq. (1). The ratio h/a determines the cavity electric energy distribution. For maximum performance and efficient propellant heating, it is desirable to have the highest concentration of field energy near the nozzle. h/a must be properly selected to minimize the field strength at the cavity’s mid-plane, while allowing for appropriate propellant flow around and through the plasma. Figure 3 shows the relationship between cavity geometry and resonant frequency, as well as some microwave frequency assignments. MET operation at higher frequencies has a clear effect on reducing the volume, and thus mass, of a thruster. An operating frequency near 8 GHz was chosen for this study because it offers higher energy density in the plasma chamber and less mass and volume than METs operating at the common 2.45 GHz.

The MET has been investigated in laboratories for nearly thirty years.2-16 Most previous investigations have focused on thrusters operating with limited geometric flexibility and magnetron-based PPU that didn’t allow for frequency and/or power tuning. The current Northrop Grumman and Penn State collaboration focused on design, fabrication, and parametric study of an 8-GHz modular thruster over the 100-350 W regime. Additionally, development was begun on coupled multi-disciplinary analysis tools for the MET, with preliminary electromagnetic modeling results completed.

II. Analyses

A. Electromagnetic Modeling Computational modeling was conducted using COMSOL, a finite element analysis software package used for solving multi-disciplinary problems. Using the COMSOL RF module, MET electromagnetic fields were modeled. Figure 4 shows typical COMSOL user interface screens. Figure 5 shows an example of a meshed MET cavity created using the COMSOL drawing interface. Materials were defined to match test hardware using sub-domain settings. Boundary conditions were defined such that metallic walls were perfect electrical conductors. The waveguide entrance was defined as a rectangular port with a 200 W input power. A predefined fine mesh size was used to solve the model. The extended mesh used a total of 75,226 degrees of freedom. The base mesh used 11,185 tetrahedral mesh shapes and 2,850 triangular boundary elements. The Direct (SPOOLES) linear solver was used. Parametric sweeps were used to analyze field distributions over a wide range of frequencies.

Figure 3. Resonant frequency vs. cavity geometry.

0

2

4

6

8

10

12

14

16

18

20

2.0 3.0 4.0 5.0 6.0 7.0 8.0 9.0 10.0 11.0 12.0 13.0

ME

T C

ham

ber

Dim

ensi

on (c

m)

Frequency (GHz)

Cavity Diameter (cm)

Cavity Height (cm)

Industrial, Scientific, and Medical

5.8 GHz

C Band Downlink3.7 - 4.2 GHz

C Band Uplink5.9 - 6.4 GHz

X Band Downlink7.25 - 7.75 GHz

X Band Uplink7.9 - 8.4 GHz

Ku Band Downlink12.5 - 12.75 GHz

Food Heating and Plasma Processing

2.45 GHz


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B. Pressure-Based Performance Calculations Figure 6 shows ammonia and hydrazine equilibrium composition as a function of temperature at a pressure of

14.5 psia. The solutions are similar because in each case, only NH3, N2, H2, N, and H are present in quantities significant enough to affect performance. The degree of ionization is relatively low at temperatures of interest, so ions are not considered for purposes of performance prediction. Over the temperature range of approximately 1800–3960 R at atmospheric pressure, NH3 is fully dissociated, leaving a mixture of H2 and N2 in a mole ratio of 3:1 for pure ammonia and 2:1 for hydrazine. Non-negligible amounts of H and N appear at temperatures of approximately 3960 R and 8100 R, respectively.

To calculate MET performance, mean chamber gas composition and temperature must be known. Hot-fire chamber temperatures are not directly measured, but rather are calculated utilizing measured pressures and flow rates. Properties are calculated with the NASA CEA code embedded in MS Excel spreadsheet via Software and Engineering Associates’ CEQUEL add-in. For a given thruster configuration, a series of measurements with different cold flows are conducted. At each cold flow point, Cd*At (Cd is discharge coefficient, At is throat area) is calculated as a function of Reynolds number (Re). Cold flow propellant specific enthalpy (at chamber conditions) is also calculated. For a given input power, the same flow rates are then run with a MET chamber plasma firing. Throat Re is calculated and a Cd*At is applied via curve-fit. Assuming choked isentropic flow through the nozzle, gas temperature is calculated using eq. (2), where the subscripts h and c stand for hot and cold, respectively, the subscript 0 stands for stagnation property, T is temperature, p is pressure, γ is specific heat ratio, MW is molecular weight, and T0c is taken to be 298 K.

