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NASA CR - 180862

N/ SA

ADVANCED ELECTRIC PROPULSION

AND

SPACE PLASMA CONTACTOR RESEARCH

[II&SA-CR- 180862} ADVANCED ELECTRIC

PROPULSION AND SPACE PLASBA CONTACTOR

RESEARCH Annual Report |Colorado State

Univ. l 163 p CSCL 21C

Prepared For

LEWIS RESEARCH CENTER

t188-14126

Unclas

G3/20 0116605

NATIONAL AERONAUTICS AND SPACE ADMINISTRATION

Grant NGR-06-O02-112

Annual Report

January 1987

Paul J. Wilbur

Department of Mechanical Engineering

Colorado State University

Fort Collins, CO 80523
1. ReDort No. I

NASA CR-180862 I

4. Title and Subtitle

2. Government Accession No. 3. Recipient's Catalog No.

ADVANCED ELECTRIC PROPULSION AND SPACE CONTACTOR RESEARCH

7. Author(s)

Paul J. Wilbur

9. _rforming Organization Nameand Address

Department of Mechanical Engineering

Colorado State University

Fort Collins, CO 80523

12. Sponsoring A_ncy Name and Address

National Aeronautics and Space Administration

Washington, D.C. 20546

5. Report Date

Jan, 1987

6. Performing Organization Code

8. Performing Organization Report No.

10. Work Unit No.

11. Contract or Grant No.

NGR-06-002-112

13. Type of Report and Period Covered

Annual

Jan. 1, 1986-Jan. 1, 1987

14. Sponsoring Agency Code

15. Supplementary Notes

Grant Monitor - Vincent K. Rawlln, NASA Lewis Research Center

Cleveland, Ohio 44135

16. Abstract

A theory of the plasma contacting process is described and experimental results

obtained using three different hollow cathode-based plasma contactors are presented. The

existence of a sheath across which the bulk of the voltage drop associated with the

contacting process occurs is demonstrated. Test results are shown to agree with a model

of a spherical, space-charge-limlted double-sheath. The concept of ignited mode

contactor operation is discussed and thls mode of operation is shown to enhance contactor

efficiency when it is collecting electrons. A detailed investigation of the potentials

in the plasma plumes downstream of contactors operating at typical conditions is

presented. Results of tests performed on hollow cathodes operating at high

interelectrode pressures (up to about I000 Tort) on ammonia are presented and criteria

that are necessary to ensure that the cathode will operate properly in this regime are

presented. These results suggest that hlgh pressure hollow cathode operation is

difficult to achieve and that special care must be taken to assure that the electron

emission region remains diffuse and attached to the low work function insert.

Experiments conducted to verify results obtained prevlously using a ring cusp ion source

equipped with a movable anode are described and test results are reported. A theoretical

study of hollow cathode operation at high electron emission currents is presented. This

study suggests that it is desirable to increase the hollow cathode body diameter as

emission capabillty is increased and the reasons why this is desirable are discussed.

Preliminary experiments conducted using the constrained sheath optics concept to achieve

ion extraction under conditions of high beam current density, low net accelerating

voltage and well columnlated beamlet formation are discussed.

17. Key Words(Suggestedbv Auth,(s))

Electrostatic Thruster

ArcJet Thruster

Hollow Cathode

Electrodynamlc Tether

Plasma Contactor

19. Security Classif.(ofthisreport)

Unclassified

18. Distribution Statement

Unclassified-Unlimited

I 20" SecurityUnclassified classif" (°f this page) I 21' N°' 173°f Pages I

For sale bythe NationalTechnicalInformationService,Springfield. Virginia 22161

22. Price*

NASA-C-168 (Rev. 10-7S) -l-
TABLE OF CONTENTS

Topic Page

Abstract .................................................... i

Electrodynamlc Tether Plasma Contactor Research ............. 1

Introduction ......................................... 1

Theory .................... --....................... 1

Experimental Apparatus'and'Procedure..... ..... 7

Results .............................................. 17

Conventional Hollow Cathode Contactor .......... 25

Ring Cusp Contactor Optimization Study ......... 32

Effect of Anode Area on Contactor Performance.. 44

Circuit Considerations at High Electron

Collection Currents ........................... 65

Characterizing Plasma Contactor Performance .... 70

Electron Current Distribution on the Anode ..... 75

Conclusions .......................................... 77

High Pressure Hollow Cathode Research ....................... 79

Apparatus and Procedure .............................. 79

Results .............................................. 91

Conclusions .......................................... 110

Verification of Ring Cusp Discharge Chamber Performance ..... 112

• Introduction ......................................... 112

Apparatus ............................................ 112

Results .............................................. 116

Conclusion ........................................... 125

High Current Hollow Cathode Research ........................ 126

Constrained Sheath Optics Research .......................... 137

Apparatus and Procedure .............................. 141

Results .............................................. 145

Conclusions .......................................... 151

References .................................................. 154

Appendix A (Space-charge-llmlted current flow in a

spherical double sheath) ......................... 157

Distribution List ........................................... 162

Table No.

I

II

LIST OF TABLES

Title Page

Expellant Flow Options .......................... 82

Parameters Varied In Hollow Cathode Study ....... 85

-ii-
LIST OFFIGURES

Figure No. Title Page

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

18

19

20

21

22

23

24

25

26

27

28

Plasma Contacting Model ........................ 3

Theoretical Model Concept ...................... 5

Schematic Diagrams ............................. 8

Alternative Plasma Contactor Schematics ........ 9

Conventional Hollow Cathode Contactor .......... 11

Ring Cusp Magnetic Field Contactor ............. 12

Variable Anode Area Plasma Contactor ........... 13

Generic Contactor Characteristics .............. 18

Generic Potential Profiles ..................... 21

Plasma Density Profiles ........................ 24

Effect of Magnetic Field on Contactor

Performance .................................... 26

Conventional Hollow Cathode

Characteristics ................................ 28

Typical Plasma Potential Profiles for

Conventional Contactor with Simulator on ....... 29

Typical Plasma Potential Profiles for

Conventional contactor with Simulator off ...... 30

Effect of Changing Point of Connection to

Simulator-Conventional Hollow Cathode

Contactor ...................................... 33

Effect of Changing Point of Connection to

Simulator-12 cm dla Anode Hollow Cathode

Contactor ...................................... 34

Typical Plasma Ion Energy Cost Curve ........... 38

Effect of Mask Aperture Diameter on

Plasma Ion Energy Cost Curves .................. 39

Effect of Mask Aperture Diameter on

Extracted Ion Fraction ......................... 40

Effect of Mask Aperture Diameter on

Extracted Ion Energy Cost ...................... 42

Effect of Flowrate on Ring Cusp Contactor

Ion Cost ....................................... 43

Effect of Mask Aperture Diameter on Ring

Cusp Contractor Performance .................... 45

Effect of Anode Diameter on Contactor

Performance .................................... 47

Perturbing Effect of Vacuum Tank Walls ......... 49

Anomalous Contactor Performance Curves ......... 50

Effect of Contactor Anode Current on

Performance at Low Flowrate .................... 52

Effect of Contactor Anode Current on

Performance at a Moderate Flowrate ............. 53

Effect of Contactor Anode Current on

Performance at a High Flowrate ................. 54

-ill-
Figure

29

30

31

32

33

34

35

36

37

38

39

4O

41

42

43

44

45

46

47

48

49

50

51

52

Title Page

Typical 12 cm Dia. Anode Contactor

Characteristics ................................ 55

Typical Plasma Potential Profiles for the

12 cm Dia. Anode Contactor-Simulator on ........ 57

Typical Plasma Potential Profiles for the

12 cm Dia. Anode Contactor-Simulator off ....... 58

Typical Potential Contour Map Observed in

the Ignited Electron Collection Mode ........... 60

Typical Comparison of Measured and

Predicted Plasma Potential Profiles ............ 62

Radius Ratio Comparison for Spherlcal

Double Sheath Model ............................ 64

Radius Ratio Comparison for Hemispherical

Double Sheath Model ............................ 66

Effect of Transition to Ignited Mode

Operation on Electron Current Balance .......... 67

Effect of Simulator-to-Tank Connection

Point on Electron Collectlon Current Balance

(Condition A) .................................. 69

Effect of Simulator-to-Tank Connection

Point on Electron Emission Current Balance

(Condition E) .................................. 71

Contactor Anode-to-Ambient Plasma

Potential Difference vs. Emission

Current-Low Anode Current Case ................. 73

Contactor Anode-to Ambient Plasma

Potential Difference vs. Emission

Current-High Anode Current Case ................ 74

Anode Current Distribution ..................... 76

Hollow Cathode Test Apparatus Schematic ........ 81

Vertically Oriented Test Apparatus ............. 84

Boron Nitride Body/Tur_sten Orifice Plate

Hollow Cathode ................................. 86

Graphite Orifice Cap Hollow Cathode ............ 88

Hollow Cathode wlth Quartz Tube

Insulation ..................................... 89

Typical Voltage vs. Interelectrode

Pressure Characteristic ........................ 92

Typical Plasma Potential Profile ............... 93

Effect of Flow Direction on Cathode

Interior Pressure .............................. 95

Effect of Orifice Diameter on Pressure

Distribution for no Discharge Case ............. 97

Effect of Orifice Diameter on

Voltage/Pressure Characteristic ................ 98

Effect of Discharge Current on Operating

Pressures at Constant Reverse Flow ............. 100

-iv-
Figure

53

54

55

56

57

58

59

60

61

62

63

64

65

66

67

68

69

70

71

72

7Z

74

75-

76

77

Title Page

Effect of Discharge Current on Anode

Voltage vs. Interelectrode Pressure ............ 102

Voltage/Pressure Characteristic for

Normal Flow Case ............................... 103

Interior Cathode Pressure vs.

Interelectrode Pressure ........................ 105

Anode Voltage vs. Interelectrode Pressure ...... 107

Effect of Orifice Cap Configuration on

Voltage/Pressure Curve ......................... I09

Movable Loop Anode Discharge Chamber ........... 113

Ring Cusp Ion Thruster Electrical

Schematic ...................................... 115

Typical Plasma Ion Energy Cost Curve ........... 117

Effect of Ring Cusp Location on Plasma

Ion Energy Cost Curves ......................... 119

Effect of Ring Cusp Location on Extracted

Ion Fraction ................................... 120

Effect of Ring Cusp Location on Beam Ion

Energy Cost .................................... 121

Performance Comparison at High Flowrate ........ 12S

Performance Comparison at Moderate

Flowrate ....................................... 124

Hollow Cathode Terminology ..................... 127

Effect of Heat Removal Rate and Emission

Region Length on Cathode Performance

Parameters ..................................... 129

Effect of Insert Diameter on Cathode

Interior Pressure .............................. 132

Effect of Insert Diameter on Cathode

Performance Parameters ......................... 1SS

Flowrate Correlatlon Data up to 15 A

Discharge Current .............................. 135

Equlpotentlal Contour and Ion Trajectory

Diagram ........................................ 138

Constrained Sheath Optics Schematic

Diagram ........................................ 142

Typical Current Density Profiles ............... 144

Effect of Grid Separation on Divergence

Characteristics ................................ 146

Effect of Net Accelerating Voltage on

Divergence Characteristics ..................... 149

Effect of Net-to-Total Accelerating

Voltage Ratio on Divergence

Characteristics ................................ 150

Effect of Constraining Mesh Radius on

Curvature on Divergence Characteristics ........ 152

-V-
ELECTRODYNAMIC TETHER PLASMA CONTACTOR RESEARCH

John D. Williams

Introduction

It As important in some space applications to be able to establish an

electrical connection to the local space plasma. A device designed to perform

thls functlon is known as a space plasma contactor and ideally it would clamp

a point on a spacecraft to a potential near space plasma potential regardless

of the current passing between this point and the ambient space plasma. Functioning iike a grounding wire on earth, a plasma contactor would provide a

reference potential for Instruments on board a spacecraft and prevent the

spacecraft from charging relative to the potential of the local plasma. The

per_'ormance of a plasma contactor would become more critical as the current it

could be called upon to handle increased to values An the decampere range that

are expected in future electrodynamic tether applications. 1 The objective of

the work described here is to develop an understanding of the plasma contacting process which can eventually lead to the design of a device suitable for

such electrodynamlc tether systems.

Plasma contactor performance can be characterized by plotting the contactor electron emission current vs. the contactor potential measured relative to

the local space plasma potential. The ideal plasma contactor would exhibit a

current/voltage characteristic curve appearing as a vertical llne at zero

potential if it could make ions and/or electrons available at a sufficiently

hlgh rate so negligible voltage drops would develop over the full range of
current demanded.Such performance might be approached if contactor ion/electron

production rates were sufficiently high so that each contactor produced a high

density, high conductivity plasma plume with a surface area equal to or greater

than the associated tether current divided by the ambient space current density.

