NASA CR-174623

Contract NAS7-918

Task Order RE-65, Amendment 399

Test Bed Ion Engine Development

Final Report

By

Graeme Aston

William D. Deininger

Electric Propulsion and Plasma

Technology Group

Jet Propulsion Laboratory

California Institute of Technology

Pasadena, California

March 1984

Prepared for

National Aeronautics and Space Administration

NASA Lewis Research Center

Vincent K. Rawlin, Program Manager
1 Report No

CR-174623

2 Government Accession No 3 Recipient's Catalog No

4 Title and Subtitle

Test Bed Ion Engine Development

5 Report Date

March, 1984

6 Performing Organization Code

7. Author(s) 8. Performing Organization Report No

Graeme Aston, William D. Deininger

10 Work Unit No

9 Performing Organization Name and Address

Jet Propulsion Laboratory

4800 Oak Grove Drive

Pasadena, CA 91109

11 Contract or Grant No

NAS7-918

12 Sponsoring Agency Name and Address

NASA Lewis Research Center

21000 Brookpark Rd.

Cleveland, OH 44135

13. Type of Report and Period Covered

Final Report

Sept. 1982 - Nov. 1983

14 Sponsoring Agency Code

15. Supplementary Notes

Program Manager: Vincent K. Rawlin,

NASA Lewis Research Center,

Cleveland, OH 44135

bstract A Tegt Bed Ion (XBI) engine has been developed to serve as a tool in exploring the

limits of electrostatic ion thruster performance. This report describes three key ion engine

components, the Decoupled Extraction and Amplified Current (DE-AC) accelerator system, Field

Enhanced Refractory Metal (FERM) hollow cathode and Divergent Line Cusp (DLC) discharge chamber,

whose designs and operating philosophies differ markedly from conventional thruster technology.

Significant program achievements were: high current density DE-AC accelerator system operation at

2

low electric field stress with indicated feasibility of a 60 mA/cm argon ion beam, reliable

FERM cathode start up times of 1-2 sees, and demonstrated 35 ampere emission levels, DLC discharge

chamber plasma potentials negative of anode potential and identification of an efficient high

plasma density engine operating mode. Using the performance projections of this program and

reasonable estimates of other parameter values, a 1.0 newton thrust ion engine is identified as a

realizable technology goal. Calculations show that such an engine, comparable in beam area to a J

series 30 cm thruster, could, operating on X or H , have thruster efficiencies as high as

0.76 and 0.78, respectively, with a 100 eV/ion discharge loss.

17. Key Words (Suggested by Author(s) )

Ion Engine

Accelerator System

Hollow Cathode

Discharge Chamber

19 Security Classif (of this report)

Unclassified

18 Distribution Statement

Unclassified -

Unlimited

20. Security Classif (of this pagel 21 . No. of Pages 22 Price*

Unclassified

* For sale by the National Technical Information Service, Springfield, Virginia 22161

NASA-C-168 (Rev 10-75)
ACKNOWLEDGMENT

The work presented in this report was performed at the Jet

Propulsion Laboratory by members of the Electric Propulsion and Plasma

Technology Group. Those group members who made major contributions during the

conduct of the program were, Allison Owens, Robert Toomath and Lewis Pless.

Their efforts and those of others in the group, who offered many valuable

suggestions during the preparation of this report, are recognized herein.

Graeme Aston

Supervisor

Electric Propulsion and

Plasma Technology Group
CONTENTS

PAGE

1. INTRODUCTION 1

2. PRINCIPLES OF OPERATION 3

2.1 DE-AC ACCELERATOR SYSTEM 3

2.1.1 Ion Extraction Limits 3

2. 1.2 Design Features 4

2.1.3 Electrode Biasing Configurations 7

2.2 PERM HOLLOW CATHODE 9

2.2.1 Design and Operating Features 11

2.2.2 Electron Emission Characteristics 14

2.2.3 Starting Pressure Requirements 16

2.3 DLC DISCHARGE CHAMBER 16

3. TEST BED ION ENGINE 22

3.1 SPECIFICATION AND CONSTRUCTION DETAILS 22

3.1.1 Cooling and Mechanical Design 22

3.1.2 Cathode and Langmuir Probe Location 23

3.2 SUPPORT FACILITY 28

3.2.1 Vacuum Tank and Power Supplies 28

3.2.2 Beam Probe System 32

3.3 DISCHARGE CHAMBER PERFORMANCE 35

3.3.1 V-I Characteristics 35

3.3.2 Langmuir Probe Results 37

3.3.3 Optics Mask Biasing 45

11
CONTENTS Contd.

PAGE

3.4 BEAM EXTRACTION CHARACTERISTICS 46

3.4.1 Test DE-AC Module Design 47

3.4.2 Electrode Impingement Current 47

3.4.3 Extracted Ion Current Density 51

3.5 DE-AC MODULE BEAM QUALITY 54

3.5.1 Divergence Characteristics 54

3.5.2 Beam Current Analysis 56

4. ADVANCED ION ENGINE PERFORMANCE 60

5. CONCLUSIONS 63

6. REFERENCES 65

7. APPENDIX A 66

iii
LIST OF FIGURES

FIGURE PAGE

1 Effect of R on Beamlet Shape 5

2 DE-AC Accelerator System Design 6

3 Effect of Initial Ion Velocity on Child's Law Current

Density 8

4 DE-AC Accelerator System Biasing Configurations 10

5 FERM Hollow Cathode Major Components 12

6 FERM Cathode Starting Characteristics 13

7 Tantalum Thermionic Electron Emission Characteristics ... 15

8a. Stagnation Pressure Measurements: Thick Orifice Plate, Small

Orifice 17

8b. Stagnation Pressure Measurements: Thick Orifice Plate, Large

Orifice 18

8c. Stagnation Pressure Measurements: Thin Orifice Plate, Large

Orifice 19

9 DLC Discharge Chamber and Magnetic Field Geometry 20

10 TBI Engine Mounted on the Vacuum Tank Bulkhead 24

11 Major TBI Engine Design Details 25

12 DLC Discharge Chamber Side Plate Prior to Final Assembly . . 26

13 Interior of TBI Engine 27

14 FERM Cathode in Slip Tube 29

15 Method of FERM Cathode Isolation 30

16 TBI Engine Test Facility 31

17 X-Y Beam Probe Carriage 33

18 Faraday Beam Ion Probe 34

19 TBI Engine Arc Discharge V-I Characteristics 36

20 TBI Engine Discharge Plasma Characteristics: x = 1.90 cm . 39

21 TBI Engine Discharge Plasma Characteristics: x = 4.90 cm . 40

IV
LIST OF FIGURES Contd.

FIGURE PAGE

22 TBI Engine Discharge Plasma Characteristics: x = 7.90 cm . 41

23 TBI Engine Discharge Plasma Characteristics: x = 10.90 cm . 42

24 TBI Engine Mode Change Behavior 44

25 DE-AC Accelerator System Test Module 48

26 Effect of Screen Grid Voltage on Impingement 49

27 Effect of Post Acceleration on Extracted Ion Current Density 52

28 Beam Profiles Used in Beam Divergence Determinations .... 55

29 Complete Beam Scan for Integrated Beam Current Calculations 57

30 Sample Output of Integrated Beam Current Program 58
LIST OF TABLES

TABLE PAGE

1 Effects of Optics Mask Biasing 46

2 One Newton Ion Engine Performance 61

VI
NOMENCLATURE

d screen grid hole diameter, mm

d, focusing grid hole diameter, mm

d accelerator slot width, mm

3.

d decelerator slot width, mm

E extraction gap field stress, V/mm

e electronic charge, coul.