(2)

1 12

0 0

1 10 0

2 1

2 1

h h

c c

h hh dh h h

c dc c cc c

T C p MW

T C p MW

Figure 4. COMSOL interface Figure 5. Meshed MET model

Figure 6. Equilibrium composition vs. temperature for a) NH3 & b) N2H4

0.0

0.2

0.4

0.6

0.8

1.0

0 2000 4000 6000 8000 10000

Mo

le F

ract

ion

Temperature (R)

N2

H

N

NH3

H2

p = 14.5 psi

0.0

0.2

0.4

0.6

0.8

1.0

0 2000 4000 6000 8000 10000

Mo

le F

ract

ion

Temperature (R)

N2

H

N

NH3H2

p = 14.5 psi

a) NH3 b) N2H4


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Thus, at a given mass flow, the ratio of hot-to-cold chamber pressures yields insight on chamber temperature and, hence, specific impulse. Equation (2), along with the appropriate set of equilibrium relations, can be solved simultaneously to determine the mean equilibrium gas composition and chamber temperature of the propellant. Performance can then be calculated assuming quasi-1-D flow of a perfect gas with constant specific heat. Isp, thrust, and specific enthalpy are calculated based upon measured chamber pressure and calculated chamber temperature. The MET nozzles are straight-walled conical nozzles and they suffer divergence losses due to off-axis gas flow. The applied divergence loss factor, λ, is given by eq. (3), where α is the conical half-angle of the nozzle.

1 cos 2 (3)

Absorbed power (Pabsorbed) is calculated by taking the difference between input and reflected power (Pinput – Preflected). Coupling efficiency is the fraction of input microwave power that is absorbed by the plasma. Thermal efficiency is the fraction of power absorbed by the plasma that is transferred to the propellant resulting in an increase in stagnation enthalpy. hhot and hcold are the propellant specific enthalpy during firing and during cold-flow, respectively.

100 (4)

(5)

III. Hardware Description

A. Modular MET The building block, modular MET, shown in Figure 7, has a cylindrical hole bored through its center. The block

was divided into two pieces so that the separation plate could be inserted to create two chambers. The cavity was capped by two conductive endplates, an antenna plate at one end and a nozzle plate at the other. The antenna plate contains an aperture for an antenna to protrude into the chamber and introduce microwave energy. The nozzle, designed for vacuum operation, is machined directly into the nozzle plate. Propellant was tangentially injected into the chamber using removable injectors. A window in the side of the plasma chamber allowed for optical access to the free-floating plasma. The modular MET was bolted to a bell jar, such that plasma exhausted into vacuum through the nozzle while the MET was accessible in ambient conditions. Table 1 shows the various modular components tested during this study.

Figure 8 shows a schematic of the power generation and transmission system utilized. The MET is powered by a TWTA that can provide up to 350 W over 7.9-8.4 GHz. Microwave signal is provided by an HP 8684B signal generator

Figure 7. Modular MET.

Separation plate Antenna

Table 1. Modular MET parameters. Parameter Values

Nozzle Throat Diameter

D1

D2

D3

D4

D5

D6

Injector Diameter

Small

Medium

Large

Antenna Height and Shape

Short Round

Tall Round

Short Flat

Medium Flat

Tall Flat


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with 1 mW output power. Energy is transported to the cavity via waveguide. Just upstream of the MET cavity, a waveguide-to-coaxial transition is used to allow the energy to propagate through N-type microwave connectors. A panel mount connection with a removable center connector is used to introduce the energy in to the antenna cavity. The center connector functions as an antenna.

B. Instrumentation and Control All data was collected via Labview on a Sony Vaio PC. Transient and averaged data are exported to Microsoft

Excel via csv files for additional analysis. Power is measured via HP 8481A thermistors that are connected to an ATM 112-308AD bi-directional coupler. Power readings are then taken by HP 437B Power Meter control boxes. Propellant is metered using UNIT mass flow controllers varying in range from as small as 500 sccm to as high as 30 SLM. These devices are controlled by a MKS Multi-Gas control box that can control up to four controllers at once, which is useful for combining mixtures of gasses for simulated hydrazine testing. Antenna and plasma chamber pressure are measured by Omega PX303 pressure transducers that are outputted to Omega DP25-E control boxes. The modular MET was tested with the nozzle plate bolted to a bell jar. The bell jar is evacuated by use of a Welch 1402 pump.