Although it might be desirable to achieve this level of performance, the large

contactor operating powers that could be required to produce the ion and electron current densities needed at high tether current levels could depress

overall system efflclencles to the point where this performance level would be

unattractive. Consequently, operation of the contactors at conditions where

the performance curve would depart from the vertical might be preferred. As

Parks, et al., 2 have pointed out, reduction of the contactor plasma plume size

below that which would produce the near vertical (approximately zero impedance)

current/voltage characteristic results in the development of a double sheath

between the region of the contacting plasma and unperturbed space plasma in

the manner suggested by the conceptual sketch in Fig. I. In this intermediate

region two currents, one due to particles from the contactlng plasma and the

other due to oppositely charged particles from the space plasma, are counterflowing and it is because of this that the term double sheath is used. The

total current produced by these counterflowing charged particles under a

prescribed potential difference will be limited by space-charge effects and it

is suggested that a plasma contactor may always operate at a space-chargelimited condition determined by the sizes of the boundaries between 1) the

high density plume and the sheath and 2) the undisturbed space plasma and the

sheath. The space-charge-llmlted current that would flow under ideal conditions can be found by applying a theoretical model developed during this grant

period 3 and included as Appendix A to this report. In this model it is

assumed that spherical symmetry exists and that the high density plume can be

-2-
I- E i,i

_: U.J (J

.J Q. _:

UJ Z In

E _ (n

llJ

0

0

0

0

-3-
represented by a surface of radius r i which is an infinite source of zero

velocity, charged particles of mass m i as shown in Fig. 2. Under the influence of a potential V i a particle current Ji is drawn radially outward toward

a second surface representing the low density, ambient space plasma. This

outer surface; which serves as an infinite source of zero velocity, oppositely

charged particles of mass m ° at zero potential; has a radius r o and contributes a particle current Jo flowing inward. The model assumes that the

effects of ion and electron collisions wlth either neutral atoms or other ions

and electrons can be neglected in the vacuous region between the two surfaces

both in terms of the impedance they induce and the production of additional

ion-electron pairs. The solution sought to thls physical problem is the one

corresponding to maximum current flow for both species. This occurs when the

electric fields at both r o and r i are zero i.e. the doubly space-chargelimited condition. The maln result of the paper can be given in the form of

equations describing the currents Jo and Jl which are reproduced here as Eqs.

1 and 2.

Jo = 4_tSo Vt 3/2 (2e/too)l/2 Jo (l)

Jl = Jo (a)-l(mo/ml)112 (2)

These equations describe the counterflowlng currents in terms of the mass m o

of the particles flowing from the outer surface and the mass mi of the particles flowing from the inner surface, the potential difference Vi between the

two surfaces, the constants • and so and the parameters Jo(rl/ro) and

a(ri/ro). These parameter_ which have been determined numerically, are plotted as a function of radius ratio in Appendix A. Note that the equations show the

currents Jo and Ji are implicit functions of the radius ratio (through Jo and a)

and the imposed potential difference but are independent of the actual magnitudes

-4-
POTENTIAL ZERO

CURRENT J•

PARTICLE MASS m•

Fig. 2 Theoretical Model Concept

-5-
of the radii. In this report potential profiles for theoretical solutions to the

space-charge-llmited problem will be compared to experimental results in order

to determine the conditions under which the model is valid. It is believed

that the model will provide a basic picture of the plasma contacting process

and be useful in computing currents at prescribed potential differences, when

low collision frequency conditions prevail in the double sheath region.

In order to differentiate between the two modes of operation of the

plasma contactors (i.e. whether V i is positive or negative), the term electron

collection mode will be used when the contactor is collecting electrons and

emitting ions (V i positive) and the term electron emission mode will be used

when the contactor is emitting electrons and collecting ions (Vi negative).

Hence for a system operating in the electron collection mode Jo would

represent an electron current, Ji an ion current and the two currents would

add to give the total conventional current flowing between the two surfaces.

Note from Eq. 2 that Ji is considerably less than Jo in the electron collection case because the ion mass is considerably greater than the electron mass.

Although this may suggest the ion current can be neglected, the analysis of

AppenClx A shows that the real importance of the ions is reflected in their

ability to neutralize the electron space charge and_ hence_increase the electron current over what would flow with no ions present. Since the model

predicts improved performance as the radius of the plasma plume approaches the

radius of the unperturbed space plasma boundary (i.e. rl/ro--b-1) and since the

location of the outer boundary is probably fixed for a prescribed total

current and ambient plasma condition, the ultimate low impedance plasma contactor should be realized when the high density plume (Fig. 1) is expanded to

meet the outer boundary. It appears that this could be accomplished by

increasing the ion production rate of the plasma contactor.

-6-
_perimenta_ Apparatus an_ Procedure

In order to measure the performance of a typical plasma contactor as a

functlon of its controllable operating parameters the apparatus suggested by

the electrical schematic diagrams of Fig. 3 was set up. As the figure suggests the two major elements of the system are the contactor under test and

the simulator used to generate a diffuse, simulated space plasma that can couple to the contactor. Both the contactor and simulator utilize heater power

supplies to facilitate cathode startup and maintenance of cathode temperatures

and anode power supplies that maintain electrical discharges between their

anodes and cathodes. The coupling power supply is used to bias the contactor

relative to the simulator so a current can be forced through the plasma between

these two devices.

The contactor and simulator are located within a 1.2 m dia by 5.3 m long

vacuum tank and are separated by 2.7 m as Fig. 3a shows. The simulator axis

is oriented perpendicular to the common axis of the tank and the contactor.

The simulator is a 0.64 cm dia hollow cathode plasma source that has a 0.76 mm

dia cathode orifice and utilizes an R-$OOmtreated, rolled tantalum loll

insert. It is equipped wlth a 2 cm dla flat plate anode I mm from the cathode

witn no orifice in it. Thls anode design was used because it was believed

that it would provide a diffuse, non-focused plasma which would approach a

realistic simulation of the ionospheric plasma. Although the same hollow

cathode device was used to produce the simulated space plasma throughout all

experiments, the conventional hollow cathode plasma contactor (dlstlngulshd by

the dotted boundary in Fig. 3b) was replaced with two other contactor configurations. These contactors, the ring cusp contactor and the variable area

anode contactor, are shown schematically in Fig. 4 and are also enclosed by

dotted boundaries to indicate visually that they can be interchanged with the

S

Chemical R-500 is a double carbonate (BaCO 3. SrCO 3) mixture that has been manufactured by the J.R. Baker Chemlcal Co., Philllpsburg, N.J. but is no longer made.

--7--
_- 5.:3 m

/__HOLLOW CATHODE

// PLASMA fSUPPORTfPROBE CONTACTOR ROD __i MULATOR

--1 /( o _ [/PROBEEMISSIVE

0 r .... "-,

I-- 2.7m -_I

1.2 m

o. MECHANICAL

CONTACTOR

ANODE

SUPPLY

+ ORCOUPLING

SUPPLY

HTR

SUPPLY

_ J

I

I

I

I

I

I

I

I

I

/

HTR

SUPPLY

TANK

WALL

i

m

L

SIMULATOR

ANODE

'ISUPPLY

+

b. ELECTRICAL

Fig. 3 Sobematle Diagrams

-8-
t9

I

0

zaFiE. 4 Alternatlve Plasma Contactor Sehematlas

-9-
one shown in Fig. 3b. Also shown in Figs. 3b and 4 are the locations of the

meters used to measure the currents and potentials considered important to the

tests. The symbols used to designate the measured currents or potentials are

indicated within the circles representing the meters. The power supplies

needed to sustain operation of the contactors and the coupling and bias

supplies are labeled and shown as squares on the figures.

Also shown in Fig. 3a is the rod used to support and position the emissive and Langmuir probes which were used to measure plasma properties in the

region extending to the plane 1 m downstream of the contactor. The measuring

circuits and mechanical designs for these probes are similar to those which

have been used previously and they are described by Aston 4 and Laupa, $ respectively. In the present research the emissive probe system differs from the

previous design in that the probe output is transmitted to an X-Y plotter

along with the output of a probe position indicating potentiometer so continuous traces of the plasma potential vs. axial position between the contactor

ana simulator are recorded. The emisslve probe can be swept along the tank

centerline or rotated so plasma potential profiles along paths parallel to the

centerllne are obtained. By using the data from such traces recorded at various radii, sufficient data can be collected so equipotential contours can be

generated to describe the complete potential picture in the region downstream

of the contactor.

Three different configurations of contactors were investigated during the

grant period. They included the conventional hollow cathode (Fig. 5) the ring

cusp magnetic field (Fig. 6) and the variable anode area (Fig. 7) plasma contactors. The conventional hollow cathode contactor consists of a 7 mm dia

tantalum tube with a 0.7 mm orifice in an orifice plate welded to one end of

-10-
..AT..w,.._o,.,Nsu..T,O,,- h

ROLLED TANTALUM FOIL _ ._

INSERT(R500 TREATMENT)_ I' -_ -_,| S

RI NG ANODE

5 mm

1.5 mm

"0.7 mm ORIFICE DIA

FiE. 5 Conventional Hollow Cathode Contaotor

-11-
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r_

-12-
re)

0m0

r__mm

o2m

t_L

"o0

,-.4 m

r._
the tube. The insert in the cathode is made of 0.013 mm thick rolled tantalum

loll covered with the low work function material Chemical R-500. The insert

region of the tantalum tube is heated by a swaged heater wrapped with a tantalum loll radiation shield. The anode is 1 mm downstream of the orifice and

is a loop of 1.5 mm dla tantalum wire with an outside dla of 8 mm. Because

this device was not designed to produce ions efficiently or to provide a large

anode plume boundary, it was not expected to be a superior plasma contactor,

but rather a device that could be used to study the plasma contacting process.

A contactor which should be capable of better performance than the conventional contaetor because it is designed to produce ions efficiently is the

ring cusp magnetic field contactor 6 shown in Fig. 6. The ring cusp contactor

utilizes the same hollow cathode electron source as the conventional hollow

cathode contactor, and enhanced ion production is facilitated in it because of

the cusped magnetic field indicated by the dotted lines in Fig. 6. This

field, which is produced by a ring of samarium cobalt magnets near the chamber

midsection and the magnet at the upstream face of the chamber, enhances performance by confining most primary electrons (i.e. those supplled by the

cathode) from being lost directly to the anode until after they have had ionizlng collisions. As a consequence of the electron confinement it is believed

that the ions are also confined so they can be extracted preferentially

through the open end of the chamber rather than being lost through recombination on the chamber walls. The samarium cobalt magnets have a flux density at

their surfaces of 0.27 Tesla and the discharge chamber steel construction provides a low reluctance return path for the magnetic fleld lines they induce

through the chamber. The plasma source is also designed so expellant can be

supplied through both the hollow cathode at a flowrate mc and a main flow ple-

-14-
hum located at the downstream end of the discharge chamber at a rate mb.

The variable area anode plasma contactor (Fig. 7) also utilizes the conventional hollow cathode as its electron source but it has a large diameter

segmented anode. The sustaining anode shown in Fig. 7 is always held at anode

potential, but each of the progressively larger segments shown may either

float or be held at anode potential. This design makes it possible to vary

the effective anode diameter from -1 to 12 cm by simply changing switch positions.

Normally the contactors were first tested with the simulator off by varyIng the coupling supply output voltage (Vc) from --70 to 130 V. As the contactor is biased negative relative to the tank, it will emit electrons and

when the contactor is biased sufficiently positive the electrons produced in

the discharge will be reflected from the tank walls and simulator anode

thereby allowing measurement of the small ion current (ion production current)

being drawn to these surfaces. Typically this ion production rate associated with a particular contactor operating condition can be measured when the

coupllngvoltage V c is above -30 V. Further moderate increases in Vc will not

cause this curren_ measured on the contactor emission current meter JCE shown

in Figs 3b and 4_to increase. This occurs because the sheath that develops at

the tank wall is able to reflect all electrons including those from the

Maxwelllan tall so only the ions produced are collected. Hence the ion

current collected at the tank wall is limited to the rate at which ions are

produced by the contactor at its prevaillr_ flowrate and discharge current

operating condition. It has also been observed that increasing the coupling

voltage Vc above --100 V causes the ion current measured to increase above the

stable value observed at a coupling voltage near 30 V. It is assumed that

-15-
this occurs because high potential differences cause excessive acceleration of

the ions through the sheath at the tank wall thereby causing secondary electrons to be ejected from the tank surface. These electrons are then collected

at the contactor anode thereby causing the contactor emission current JCE to

increase. Arcing from the tank walls to the contactor was also observed at

these high potentials and this supports the suggestion that the metallic tank

may have been participating in this way. To prevent these effects a coupling

voltage near 30 V was used when ion production rates were being measured.

After the ion production capability of each contactor had beem measured at a

specified operating condition, plasma potential profiles associated with particular contactor operatin6 conditions were recorded using the emissive probe.

The simulator was started next and allowed to stabilize at a flowrate

(ms) of 1.9 sccm (Xe) and a dlscharge current JSA of 0.3 A; this normally

resulted in an anode voltage VSA of -14 V. Once the simulator was operating

stably the coupling voltage V C was varied through the range _-70 to 130 V.