2

g gravitational acceleration, m/sec

I specific impulse, sec

J. discharge current, ampere

2

j. amplified Child's law current density, mA/cm

2

normal Child's law current density, mA/cm

j extracted current density, mA/cm

J0 beam current, ampere

0

H extraction gap length, mm

£ acceleration gap length, mm

3 *

£ , deceleration gap length, mm

X, accelerated ion position, mm

m ion mass, kg

m propellant mass flow rate, seem

-3

n Maxwellian electron number density, m

m

P stagnation pressure, Torr

P total thruster input power, watt

R net-to-total voltage ratio

R screen grid radius of curvature, mm

s

t screen grid thickness, mm

VII
NOMENCLATURE Contd.

t,; focusing grid thickness, mm

t accelerator electrode diameter, mm

3.

t, decelerator electrode diameter, mm

T Maxwellian electron temperature, eV

T ion engine thrust, newton

V screen grid voltage, volt

s

V,. focusing grid voltage, volt

V. accelerator electrode voltage, kV

V, decelerator electrode voltage, volt

VD discharge voltage, volt

V. extraction gap voltage, volt

V~ post acceleration voltage, volt

V..., neutralizer coupling voltage, volt

V „ mask-to-cathode voltage, volt

Vp plasma potential relative to emitter tube, volt

V cathode body potential, volt

V accelerated ion voltage, volt

v ion velocity, m/sec

W DE-AC module width, mm

X. axial position from screen grid, cm

e vacuum permitivity, farad/m

e discharge loss, eV/ion

3

P charge density, coul./m

1 propellant utilization

Vlll
NOMENCLATURE Contd.

rirr, ion engine efficiency

T beam divergence and double ion loss factor

0 electronic work function, eV

ix
SECTION 1

INTRODUCTION

Development of high thrust density ion engines is an increasingly

important technology need as more energetic and reduced trip time electric

propulsion missions are considered. Specific impulse mission constraints

usually impose an upper limit to the ion beam energy, leaving increased beam

current density as the desired approach to raising ion engine thrust density.

This beam current increase can be achieved by making the discharge chamber

plasma density profile more uniform so that more ions are available for

extraction. Presently, ion thruster discharge chambers have relatively peaked

plasma flatness parameters in the range 0.45-0.55. In addition, beam current

can, in principle, be further increased by raising the accelerator system

electric field stress. However, present ion thruster grid system designs

generally cannot be operated at field stresses approaching their breakdown

limit. This results from mechanical considerations which dont allow such

close electrode spacings over the relatively large thruster diameters. As a

consequence of these restrictions, there is an upper limit to the beam current

that can be extracted with these accelerator systems.

This report presents the results of a technology development

program to significantly increase ion engine thrust density. The major

motivation for this work was the recent development of an ion accelerator

system which showed the capability for high beam current density ion engine

operation. While originally initiated as an accelerator system

investigation, the program was expanded to include testing of an oxide-free

metallic emitter hollow cathode and a divergent line cusp discharge chamber.

These latter components were developed as a working plasma source to provide
very high ion flux levels for the accelerator system beam extraction tests.

However, discharge chamber performance was so encouraging as to warrant the

eventual construction of a Test Bed Ion (TBI) engine incorporating the

accelerator system, divergent line cusp plasma containment scheme and metallic

emitter hollow cathode. It is this TBI engine, a description of its component

technologies and its application to testing the limits of high thrust density

operation, that comprises this report.
SECTION 2

PRINCIPLES OF OPERATION

This section discusses the theoretical basis and operating

principles of the Decoupled Extraction and Amplified Current (DE-AC)

accelerator system,,, the Field Enhanced Refractory Metal (FERM) hollow cathode

and the Divergent Line Cusp (DLC) discharge chamber.

2.1 DE-AC ACCELERATOR SYSTEM

As the name implies, the DE-AC ion optical system separates the ion

extraction process from the acceleration and focusing processes. This

separation of functions results in much higher beam current densities, at the

same electric field stress, when compared to conventional ion thruster

accelerator systems.

2.1.1 Ion Extraction Limits

During operation of an ion thruster extraction system the

indication a beam current limit has been reached is sudden and direct

accelerator grid ion impingement. This impingement results from space charge

forces between the ions overcoming the externally imposed electric fields

which guide their motion. However, it is possible to rearrange the imposed

field distribution so that this beam blow up and subsequent impingement occurs

at larger ion current densities. With a conventional ion thruster grid set

this effect may be observed by increasing the net-to-total accelerating

voltage ratio, R, while holding the total voltage, V , constant. Under

these conditions it has been shown that the impingement limited beam current
2 3 increases. ' This phenomenon is depicted in Fig. 1. For any value of R

less than 1.0 the diverging electric fields in the deceleration region act to

enhance space charge blow up of the ion beamlet in the vicinity of the

accelerator aperture. At low R values this enhancement is great (Fig. la) but

decreases as R approaches 1.0 (Fig. lb). However, for R >1.0 an abrupt change

occurs. The downstream electric fields further accelerate the ions and impart

forces that neck down the beam (Fig. lc) preventing impingement. In practice,

conventional ion thruster grid sets experience electron backstreaming at R ~

0.9. Indeed, by definition, operation at R >1.0 for these grid sets is

impossible.

2.1.2 Design Features

The DE-AC accelerator system shown in Fig. 2 operates with its

first, or ion extraction stage, at R >1.0. This is possible because the ions

are post accelerated as they leave the focusing electrode. In the second, or

acceleration and focusing stage, the ion beamlets from the multiple hole

extraction stage are brought to a common focus within the large slotted

accelerator electrode. Beyond the accelerator electrode these ions enter the

third, or decelerator stage, where they are returned to ground potential.

Injecting ions with large initial velocities into the acceleration stage

significantly increases the influence of the externally imposed electric

fields on ion motion within this stage. The space charge beam current limit

of the acceleration stage is thereby increased and is described by a modified

Child's law current density expression which is derived in Appendix A and

which has the following form:
ACCELERATOR

GRID

R =

(a)

1

R«1.0

VA

R>1.0

(c)

Fig. I. Effect of R on Beamlet Shape

5
ACCELERATOR

ELECTRODE

SCREEN

GRID

DECELERATOR

ELECTRODE

BEAM PROBE Y

DIRECTION

PERPENDICULAR

TO PAGE

BEAM

PROBE X

DIRECTION

1. EXTRACTION

STAGE

2. ACCELERATION

AND FOCUSING

STAGE

3. DECELERATION

STAGE

Fig. 2. DE-AC Accelerator System Design

6
_JAO

The quantity j,/JAO is the ratio of current density for a non-zero

initial ion velocity to the normal Child's law current density with zero

initial ion velocity. Here, V. is the injected ion energy in eV and V

the final ion energy, in eV, after post acceleration. Appendix A defines V.

and V in terms of DE-AC accelerator system electrode potentials. Figure 3

shows a plot of JA/JAO

 as a function of the acceleration-to-extraction gap

voltage ratio

 vo^l " Hence, in the DE-AC accelerator system, where

typically ^ /V ~ 10, the current density capability of the accelerator

stage is typically enhanced by a factor of 1.8 because of the pre-accelerated

ions injected from the extraction stage. Consequently, a greater flux of ions

can pass through this stage before their own space charge significantly

affects the externally imposed fields guiding their motion. A further

advantage of this arrangement is that the accelerator and decelerator

electrodes are quite large and so can easily be shaped to minimize direct ion

interception. This latter capability is especially important for the

decelerator electrode because the slowed ions have a high number density and

significant ion space charge as they pass through this electrode.

2.1.3 - Electrode Biasing Configurations

The DE-AC accelerator system, as shown in Fig. 2, has both the

focusing and decelerator grids grounded and therefore requires only two power
100 1000

Fig. 3. Effect of Initial Ion Velocity on Child's Law Current Density
supplies. This system represents the simplest arrangement and was used

throughout this test program. The specific impulse range possible with this

biasing configuration is limited because the screen grid potential, when added

to the discharge plasma potential, determines beam energy. However, with the

addition of a focusing grid power supply, the DE-AC accelerator system can, in

principle, be operated over a much wider specific impulse range as is shown in

Fig. 4. With this more general biasing arrangement, independent control of

specific impulse from 1500 to 25000 sees, (for argon) is possible at any beam

current level, with no accelerator system geometry changes. The low specific

impulse limit is reached when the screen grid potential equals the discharge

plasma potential, while insulator size and X-ray generation considerations

help define the practical upper specific impulse limit.

To specify perveance, for the DE-AC accelerator system, only the

extraction stage potential difference and beam current are used. This is so

because the potentials from the far removed 'accelerator or decelerator

electrodes cannot reach back through the closely coupled extraction stage

grids to significantly affect the ion engine discharge plasma. Such behavior

is different from conventional two or three-grid ion thruster accelerator

systems where ion extraction and acceleration occurs within a common gap

defined by the screen hole plasma sheath and accelerator grid.

2.2 PERM HOLLOW CATHODE

Ion thruster hollow cathodes presently employ oxide impregnated

porous tungsten inserts as the electron emitting structure. These cathodes

have demonstrated very long life (tens of thousands of hours), are well

understood and can be scaled to small and large power levels. However, by the

nature of their design, these cathodes do have certain operating limitations.
•±r DECEL

LOW Isp

INTERMEDIATE Isp

HIGH Isp

Fig. 4. DE-AC Accelerator System Biasing Configurations

10
Start up times are generally of the order of several minutes with arc ignition

by a random keeper coupling process or by drawing a spark from an external

tickler electrode. Additionally, present ion thruster cathodes are fairly

complex devices requiring several carefully manufactured parts. In

particular, careful monitoring of the insert manufacturing and oxide

inpregnation processes is required to ensure repeatable cathode operating

characteristics.