IV. Results and Discussion

A. Electromagnetic Modeling While a fully coupled multidisciplinary MET analysis model is the ultimate goal, the initial objective was to

model the electromagnetic fields within an empty MET cavity using COMSOL. The first configuration investigated was same as the MET constructed for this study. Relevant geometries and material properties matched those of test hardware. Figure 9 shows a contour plot of the E-field magnitude along with the field lines as well as a plot of the E-field strength at the nozzle inlet. As expected, the regions of a high E-field are concentrated at the antenna and the entrance to the nozzle. This distribution is typical of the electric field configuration. The empty cavity prediction of a resonant frequency of approximately 8.23 GHz is ~0.04 GHz lower than the input frequency that produced maximum performance for the test article. The difference is likely explained by changes in electromagnetic characteristics of the system with high temperature plasma, which is not modeled.

Figure 8. Laboratory MET power generation and transmission system.

Power Meters

Dummy Load

Reflected Power

Power Supply

Forward Power

Three-Port Circulator

Dual Directional Coupler

MET MET

Power Meters

Dual Directional Coupler

3-Port Circulator

Dummy Load

Forward Power

Reflected Power

Signal Generator /

TWTA Power Supply

0 – 350 W7.9 – 8.4 GHz

Waveguide – N Transition

Figure 9. a) MET E-Field at 8.23 GHz & b) Nozzle inlet E-Field vs. frequency.

a)

b)


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B. Pressure-Based Performance Calculations Tests were conducted to optimize the modular MET by parametrically varying parameters including antenna

shape and depth, injector diameter, nozzle throat diameter, and propellant flow rate. Ammonia was the initial propellant chosen because it has relatively low molecular weight and self-pressurizes. Figure 10 shows the modular MET firing into an evacuated bell jar. The first parametric optimization was conducted by varying the shape of the

antenna tip and the antenna’s depth into the antenna chamber. Hemispherical and flat antennas were investigated. Table 2 shows the maximum ratio of plasma chamber pressure (hot-fire vs. cold flow) achieved for 150 W ammonia tests with a D3 nozzle and medium diameter injectors for different antenna configurations. In general, utilizing a shorter

antenna improved performance, as did using a hemispherical tip. However, once the flat antenna length was shortened to a certain point, no additional performance increase was seen. The medium height flat antenna was used for following tests.

The performance is thought to have increased with shorter antenna because the opposing EM node moved towards the nozzle inlet, resulting in plasma located closer to the nozzle inlet that heated propellant more efficiently.

A series of tests using different pairs of injectors were conducted. The data shows that maximum coupling efficiency, maximum thermal efficiency, and maximum Isp are not greatly affected by injector diameter. Typical effects of this variation can be seen in the results shown in Table 3. Plates with embedded converging nozzles of various throat diameters were bolted to the plasma chamber. Table 4 summarizes results of the nozzle investigation. All nozzles attained approximately the same maximum coupling efficiency, regardless of input power or flow rate. Maximum thermal efficiency and specific impulse increases with increasing nozzle diameter. This is likely because the larger nozzles have a smaller fraction of the flow that slips by the plasma and is poorly heated.

The optimized configuration, with a D5 nozzle, medium diameter injectors, and a flat-top medium protrusion antenna was tested over a range of conditions. With input power constant, mass flow was increased incrementally and power and pressure measurements were taken when the system reached steady state. Mass flow was increased until the plasma was extinguished. Coupling efficiency, thermal efficiency, and Isp are shown as a function of plasma chamber pressure in Figs. 11-13.

Table 2. Antenna comparison.

Antenna Height and Shape Phot/Pcold

Tall Round 2.33

Short Round 2.76

Tall Flat 2.65

Medium Flat 2.86

Short Flat 2.85

Table 3. Typical injector comparison results (with D3 nozzle).