At each operating point all of the currents and voltages indicated in Fig. 3b

were measured and plasma potential profiles could be recorded using the emissire probe. Typically these data were used to generate a curve of contactor

electron emission and collection current (JCE) as a function of a potential

difference consldered representative of the potential difference between the

contactor and simulator plasmas. Such a curve is referred to here as the

characteristic curve of the contactor. The potential difference used to generate the curve was initially the one between the contactor anode and the

simulator anode (V C + VCA). This potential difference was chosen because the

potentials of the plasmas at the simulator and contactor were expected to

close to their respective anode potentials so it was believed this potential

difference would be representative of the potential drop through the

-16-
intervening plasma. It should be noted, however, that the contactor potential

(V c + VCA) defined in this way is actually the potential drop across both the

contactor and simulator and it was therefore not the preferred measure of contactor performance as tests conducted near the end of the grant period eventually demonstrated. All tests were performed using xenon as the expellant over

a wide range of contaotor anode currents and voltages. The tank pressures

(Pt) measured during tests ranged from Ix10 -6 Torr at low contactor flowra_es

to 1.2x10 -5 Torr at high ones.

Resul;_

Generic characteristic curves obtained showing the operating features

observed in essentially all hollow cathode-based plasma contactors with the

simulator on and off are shown in Fig. 8. It shows that a contactor operating

in the electron emission mode (2nd quadrant) exhibits a curve that is close to

the ideal (vertical) one. In this region essentially all hollow cathode con-"

tactors are observed to emit up to several amperes of current at contactor

potentials of ~ -50 V independent of whether the simulator is on or off. This

independence of the electron emission portion of the characteristic curve from

the simulator operatinE condition is a consequence of the fact that electrons

emitted by a contactor are generally collected on the vacuum tank wall whether

the simulator is operating or not. The tests described herein were conducted

under conditions where the simulator was operatlng at a low power level to

produce plasma densities of the order of 106/cm -3 near the midpoint of the

region between the contactor and simulator. Operation of the simulator to

produce a higher plasma density might be expected to induce a still steeper

curve but it would be less representative of the performance expected in

space.

-17-
ELECTRON

EM I SS ION CURRENT

-50

ELECTRON EMISSION

(SIMULATOR ON OR OFF)

5O

I

UNIGNITED _%_

ELECTRON COLLECTION '

(SIMULATOR ON)

CONTACTOR POTENTIAL

[v¢. v_,] (v)

I001 1501

LION SATURATION CURRENT

(SIMULATOR OFF)

I

I

I

I /IGNITED

/ELECTRON

_ / COLLECTION

_ MULATOR

ON)

Fig. 8 Generic ContaotoP Charaeteristles

-18-
In general, the contactor performance curve in the electron

collectlon/ion emission mode (4th quadrant of Fig. 8) when the simulator is on

shows much more structure than the electron emission curves or the ion emission only curve (4th quadrant, simulator off). In this mode the electron collection current increases in both magnitude and absolute gradient as the contactor potential is increased from low positive values. Near 60 V Fig. 8

shows an inflection point in the electron collection curve beyond which the

gradient of electron collection current wlth contactor potential decreases and

the curve levels out due to a limiting phenomenon which has not been identified. At contactor potentials (V C + VCA ) of the order of 100 V the contactor

has been observed to undergo a transition from what might be termed the unlgnited mode to the ignited mode of electron collection. This transition, which

is accompanied by the appearance of a visible plume within which expellant

excitation and presumably ionization are occurring, is most likely to occur in

contactors utilizing large diameter anodes. The visible plasma plume associated with ignited mode operation usually extends downstream of the anode -5 to

20 cm. Again the point at which transition to ignited electron collection

occurs depends on hollow cathode operating conditions and ambient pressure.

For example this transition may occur before or after the unlgnlted electron

collection curve begins to level out depending on these conditions.

The ion saturation current, measured with the simulator off, is

represented as the vertical separation between the solid curve and the horIzonal axis in the 4th quadrant of Fig. 8. This curve remains relatively flat

for contactor potentials greater than -30 V at an ion production current which

depends upon the particular contactor being tested, the expellant flowrate,

and discharge power.

-19-
Whena contactor is operated in the electron collection mode with the

simulator on, the plasma potential profile along the centerline joining the

contactor and simulator has the general shape shown by the solid llne in Fig.

9a. This fiEure shows that the plasma potential at the contactor, measured

relative to tank/simulator anode potential, is near contactor anode potential

and that it remains at this potential as one moves downstream to the point ru.

Between the points ru and rd the potential drops by VS because of a double

sheath that develops at the prevailln6 current condition. Beyond the location

rd the potential chanEes very slowly until it drops rapidly to simulator anode

potential throuEh an electron extractlon/ion collection sheath that develops

adjacent to the simulator. The region between the double sheath and the simulator in Fig. 9a miEht show a small Eradual potential drop that could develop

because of collislonal effects or maEnetially induced impedance, but slight

potential increases and neEliEible potential gradients have also been observed

in this reglon.

For tae case of ion emission only (simulator off) the dotted plasma

potential profile in FiE. 9a indicates that a small potential sheath develops

between the contactor anode and the plasma near the contactor. The potential

directly downstream of the contactor remains constant near anode potential

until a sheath, which was not measured directly but which probably develops

near the tank wall and cold anode of the simulator collects the ions and completes the circuit. It is noted that the potential profiles shown in Fig. 9

cannot be measured continuously uslng the emlssive probe because it can only

be swept 1 m downstream of the contactor. The solid line (simulator on) profile shown in FiE. 9a was actually constructed by measurlng profiles while

operatlng the contactor in the electron collection mode and then in the electron emlsslon mode and combining the results to show the expected potential

-20-
ANODE

POTENTIAlCATHODE

POTENTIA'E

/

o

FZ

LI.I

I--

O

Q.

Dm m|

Vs

Vc + VCA

SHEATH

REGION ;IMULATOR ON

CONTACTOR ru rd POSITION

LOCAT IO N

a) ELECTRON COLLECTION/ION EMISSION

-SIMULATOR OFF

t

I

I

I

I

I

I

I

I

SIMULATOR

LOCAT I0 N

CONTACTOR

LOCATION

SIMULATOR

LOCATION

CONTACTOR

ANODE

CONTACTOR

CATHODE:

vc+VcA

_Vp

_"'-- S I MULATOR

ON OR OFF

b) ION COLLECTION/ELECTRON EMISSION

J

SIMULATOR

ANODE

CATHODE

SIMULATOR

• ANODE

---SIMULATOR

CAT H ODE

FiE. 9 Generic Potentlal Profiles

-21-
variation along the centerllne between the contactor and simulator. Figure 9b

shows the generalized potential profile for a contactor operating in the electron emission mode when the simulator is either on or off. In these cases the

plasma adjacent to the contactor is near anode potential and the potential

changes only slightly with position until the bulk of the potential rise

occurs near the simulator anode surface.

In order to achieve a good simulation of the plasma contacting phenomena

expected in space it is necessary that the simulation environment reproduce

the environment expected in space. This condition would be satisfied if the

following variables could be controlled at values similar to those expected in

space :

ambient ionic and neutral gas species and their densities,

ambient electron density and temperature,

ambient magnetic field strer_th.

No attempt has been made in these tests to keep the ionic and neutral gas

species at the concentrations expected in space. In fact, xenon levels in the

tank required for proper contactor operation induce vacuum tank pressures that

exceed those expected in the space environment. Also the plasma densities in

the dilute, simulated space plasma have not been measured during all of the

tests. Actually, with the present experimental set up the Langmulr probe can

only be swept through the region extending to 1 m downstream of the contactor

so it cannot probe the plasma adjacent to the simulator. The plasma in the

region immediately downstream of the conventional contactor has been probed

over the range of conditions investigated in this report however, and typical

-22-
plasma density profiles measured in these tests are shown in Fig. 10. They

suggest electron densities in range near 105 - 106 cm -3 exist in the region

intermediate between the contactor and simulator. This range of densities is

considered to be reasonably close to the values expected in the space environment where tether tests are planned. Electron temperatures measured in the

tests were about 1 eV which is higher than the 0.2 eV temperature associated

with space plasma.

The symbols used in Fig. 10 to represent voltages (V) and currents (J) have

been defined in Figs. 3 and 4. The other symbols (i) and (Pt) represent flowrate

in standard cubic centimeters per minute (sccm) of xenon (Xe) and vacuum tank

background pressure in Tort respectively. The background pressures have been

corrected to account for the fact that the cold cathode gauge used to measure them

has a different sensitivity for xenon than it does for the nitrogen on which its

calibration is based.

In order to investigate the importance of the ambient magnetic field,

tests were conducted in which the ambient field was changed to determine its

effect on the contactor performance. The test facility in which the tests

were conducted is stainless steel so the 0.5 gauss geomagnetic fleld

penetrates the tank. This field is inclined at -70 ° to the axis Joining the

contactor and simulator throughout the region of plasma interaction . In

order to investigate the effect of magnetic field on the test results, a

Helmholtz coll was installed so that it could be energized to null the geomagnetic field. With the currents through the Helmholtz coll adjusted to their

optimum values the magnetic field was measured alone the axis of the tank and

determined to be within 0.I gauss of zero over the plasma interaction region.

-23-
108

I

E

0

>-

Ira

107

Z

W

a

Z

o

II.u

.J

W

z IOe

J

.J

LU

X

=E

CONVENTIONAL HOLLOW CATHODE CONTACTOR

0

FLOW CASE

rh = 1.4 sccm (Xe)

JCA = 0.3 A

VCA = 30 V

JCE = -6 mA

Vc = 80 V

Pt - 4 x I0 -e Torr

105

0 25

_HIGH FLOW CASE

rn = 6.9 sccm (Xe)

J CA = 0.3 A

VCA=17V

0 . _ JCE = -52 mA

V C = 73 V

Pt = 7 x 10.6 T0rr

I I I I

50 75 I00 125

AXIAL LOCATION (cm)

Fig. 10 Plasma Density Profiles

-24-
These tests were conducted early in the grant period so the simulator used was

actually a hot filament source of electrons 5 rather than a hollow cathode.

Still the test results shown in Fig. 11 indicate that the geomagnetic field

has no significant effect on the performance of the plasma contactor. Hence

for all of the tests described herein the geomagnetic field was not nulled and

a 0.5 gauss field existed throughout the interaction region.

The approach used in conducting this test has been to involve the vacuum

tank walls in such a way that the most realistic simulation of the space

environment would be realized. For a contactor operating in the electron

emission mode the tank walls surrounding the contactor and simulator have been

used in conjunction with the simulator to collect the electrons emitted by the

contactor. This is Justified on the basis that the contactor will be completely surrounded in space by a plasma which can collect the emitted electrons. If the simulator anode were forced to collect all of the emitted electrons in these ground tests, they would have to flow to the simulator anode, which

is very small. Consequently, it was considered more appropriate to achieve better

geometric simulation and use the surrounding tank walls to collect the electrons. For the case of electron collection by the contactor the tank walls

cannot realistically serve as a source of electrons so the approach has been

to minimize the tank wall participation and to require that all electrons collected by the contactor be supplied from the simulator. Any significant

discrepancy between the simulator emission current (JsE in Fig. 3b) and the

contactor collection current (JcE in Figs. 3b and 4) suggests the tank is

affecting the results and was regarded as Justification for rejecting the

data.

Conventional Hollow Cathode Contactor

Actual characteristic curves measured using the conventional hollow

-25-
Fig. 11

-5O

I

ELECTRON

EMISSION CURRENT

-I000

CONVENTIONAL HOLLOW CATHODE CONTAC'I

FILAMENT SIMULATOR (0.91m FROM CONT/_

JCA = 0.3 A

VCA=17 TO 28 V

-750 rh = 1.4 sccm (Xe)

Pt = 6x I0 -s Torr

-5

-I0

-5OO

- 250

0 NATURAL GEOMAGNETIC FIELD (~0.5 G(

o NULLED MAGNETIC FIELD

CONTACTOR POTENTIAL

[Vc,v_,] (v)

/.o.7.... _,o ,oo ,5o,

NOTE

SCALE

o

Effect of Magnetic Field on Contactor

Performance

-26-
cathode are shown in Fig. 12. Note that although the contactor performs well

in the electron emission mode with the simulator on or off, it does not

undergo a transition into the ignited mode of electron collection when the

simulator is on, conditions are as defined on the figure and contactor potentials

are limited to 150 V; consequently, it does not collect large currents from the

simulator. These low electron collection currents are plotted on a different

scale than the electron emission data as indicated on the figure. The ion

emission current data represented by the square symbols (simulator off) in the

4th quadrant show this current remains relatively constant at -1 mA before it

increases slightly at high contactor potentials. This increase in current is

presumably due to collection of electrons ejected from the tank wall by high

energy ions (-140 eV) rather than any improvement in the ion production capability of the contactor at high coupling potentials.