2.2.1 Design and Operating Features

The Field Enhanced Refractory Metal (PERM) hollow cathode was

developed as an alternative to conventional oxide hollow cathodes and exhibits

unique design and operating characteristics which make it attractive for ion

thruster applications. Figure 5 shows the FERM cathode design developed for

and tested with the TBI engine. No insert or resistance heaters are used with

the device. Instead, a voltage, applied between the central emitter tube and

surrounding cathode barrel, breaks down the injected propellant gas to form a

glow discharge. This glow discharge rapidly heats the emitter tube end to

thermionic electron emitting temperatures. The much larger heat capacity of

the cathode barrel prevents this component from heating significantly during

the starting process. When the emitter tube begins emitting thermionically, a

rapid transition from a glow to an arc discharge occurs, resulting in a low

coupling voltage between the emitter tube and surrounding cathode barrel.

Once this arc is established the cathode is on and is a stable, high density

plasma, electron source. To extract these electrons requires only a few tens

of volts positive bias on the ion engine discharge chamber anode.

Figure 6 shows typical volt-ampere starting characteristics for the

FERM cathode shown in Fig. 5. The 1-2 second starting times depend primarily

11
STAINLESS STEEL

PROPELLANT

TUBE THREADED

INTO EMITTER

TUBE BASE

EMITTER TUBE

ISOLATION

FEEDTHROUGH

'

BORON

NITRIDE

INSULATOR

MOLYBDENUM

CATHODE BARREL

WITH TUNGSTEN

ORIFICE PLATE

STAINLESS STEEL

COUPLING FLANGES

TANTALUM

EMITTER TUBE

Fig. 5. FERM Hollow Cathode Major Components

12
400

o 300

LJJ

200

QQ

D

IDC

LU

100

OJ

ffi

0

5.0

4.0

E 3.0

QC

D

O

UJ

59 2.0

cc

LLJ

LU

1.0

0

0.5 1.0

STARTING TIME, sec

1.5 2.0

Fig. 6. FERM Cathode Starting Characteristics

13
upon the heat capacity of the last few millimeters of the emitter tube. Once

started, the minimum power required to maintain a stable arc is approximately

100 watts.

2.2.2 Electron Emission Characteristics

Electron emission currents up to 35 amperes have been extracted

from the tantalum emitter tube, through the 1.0 mm diameter tungsten orifice,

during TBI engine operation. Arc attachment evidence on the emitter tube end

shows that only the last 4-5 mm are involved in the electron emission

process. Assuming emission from the inner and outer surfaces of this tube,

2

the electron current density is approximately 50 amp/cm'" at 35 amperes.

Normally, such large sustained current densities, with metallic emitting

surfaces, would result in melting and rapid evaporation because of the very

high surface temperature required. However, within the environment of the

dense hollow cathode plasma, the strong plasma sheath electric fields adjacent

to the inner tube surfaces effectively lower the work function of the tube

4

metal. This process has been suggested by Seigfried as important in

conventional oxide thruster hollow cathodes and is even more important in the

PERM cathode. The effect of even slight reductions in work function, on

electron current density, is substantial. Figure 7 plots current density for

tantalum, as predicted by the Richardson-Dushman equation, against tube

surface temperature with the surface work function as a parameter. In

addition to the work function reduction, significant tube cooling occurs

because of the energy removal associated with the electron emission process.

Both of these factors are important for successful FERM cathode operation at

high electron emission levels.

14
1000

100

CN

a.

E

CO

z

LU

Q

LU

cc

DC

D

O

z

o

cr

hO

LU

10

0.1

= 4.1 eV

0.950

0.900

0.850

2000 2200 2400 2600 2800 3000

TEMPERATURE, °K

Fig. 7. Tantalum Thermionic Electron Emission Characteristics
2.2.3 Starting Pressure Requirements

An important parameter in successfully starting the PERM cathode is

ensuring that the internal stagnation gas pressure is high enough to be in the

correct region of the Paschen break down curve. For the PERM cathode barrel

and emitter tube geometry, low Paschen break down voltages are realized, for

most gasses, with a pressure distance product of about 1.0 Torr-cm.

Assuming a typical emitter tube-to-orifice plate (or cathode barrel) distance

of 0.5 cm, the required starting PERM cathode gas pressure is approximately

2.0 Torr. Calculations were performed to determine the PERM cathode

stagnation gas pressure in terms of the orifice geometry and gas specie.

However, the results were unsatisfactory because of difficulty in accurately

accounting for the effects of machining irregularities in forming the

orifice. An accurate determination of stagnation gas pressure was achieved by

conducting a series of experiments measuring this parameter for different

orifice diameters, plate thicknesses and gasses. A Kinney Measuvac Type TDI

MK III diaphragm mercury McCleod gauge was plumbed directly into the cathode

barrel to take these measurements. Figures 8a, b, and c plot the recorded

stagnation pressure against flow rate, with a range of gasses, for each

orifice plate tested.

2.3 DLC DISCHARGE CHAMBER

Substantial reductions in ion thruster discharge loss levels have

recently been realized by using high magnetic field strength plasma

containment schemes. The Divergent Line Cusp discharge chamber, shown in

Fig. 9, borrows heavily from this prior work but departs from traditional ion

thruster designs by its pyramid shape, which obviates the need for a back

plate. In addition, the line cusp samarium cobalt magnets are located

16
6.0

P0 ERROR IS± 15%

OF INDICATED VALUE

6 8 10

GAS FLOW RATE, m ~ SCCM

14 16

Fig. 8a. Stagnation Pressure Measurements: Thick Orifice Plate, Small Orifice

L7
o

I

0°

CO

CO

UJ

DC

QO

h-

<

CO

LLJ

Q

O

X

<

O

OF INDICATED VALUE

0

6 8 10

GAS FLOW RATE, m ~ SCCM

14 16

Fig. 8b. Stagnation Pressure Measurements: Thick Orifice Plate, Large Orifice

IS
* 2.0 -

LJJ

QC

D

CO

CO

LJJ

QC

CL

CO

LU

Q

O

X

CJ

OF INDICATED VALUE

0

6 8

GAS FLOW RATE.m

16

Fig. 8c. Stagnation Pressure Measurements: Thin Orifice Plate, Large Orifice

19
HOLLOW

CATHODE

HOUSING

(REAR)

PYRAMID

SHAPED

DISCHARGE

CHAMBER

CROSS SECTION B

CROSS

SECTION A

OPTICS

PLANE

(FRONT)

Fig. 9. DLC Discharge Chamber and Magnetic Field Geometry

20
external to the discharge chamber. The magnetic field distribution along the

chamber inner surface is such that linear cusps are formed between the cathode

location and the chamber exit plane. The magnetic field intensity, at the

cusps, is intense at the cathode plane and decreases toward the exit plane.

By tapering the line cusps in this manner it was also felt that easy cathode

to anode coupling (the entire DLC chamber is at anode potential in Fig. 9)

would be assured because of the large magnetic impedance variation to electron

flow.

21
SECTION 3

TEST BED ION ENGINE

The DE-AC accelerator system, FERM hollow cathode and DLC discharge

chamber consitute the Test Bed Ion Engine. This section describes the unique

engine design and construction methods which resulted in a minimum of

structural components yet retained the versatility of being able to change

critical cathode and accelerator system parameters throughout the test

program. Results of extensive performance tests are presented which provide a

basis for determining the maximum thrust density of this technology.

3.1 SPECIFICATIONS AND CONSTRUCTION DETAILS

3.1.1 Cooling and Mechanical Design

One of the most important considerations in the mechanical design

of the TBI engine was to provide adequate cooling. This requirement was

motivated by a need to operate the engine at high discharge power levels (>1.0

kW) while keeping the samarium cobalt magnets well below their demagnetizing

threshold (~250 C). Further compounding the cooling problem was an

inability to radiate any substantial heat flux through the front of the

engine. This was so because most of the discharge chamber face was to be

masked off so that only small, easily modified, DE-AC accelerator system

modules could be tested. Finally, manufacturing cost considerations dictated

aluminum as the preferred TBI engine fabrication material. All of these

constraints ultimately lead to water cooling of the entire TBI engine as the

most satisfactory way to guarantee low engine operating temperatures. Since

the engine was a technology test bed, water cooling was an acceptable solution.