Injector Diameter Input Power (W) Maximum % Coupling

Efficiency Maximum %

Thermal Efficiency Maximum

Isp (s)

Medium

100 98.5 70.5 321

150 96.6 65.3 371

200 98.8 65.8 404

250 99.1 66.8 434

Large

100 98.3 73.5 325

150 98.8 68.3 373

200 98.6 67.2 405

250 98.9 66.9 434

Table 4. Nozzle comparison results (200 W, medium injectors).

Nozzle Maximum %

Coupling Efficiency Maximum %

Thermal Efficiency Maximum

Isp (s)

D1 98.7 6.9 308

D2 99.0 25.8 363

D3 98.5 51.7 371

D4 98.8 65.8 404

D5 98.5 70.1 407

D6 99.3 74.7 359

Figure 10. Modular METammonia hot-fire


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All cases achieve > 95% maximum coupling efficiency, independent of input power. To throttle the MET, power and flow rate must change in concert for a fixed nozzle to maintain high coupling efficiency. Theoretically, coupling efficiency is proportional to plasma conductivity, which is maximized when the electron-neutral collision frequency is equal to the microwave excitation frequency. Since collision frequency is proportional to pressure, an increase in conductivity and, thus coupling efficiency, is observed initially as flow is increased following low pressure ignition. However, as pressure increases past a point at which coupling efficiency is maximized, electron-ion recombination losses increase until the plasma is eventually extinguished. Thermal efficiency increases with increasing flow and pressure for a given input power. It is believed that the increase in flow rate and pressure forces the plasma closer to the nozzle entrance, which decreases the amount of propellant that slips past the plasma without being heated to high temperature. Maximum thermal efficiency does not appear to be a function of input power.

Calculated 250:1 expansion ratio Isp increases with increasing input power. Chamber pressure corresponding to maximum Isp varies with input power, increasing with higher input powers. Following low pressure ignition, an increase in Isp with increasing flow rate and pressure is observed due to the increase in efficiency. However, the increase in flow rate at a given input power results in a decrease of propellant specific stagnation enthalpy, so eventually the Isp begins to decrease even though the heating is becoming more efficient.

The configuration of the modular MET used for simulated 50% decomposed hydrazine tests utilized the optimal antenna depth and injector size determined from the ammonia testing. The three nozzles used for this set of testing were the largest nozzles used in ammonia testing. Figures 14-16 summarize the results of the simulated 50% decomposed hydrazine optimization testing, with coupling efficiency, thermal efficiency, and specific impulse shown at several power levels for the D5 nozzle, which again resulted in the highest specific impulses. The coupling efficiencies, thermal efficiencies, and Isp in these results behave in the same manner as the ammonia data. Once again, Isp increases as input power increases. Maximum thermal efficiency and maximum specific impulse observed for the simulated hydrazine tests are less than the maxima observed from the ammonia tests. There is greater uncertainty in simulated hydrazine calculations because

Figure 11. Optimized configuration NH3 coupling efficiency vs. plasma chamber pressure.

70%

75%

80%

85%

90%

95%

100%

2 4 6 8 10 12 14 16 18 20 22 24

Mic

row

ave

Co

up

ling

Eff

icie

ncy

Chamber Pressure (psia)

100 W 150 W 200 W

250 W 300 W 350 W

Figure 12. Optimized configuration NH3 thermal efficiency vs. plasma chamber pressure.

25%

35%

45%

55%

65%

75%

2 4 6 8 10 12 14 16 18 20 22 24

Th

erm

al E

ffic

ien

cy


100 W 150 W 200 W

250 W 300 W 350 W

Figure 13. Optimized configuration NH3 theoretical 250:1 Isp vs. plasma chamber pressure.

250

300

350

400

450

500

550

2 4 6 8 10 12 14 16 18 20 22 24

Va

cuu

m Is

p (s

)


100 W 150 W 200 W250 W 300 W 350 W

Figure 14. Optimized configuration simulated N2H4 coupling efficiency vs. chamber pressure.

70%

75%

80%

85%

90%

95%

100%

2 4 6 8 10 12 14 16 18 20 22 24

Mic

row

ave

Co

up

ling

Eff

icie

ncy


150 W 200 W 250 W


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gas mixture properties are challenging to compute. Also, the data presented is for cold simulated hydrazine, whereas actual decomposed hydrazine will likely enter at some higher temperature.