At each of the solid data points identified by letter in Fig. 12 plasma

potential profiles were measured using the emisslve probe and these profiles,

identified by the same letters, are shown in Figs. 13 and 14. Under electron

emission conditions (Curves C, D° and I) the potential profiles are seen to be

quite flat except immediately adjacent to the contactor, i.e.° within about $ cm,

where a voltage drop indicative of a sheath exits. This must be a double

sheath which develops as the contactor emits electrons to and draws ions from

the simulated plasma, but the sheath is located close to the contactor in a

region that is difficult to probe. If a high plasma density, constant potential region exists adjacent to the contactor is not apparent from these measurements. Curve H in Fig. 14 shows a less severe potential gradient near the

contactor than the other electron emission profiles presumably due to its

lower emission current, (500 mA as opposed to 650, 1000, and 1000 mA). Under

conditions of electron collection (Curves A and B) the profiles are observed

to be quite different. In general they show a region of relatively constant

-27-
H

ELECTRON

EMISSION CURRENT

( JCE ) (mA)

-I000

750

500

250

JcA=O.5 A

VCA=I5 TO 22V

rh =2.7 sccm (Xe)

Pt =3 x I0 -s Torr

SOLID SYMBOLS DESIGNATE

CONDITIONS WHERE POTENTIAL

PROFILE DATA WERE TAKEN

Fig. 12

SCALE

CHANGE

-5O

Conventional Hollow Cathode

Charaeterlstlos

5O

CON TACTOR POTENTIAL

(V c+ VCA) (V)

I00 150

I

G F E

SIMULATOR OFF

B

SIMULATOR ON A

-28-
>

._1

<_

m

jZ

hi

0

=E

(/)

..J

O.

140

120

I00

80

JCA = O. 3 A

VGA= 15 TO 22V

rh¢ = 2.7 sccm (Xe)

Pt = 3x I0 -s Torr

A (JCE=- 24 mA)

6O

40

2O

B (JcE =- 18 mA)

ELECTRON

COLLECTI ON /

ION EMISSION

AXIAL POSITION (cm)

0 I I I I I

20 40 60 80 I00

f ELECTRON EMISSION /

-20__

40V _D (JcE =1000 reAl

r

-60 -

Fig. 13 Typloal Plasma Potential Profiles for

Conventional Contaotor with Simulator on

-2P-
140,-

_>

_1

<C

m

I--

Z

bJ

F0

O.

co

.J

a.

120

IO0

8O

60

40

20

0

-20

-4O

- 6O

-8O

ION EMISSION

ONLY

JCA = 0.3 A

VCA = 15 TO 22V

rh =2.7 sccm (Xe)

Pt =:SxlO-e Tort

I

2O

AXIAL POSITION (cm)

I I I I

40 60 80 I00

, f--H (JcE =500 mA)

-I (JcE=IO00 mA)

ELECTRON EMISSION ONLY

Fig. 14 Typical Plasma Potential Profiles for

Conventional contactor with Simulator off

-30-
potential plasma adjacent to the contactor followed by a region of steep

potential variation and a final region of near constant potential (i.e. they show

the double sheath structure identified by the solid llne in Fig. 9a). Curves

E, F, and G show ion emission only profiles which remain relatively constant

at a plasma potential ~20 V below the contactor anode throughout the 1 m long

region t_at was probed. These curves must drop to zero potential at the tank

wall so it is suggested that a large sheath develops near the tank wall and

simulator anode which collects the ions being produced in the contactor

discharge.

An examination of the data of Fig. 13 suggests that the contactor potential (V C + VCA) , used to characterize plasma contactor performance in curves

llke those in Fig. 8 and 12, reflects a higher potential drop than the one

actually observed between the contactor and the simulated space plasma. For

Curve A in Fig. 13 for example, the contactor potential (i.e. the contactor

anode-to-tank potential difference) is 140 - 0 = 140 V while the contactor

anode to simulated space plasma potential drop (V s) identified in Fig. 7 is

only 140 - 50 - 90 V. Similarly for Curve D the magnitude of the contactor

potential would be 58 - 0 = 58 V while the more correct magnitude (Vs) would

be 58 - 37 - 21 V. Use of this contactor sheath potential drop (V_ removes

the potential drop experienced across the sheaths at the tank wall and simulator from the characterizing potential describing plasma contactor performance.

Unfortunately most of the data collected during the grant period did not

include ambient plasma potential so most of the performance data are presented

hereln in terms of eontaotor potential. Where the contactor potential (V c +

VCA) has been used in this report as the characterizing potential instead of

the observed sheath potential drop across the contactor, the resultlng curves

are pessimistic representations of the performance of a particular contactor.

However, since the simulator operating conditions have been kept fixed

-31-
throughout these tests, trends identified by comparing performance curves

plotted in terms of contactor potential can be assumed to be correct.

Tests conducted early in the grant period showed that connecting the

simulator cathode to ground as opposed to connecting its anode to ground

shifted the electron collection portion of the contactor characteristic curve

but left the curves otherwise unchanged. This effect is shown in the data of

Fig. 15. The potential offset associated with making this change is similar

to the change in reference potential introduced by changing the connection

(i.e. about equal to the simulator anode voltage). It was expected however,

that changing the point of tank connection to the simulator would make little

difference in the characteristic curves if the contactor sheath potential (Vs)

were used to plot the curves instead of contactor potential (Vc + VCA). This

has in fact been observed and consequently, it is argued that changing the

simulator reference point shifts the characteristic curve along the horizontal

axis when the contactor potential (VC + VCA) is used, but does not change the

coupling process occurrlr_ at the contactor. When this test was repeated

using the variable area anode contactor and the data were presented in terms

of the contactor-to-simulated space plasma potential difference rather than

the contactor anode-to-slmulator anode potential difference the data of Fig.

16 were obtained. These results show that changing the point of connection

between the tank and the simulator from the simulator anode to the simulator

cathode (i.e. changing the simulator reference point) does not change the

contactor's performance characteristics. This result was expected and it

again suggests that one should use the contactor-to-amblent plasma potential

difference to characterize contactor performance.

Ring Cusp Contactor Optimization Study

Experiments conducted on ring cusp contactors at both NASA Lewis ? and

-32-
ELECTRON

EMISSION CURRENT

[JcE] (mA)

-I000

-750

QUALIT Y

JCA= 0.5 A

VCA= 18 TO 33 V

rnc = 1.4 sccm (Xe)

Pt =4 x I0 -s Torr

- 500

250

Fig. 15

-5O

I

C.,NO,

SIMULATOR ANODE '<_

- 25 CONNECTED TO TANK _

b

-50

Effeet of Changing Point of Conneetlon to

Simulator-Conventlonal Hollow Cathode

Contaetor

-33-
ELECT RO N

EMISSION CURRENT

'JcE ] (mA)

-I000

- 750

JCA = 0.5 A

VCA=IOTO 14V

= 2.7 sccm (Xe)

Pt = 3.5 x I0 -e Tort

5OO

0 -SIMULATOR ANODE

GROUNDED

0 -SIMULATOR CATHODE

GROUNDED

250

CONTACTOR ANODE TO PLASMA

POTENTIAL DIFF. (V)

-250

-500

150

I

Fig. 16 Effect of Changing Point of Connection to

Simulator-12 em dla Anode Hollow Cathode

Contactor

-34-
Colorado State University have demonstrated that these contactors are more

efficient ion sources than conventional hollow cathodes. This sugEests, in

light of the model in Appendix A, that they should be better plasma contactors.

This conclusion is a consequence of the following logic sequence: 1) Higher

ion production rates by a contactor operating at a prescribed power and electron collection current induce an increase in the size of the high density

plasma region (r i in Fig. 2 increases); 2) Because the electron collection

current and the ambient plasma density are prescribed, the radius of the outer

boundary of the double sheath (r o in Fig. 2) remains unchanged at a value

required to satisfy the condition of equal current densities across this boundary_ 3) As a consequence of I and 2 the radius ratio rl/r o increases and in

accordance wlth the model of Appendix A the voltage drop across the double

sheath decreases at the prescribed current level. Hence one concludes that

the contactor with the higher ion production would hold a spacecraft closer to

ambient plasma potential than a conventional contactor operating at the same

current level.

Experiments were conducted on the rlr_ cusp contactor shown in Fig. 7 to

optimize its ion production efficiency by adjusting its plasma extraction

aperture area. An optimum extraction area was expected because masking down

the extraction aperture was expected to induce two competing effects, namely,

1) improvement of neutral atom confinement and hence an increased rate of

plasma ion production and 2) reduction of the fraction of these ions that

would escape through the smaller extraction aperture. Because of these competing effects an optimum ion production rate was expected at an aperture

diameter somewhere between the open ended (no mask) condition of 6.5 cm and

zero. Thls optimum aperture contactor would then be expected to outperform

the unoptlmlzed configurations due to its larger extractable ion current at a

given operating condition.

-35-
The optimum extraction area test conducted on the ring cusp contactor,

accomplished using the test apparatus schematic shown in Fig 4a, involved the

installation of a series of masks (each having a different aperture area) over

the 6.5 cm dla ion source extraction aperture. The masks were moved into

place one after the other while the ion source was operating so the effect of

changing only the aperture area could be studied. Masks having 2.7, 4.4, 5.5,

and 6.5 cm diameter apertures were selected for the study so the aperture area

could be changed from 5.7 cm 2 to 33 cm 2 in equal area increments. As the

arrar_ement of Fig. 4a suggests, the mask and the body of the ion source were

connected through a switch to a power supply that biased them sufficiently

negative so any ions arriving at their surfaces (J - ion current to mask and

m

Jb - ion current to body) could be measured. Using these measured currents

together with the ion beam current extracted from the source to the vacuum

tank, which could also be measured with the simulator off, it was possible to

determine the plasma ion energy cost and the extracted plasma ion energy cost

over a range of operating conditions. From these measured quantities the

parameters defined by Brophy $ to characterize discharge chamber performance,

namely the primary electron utilization factor CO and the baseline plasma ion

cost ap were computed.

In presenting the data associated with these tests, atomic flowrates

will be expressed in milliamperes equivalent (mA eq.). This unit represents

the current of atoms that would flow if each atom carried the charge of an

electron. It is useful in describing ion source performance because it

expresses flowrate in the same units as the ion current being extracted from

the ion source. Hence the ratio of ion current to flowrate (the expellant

utilization) can be computed directly with no conversion factor if multiply

charged ion production is small. Because the sccm unit is more familiar to

plasma contactor researchers it will be used to express flowrates for all

-36-
contactors except the ring cusp ion source. Conversion between the two sets

of units can be made readily be recognizing that a flowrate of 1 sccm of any

gas corresponds to 67 mA eq.

The typical plasma ion energy cost curve shown in Fig. 17 pertains to a

4.4 cm dla ion extraction aperture. This curve was obtained by selecting

values of C and e by trial and error and then using them in the discharge

o p

chamber model to determine a best fit curve to the measured data points

shown. The data were collected by holding the cathode flowrate, mc' at

93 mA eq (Xe) and varying the flowrate through the body (main flow plenum),

mb' from zero to 280 mA eq. (Xe). The scatter in the data about the curve

shown in Fig. 17 induced by the changes in flowrate typifies that observed

with all of the masks. The systematic scatter about the curve at each

flowrate is symptomatic of a minor deficiency in the basic model being

applied to describe the curve. This scatter is probably due to changes in

baseline plasma ion energy costs that are believed to occur at low discharge

voltage conditions, like those being encountered in these tests, that are not

considered in the model S. Over the range of flows and discharge currents

investigated, the discharge voltage varied but it was typically found

to be within a few volts of 25 V as indicated in the legend of Fig. 17.

Figure 18, which shows a comparison of curve fits obtained for the various

aperture diameters investigated, indicates that the smallest diameter aperture

(2.7 cm) gives the lowest plasma ion energy cost curve and so from the pointof-vlew of producing ions only, it would be preferrable. Figure 19 shows the

effect of mask aperture diameter on the other important factor, namely the

fraction of ions produced that are extracted from the discharge. These data

-37-
t.O

(5

II I I

0

)

I I

0 0

0 0

CO tO

I I I I I I I

0 0 0 0 0 0 0

0 0 0 0 0 0 0

N 0 cO _0 _1"

(uo, ouJsold/Ae) [ d_ ] 1S00 Ag_I3N3 NOI VlNSV7cl

:U3

(5

<

-o

0

,,=,

r_ P-

• IJJ

o

<_

n-

<_

{1.

>-

pro

o3

¢_J Z

"0 w

J

E

LtJ

Z

-Q

0

-38-
I I I I

0 0 0 0

0 0 0 0

( uom otUSOld/Ae)

I I I I I

0 0 0 0

0 0 0 0

0

d

d ,

m

Q=

t_

W

d _

>-

tO

Z

• W

0 a

..I

n.-

I--

LU

Z

0

0

g.

4.)

•_ 0

L(.)

L

mr=]

0

O_

-3 9-
+

E

'=3

÷

I.iJ

o

"3

uJ

(.)

,-j

II

1,1.1

(,_

i

0

m

I--

¢J

E

i,

Z

O

a

i,i

O

E

X

hi

0.6-

0.5

0.4

0.3

0.2

0.1 "

0

rhc = 9:5 mAeq (Xe)_ /_ rnb=370mAeq (Xe) /

/

/

/

/

/

=9:3 mA eq (Xe)

rhb=0

7 cm DIA RING CUSP CONTACTOR

Vow, 25V

I I I

2 4 6 8

MASK APERTURE DIAMETER (cm)

Fig. 19 Effect of Mask Aperture Diameter on

Extracted Ion Fraction

-40-
indicate that the larger the mask aperture the greater the fraction of ions

produced that escape through the aperture. Note in regard to the data of Fig.