22
Figure 10 shows the complete TBI engine mounted on the vacuum

facility bulkhead. Although only a small DE-AC accelerator system is shown,

the area available for useful ion extraction, with this engine, is a square

approximately 25 cm x 25 cm. Detailed drawings of the TBI engine are shown in

Fig. 11. The samarium-cobalt magnets were cut and ground to a length of

5.72 cm a height of 1.270 cm and a width of 0.508 cm. These magnets were

magnetized parallel to their 1.270 cm dimension. Cooling water passages were

milled into the discharge chamber side plates and routed to pass by each

magnet location slot. An aluminum thickness of 0.076 cm separated each magnet

pole face from the discharge chamber plasma. With this spacing the cusp

magnetic field flux density was 2500 gauss. The DLC discharge chamber was

assembled from four side plates identical to that shown in Fig. 12. These

plates, the optics mounting flange and the water cooled cathode mounting

block, were inert gas welded into one integral gas tight structure. The

sheathing covering the side plates to seal off the milled water passages and

the water inlet and outlet ports for each plate were salt dip brazed in one

simultaneous operation.

3.1.2 Cathode and Langmuir Probe Location

Since the magnets were external to the discharge chamber the

interior of the TBI engine contained only the PERM cathode as shown in Fig.

13. For diagnostic purposes, a Langmuir probe was also placed in the

chamber. This probe could be moved, axially, from a position slightly to the

rear of the cathode orifice plate to just upstream of the DE-AC accelerator

system. Arc attachment lines on the DLC discharge chamber, in Fig. 13,

clearly define the magnetic cusp placement. Also, as can be seen from this

figure, no cathode keeper or external tickler electrode is required when using

23
SOFT IRON

RETURN

CIRCUIT

SAMARIUM

COBALT

MAGNETS

GRAPHITE

DISCHARGE

CHAMBER

MASK

SMALL DE-AC

ACCELERATOR

SYSTEM

SIDE PLATE

MAGNET

ASSEMBLY

RETAINER

BRACKETS

Fig. 10. TBI Engine Mounted on the Vacuum Tank Bulkhead

24
FOUR MAGNETS

PER LINE CUSP

PATH OF COOLING

WATER CHANNELS

AROUND MAGNETS

20 LINE CUSPS AROUND

OUTSIDE SURFACE OF

PYRAMID SHAPED CHAMBER

SAMARIUM COBALT

MAGNETS, 0.50 cm x 1.27 cm

x 5.72cm LONG

0.16cm THICK

IRON RETURN

CIRCUIT

ALUMINUM

PLATE WELDED

OVER COOLING

CHANNELS

COOLING WATER CHANNELS

MILLED IN 0.635 cm THICK 6061

ALUMINUM ANODE PLATE

DETAIL A

Fig. 11. Major TBI Engine Design Details

25
MAGNET

RETAINER

SLOTS

Fig. 12. DLC Discharge Chamber Side Plate Prior to Final Assembly

26
Fig. 13. Interior of TBI Engine

27
the FERM hollow cathode. Figure 14 shows the cathode removed from the engine

and inside a stainless steel slip tube which permitted the entire assembly to

be positioned at different axial locations within the discharge chamber. The

FERM cathode is shown in Fig. 15 removed from its slip tube but still bolted

to the boron nitride block which isolated the cathode body from engine anode

potential.

3.2 SUPPORT FACILITY

3.2.1 Vacuum Tank and Power Supplies

All TBI engine tests were done in a 0.9 m diameter x 2.1 m long

non-magnetic stainless steel vacuum tank pumped by two 0.25 m diameter silicon

oil diffusion pumps. The cryo traps of these pumps were filled with alcohol

and cooled to -50 C by refrigerated immersion probes. Typical no load tank

pressure was 1.0-2.0x10 Torr with an 8.0-10.0x10 " Torr pressure range

during TBI engine operation. Figure 16 shows the vacuum tank, control console

and accompanying engine power supply rack and flow control system.

Four power supplies were used at different times to start and

operate the TBI engine. The screen and accelerator supplies were Hippotronic

unregulated 0-1 kV and 0-20 kV, 1 kW units respectively. A 0-500V, 5A low

duty cycle unregulated supply was used to start the FERM cathode and was turned

off when the anode discharge was established. The anode power supply was a

laboratory assembled constant current unit with an 80 volt compliance voltage

and a 0-35A current range. No neutralizer was used with the TBI engine since

refractory filament neutralizer tests, made at the start of the test program,

showed no change in ion beam quality or engine performance when the beam was

self neutralized by secondary electron emission from the vacuum tank walls.

28
MOLYBDENUM

CATHODE

BARREL

TUNGSTEN

ORIFICE PLATE

CATHODE AXIAL

POSITIONING

SLIP TUBE ,

' HOLLOW BOLT

FOR LANGMUIR

PROBE ENTRY

IN PUT POWER

LEADS

Fig. 14. FERM Cathode in Slip Tube

29
BORON NITRIDE

INSULATOR

Fig. 15. Method of FERN Cathode Isolation

30
Fig. 16. TBI Engine Test Facility
3.2.2 Beam Probe System

An x-y probe carriage for beam diagnostics was installed in the

vacuum facility as shown in Fig. 17. The probe carriage base consisted of a

square aluminum frame with two parallel aluminum tracks rigidly attached to it

on opposite corners. A motor driven lead screw was attached across the

diagonal of the probe carriage base, parallel to the two tracks. Movement

parallel to this lead screw was defined as motion in the y direction. An

aluminum track was attached to the pillow block on the base lead screw. This

track was orthogonal to the base lead screw and the two tracks on the base.

Bearings attached to this track allowed it to be guided along the parallel

tracks on the base, by the base lead screw. In addition, a lead screw and

motor assembly, similar to those on the base, were attached to this track. A

probe platform was attached to the pillow block on this lead screw and

bearings attached to the platform guided it along the track. Movement

parallel to this lead screw was defined as motion in the x direction. The

whole apparatus was hung in the vacuum facility such that the lead screw

attached to the base was vertical. Gear assemblies attached to each lead

screw drove ten turn potentiometers which indicated the probe position as a

function of output voltage. There was 56.8 cm of travel in the x direction

and 30.0 cm of travel in the y direction.

A single Faraday probe was used for all beam current density

measurements. This probe can be seen within the center of the large disc

evident in Fig. 17. Probe fabrication details are shown in Fig. 18. The

probe ion collection surface and surrounding guard ring were biased at +18.0

volts to repel low energy, charge exchange, ions and prevent the loss of

secondary electrons from the probe surface. An 81% open area tungsten screen

biased at -18 volts, with respect to facility ground, prevented beam plasma

32
Fig. 17. X-Y Beam Probe Carriage

33
PROBE

SURFACE

+18 V

ELECTRON

SUPPRESSION

SCREEN

-18V

Fig. 18. Faraday Beam Ion Probe
electrons from reaching the probe. The large disc attached to this negatively

biased screen and probe body ensured uniform beam neutralization within the

immediate vicinity of the probe.

3.3 DISCHARGE CHAMBER PERFORMANCE

Early in the test program it became apparent that the plasma

property characteristics of the PERM cathode and DLC discharge chamber

combination were unusual. To better understand the observed behavior, a

considerable effort was expended on documenting those plasma property values

most useful in describing the plasma generation and containment processes.

3.3.1 V-I Characteristics

Figure 19 shows typical arc discharge current and voltage

characteristics for the TBI engine. For these data, all of the indicated

argon gas flow was passing through the PERM hollow cathode, which was operated

in the self heating mode. Also, as mentioned earlier, the downstream end of

the TBI engine was almost completely masked over, for this test program, so

that low gas flow rates corresponded to high neutral gas pressures within the

engine.

The V-I characteristics in Fig. 19 illustrate a trend of low

discharge voltages at high discharge currents as the discharge chamber neutral

number density increases (i.e. increasing flow rate). Of equal interest is

the low potential the floating FERM cathode body assumes relative to the inner

electron emitting tube, V . Such low values of V should cause little

backstreaming ion sputter erosion on the electron emitting tube. The V-I

characteristics of the TBI engine discharge plasma in Fig. 19, while very

35
70

o

O

UJ

X

o

CO

±?

o

LLJ

Io

Q.