V. Recent and Future Work Collaboration between Northrop Grumman and Penn State is ongoing. Upcoming testing will be conducted with

a new lightweight MET and vertical thrust stand that have been constructed. Performance will be directly measured for ammonia and hydrazine propellant from 100-350 W over a range of flow rates at the frequencies that provides for greatest Isp. Additionally, the coupled multi-physics model development will continue, to provide future designs with physics-based design criteria.

A. Lightweight MET After successful parametric optimization of the modular MET cavity, the next step was to create a lightweight,

performance-optimized cavity. The motivation for this was twofold. First, in order to get a direct thrust measurement with adequate resolution, a cavity lighter than the modular version was needed. Second, this lightweight cavity would be constructed in a manner more consistent with a flight design, including the choice of materials. Figure 17 shows a schematic of the lightweight cavity. The lightweight MET hardware was designed for operation completely in vacuum, rather than bolted to a bell jar.

Critical design dimensions maintained were the height and diameter of the cavity, the location of the injectors, location of the separation plate, and the location of the antenna aperture. A cylindrical geometry was the basis for the lightweight design. Separate plasma chamber and antenna chambers bolted together with a separation plate sandwiched between were utilized. Brazing or other permanent joining methods could have been used to reduce weight and parts count, but the flexibility to change-out the separation plate was desired. An integrated converging/diverging nozzlewith an expansion ratio of 40:1 and the same throat diameter of the D5 nozzle was utilized. The total MET cavity mass is just over 500 grams. For a flight design, further reductions in mass are possible.

B. Vertical Thrust Stand When operated horizontally in the laboratory, the MET plasma rises due to buoyancy. With an off-axis plasma,

propellant heating is diminished and performance is reduced. Therefore, to get an accurate measure of true MET performance, a vertical thrust stand is required. A simple design of mounting a strain gauge load cell to a stand and placing the thruster directly on top of it was chosen, as shown in Fig. 17. To allow force measurement, a flexible waveguide is used to connect fixed waveguide to the waveguide-to-coaxial transition.

Figure 15. Optimized configuration simulated N2H4 thermal efficiency vs. chamber pressure.

20%

25%

30%

35%

40%

45%

50%

55%

60%

65%

70%

75%

2 4 6 8 10 12 14 16 18 20 22 24

Th

erm

al E

ffic

ien

cy


150 W 200 W 250 W

Figure 16. Optimized configuration simulated N2H4 theoretical 250:1 Isp vs. chamber pressure.

250

300

350

400

450

500

550

2 4 6 8 10 12 14 16 18 20 22 24

Vac

uu

m Is

p (s

)


150 W 200 W 250 W

Figure 17. Lightweight MET.

Separation Plate

Antenna Chamber

Plasma ChamberPropellant Injection Port

Antenna Transition Mount

40:1 Expansion Nozzle

Plasma Chamber Pressure Port

Antenna Chamber Pressure Port


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In order to obtain thrust measurements, a LCAE Platform 1.5 kg Omega bar strain gauge load cell with a resolution of 2 mN was chosen. The strain gauge is hooked up to an Omega DP25-S strain gauge controller. Mounting of the thrust stand onto the vacuum chamber was done via optical breadboard that was placed and leveled inside the chamber. Two versa-strut beams were used to mount a shelf to that the strain gauge was attached to. The versa-strut offered flexibility to reposition the height of the strain gauge if the configuration needed to be modified.

VI. Conclusion The objective of the 8-GHz MET research was to optimize the thruster configuration and characterize the

performance of the optimized thruster under vacuum conditions using ammonia or cold simulated hydrazine decomposition products as propellants. The finite element software package COMSOL was used to model the electromagnetic field within the thruster to determine the optimal chamber geometry and dimensions and the optimal method for coupling the microwave energy into the chamber. Variations in antenna depth, injector size and nozzle throat area were experimentally examined to produce an optimal configuration. This configuration produced the highest specific impulse ever attained by any MET configuration using ammonia propellant, 540 s with 350 W of input power, which is approximately 33% higher than the maximum specific impulse achieved by the 2.45-GHz MET. Thermal efficiencies up to 75% were achieved as well. Electrical coupling efficiencies greater than 98% were achieved for all configurations.