19 that the extracted ion fraction is affected by flowrate when the mask diameter is large. Discharge power did not cause the extracted ion fraction to

change, however, and this was expected based on the discharge chamber model. 8

Considering the data of Figs. 18 and 19 together it is apparent that

increasing the mask aperture diameter has the beneficial effect of increasing

extracted ion fraction and the detrimental effect of increasing plasma ion

energy cost. Hence there will be a diameter at which these effects will balance and contactor performance will be optimal. By plotting the extracted ion

energy cost versus expellant utilization at a particular flowrate in the

manner suggested in Fig. 20, the optimum mask aperture diameter is found to be

5.5 cm. This plot demonstrates that reducing the aperture area from 6.5 cm to

2.7 cm in increments causes the curves to move down and to the right and then

back up and to the left again. Since the preferred curve is the one lowest

and furthest to the right Fig. 20 suggests that the $.5 cm dia aperture yields

the highest extracted ion production rate per unit energy expenditure. All

flowrates investigated yielded the same 5.5 cm dla mask aperture as the

optimum, however, increasing the flowrate through the plenum was found to

improve the ion production performance of the source. This result, shown in

Fig. 21, is a reflection of effect of flowrate on extracted ion fraction

indicated by the data of Fig. 19.

It is also of interest to determine whether changing the extraction aperture diameter of the ion source affects its ability to emit electrons to or

collect electrons from a simulated plasma. In order to determine this, the

-41-
2500 -

c

0

°_

> 2000-

t"----"l

I.l.I

'-3

,_ 1500 -

0

.=_

1--

03

0

tO

>-

t,.9

I000-

i,i

Z

i,i

Z

0

1

a

i,i

tO 500 -

I--

X

W

0

7 cm DIA RING CUSP ION SOURCE

rhc = 93 mA eq (Xe)

rhb = 186 mA eq (Xe)

/__---2.7 cm DIA

MASK APERTURE

,P

¢

NO MASK_ / ,'d

• r7 5.5 cm DIA

ID----''ET "-4.4 cm DIA

MASK APERTURE

I I I I

0.1 0.2 0.3 0.4

EXPELLANT UTILIZATION [JcE/rh] (mA/mA eq)

I

0.5

Fig. 20 Effect of Mask Aperture Diameter on

Extracted Ion Energy Cost

-42-
2500 -

c

0

.m

2000

tO

:>_o1500

i

I000

,.=,

o_=

1-

500

lX

ILl

RING CUSP PLASMA CONTACTOR

5.5 cm DIA MASK APERTURE

the = 95mA eq (Xe) ALL CASES

mb = 93 mA eq (Xe)

rhb= 186 mA eq (Xe)

I I I I

0 0. I 0._ 0._ 0.4

EXPELLANT UTILIZATION [JB !rh]

rhb=372 mA eq (Xe)

I

0.5

Fig. 21 Effect of Flowrste on Ring Cusp Contactor

Ion Cost

-43-
simulator was started and the ring cusp contactor was biased using the coupling supply to effect electron emission and collection through the various

mask aperture diameters investigated. Figure 22 shows the characteristic

curves measured in this way at a contactor discharge power of -20 W. The

curves suggest that installing masks with aperture diameters varying from 4.4

to 6.5 cm results in essentially the same characteristics although the one for

the 5.5 cm dia mask does seem to be the best. Note that the very small mask

aperture (2.7 cm dla) appears to cause the ion source to perform substantially

poorer. Since the spherical double sheath model 3 suggests that reducing the

effective radius of the plasma surface from which ions are drawn in the electron collection mode or electrons are drawn in the electron emission mode

causes the sheath voltage to increase at a prescribed current level a degradation of performance with this low aperture area was expected. The conclusion of this work is that there is an optimum aperture which should" effect the

most efficient operation of a plasma contactor. For the 6.5 cm dia ring cusp

contactor this aperture diameter is approximately 5.5 cm.

Effect of Anode Area on Contactor Performance

A comparison of typical performance curves observed with the conventional

hollow 8athode contactor with those associated with the ring cusp plasma contactor (e.g. Fig. 12 vs. Fig. 22) shows that the ring cusp contactor performs

better than the conventional one , especially in the electron collection

region. It has been suggested in this report that this could be because the

ring cusp contactor has a hlgner ion production rate than the conventional

contactor, and it has also been suggested that it could be because the effective area of the ring cusp contactor for contacting the ambient space plasma

-44-
?

I

!

t

ELECTRON

EMISSION CURRENT

[_c_](mA)

I000

-20

I

%

\

800

600

400

200

-200

-400

7 cm DIA RING CUSP CONTACTOR

mc -- 93 mA eq (Xe)

mb= 186mA eq (Xe)

DISCHARGE POWER

~ 20 Watt s

CON TACTOR POTENTIAL

[v_+vo.] (v)

40 60 80 I00

I I I I

o

4.4cm DIA MASK APERTURE

MASK APERTURE •

% NO MASK

5.5 cm DIA

MASK APERTURE

b

Fig. 22 Effect of Mask Aperture Diameter on Ring

Cusp Contractor Performance

-45-
is larger than that for the conventional cathode (6.5 cm for the ring cusp

contactor vs. 8 mm for the conventional contactor). In order to investigate

the effect of anode area on the performance of a contactor, the apparatus

shown in Fig. 7 was constructed. It consists of a conventional hollow cathode

with a sustaining anode that is identical to the conventional contactor with

three added (auxiliary) surfaces that can either be held at anode potential or

allowed to float. When floating, it is assumed that these surfaces do not

participate slgniflcantly in the contacting process and since they can be

switched between the floating and the anode potential conditions during operation using the circuity shown in Fig. 4b, the effect of anode area on contactor performance can be examined readily. Curves showing the effect of changing the anode diameter on contactor performance are presented in Fig. 23.

Data points have been left off of these curves because of the large number of

curves and the tendency for the data points to overlap and clutter the figure,

however, the individual curves are representative of the actual performance of

the contactor observed. These curves show that with either the sustaining

anode or the sustaining anode and anodes 1 or 2, the performance is similar to

the performance of the conventional contactor in the electron collection

region, i.e. the electron collection currents are in a range less than 100

mA. However, when anode 3 or all of the anodes are connected the contactor

undergoes a transition to what has been termed the ignited mode of electron

collection. In this mode a substantial increase in the electron collection

current to values over a few 100 mA at potentials less than 100 V are

observed. Visual observation of the contactor discharge during these transitions has shown that the transition to the ignited mode is accompanied by the

development of a luminous plume having a size similar to the outer diameter of

-46-
Fig. 23

ELECTRON

EMISSION CURRENT

[JCE] (mA)

t -I000

I

I

I

I

I

I

I

I

I - 750

I

I

/

- 500

rh =4.1 sccm (Xe)

Pt = 4 x I0 -e Torr

JCA = 0.3 A

VCA =-28 TO 18 V (IGNITED)

VCA = 12 TO 18 V (UNIGNITED)

250

CON TACTOR POTENTIAL

[Vc*Vc,] (v)

-50 50 I00 150

I L

SUSTAI N ING ANODE

,- _. (S) ONLY (Icm DIA)

I

II JODE S AND I

IGNITION _ i (3 cm DIA)

TRANSITION I ANODES S AND 2

_ - (7cm MAX DIA)

250 REGION // L

I- ; -"'ANODES S AND :3

_N L (12 cmMAx DIA)

0 ES S, I, 2 AND 3

-500 (12 cm DIA)

Effect of Anode Diameter on Contactcr

Performance

-4?-
the Srd anode. Since the observed luminousity is due to atomic or ionic excitation it is suggested that ionization is also occurring and that it could be

the cause of the improved contacting performance accompanying the transition

to ignited mode electron collection. Figure 24 is a typical comparison of the

electron current collected by the contactor (JcE) compared to the electron

emission current from the simulator (JsE) . These curves are similar and

because there is no substantial change in the degree of agreement between

them it is argued that the transition to the ignited mode indeed results from

plasma phenomena, and not from electron currents being drawn from the vacuum

tank walls.

Figure 25 shows a series of plots obtained using the 12 cm dia anode

contactor operating under conditions similar to those associated with Fig. 23

but where the contactor discharge did not undergo transition to the ignited

mode. The data of Fig. 25 were collected soon after the ignited mode of

electron collection was first observed and in this particular test transition

to ignited mode operation with the large diameter anode could not be

effected. It was believed at the time that transition to the ignited mode

was inhibited by an insulating layer of fractionated diffusion pump oil on

the anodes that prevented them from functioning properly. When the anode

discoloration was removed after the test, transition to ignited mode operation was achieved readily. Since that time, however, the varlble anode area

contactor has been operated for long periods of time and has become coated to

the point where its appearance was worse than it was during this early test.

Yet, transition to ignited mode operation has been initiated regularly with no

difficulty. This suggests that the coated anode was not the reason for the

inability to induce the transition. More recent tests indicate that transition to ignited mode operation may be affected by small changes in simulator

-48-
o

oo

÷

o

_ot0

tZ

0

0

0

c_J

0

i

_0

QW

Li.

lib.

OW

Z

Wv

ZC3 ,,, n'z

ow _ _ w._

I--I-- _,-_ b. l'- r o__. ',,_

r

\ z__w "_

0

0:: 0 0

Z

O_ _C_ _:] _

"'_ d d \--_

q

"tl z_ r,.

,'I 2

0 0

x_

ET w

u 0 Q u

_ 0

_ X Z

II II II

u

_.E_

I I I I I

0 0 0 0 0

0 0 0 0 0

I I I I I

(V u_) IN3_I_NO N0_II0373

o

e.l

e4

",4

-49-
ELECTRON

EMISSION CURRENT

[JcE] (mA)

-I000

Fig. 25

-750

JCA =0.9 A

VCA=I2 TO 19V

rh =4,1 sccm (Xe)

Pt =4 x 10-6 Ton"

-5OO

-25O

CONTACTOR POTENTIAL

[vc+vc.](v)

50 I00 150

I _ I I I

.O;T_

SCALE _ _ _ _ANODE S ONLY

CHANGE "_\\ \.. _( (I cm DIA)

\_ '\ _--ANODES SAND I

5 0

/ "..._-_ ANODES S, I AND ?-

/ -"_'_,_ \. \ (7 cm DIA)

ANODES S,',2, -_'_."% N

_NO_(,__ o,_)".:_.\

-I00 - ___

Anomolous Contaetor' 9erfor'mance Cur'ves

-50-
operating conditions. At this point it is uncertain why occasional

difficulties are encountered in achieving ignited mode operation.

Figure 26 shows the effect of contactor anode current JCA on the transition into ignited mode operation for an anode dia. of 12 cm. These data suggest that increases in the contactor anode current caused the transition to

ignited mode operation to shirt to progresslvely lower contactor potentials.

If the contactor flowrate is increased from 1.4 sccm to 4.1 sccm the contactor

potential at which this transition occurs becomes relatively unaffected by the

contactor anode current as shown in Fig. 27. Comparison of Figs. 26 and 27

shows, however, that the increased _enon flowrate itself induces a substantially decreased contactor potential at which the transition occurs. A

further increase in contaetor flowrate to 6.8 sccm induces an additional

decrease in the contactor potential at which this transition occurs as the

data of Fig. 28 indicate.

Typical voltage/current characteristic curves for the 12 cm dla anode

contactor operating with and without the simulator on are shown in Fig. 29.

The data on these two curves can be compared with those of Fig. 12 to see the

effect of increasing the contactor anode diameter on the performance curves.

This comparlson shows that curves obtained with the simulator off are essentially identical while those associated with the simulator on exhibit a

dramatic increase in the electron collection current beyond the point of transition into the ignited mode of electron collection. Comparison of the data

shows that the contactor anode voltages (VcA) observed over the full range of

electron emission currents are the same (i.e. this voltage is Independant of

the anode diameter). At each of the solid data points identified in Fig. 29,

plasma potential profiles along the tank cen_erllne downstream of the contac-

-51-
!

-50

I

ELECTRON

EMISSION CURRENT

-I000

-250

750

Fig. 26

0

Ill

! ! I !

JCA VCA

0.3A 16 TO 27 V

0.6A 14 TO 27 V

0.9A 19 TO 28 V

,5OO

rh = 1.4 sccm (Xe)

Pt = 2.2 x I0 -6 Torr

SURFACES S, t, 2, 3

AT ANODE POTENTIAL

250

CONTACTOR POTENTIAL

[Vc*Vc.] (v)

50 I00 150

\\\

- _

• IGNITED

-500 -

Effect of Contaator Anode Current on

Performance at Low Flowrate

-52-
Fig. 27

-5O

I

-250

-500 -

-750

JCA VCA

0.3 A -28 TO 17 V

0.6 A 12 TO 19 V

0.9 A 14TO 23 V

-250

rh = 4.1 sccm (Xe)

Pt = 4 x I0 -e Torr

SURFACES S, I, 2, :5

AT ANODE POTENTIAL

5O

CONTACTOR POTENTIAL

[Vc+VcA] (V)

I00 150

I I

IGNITED

ELECTRON

COLLECTION

Effect of Contactor Anode Current on

Performance at a Moderate Flowrate

-53-
Fig. 28

ELECTRON

[JcE] (rnA)

-I000

750

500

JCA VCA

-250

!