Q

O

CO

UJ

Q

O

X

CJ

10 -

10 15 20 25 30

DISCHARGE CURRENT, JD ~ amperes

Fig. 19. TBI Engine Arc Discharge V-I Characteristics

36

35
encouraging, do need verification with an unmasked discharge chamber before

meaningful comparisons with other ion engine designs can be made.

No life tests of the FERM cathode or DLC discharge chamber were

performed. Nevertheless, a few tens of hours total operating time were

accumulated on these components and some comments can be made. Erosion of the

FERM cathode orifice appears similar to that of present ion thruster oxide

hollow cathode designs wherein the downstream hole edges quickly become

rounded with little further scalloping taking place. Cathode operation at

high discharge power levels (>1.0 kW) did cause local melting of the emitter

tube end. This component was made of tantalum and the substitution of

tungsten is expected to correct this problem. Although arc discoloration

lines clearly defined the samarium-cobalt magnet locations around the DLC

discharge chamber (Fig. 13), there was no evidence of erosion or degradation

of any portion of this aluminum structure.

3.3.2 Langmuir Probe Results

Approximately one hundred Langmuir probe traces were taken, within

the TBI engine discharge chamber, throughout the duration of the test

program. Three different cylindrical Langmuir probes were used, with tungsten

rod diameters of 1.02, 0.51 and 0.25 mm. All of the data presented in the

next few pages were obtained with the 0.25 mm dia. probe although consistant

results were obtained with the other probes. Power for the probe bias circuit

was provided by an isolated supply made up of 125 alkaline D cells. This

supply provided a probe sweep range of from -50V to + 100V with respect to the

cathode electron emitting tube.

Probe trace analysis generally followed a procedure outlined by

Q

Seattle and was done in several parts. First, the coefficients of a linear

37
plus exponential equation were obtained by curve fitting to the data in the

electron retarding region. The intersection of this retarding region and the

probe trace linear electron saturation region was determined iteratively using

the Newton-Raphson method. Finally, using the thin sheath cylindrical probe

equations of Langmuir, the Maxwellian electron temperature, plasma potential,

Maxwellian electron density, primary electron temperature and primary-toMaxwellian electron density ratio were determined for each probe trace.

For all the probe traces analyzed no significant primary electron

population was found, even with the movable probe positioned close to the

cathode orifice. Clearly, a primary electron population must exist in the TBI

engine but it appears to be less than 5%, of the Maxwellian electron

population, in those regions probed. It was felt that the small primary

electron populations detected were within the error range of this difficult to

determine parameter and so are not presented in this report. For all TBI

engine operating conditions tested the Maxwellian electron densities presented

can be assumed to be equal to the entire plasma electron density.

Figures 20 through 23 show the variation in plasma potential,

Maxwellian electron density and Maxwellian electron temperature with increasing

discharge current as a function of flow rate for four different Langmuir probe

axial positions. For these data the axial position x=0 corresponded to the

location of the DE-AC accelerator system screen grid, while the front of the

cathode orifice plate was located at an axial postion of x=10.9 cm. Data

collection and analysis uncertainties indicated a plasma potential error range

of ±1.0 volt, a Maxwellian electron density error range of ±20% of the

value indicated and a Maxwellian electron temperature error range of ±1.0 eV.

Several things are immediately apparent with the TBI engine plasma

property data in Figs. 20 through 23. The measured plasma potential was

38
I—Nr

J

 m

<8

o

LU

O H LU

DL LU O

CO <

< _J

^

a. c^c

LU I— co o

^ O z 5=

$ LU UJ v

LU £<

n cc

0)

5 10

DISCHARGE CURRENT. JD ~ ampere

15

Fig. 20. TBI Engine Discharge Plasma Characteristics: x = 1.90 cm

39
<8a

£11

>

I

0

-2

-4

-6

-8

-10

8.0

7.0

6.0

5.0

4.0

3.0

2.0

1.0

0.9

0.8

0.7

0.6

0.5

8

6

4

2

0

—N

0

I

FLOW

SCCM

O 5.72

D 7.15

A 9-30

5 10 15

DISCHARGE CURRENT, JD - ampere

Fig. 21. TBI Engine Discharge Plasma Characteristics: x = 4.90 cm

40
^SQo 2

H 2 "o

Z < >

O£ 0r~~ 1^U 1 1 J

Q_ LU d

sPh

£<0

-1 Qj >

Q- QC

Z n

< z . 'E

HO^r ^

_l QC t TLU 1— V) o $ o z "

^ UJ LU y.

<LllQ E

c

-^ LU"

<z§

no2 >

j QC ^ <u

l^I^S>

•* 1-

0

-2

-4

-6

-8

-10

8.0

7.0

6.0

5.0

4.0

3.0

2.0

1.0

0.9

0.8

0.7

0.6

0.5

8

6

4

2

0

FLOW

SCCM

O 5.72

D

 7

-

15

A 9-30

I

5 10

DISCHARGE CURRENT, JD ~ ampere

15

Fig. 22. TBI Engine Discharge Plasma Characteristics: x = 7.90 cm

41
<os

I- Z 75

z < >

0- LU O

Z

>

 E

^ — T

m

, i" "x

0 5 10

DISCHARGE CURRENT, JD ~ ampere

15

Fig. 23. TBI Engine Discharge Plasma Characteristics: x = 10.90 cm

42
always negative, relative to anode potential, with two modes of engine

operation evident. At low discharge currents, plasma potential was negative

of anode potential by 8-10 volts and this difference decreased to about 1-2

volts, negative, with increasing discharge current (it should be noted that

the actual value of the anode potential can be found by referring to the V-I

curves in Fig. 19). The swiftness of this mode change decreased with

increasing propellant flow (or increasing discharge chamber neutral number

density). A rapid increase in plasma density accompanied this mode change

while electron temperature remained generally unchanged. The low electron

temperatures measured are evidence of efficient electron containment within

9

the DLC discharge chamber. Goebel has predicted plasma potentials

negative, relative to anode potential, for low electron temperatures. Under

these conditions, the electron flux to the anode is increased, by an

accelerating potential difference, to maintain the anode power supply current

level.

Figure 24 shows the variation in plasma properties with axial probe

location for two values of the discharge current, one before and one after the

mode change, for the lowest propellant flow investigated. Under all conditions

tested it was observed that the electron density gradient, from the peak value

to the value near the screen grid, always decreased after the mode change

occurred. This decrease is evident in Fig. 24 and suggests a more axially

diffuse ion production region. Because of the TBI engine diverging shape,

this would significantly increase the available anode area. This additional

area would satisfy the anode power supply current requirement and prevent the

establishment of a negative plasma-to-anode potential difference which would

otherwise be required to increase electron arrival rate at the anode. It is

clearly evident, from Fig. 24, that while plasma potentials can be several

43
FLOW = 5.72 SCCM

<z

-10

> o

§ LU

2 -"

<. LU

LU

OC

LU

Q.

LU

"\

10 8 1 i

CATHODE

ORIFICE PLATE

LANGMUIR PROBE

POSITION, X~c m

0

SCREEN

GRID

Fig. 24. TBI Engine Mode Change Behavior

44
volts negative of anode potential, with the TBI engine, much more efficient

operation occurs when the plasma potential and anode potential are more

equal. The merit of this mode change effect in the TBI engine can be

appreciated by observing, in Fig. 24, that a 50% discharge current increase

resulted in approximately a 300% plasma density increase.

3.3.3 Optics Mask Biasing

The large graphite plate, to which the DE-AC accelerator system

module was mounted (Fig. 10), was isolated from the anode biased DLC discharge

chamber. Throughout the test program a 50.0 ohm resistor was used to

electrically connect this mask to the chamber. This biasing configuration was

rhosen because it produced the largest extracted ion current densities. To

understand why this occurred a series of tests were performed with the mask

(and the attached DE-AC module screen grid) biased, in several different ways,

to the TBI engine. Table 1 shows the variation in discharge voltage V ,

mask-to-cathode potential difference V , plasma potential measured relative

y[\j

to the cathode emitter tube V , Maxwellian electron temperature T and

P e

Maxwellian electron number density n , recorded at a Langmuir probe position

3 cm upstream of the screen grid. For these data identical discharge current

and propellant flow rates were maintained.