Although specific impulse values attained with cold simulated 50% decomposed hydrazine tests, approximately 350 s with 250 W of input power, are greater than those obtained by previous MET investigations, performance was lower than that of ammonia. There is greater uncertainty in simulated hydrazine performance calculations because gas mixture properties are challenging to compute. Also, the data presented is for cold simulated hydrazine, whereas actual decomposed hydrazine will likely enter at some higher temperature.

A lightweight, performance-optimized MET has been constructed to further evaluate the capabilities of the device. This thruster will be tested over a range of operating conditions using a new thrust stand that has been designed and fabricated for the first ever direct vertical thrust measurements of a MET.

Acknowledgments This effort was funded using Northrop Grumman internal research and development resources.

References 1Balanis, C. A., Advanced Engineering Electromagnetics, John Wiley & Sons, Inc., Hoboken (1989). 2Whitehair, S., Asmussen, J., and Nakanishi, S., “Microwave Electrothermal Thruster Performance in Helium Gas,” Journal

of Propulsion 3(2), pp. 136–144 (1987). 3Micci, M. M., “Prospects for Microwave Heated Propulsion,” AIAA Paper 1984-1390 (1984). 4Balaam, P., and Micci, M. M., “Investigation of Free-Floating Resonant Cavity Microwave Plasmas for Propulsion,”

Journal of Propulsion 8(1), pp. 103–109 (1992). 5Balaam, P., and Micci, M. M., “Investigation of Stabilized Resonant Cavity Microwave Plasmas for Propulsion,” Journal of

Propulsion and Power 11(5), pp. 1021–1027 (1995). 6Sullivan, D. J., and Micci, M. M., “Performance Testing and Exhaust Plume Characterization of the Microwave Arcjet

Thruster,” AIAA Paper 1994-3127 (1994). 7Sullivan, D. J., Kline, J., Philippe, C., and Micci, M. M., “Current Status of the Microwave Arcjet Thruster,” AIAA Paper

1995-3065 (1995). 8Chianese, S. G., and Micci, M. M., “Microwave Electrothermal Thruster Chamber Temperature Measurements and

Performance Calculations,” Journal of Propulsion and Power 22(1), pp. 31–37 (2006). 9Nordling, D., and Micci, M. M., “Low Power Microwave Arcjet Development,” IEPC Paper 1997-089 (1997). 10Nordling, D., Souliez, F., and Micci, M. M., “Low-Power Microwave Arcjet Testing,” AIAA Paper 1998-3499 (1998).

Figure 17. MET thrust stand.

Flexible Waveguide

Load Cell

Mounting Plate

Optical Breadboard

Versa-StrutLightweight MET

Adjustment Plate

Waveguide – N Transition


11

11Souliez, F. J., Chianese, S. G., Dizac, G. H., and Micci, M. M., “Low-Power Microwave Arcjet Testing: Plasma and Plume Diagnostics and Performance Evaluation,” Micropropulsion for Small Spacecraft, M. M. Micci and A. D. Ketsdever (eds), Vol. 147, Progress in Astronautics and Aeronautics, AIAA, Reston, VA, pp. 199–214 (2000).

12Sanfillipo, J., Bilén, S. G., Clemens, D. E., Welander, B. A., and Micci, M. M., “Measurement of Electromagnetic Interference from a Microwave Electrothermal Thruster,” Journal of Propulsion and Power, Submitted for Publication (2007).

13Roos, C. J. A., “Vertical-Deflection Thrust Stand Measurements of a Low Power Microwave Arcjet Thruster,” Master of Science Thesis, Department of Aerospace Engineering, The Pennsylvania State University (2001).

14Diamant, K. D., Zeigler, B. L., and Cohen, R. B., “Microwave Electrothermal Thruster Performance,” Journal of Propulsion and Power 23(1), pp. 27–34 (2007).

15Clemens, D. E., Micci, M. M., and Bilén, S. G., “Microwave Electrothermal Thruster Using Simulated Hydrazine,” AIAA Paper 2006-5156 (2006).

16Clemens, D. E., Micci, M. M., Bilén, S. G., and Chianese, S. G., Microwave Electrothermal Thruster Performance using Nitrogen, Simulated Hydrazine, and Ammonia Propellants,” JANNAF SPS Meeting, Orlando, FL, December 2007.

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