0

D

-250

-500 -

0.3A -34 TO 13 V

0.6A II TO 14V

O.9A II TO 13V

rh = 6.8 sccm (Xe)

Pt = 5x I0 -s Torr

SURFACES S, I, 2, 3

AT ANODE POTENTIAL

CONTACTOR POTENTIAL

[Vc+vc_] (v)

I00 150

I I

I

IGNITED

ELECTRON

COLLECTION

Effect of Contactor Anode Current on

PerformanQe at a High Flowrate

-54-
Fig. 29

H

G

_E

ELECTRON

[JcE] (mA)

-I000

_D

-50

- 250

- 750

5OO

250

JCA =0.3 A

VCA =14 TO ?.3 V

=2.7 sccm (Xe)

Pt =3 x I0 -6 Torr

SOLID SYMBOLS DESIGNATE

CONDITIONS WHERE POTENTIAL

PROFILE DATA WERE TAKEN

CONTACTOR POTENTIAL

[Vc+Vc,] (v)

50 I00 150

"oo_o-_@...k_..o _ 'p =

"_,C {' F

SIMULATOR ON > t\ "SIMULATOR OFF

%

- 500 -

Typical 12 c= Dia. Anode Contactor

Cbaraeteriatles

-55-
tor were measured using the emisslve probe. These profiles are identified in

Fig. 29 by letter and also in Figs. 30 and 31 by the same letter for the two

situations being considered (simulator on and off). Under conditions of the

higher electron emission currents (Curves D, E and H) relatively flat potential profiles are observed except within about S cm of the contactor, where a

voltage drop indicative of a sheath exists. In case G where the simulator is

off and the contactor emission current is the lowest, this sheath extends considerably further downstream. It is noted that the same behavior was observed

in the curves obtained in Figs. 13 and 14 for the conventional contactor. The

similarity of these curves indicates that the anode size has no significant

effect on contactor performance when it is operating in the electron emission

mode. Note also that the shapes of all of these curves suggest the existence

of a sheath. It would be a slr_le sheath when the contactor is emitting electrons to the tank in the case where the simulator is off (Curves G and H) and

it would be a double sheath in the case of Curves g and D where electrons are

being emitted and ions are available to be drawn from the simulated space

plasma.

Under conditions of electron collection/ ion emission (Curves A, B, C and

F) the plasma potentials associated with the profiles are positive relative to

the tank. With the simulator off (Curve F) the profile is quite flat and is

almost identical to the profile observed with the conventional contactor at an

ion emission current near 1.4 mA. In general the curves obtained with the

simulator on (Curves A, B and C) show a region of relatively constant plasma

potential adjacent to the contactor followed by a region of steep potential

variation and a final region of relatively constant potential. These profiles

are also slmllar to those obtained in the electron collection mode with the

conventional conatactor (Fig. 13).

-56-
>

.J

I

Z

hi

Q

Q.

if)

.J

140 -

120

I00

8O

6O

4O

2O

0

-6O

_ _r _-A

(JcE =- 300 mA)

ii=_]

i

i

JCA =0.3 A

= 2.7 sccm (Xe)

Pt =3 x I0 -s Torr

12 cm ANODE DIA

_t _C (JcE= -29 mA )

_-- - --.z..... ELECTRON COLLECTION /

_ _ ION EMISSION

2O

AXIAL POSITION (cm)

I I I

40 60 80

I

I00

I

--,- __ fo

ELECTRON EMISSION /

ION COLLECTION

Fig. 3 0 Typlcal Plasma Potential Profiles for the

12 cm Dia. Anode Contactor-Simulator on

-5?-
D,

J

Z

_J

0

Q.

:E

J

0.

120

I00

8O

60

2O

0

___.=.._f__F (JcE =-1"4mA)

ION EMISSION _

JCA =0.5 A

VCA =14 TO 23 V

rn =2.7 sccm (Xe)

Pt =3 x I0 -s Torr

AXIAL POSITION (cm)

I I I

20 40

I

6O 8O

G (JCE =500 mA)

I

I00

ELECTRON

EMISSION

H (JcE = I000 mA)

Fig. 31 Typical Plasma Potential Profiles for the

12 cm Dia. Anode Contactor-Simulator off

-58-
When several potential profiles sweeps are made with the emlsslve probe

at various radii from the contactor centerline a potential contour map like

the one shown in Fig. 32 can be constructed. These data, obtained with the 12

cm dia anode contactor operating in the ignited electron collection mode, are

typical and show that the double sheaths suggested by Curves A, B and presumably C of Fig. 30 are really relatively spherical in shape. It is noted that

in the case of the Fig. 32 data the contactor cathode was 133 V positive of

tank potential and the contactor anode was 156 V positive of that potential.

This resulted in an electron collection current of 175 mA. These contours

again suggest that the potential is relatively uniform in the region very

close to the cathode (i.e. the hlgh density plume region) and that it then

drops relatively rapidly from about 140 Y down to 50 V. At 50 V a potential

well develops about 17 cm downstream of the contactor anode beyond which the

potential rises and and then levels off at a value between 50 and 60 V. Data

not included in Fig. 32 indicate that the potential remained approximately

constant from the outermost contour line to a plane 1 m downstream of the contactor. The fact that the contours take on a somewhat hemispherical shape

suggests that electrons are flowing nearly radially inward and ions are flowing nearly radially outward from the hlgh density plasma located within the

140 V contour. It should be noted that electrons could also be flowing to

the upstream surface of the-anode shown in Fig. 32 so the electron collection

process could also be occurring over a full spherical surface.

Since the typical sheath contours of Fig. 32 are essentially spherical

and the potential difference across the sheath can be measured using the emlssire probe the experimentally measured data can be compared to the predictions

of the simple spherical double sheath model mentioned in the introduction to

this section and developed fully in Ref. 3 (reproduced in Appendix A). In

order to determine the extent to which the model describes the physical

-59-
I

O

m

I

0

E

(.1

Z

0

0 >

.J

× _

0

o_

0

Pt

0

0 _

g,j _

I

-6O-
behavlor of the plasma contacting process as represented by potential profiles

llke those in Figs. 1S and 30 a number of these profiles were compared to the

predictions of the model. The method by which this has been done can be

understood by reconsidering Fig. 9a. In order to apply the model, the quantities that must be measured are the experimental voltage drop across the sheath

(Vs) in Fig. 9a, and the experimentally determined radii associated with the

upstream and downstream boundaries of the double sheath region. In Fig. 9a

these are designated ru and rd. These points were determined from the experimental profiles by defining the upstream and downstream boundaries as the

points of minimum and maximum second derivative of the potential with respect

to position respectively. Using the measured potential drop V s together with

the measured electron collection current JCE the normalized current from the

outer surface Jo could be computed by substituting Vi = Vs and Jo = JCE into

Eq. (1). This value of Jo could then be used to enter Fig. 4 of the paper in

Appendix A to determine the theoretical sphere radius ratio rl/r o predicted

for the collection of the current JCE at a sheath potential drop of magnitude

Vs . This theoretically- based radius ratio could then be compared to the

measured radius ratio ru/rd .

A typical example of the comparison of a measured potential profile with

one computed using the equations from Appendix A is shown in Fig. SS. In this

case, the electron current from the outer sphere JCE = Jo was 250 mA. The

potential drop across the double sheath, Vs - V i was 10S - 48 = S$ V and the

normalized electron current from the outer sphere Jo = 8.96 was computed using

gq. 1. From Fig. 4 in Appendix A this value of Jo corresponds to a theoretically based radius ratio (rl/ro) of 0.7S. The experimentally determined value

of the radius ratio ru/r d determined from the solid llne profile in Fig. SS is

7.8/12.8 = 0.61. Hence theory predicts a profile that is steeper than the

experimentally measured one as Fig. SS shows.

-61-
I00

80

60

=o

<{

2O Jo= 8.96 = o82/

0 5

I I I

I0 15 2O

AXIAL POSITION (cm)

I

25

Fig. 33 Typical Comparison of Measured and

Predicted Plasma Potential Profiles

-62-
Comparlson of the experimentally and theoretically-determlned radius

ratios obtained in the described manner over a wide range of operating conditions is presented in Fig. 34. In obtaining these data the anode diameter of

the contactor was varied from 1 to 12 cm, the hollow cathode discharge power

was varled from zero to 17 W, the xenon flowrate was varied from 1.2 to 6.8

sccm and this also resulted in tank pressure variations from 2 to 10 x 10 -6

Torr. The extent to which the theoretical model describes the physical

results is reflected in the proximity of the data points shown to the solid,

perfect fit llne of Fig. 34. Examination of the data in this figure suggests

that the predicted radius ratios agree with the measured ones to within about

2_ and the predicted ratios are systematically greater than the measured ones. It

is noted that points that fall below the -2R_ error boundary are ones that were

obtalned elther under conditions where slgnlflcanc electron current was being

drawn from the tank (approximately lO&) or extra expellant other than that

through the hollow cathode orifice was being fed into the vacuum chamber to

increase the vacuum tank back pressure. Operation at either of these conditions could be expected to cause the assumptions of the theoretical development to break down because both electron emission from any point other than

the outer sphere in the theoretical model and the effects of ionizing colllsions in the double sheath have been neglected in model development.

Reconsideration of the potential contour map of Fig. 32 suggests that it

might be more appropriate to consider the current flow process to be occurring

between hemispherical rather than spherical surfaces. In order to investigate

this, the equations developed in Appendix A were redeveloped to yleld the

current flowing from an outer hemisphere:

-63-
o

pj/nj] OI.L_N _nlC]_l -]_.LN31NI_3dX3

-64-
Jo = 2n8o Vi3/2 _/ 2_._ee Jo

mo

(3)

Application of this equation in place of Eq. 1 to the same data as that used

to obtain the results presented in Fig. 34, yielded the results shown in Fig.

35. Since these data appear to be skewed further from the perfect fit line in

the results of Fig. 34, it was concluded that the spherical case represents a

better flt to the experimental data than the hemispherical one.

Circuit Considerations at High Electron Collection Currents

In early experiments conducted on plasma contactors during the grant

period the electron collection currents were very small compared to the contactor and simulator anode currents (JCA and JSA) identified in Fig. 3. When

ignited electron collection operation occurred this was no longer the case. As a

result some operational constraints developed which can be best understood by

looking at detailed current balances associated with operation at points B and

C in Fig. 29. Electron current balance diagrams associated with these two

operating conditions are presented in Fig. 36. In the unlgnited situation

(Fig. 36a) 38 mA of electron collection current is picked up at the eontactor

anode and this combines with 262 mA of electron emission from the contaetor

cathode to the contactor anode to produce the 300 mA contactor anode current

which was being maintained by the anode power supply during the test. At the

simulator where the 38 mA electron current collected at the contactor was

produced, a 262 mA electron current had to be collected at the simulator anode

because a 300 mA electron emission current from the simulator cathode was being

maintained by the simulator anode power supply. In this particular case the

electron current drawn to the contaetor anode from the simulator cathode was small

compared to the anode and cathode currents being maintained at the contactor and

-65-
I

0

0

m

0

\

0 0

-66-
+14V

[-]

+39V

300 mA

262

[

CONTACTOR

300 mA

38 mA L

a. CONDITION C (UNIGNITED)

58 mA

~0

TANK

SIMULATOR

__.,_ 262 mA

II.--

CONTACTOR

+8.5V

[ ]

"--300 mA

_.O..OmASOMA_.) i

.94vT125°mA

250 mA

b. CONDITION B (IGNITED)

250 mA

AN K

Flg. 36 Effect of Transition to Ignited Mode

Operation on Electron Current Balance

SIMULATOR

-67-
simulator, respectively.