From Table 1 it is clear that biasing the mask through a 50.0 ohm

resistor to the anode would result in larger beam currents because the

upstream plasma density is greatest. It appears that this configuration

allows some of the higher energy plasma electrons to approach closer to the

mask and so increase ion production in this region. By contrast, with the

mask acting directly as another anode there is a large electron loss rate to

the mask surface. This is evidenced by a plasma potential which becomes

45
Table 1. Effects of Optics Mask Biasing

Bias

Method

Directly to

Cathode

50 ohm to

Cathode

50 ohm to

Anode

Directly to

Anode

VD

(volt)

38.9

39.0

37.2

32.4

VMC

(volt)

0

20.9

24.1

32.4

vp

(volt)

35.7

35.7

36.6

36.4

TAe

(ev)

2.3

2.8

2.6

1.9

n

m 0

(nT3

2.4 x

2.5 x

2.9 x

2.5 x

)

10"

10"

10"

10"

positive of anode potential in an effort to curb the loss rate. In addition,

the large loss of energetic electrons has depressed the local Maxwellian

electron temperature. The small differences noted between the two modes of

mask-to-cathode bias are caused by a small electron flux matching the ion flux

to the mask in the 50.0 ohm bias configuration. This added electron flux to

the mask creates a small increase in ion production and slightly increases

electron temperature upstream of the mask. In summary, it is felt that these

different optics mask biasing effects may explain why increased beam currents

have been observed when ion thruster screen grids have been biased between

cathode and anode potential.

3.4 BEAM EXTRACTION CHARACTERISTICS

In conducting the DE-AC accelerator system beam extraction tests,

obtaining well documented performance trends was of primary importance. These

trends would then serve as a technology base for further DE-AC grid system

designs. To this end tests were performed on the impingement variation,

46
extracted current density scaling behavior and beam divergence characteristics

of the DE-AC accelerator system.

3.4.1 Test DE-AC Module Design

The small DE-AC accelerator system module tested during this

program is shown, in Fig. 25, positioned in the center of the large graphite

plate used to mask down the TBI engine discharge chamber. A hexagonal array

pattern of 65 holes, each 1.905 mm in diameter, were match drilled normal to

the curved screen and focusing grid surfaces. An upstream center-to-center

hole spacing of 2.362 mm was used on the screen grid for an open area fraction

of 0.59. Depending on the screen grid radius of curvature, the open area

fractions of the downstream screen grid face and upstream and downstream

focusing grid faces were successively larger. Both the screen and focusing

grids were machined, using conventional milling techniques, from graphite

plate stock. Early in the test program several complicated geometries, of

both the accelerator and decelerator electrodes, were investigated. However,

superior performance (i.e., lowest ion impingement and best beam quality) was

achieved by using the simple stainless steel rod design shown in Fig. 25. The

insulator mounting arrangements, shown in Fig. 25, were used because they

permitted easy DE-AC grid system geometry changes.

3.4.2 Electrode Impingement Current

Figure 26 shows the impingement current variation for the focus,

accelerator and decelerator electrodes, of the DE-AC accelerator system, as a

function of screen grid voltage. For these data the TBI engine discharge

current and voltage were held constant at 15.0 amperes and 41.0 volts

respectively. The rapid rise in accelerator grid current at lower screen grid

47
BEAM DIRECTION

ACCELERATOR

ELECTRODES DECELERATOR

ELECTRODES

I

ELECTRO

INSULATOR

SUPPORTS

SCREEN AND

FOCUSING GRIDS

Fig. 25. DE-AC Accelerator System Test Module
60

50

cc

DC

D

O

40

Q

O

DC

O

LU

_J

LU

or

cc

o

<

DC

IX

30

20

10

0

-10

«e = 0.56 mm

C = 4.83 mm

Ca = 9.40 mm

Rs = 12.7 mm

ds = df = 1.905 mm da = 6.55 mm

dd = 8.31 mm ts = tf = 0.40 mm

ta = td = 6.35 mm

VA= 12.5kV

FLOW= 8.6SCCM

ACCEL GRID

FOCUS GRID

DECELGRID

14

13

200 400 600 800 1000

SCREEN GRID VOLTAGE, Vs~ volt

Fig. 26. Effect of Screen Grid Voltage on Impingement
potentials indicates direct ion impingement. This behavior shows that the

energy and trajectory of the ions leaving the extraction region are important

in ensuring proper ion collimation in the acceleration section. The increased

impingement current at low screen voltages, with the focusing and decelerator

electrodes, is direct ion impingement. This results from space charge blow up

in these regions because of the high current density and low ion velocity.

Good DE-AC grid system operation was achieved with a screen grid

radius of curvature that was centered within the accelerator electrode. Three

extraction grid sets were tested with the screen grid radius of curvature

having values of R = 1.27, 1.91 and 3.81 cm but using the same module

width, w, and multiple hole pattern. The lowest impingement currents, best

beam quality and highest beam currents were obtained with R =1.2 7 cm. The

onset of direct accelerator electrode ion impingement, for a given extracted

current density, was very sensitive to the screen grid radius of curvature.

It was felt that much lower impingement currents could have been obtained with

a screen grid radius of curvature less than 1.27 cm. Time and resources fixed

R at 1.27 cm in order that the DE-AC module extraction stage characteristics

could be adequately investigated. The other DE-AC grid system geometrical

parameters noted in Fig. 26 were easier to select than the screen grid radius

of curvature and are close to their optimum values. The extraction gap

length, £ , was set at 0.56 mm throughout the test program.

For all beam extraction tests described in this report the baseline

impingement currents were very high because of a large neutral gas flow through

the accelerator system. This behavior was a result of performing beam

extraction tests with a masked down discharge chamber. Because the PERM

cathode requires a minimum gas flow to operate (3-5 seem) the gas flow rate

2

through the 1.8 cm DE-AC module was always very high. As an example, for

50
the results presented in Fig. 26, the milliampere equivalent neutral flow rate

was 32 times the ion current extracted from the plasma. Calculations showed

that the charge exchange ion current, created by this excess gas flow,

accounted for the large baseline accelerator electrode impingement current

shown in Fig. 26. Similarly, the negative focusing and decelerator grid

currents were secondary electrons, created by this large charge exchange ion

flux, bombarding the accelerator electrode.

3.4.3 Extracted Ion Current Density

Figure 27 shows how post acceleration of the ions results in higher

extracted ion current densities for a given extraction gap field stress.

These data were obtained by operating the TBI engine at a constant discharge

current and voltage and by reducing the screen grid potential while monitoring

the change in focusing grid impingement current. The knee of the impingement

curves, obtained in this way, defined the onset of space charge blow up and

were used to determine the field stress values recorded in Fig. 27. These

field stress values were calculated by dividing the screen grid potential by

the extraction gap length £ . Similarly, the values of extracted current

density were calculated by dividing the observed screen grid current by the

combined area of the 65 screen grid holes in the DE-AC accelerator system test

module.

The data curves in Fig. 27 follow a trend where the extracted

3/2

current density varies approximately as E for all V. values.

e A

Typically, post accelerating the ions to 12.5 kV requires about 55% of the

extraction gap field stress that would be required by a conventional ion

thruster accelerator system, with no post acceleration (V =0), to extract a

given current density. If an extraction gap field stress, E , of 2000V/mm

51
12.5

10.0

oc

LU

7.5

QC

(D

oc 5.0

O

o 2.5

0

if)

-z.

LU

O

111

QC

QC

O

Q

LU

O

QC

IX

20

CN

E

,0

<

I

•-* 15

10

0

0

Je

 VD JD

mA/cm2 volts amperes

O 2.3

D 9-4

A 20.0

41.0

40.5

48.5

FLOW = 8.6SCCM

O

10.0

15.0

18.0

= 4.83 mm RS = 12.7 mm

ee = 0.56 mm Ca = 9.40 mm

Cd = 4.83 mm RS =

ds = df = 1.905 mm

dg = 6.55 mm

dd = 8.31 mm

ts = tf = 0.40 mm

to

 = t,j = 6.35 mm

1200

Fig.

200 400 600 800

EXTRACTION GAP FIELD STRESS, Efi

27. Effec t of Post Acceleration on Extracted Ion Current Density

1000

V/mm
is assumed for the DE-AC accelerator system, projecting the data trends of

2

Fig. 27 shows that an extracted current density of 100 mA/cm , for argon,

could be realized with V =12.5 kV. Including a conservative screen grid

open area fraction of 0.6, the projected beam current density would still be

2

60 mA/cm . The open area fraction for the DE-AC accelerator system may have

an effective value much larger than 0.6. This is because the curved screen

grid surface (Fig. 2) presents an ion extraction area to the discharge plasma

that is larger than the area inferred by the module width, W. The largest

electric field stress region in the DE-AC accelerator system, at these high

current densities, is the deceleration region where the post accelerated ions

are returned to an energy equal to the sum of the screen grid and plasma

potential. However, the rod or tubular design of the accelerator and

decelerator electrodes permits reliable decelerator gap operation at field

stress levels considerably in excess of the extraction gap field stress.