When the contactor underwent transition to the ignited mode, however, the

situation changed (Fig. 36b) because a 250 mA electron current was conducted

between the simulator and the contactor and this current was comparable to the

currents being maintained to the contactor anode and simulator cathode. Under

this operating condition a limitation on the contactor anode and simulator

cathode currents effectively limited the electron current flowing between the

simulator cathode and contactor anode. This effect is illustrated even more

dramatically at operating Condition A defined in Fig. 29 which corresponds to the

current balance situation of Fig. 37a. In this case limiting the simulator

cathode and contactor anode currents results in a limitation on the current that

can flow through the plasma to the extent that no current is collected at the simulator anode, 18 mA is drawn from the tank wall, and 30 mA of electrons are

collected at the contactor cathode. In order to accommodate this operating

condition the simulator anode is actually negative of the simulator cathode

and the contactor anode is only 1.5 V positive of the contactor cathode. This

situation can be rectified so a prescribed current can continue to flow

through the simulator anode/cathode loop if the point of connection between

the vacuum tank and the simulator is transferred from the simulator anode to

the simulator cathode in the manner suggested in Fig. 37b. When the connection is made in this way, it is possible to sustain the desired simulator

anode/cathode loop current (300 mA in this case) while supplying any contactor

collection current that may be demanded. Further, the experiment-disturbing

electron current drawn from the vacuum tank wall drops to zero as the data of

Fig. 37b indicate. In Fig. 37b, however, one notices that the current that

can be collected at the contactor is still limited by the contactor anode

current which is being held at a specified value (_330 mA). Switching the point

of coupling power supply connection from the contactor cathode to the contac-

-68-
+1,5 V

D

CONTACTOR

mA

Sl MULATOR

O. GROUNDED SIMULATOR ANODE

+11V

F__]

CONTACTOR

,iB---------

330 m A

--I

vT!_o mA +121 t

550 mA

,,,0

___TAN K

b. GROUNDED SIMULATOR CATHODE

SIMULATOR

-1

Fig. 37

16 V

CONTACTOR

300 mA_-_

+e3vrl 4°°mA

4_00 mA

380 mA

SIMULATOR

b

-I0 V

]

-20 mA $8_

I TANK

@

IDEAL CONNECTION CONFIGURATION

Effect of Simulator-to-Tank Connection

Point on Electron Collectlon Current

Balance (Condition A)

-69-
tor anode alleviates this situation when the contactor is operating in the

electron collection mode. By doing this the contactor operating current (i.e.

in the loop between the contaetor cathode and anode) can be controlled

independent of the current collected from the simulated space plasma at the

contactor anode. A typical current balance diagram showing the currents and

voltages measured with this preferred connection arrangement in presented in

Fig. 37c.

When the contactor is being operated in the electron emission mode (condition E in Fig. 29) and the simulator anode is grounded the current balance

diagram shown in Fig. 38a is observed. In this case, the system handles any

current demanded in the electron emission process while holding the anode

current at the contactor and cathode current at the simulator constant. Switching

the point of vacuum tank connection from the simulator anode to the simulator

cathode results in the electron current balance shown in Fig. 38b. In this case

controlling the contactor anode current enhances the electron collection current

at the simulator anode to a small extent. However, the bulk of the electron

current is being collected on the tank in this mode of operation and as a

result having the tank connected to the simulator cathode is not particularly

detrimental. In general, however, the results presented in Figs. 36 to 38

indicate the logical fact that electron emission occurs from a cathode and

electron collection occurs at an anode. Hence one can conclude that a contactor

operating in the electron emission mode should be connected with its cathode

to the coupling supply and it would be preferable in this case for the simulator anode to be connected to the tank. In the electron collection mode on the

other hand, the coupling supply should be connected to the contactor anode and

the tank should be connected to the simulator cathode.

Characterizing Plasma Contactor Performance

Throughout this report contactor potential has been used as the

-70-
+lOV

E]

-63

CONTACTOR

300 rnA

I

, ooyT

[

I1000 mA

l

I000 mA

]

75 mA

925 mA

a. GROUNDED SI MULATOR ANODE

TANK

SIMULATOR

"_Om A[_O V

J 375 mA

A I= ,=r

Fig.

+9.7 Vl

CONTACTOR

E]

300 mA

-62V

[

lO00 mA

]

lO00mA

9?.0 mA _L TANK

SIMULAT( )R

I

t

b. GROUNDED SIMULATOR CATHODE

38 Effect of Simulator-to-Tank Connection

Point on Electron Emission Current

Balance (Condition E)

/_'20 mA

[

J_oooy,

-17V

]

80mA )

80

mA

-71-
independent variable to characterize contactor performance on plots of electron emission current. Contactor potential has been defined as the potential

difference between the anode of the contactcr and the anode of the simulator

which has also been connected to vacuum tank ground. This potential difference actually represents the potential difference across two contactors,

namely the contactor under test and the simulator. It had been hoped that the

potential drop associated with the simulator might be small and that the

plasma it produced would have a potential relatively near simulator anode

potential and it was for this reason that the tank connection was initially

made to the simulator anode.

Plasma potential measurements like those shown in Fig. 13, 14, 30 and 31

have implied that a substantial potential drop exists at the simulator and that

contactor potential is not a particularly good indicator of contactor performance. By using the emlsslve probe it is possible to measure the potential

difference between the contactor anode and space plasma potential beyond the

sheath region (i.e. in the constant potential regions of Figs. 13 and 30 at a

locatlon about 1 m downstream of of the sheath). When one does this, a curve

llke the one shown in Fig. 39 is obtained. This particular curve corresponds

exactly to the one represented by the circular symbols in Fig. 29 so comparison of these curves shows the difference obtained when one uses the contactor

anode-to-plasma potential difference rather than the contactor potential in

plotting a typical contactor performance curve.

Because the data in Fig. 39 were collected using the electrical circuit

configuration of Fig. 37b the electron current that could be collected

from the simulator-generated plasma was limited to 300 mA because the

contactor anode current was limited to 300 mA. When the contactor anode

current was increased to 0.9 A, the data of Fig. 40 were obtained. In

-72-
Flg. 39

-50

I

ELECTRON

EMISSION CURRENT

) -I000

-750

-500

-250

-250

-500 -

Contactor Anode-to-Ambient Plasma

Potential DiffePenoe vs. Emission

CurPent-Low Anode Current Case

JCA =0.5 A

fn =2.7 sccm (Xe)

Pt =3 x I0 -s Torr

12 cm ANODE DIA

VCA=I4 TO 23V

CONTACTOR ANODE- TOPLASMA POTENTIAL DIFF (V)

I00 150

I I

_..I-- TRANSITION TO

IGNITED MODE

-73-
Fig. 40

-5O

I

ELECTRON

I000

)

750

-500

-250

-250

-500

-750

JCA =0.9 A

VCA =4 TO 14V

rh = 2.7 sccm (Xe)

Pt = 3.5 x I0 -s Torr

12 cm DIA ANODE

-I000

Contactor Anode-to Ambient Plasma

Potential Difference vs. Emission

Current-High Anode Current Case

CONTACTOR ANODE TO

PLASMA POTENTIAL DIFF. (V)

50 I00 150

I I I

TRANSITION TO

I IGNITED MODE

t

-74-
this c:se it is apparent that the contactor electron collection current can be

increa ed until the higher contactor anode current being maintained (0.9A) is

reache_. It is also apparent that the curve takes on a shape very much like

the ideal contactor curve, i.e. one that is essentially vertical. Finally it

is noted that the data of Figs. 39 and 40 were obtained with the simulator and

contac_or connected in the manner as shown in Fig. 37b. This facilitated

electr< n emission from the simulator at any level demanded while the simulator

discha: ge current was being maintained at 300 mA, however, it was necessary to

increae the contactor anode current to 0.9 A in order to achieve the 0.9 A

electr n collection current. This suggests that it is probably desirable to

modify the system circuitry so that either the contactor anode or cathode can

be con acted to the coupling supply and so the connection can be switched during co factor operation. Future plans call for the simulator cathode to be

connec _,d to the tank and the contactor anode connected to the coupling power

supply when electrons are being collected by the contactor. When electrons

are be cg emitted from the contactor the contactor cathode will be connected

to the coupling supply and the simulator anode will be connected to tank

ground.

Electr, n Current Distribution on the Anode

The variable area anode is instrumented so the current from each segment

of anod_ can be monitored. Hence the current distribution to a 12 cm dia

anode can be determined using the apparatus of Fig. 4b with the contactor and

simulator connected as shown in Fig. 37b. When this is done at the contactor

operating conditions that produced the data of Fig. 40, the current

distrihutlon shown in Fig. 41 is measured, in Fig. 41 the vertical distance

between the horizontal axis and the solid llne containing circular data sym-

-75-
<>

I

I

I

I

I

I

I

I

I

I

I

I

I

6

I

I

I

.LN3_tno 3C]ONV 40 NOI.LOV_I-I

0

L

ul

e.

e.

0

G

,,,-t

-76-
bols represents current drawn to the sustalnlng anode while the vertical

separation between the diamond and triangular data symbols represents for

example, the current to the outermost anode segment (J3- see Fig. 7). All

currents are normalized using the total contactor anode current and the data are

plotted as a function of this total electron current being collected at the anode.

It should be recalled in vlewlr_ these data that the contactor anode current

is limited to 0.9 A, so this is the maximum current that can be collected.

The data of Fig. 41 suggest that the sustaining anode and anode 1 (see

Fig. 7) collect the bulk of this current when the electron current being

collected from the simulated space plasma is small. As the electron collection

current is increased, however, less current flows to these anodes and the bulk

of the current begins to flow to the anode with the largest surface area

(anode 3). This result is particularly interesting when viewed in the light

of the equlpotentlal contour map shown in Fig. 32. These data suggest that

electrons are drawn radially in toward the cathode and this in turn suggests

that they might be drawn preferentially toward anodes $ and 1. The fact that

the data of Fig. 41 show that the opposite is true indicates that electrons

being drawn in through the sheath surroundlng the contactor experience substantlal scattering probably in the high density plume region. As a result

they are able to change their direction from radial traJectorles to random

ones so they can be collected on the anodes in what appears to be direct proportion to the anode surface areas.

Based on the results collected during this grant period it is concluded

that changes in magnetic field in the range from near zero to 0.5 gauss do not

induce substantial changes in the characteristic curves that describe plasma

-77-
contactors performance. Increasing the contactor anode area facilitates transition into the ignited mode of electron collection. This mode is characterized by suDstantlally lower voltage drops than those observed for contactors

collecting electrons at a given current in the unignited mode. Double sheaths

appear to develop in both the electron emission and collection modes of operation but they are generally difficult to probe at high currents in the electron emission mode because sheath boundaries are located close to the contactor. In the electron collection mode a structure develops which includes a

high density constant potential plasma region immediately adjacent to the confactor, a double sheath region and a low density plasma region from which electrons are collected. A simple model of the double sheath region yields radius

ratios associated with the inner and outer boundaries of the region that agree

with measured radius ratios to withln about 25_.

At the high electron collection currents associated with the ignited mode

of operation it is preferable to connect the simulator cathode to the vacuum

tank and the contactor anode to the coupling power supply that is used to bias

the contactor relative to the simulator in order to induce current flow. When

the contactor is operating in the electron emission mode its cathode should be

connected to the tank. Hence it is desirable to switch the connection points as

operation is switched from the emission to the collection mode of operation.

It is believed that it will also be desirable to operate contactors in space

so the one emitting electrons has its cathode connected to the tether and the

one collecting electrons has its anode connected to the tether.

-78-
HIGH PRESSURE HOLLOW CATHODE RESEARCH

Craig Luebben

Arcjet thrusters are attractive because their specific impulse may be close

to optimum for many near-earth missions. These thrusters typically use a solid

thoriated tungsten cathode that is prone to overheat during operation and melt

at the small spot on its tip where electron emission occurs. Thus, these

cathodes may not be suitable for long term space applications. The oriflced

hollow cathode used in ion thrusters emits electrons from a relatively large

surface area on its cylindrical insert so it is not as prone to overheating.

But hollow cathodes typically operate at lnterelectrode pressures (i.e.

between the anode and cathode exterior) of 10 -S Torr, whereas arcJet thrusters

operate at pressures of 1000 Torr or greater. The major objective of this research effort has been to develop a hollow cathode that operates satisfactorily

on ammonia at high pressures and to identify the factors that facilitate this

high pressure operation. Ammonia has been selected for the study because it

is an attractive propellant for many arcJet propulsion missions. The approach

used in the study has been to increase the operating pressure range gradually

from the low values where hollow cathodes operate satisfactorily and their characteristics are known and understood to the high pressures where arcJets operate. When conditions limiting the extent to which pressure could be increased

have been encountered, the reason for the limitation has been investigated and

the test apparatus and/or procedure has been modified in an effort to overcome

the limitation and understand the physical reason for its occurrence.

Aooaratus and Procedure

A conceptual diagram illustrating the features of the apparatus used in

-79-
this study is shown in Fig. 42. This apparatus is located within a vacuum

bell jar in such a way that the interelectrode pressure can be controlled over

a pressure range from-10 -2 to 103 Torr. The principal elements of this test

apparatus are the hollow cathode and the anode. The hollow cathode, which had

been fabricated initially in the conventional way using a tantalum tube body

electron beam welded to a thorlated tungsten orifice plate was replaced by

cathodes fabricated from insulating materials. This allowed the insert, which

is located within the cathode body, to be electrically isolated from the

cathode body and the orifice plate. Various orifice plate materials were

tested, including boron nitrlde, tantalum and graphite. The typical insert

was a cylinder with a diameter of a few millimeters made by rolling several

layers of 0.013 mm thick tantalum foil treated with the low work function

s

Chemical R-SO0 . This type of insert has been used because a previous

study 9 showed that this insert, which is easy to fabricate, performs as well

in short term tests as other more costly inserts. The anode used in the

tests had a diameter that ranged from 2 cm to 5 cm and was made of tantalum

when it was radiation cooled and of copper when it was water cooled.