2

Results for a maximum extracted ion current density of 20 ma/cm

are shown in Fig. 27. Arc breakdowns between the shadow shields and metallic

(sputter) coated accelerator electrode support insulators (Fig. 25) prevented

attainment of higher extracted current densities. The high baseline fraction

of accelerator electrode ion impingement, discussed earlier, created a

constant source of sputtered material that rapidly coated nearby insulator

surfaces. Since this problem was unavoidable, the most intensive data

collection was at lower extracted current density levels to ensure that the

DE-AC accelerator system performance would be well characterized. From

voltage handling, X-ray limits and insulator size considerations, it was felt

that the accelerator electrode voltage should not exceed about 15 kV. Taking

this as a limit, the required accelerator and decelerator gap field stress

53
2

increases, to handle a 100 mA/cm extracted current density, could be

realized by reducing the overall DE-AC accelerator system size.

3.5 DE-AC MODULE BEAM QUALITY

3.5.1 Divergence Characteristics '

Ion beam quality for the DE-AC accelerator system, shown previously

in Fig. 25, was generally good. Faraday probe ion current plots of the beam

taken 29 cm downstream from the DE-AC grid module in both the x and y

directions, are shown in Fig. 28. For these traces the measured x and y beam

divergence angles were 14.0 and 18.0 , respectively. The larger beam

spread in the y direction is primarily a result of using open slot accelerator

and decelerator electrodes which do not terminate the electric field lines in

this direction. With this configuration, ion space charge forces spread the

beam as it passes through these electrode gaps. Electrodes were built with

bounded slots and beam spreading, in the y direction, was reduced considerably.

However, the versatility of the open slot structure, shown in Fig. 25, in

terms of being able to alter the electrode slot width, favored its use

throughout most of this test program.

For a given extracted ion current density and post acceleration

voltage, there was an extraction gap voltage (i.e. screen grid potential)

which minimized beam divergence. Similarly, at a fixed screen grid voltage

and extracted current density there was an accelerator electrode potential

which miminized beam divergence. With the screen and accelerator grid

potentials fixed, extracted ion current density variations did not

substantially alter the ion beam divergence. This latter behavior was not

54
Vs = 1000 volt

VA = 12.5 kV

JB = 19.3 mA

X DIRECTION

FARADAY PROBE

BEAM PROFILE

Y DIRECTION

FARADAY PROBE

BEAM PROFILE

Fig. 28. Beam Profiles Used in Beam Divergence Determinations

55
true if either the extraction, post acceleration or deceleration gaps, in the

DE-AC grid system, were operated close to their perveance limits.

3.5.2 Beam Current Analysis

Integrated beam current was an important parameter in evaluating

the performance of the DE-AC accelerator system because both the screen and

accelerator electrode power supplies were chassis grounded at their common

voltage. This prevented a single meter reading of beam current. Instead, a

metered beam current was determined by subtracting the monitored focusing,

accelerator and decelerator impingement currents from the recorded screen

current. The Farady probe and x-y carriage, described previously, were used

to determine the ion beam current distribution. A complete set of beam

profiles is shown in Fig. 29 for a given DE-AC grid system operating

condition. Each profile was generated by sweeping the probe in the x

direction at different y ordinate values. As a result, horizontal

displacement, in Fig. 29, corresponds to motion along the x axis while

vertical displacement records the collected beam current. The numerical

computation of a volume integral, given a closed surface, is -a well known

process and a computer program was written to carry out this computation. The

analysis used linear interpolation between adjoining beam profiles to generate

the closed surface of the volume. The base of the volume corresponded to the

x-y probe carriage sweep area which measured 30.0 cm by 56.8 cm. A step size

of 1 cm was used for incrementing in the x and y directions and was based on

2

the probe area of 1.0 cm . The total integrated beam current was the

cumulative current contributions from each increment.

Typical input and output of the beam analysis program is reproduced

in Fig. 30 for the 9 beam traces shown in Fig. 29. The result of this

56
SUCCESSIVE FARADAY PROBE BEAM

PROFILES AT DIFFERENT Y LOCATIONS

Fig. 29. Complete Beam Scan for Integrated Beam Current Calculations

57
Run option 5 chosen

Z-values in microamps, and Y-values in cm

numcLirve

vval Lies

*

•

*

1

0.00

'-k

jL.

3.43

•_•

7.00

4

1 0 . 57

5

14. 14

6

17.71

7

21.28

8

24.85

9

3O . OO

#*#*##?*##*&#*£*#*#$*#**?t#######*$##*#t*^

curve

points

1

2

~*%

4

5

6

7

8

9

1O

11

12

13

14

15

16

17

18

19

20

21

X

val Lies

O.OO

19.72

24 . 6O

25.66

26.71

27.77

28. 82

29.88

30.93

31.99

33.O4

34 . 1 0

35. 15

36.21

37.26

38.32

39.38

4O.43

41.49

45. 10

56.83

*

*

*

*

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z

0 . OO

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0 . 6O

1 . 40

4 . 00

4.40

4 . 00

3 . 6O

4 . 80

6. OO

6.4O

6.6O

3.60

2. OO

2.20

3.20

3.20

1 . 40

O. SO

O. OO

O . OO

-

O . OO

0.00

1 . OO

2 . 6O

6.40

9 . 'It'J

1 0 . OO

9.4O

9.6O

1 ~"i ""O

15.4O

14. 2O

9 . OO

4 . 'DO

6 . 4O

6.6O

5.00

2. . 2O

1 . 4O

O.4O

O . OO

V

O.OO

0.00

2 . OO

3.80

8.00

14.00

2O. 4O

30.00

3O. OO

32. 00

39.00

36.80

22.00

18.40

1 8 . OO

12.60

8.00

5. OO

2.80

0 . 20

0.00

O . OO

0.00

3.60

6.60

1 1 . 20

25 . OO

14.20

58. OO

74.0O

8' t 'DO

a

0. OO

0. OO

O.8O

3. 20

8.00

2O. 'DO

38. 00

49.60

68. OO

79.0O

96.OO1OO.OO

78. OO

62.80

56. 5O

44.8O

30.00

15.00

7. 00

3.20

O . OO

0.00

75.OO

68.00

60.00

47.OO

32 . OO

15.00

7.00

4 . 00

O.6O

0.00

1

0 . OO

O . OO

1 . 80

3. 20

5 . 2O

1 1 . 00

15.60

22. OO

22. OO

3 1 . (DO

49.00

42 . (DC)

26.00

29. OO

26 . 6O

16.00

8.4O

4 . 20

3 . 6O

0 . 6O

O. 00

LI

0.00

O.OO

1.20

1 . 60

4 . 6O

7.4'D

8 . OO

6.40

8.00

1 4 . OO

17.60

1 7 . 00

12.00

9 . OO

10.80

9.20

5 . SO

2.4O

1 . 60

0 . 60

O.OO

e

O . OO

O. OO

0.80

1 . 4O

3. 4O

4.4O

4 . OO

2.4O

3. 8O

8. OO

6.4O

8.00

7 . OO

3.4O

'4 . 00

5.6O

4.60

1 . 60

0.80

0.00

O. OO

s

O.OO

O.OO

O.20

1 . 2O

0.80

0.20

0 . 2O

O.8O

1.40

1 . 60

1 . 40

2.0O

1 . 80

0.80

0 . 20

1 . 1 0

1.60

1 . 2O

0.00

O . OO

O. 00

*** FINAL RESULT ***

The total collected beam current 1-5 1.0671E-002 amperes.

Fig. 30. Sample Output of Integrated Beam Current Program

58
analysis indicates an integrated beam current of 10.7 mA. The measured beam

current, recorded by the metered DE-AC grid system electrodes, for this

example was 19.6 mA. Further analysis showed that this factor of two

difference in beam current determination was due to ion charge exchange along

the beam line. As described in an earlier section, the neutral propellant

loss rate through the DE-AC test module was typically 30 times the metered

beam current. Charge exchange production rate calculations were performed by

modeling this neutral flow stream as a succession of expanding hemispherical

shells and also by a sharp-edged orifice, neutral density distribution, flow

as used by Kaufman . In both cases the calculated beam current, reaching

the Faraday probe target plane, had been reduced from 19.6 mA to approximately

13.6 mA. The 25% difference between this result and the measured 10.7 mA is

believed attributable to a lesser neutral density drop off along the beam line

than was assumed for the calculations. This is believed to result from the

large depth of the DE-AC accelerator system (~2 cm) and a downstream ground

shield shaped like a 3 cm long beam channel, both of which would create a more

collimated neutral stream.

59
SECTION 4

ADVANCED ION ENGINE PERFORMANCE

2

An argon ion beam current density of 60 mA/cm , using the DE-AC

accelerator system, appears to be feasible. At this beam current density it

becomes possible to seriously contemplate development of a one newton thrust

ion engine. In the following analysis, performance levels of such an engine

were determined based upon the assumptions listed below:

(i) Ion beam area comparable to a 30-cm J series thruster, 625

2

cm , with a beam flatness parameter of 1.0

(ii) Use of a PERM hollow cathode, in the self heating mode, for

the main and neutralizer cathodes, with no heater,

vaporizer, or keeper, power supplies.

(iii) DE-AC accelerator system extraction gap field stress of 2000

V/mm.

Ion engine performance was described by the following set of

equations.

Thrust, T: T =-5

en

u

Total Power, P:

P = JB

/gl \2

m / sp \

+ e + V,

2e U Y / NC

u

60
Thruster Efficiency, ri_,:

2

Y n

'n ,N2

m \ gl

\ sp

The solutions of these equations for a constant thrust, T = 1.0

newton, are presented in Table 2 as a function of discharge loss e.

Table 2: One Newton Ion Engine Performance

JB ISP

(eV/ion) (ampere) (sees.)

Argon,

100 37

80 37

60 37

Xenon,

100 20

80 20

60 20

Mercury

100 16

80 16

60 16

assumes Y

.5

.5

.5

assumes Y

.3

.3

.3

, assumes Y

.8

.8

.8

= 0.95,

5860

5860

5860

= 0.95,

3340

3340

3340

= 0.95

2815

2815

2815

T P nT

(newton) (watts)

1

1

1

90, VNC = 30

.000

.000

.000

\ - °-95' v =

 l5

1

1

1

, nu = 0

1

1

1

.000

.000

.000

•

97> VNC - l

.000

.000

.000

volt

40191

39441

38691

volt

21426

21020

20613

2 volt

17628

17292

16956

0.713

0.727

0.741

0.764

0.779

0.794

0.782

0.797

0.813

61
Table 2 shows clearly the power penalty in using a light atomic

weight propellant, such as argon, in a high thrust density ion engine.

Mercury, because of the superior performance afforded by its high atomic

weight, low ionization potential and minimal storage requirements, is the

preferred propellant for interplanetary propulsion. The major technical

challenges associated with the development of a one newton thrust mercury ion

engine, as defined in Table 2, are listed below:

2

(i) Design and construction of a full size (625 cm ) DE-AC

accelerator system and demonstration of full power beam

extraction for extended periods of time,

(ii) Reliable, long life, hollow cathode operation at 73 amperes

emission current (assuming a 100 eV/ion discharge loss)

using either a scaled up version of present thruster hollow

cathodes or the FERM cathode discussed in this report,

(iii) Development of an efficient discharge chamber whose magnetic

field structure provides a beam flatness parameter

approaching 1.0

(iv) Passive control of discharge chamber heat loss (estimated to

be 2184 watts for a 1.0 newton mercury ion engine) while

ensuring that the samarium cobalt magnet temperatures do not

exceed the demagnetization limit.

62
SECTION 5

CONCLUSIONS

A Test Bed Ion (TBI) engine has been developed to serve as a tool

in exploring the limits of electrostatic ion thruster performance. A

preliminary investigation, using the TBI engine and its component

technologies, has shown that substantial thrust density increases may be

achieved over present ion thrusters. Principally, these thrust density

increases can be realized by using the Decoupled Extraction and Amplified

Current (DE-AC) accelerator system. Reduced electric field stress

requirements enable this accelerator system to extract very high ion current

2

densities. An argon ion beam current density of 60 mA/cm was shown to be

technically feasible.

Tests of a water cooled, Divergent Line Cusp (DLC), discharge

chamber have demonstrated that the TBI engine can be operated at discharge

powers of several kilowatts without damage to its samarium cobalt permanent

magnets. Extensive Langmuir probe characterization, of the plasma generated

within the pyramid shaped DLC chamber, identified plasma potentials negative

of anode potential. In one engine operating mode, this negative potential

difference was as large as 8-10 volts, with low plasma density. In a second

mode, with plasma potential 1-2 volts negative of anode potential, very high

plasma densities were obtained.

Electron currents of 35 amperes were demonstrated using the Field

Enhanced Refractory Metal (PERM) cathode. This simple, rugged device, used a

metal tube electron emitter and required no oxide insert resistance heaters.

TBI engine start up times, using the PERM cathode, were typically 1-2 sees.

In addition, the FERM cathode is easily scaled to higher currents thus making

63
the TBI engine capable of generating the large plasma densities required to

explore the maximum projected beam current levels of the DE-AC accelerator

system.

A 1.0 newton thrust ion engine, with beam area comparable to the

present J series 30 cm thruster, appears to be a realizable technology goal.

This conclusion is based upon the performance projections and component

designs of the Test Bed Ion engine, in addition to, reasonable estimates of

parameter values not evaluated during this program. Calculations show that

such an engine, operating on Xe or Hg propellant, could have thruster

efficiencies as high as 0.76 and 0.78, respectively, with a 100 ev/ion

discharge loss.

64
SECTION 6

REFERENCES

1. Aston, G., "High Efficiency Ion Beam Accelerator System," Rev. Sci.

Instrum., Vol. 52 No. 9, Sept. 1981, pp. 1325-1327.

2. Aston, G. and Kaufman, H. R., "The Ion-Optics of a Two-Grid

Electron-Bombardment Thruster," AIAA Paper No. 76-1029, Nov. 1976.

3. Rovang, D. C. and Wilbur, P. J. , "Ion Extraction Capabilities of

Very Closely Spaced Grids," AIAA Paper No. 82-1894, Nov. 1982.

4. Seigfried, D., "Studies on an Experimental Quartz Tube Hollow

Cathode," AIAA Paper No. 79-2056, Nov. 1979.

5. Meek, J. M. and Craggs, J. D., "Electrical Breakdown of Gases,"

Oxford University Press, 1953.

6. Sovey, J. S. , "Improved Ion Containment Using a Ring Cusp Ion

Thruster," AIAA Paper No. 82-1928, Nov. 1982.

7. Poeschel, R. L., "Development of Advanced Inert-Gas Ion Thrusters,"

NASA CR-168206, July 1983.

8. Beattie, J. R. , "Numerical Procedure for Analyzing Langmuir Probe

Data," AIAA Journal, Vol. 13, No. 7, July 1975, pp. 950-952.

9. Goebel, D. M., "Ion Source Discharge Performance and Stability,"

Phys. Fluids, Vol. 25, No. 6, June 1982, pp. 109.3-1102.

10. Kaufman, H. R., "Charge-Exchange Plasma Generated by an Ion

Thruster," NASA CR-134844, June 1975.

65
SECTION 7

APPENDIX A

Child's Law with Non-Zero Initial Ion Velocity

DISCHARGE

PLASMA

1=0vl

Po

From Poisson's equation in one dimension

d V

66
Assuming a uniform ion current density, j , in the post acceleration section

A.

p =T

From energy conservation the ion velocity, v, is

v =

Substituting eqs 1 and 2 into eq. 3 gives

d V 3A / m

2 e l2eV

o

where e is the vacuum permitivity, m is ion mass and e the electronic

o

dV

charge. Multiplying both sides of this equation by ^jT

 and integrating

results in

..

I

J

9.J

 „

dV

Since dJl is small within the focusing grid holes, U=0), we may approximate

by

- ••

Further integrating

67
J 2 2 \ 2

V - V,

This becomes

zK^-v' Z | 3 \ v

?

 Vl /V ^ + ZV zvMv'-vTJ 1 \V

9

 V 1 / / ^l - 2&e ' | |=- ' n 'i y ,'

\ O,

Rearranging the LHS this reduces to

fef-|(OJ-»-

Further manipulation gives

J _ 1 2

//V \2 \ y^V \2 \

Setting j _ as the usual Child's Law current density the effect of a non-zero

initial ion velocity is to increase this current density by the ratio

,

Where, in terms of familiar DE-AC electrode potentials, measured with respect

to facility ground, V = V - V and V = I V I + V .

J- S L <£. I jrV I S

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