All tests were conducted using ammonia as the working fluid (expellant)

and the valving arrangement shown in Fig. 42 allowed this fluid to be

directed in two different flow patterns. Normal flow, which is used in ion

thruster hollow cathodes, involves flow through the hollow cathode from the

cathode interior region to the interelectrode region and then out of the

system through the vacuum pumps. The normal flow configuration is achieved

by positioning the valves shown in Fig. 42 in the manner designated in the

first row of Table I.

e

Chemical R-500 is a double carbonate (BaCO 3. SrCO 3) mixture that has been manufactured by the J.R. Baker Chemlcal Co., Philllpsburg, N.J. but is no longer made

-80-
1.1.1,1.1

O_

)-0

0

..I

W_

WW

I-'11 _

m

W f

Wn_

0

"T't_

O_

_/

o

o

4-)

0

r._

-$1-
Table I

Expellant Flow Configurations

Pattern

Designation

Normal

Reverse

Valve 1

Closed

Open

Valve Positions

Valve 2

Open

Closed

Valve 3

Closed

Open

During preliminary testing it was determined that high interior cathode

pressures resulted in very small emission spots on the insert which eventually

resulted in insert melting and failure. 9 Hence it was determined that the

cathode interior pressure should be held low relative to the high interelectrode pressures and the reverse flow pattern was found to be preferable to the

normal one. With reverse flow the expellant was fed through the interelectrode region into the cathode interior region through the cathode orifice and

then out through the pumping system. This flow pattern was achieved by posltlonlng the valves as shown on row 2 of Table I. It should be noted that

while a vacuum pump had to be connected to the cathode interior region to

achieve reverse flow in the laboratory, this could be achieved in space by

simply venting the cathode interior to space. The normal flow configuration

was generally used, at low Interelectrode pressures, to initiate the

discharge. Once the discharge had been started and stabilized the flow valves

could be repositioned to permit testing in the reverse flow configuration.

The pressure gauges shown in Fig. 42 were provided to measure the cathode

interior and interelectrode pressures during operation.

-82-
The schematic of Fig. 42 shows a horizontal orientation for the apparatus

and this orientation was indeed used during the early test period. At high

interelectrode pressures, however, heated expellant appeared to rise in this

apparatus thereby enhancing the probability of current flow from the cathode

orifice to the upper anode surface and possibly perturbing the test results.

In order to minimize these effects the apparatus of Fig. 4S was built. This

apparatus has two features that help to prevent the expellant buoyancy problem

Just mentioned, namely a vertical orientation of the cathode axis and vortex

inducing tubes suggested by Hardy I0 to facilitate the formation of a swirling

flow in the interelectrode region. The apparatus in Fig. 4S, however,

includes the same basic features identified in the discussion of Fig. 42. The

vortex inducing tubes shown in Fig. 4S are located approximately 2 mm above

the face of the orifice plate, and the ammonia injection ports into the

interelectrode region have a diameter of 0.1 mm and an inclination of 45 °

relative to the tangent to the anode. The anode itself is a 2 cm dla copper

tube which has copper cooling coils silver soldered to its exterior surface.

The water, which circulates through these coils, is supplied through polyflo

tubing which serves to isolate the anode electrically from the water source.

This circulating water cools the cathode indirectly by conducting heat from it

to the anode through the boron nltride cathode body. The entire apparatus

shown in Fig. 4S is enclosed within a vacuum bell Jar. Flow through the vortex inducing tubes therefore increases the Interelectrode region pressure and

escapes through the cathode orifice and through anode structure Joints into

the bell Jar.

The apparatus suggested in Figs. 42 and 4S is designed to have a great

deal of flexibility so changes, introduced to overcome the problems encountered during operation at ever increasing interelectrode pressures, can be

made. A summary of the hollow cathode design and operational parameters that

-8S-
INTERELECTRODE

REGION

\

cOPPER ANODE _._"_

COOLING _1 l

TUBES -

......... p'5

i/

1"5

""I_l INTERELECTRODE

_H PRESSURE /

j

, ANODE

o VORTEX INDUCING

Io

!o TUBES

PLAN CROSS SECTION

OF ANODE

TANTALUM ORiFICEFOIL INSERT _'_:i!"_ i'_" i_!i o INTERELECTRODEFLoW

, CATHODE POWER LEAD BORON :.:....,. ....:.,-...,.....;; ......:....,..:,:.:_

NITRIDE .........._:_}"'.:.,._..-.-, :':::. ........... .•..................

CATHODE ! _C •BODY ATHODE INTERIOR REGION

_ INTERIOR FLOW PORT

TICKLER --.- J

ELECTRODE

_x

CATHODE INTERIOR PRESSURE PORT

CATHODE SUPPORT STRUCTURE

FIE. 43 Vertlea11F Oriented Test Apparatus

-84-
were varied in the study is presented in Table II. Many of these parameters

are self-explanatory, however, some explanation of the cathode configurations

associated with the various orifice plates is needed• Boron nltrlde orifice

plates were made by simply machining round stock to produce the cathode body

and orifice configuration suggested by Fig. 43.

Table II

Parameters Varied in Hollow Cathode Study

Cathode tube material

Cathode orifice diameter

Cathode orifice region material

Orifice region configuration

Anode/cathode cooling

Expellant flow

Discharge current

Cathode/anode axis orientationanode dia.

Chan_es introduced

• boron nitrlde

• tantalum w/quartz insulator

• open-ended (5.0 ram)

• 0•25 to 1.8 mm

• boron nitride

• tungsten

graphite

• plate, orifice on centerline

• cylindrical cap, orlflce on center

• seml-conlcal cap, orifice on center

• semi-conical cap, offset orifice

. radiation cooled

• water-cooled

• normal flow

• reverse flow

• 4 A to 14 A

• horizontal-5 cm dia.

• vertical (with vortex flow

stabilization)-2 cm dia.

When this orifice plate was replaced by a tungsten on_ a cavity was machined

in the boron nitrlde body and the tungsten orifice plate was cemented in place

using ceramic adhesive assuggested in Fig. 44. In this apparatus the insert

-85-
TUNGSTEN

ORIFICE

PLATE

BORON NITRIDE

HOLLOW CATHODE

BODY

i_ii__ .-_::_.._.:-.__":_!

CATHODE

POWER

LEAD

ANODE

E BORON NITRIDE

. __CE_AMIC ADHESIVE

TUNGSTEN

"_ _ORIFICE PLATE

Fig. 44 Boron Nitride Body/Tungsten Orifice Plate

Hollow Cathode

-86-
was displaced upstream of the orifice plate by a few millimeters so it would

remain isolated electrlcally from the plate. Graphite orifice caps with various configurations were also used. They were installed in the manner

suggested in Fig. 45 so the insert and the graphite orifice cap remained isolated from each other. The graphite caps were relatively easy to machine so a

variety of them having the shapes and orifice hole positions indicated in Fig.

45b were investigated.

Experiments conducted during the previous grant perlod 9 had shown the

necessity of keeping the insert Isolated from the orifice plate to prevent

emission from occurring at the orifice plate at high interelectrode pressures.

Installatlon of the tantalum cathode body configuration of Fig. 46 therefore

necessitated provision for insulation between the insert and the tantalum

tuOe. This was accompllshed by InstalllnE a quartz tube between them as suggested in FiE. 46. It was hoped that the quartz tube would facilitate radiation from the hot insert to the cathode body, so thermal conduction from the

insert to the quartz tube would be small, anode water-coollr_ coils would

remove heat through the boron r_trlde cathode body at a sufficiently rapid

rate and the quartz tube would not melt.

In order to conduct a test, ammonia was fed in the normal flow configuration at a sufficiently high rate to produce a cathode interior pressure of

approximately i0 Torr and the anode was biased to about 400 V. Frequently

this condition was sufficient to result in electrlcal breakdown initiating a

discharge between the cathode and anode. However, on some occasions breakdowns occurred at iocatlons away from the anode and cathode and the tickler

electrode shown in Fig. 43 was therefore installed to facilltate startup. In

order to use this electrode the anode was kept inltlaily at a potential near

-87-
_DE

pRIFICE

--GRAPHITE ORIFICE CAP

CERAMIC ADHESIVE

BORON NITRIDE CATHODE BODY

R-500 TREATED Ta

FOIL INSERT

(:1. GRAPHITE ORIFICE CAP CONFIGURATION (WITH

CYLINDRICAL CAP HAVING ORIFICE ON CENTER)

_, __SEMI-CONICAL

_/1 f CAP WITH

_OFFSET ORIFICE

SEMI-CONICAL

CAP WITH

CENTERED

ORIFICE

b. ALTERNATIVE CAPS

Fig. 45 Graphite Orifice Cap Hollow Cathode

-88-
QUARTZ TUBE

INSULATOR

TUNGSTEN ORIFICE

PLATE

TANTALUM CATHODE BODY

IC ADHESIVE

TANTALUM

FOIL INSERT BORON NITRIDE

CATHODE BODY

POWER

I'ANTALUM INSERT LEAD

_ BORON NITRIDE

_RAMIC ADHESIVE _i___'_ TU NGST EN

_ORll FI CE

"_i_i _ PLATE

Fig. 46 Hollow Cathode wltb Quartz Tube

Insulation

-89-
that of the cathode insert and a potential of about 500 V was applied between

the tickler and the insert. This typically caused an electron current of

approximately 4 to 5 mA to flow. The plasma produced by this low current

discharge generally supplied sufflclent electrons so that biasing the main

anode to a few hundred volts positive of the insert resulted in a discharge

between the anode and the cathode insert. Once the main discharge had established itself in the current range between 4 and 14 A the tickler discharge

was shut down. After the discharge had stabilized for several minutes, Valve 3

(Fig. 42) was opened to allow expellant to flow through the vortex inducing

tuDes. Valve 1 was then opened and Valve 2 was closed to facilitate transfer

to the reverse flow configuration. Once operation had stabilized in the

reverse flow configuration the ammonia flow into the Interelectrode region

could be increased to increase the interelectrode pressure. Typically the

test then proceeded as the interelectrode pressure was increased in steps

while the discharge current was held constant and anode voltage and Interelectrode and cathode interior pressures were recorded. While At was also possible to throttle the flow through Valve 1 early tests showed that the cathode

interior pressure should be kept as low as possible to prevent damage to the

insert. As a result Valve 1 was generally kept at its full open position.

The insert lead shown connecting the insert to the power supply in

Figs. 44 through 46 is an 0.8 mm thick by 4 mm wide tantalum ribbon.

This lead has a sufficiently large cross-sectional area so the voltage

drop along its length remains small at the maximum current levels expected in the experiment. As a result it does not overheat.

-90-
Results

Previously reported data have shown some important characteristics of

conventional hollow cathode discharges that develop as interelectrode pressure

9

is increased. Of primary importance is the fact that interior cathode pressure

increases and as a result the area of the electron emission spot on the insert

decreases with increasing interelectrode pressure. Not only has this been

observed experimentally on a number of occasions, but it is consistent with

the model developed by Siegfrled. 11 Such decreases in spot size can lead to

insert failure due to localized overheating. It has also been observed that

increases in interelectrode pressure cause a discharge localization between

adjacent points on the cathode and anode which eventually develops into a very

fine spoke or filament of current between these points. This is an undesirable current structure for arcJet appl_catlons because it would result in

relatively poor heating of the arcJet propellant. The occurrence of these

phenomena result in two effects illustrated by the data of Fig. 47 and the

associated visual observation of the discharge. First, the anode voltage

tents to increase as interelectrode pressure is increased. Second, the site

of electron emission tends to occur from a small spot on the cathode nearest

the anode, even if there is a lower work function site on the cathode somewhat

further from the anode. In the case of the data of Fig. 47, this emission

spot migration occurred near Interelectrode pressures of 10 Torr.

Typical data showing the voltage drop distribution between the cathode

and the anode are presented in Fig. 48. These data were obtained using an

emisslve probe like the one described in Ref. 9 together with a hollow cathode

having a boron nitride body that was open ended (no orifice plate) and a 5 cm dia

tantalum anode. These data are relatively typical and have been seen with many

-91-
_g

m

=o

_:_ _m \_z

_,_z_ \_z_

E P hi I, "r

Z_ I" ,_ O- . 0

-- X _ (/)

uO I- m 0::

0 o" uO Z Z '_"

L l ,I, I

0 0 0 0

0 Io 0

oo

o

S

z_

oo

(A) 39VJ.-IOA30ONV

-92-
175

150

125

>

"" I00

ANODE VOLTAGE

-ENDED BORON NITRIDE HOLLOW CATHODE

R-500 TREATED TANTALUM FOIL INSERT

5 ¢m DIA TANTALUM ANODE

NH 3 E XPELLANT

NORMAL FLOW

4A DISCHARGE CURRENT

13

z / ANODE

75 F/_LL / _" NEAR

0 | / ANODE

50 _ SURFACE

I / BULK PLASMA

I / VOLTAGE DROP

25-I o o _f CATHODE, . o EX,T

0 "'

0 CATHODE 0

FALL

I

I I I I I I I

0 5 I0 15 20 25 30

INTERELECTRODE PRESSURE (Torr)

Fig. 48 Typical Plasma Potential Profile

-93-
different anode and cathode configurations. They show that the cathode fall,

i.e. the voltage drop between the cathode and the point in the plasma immediately adjacent to it, remains relatively constant as interelectrode pressure

increases. The voltage drop through the voltage plasma and at the anode, however, increases with interelectrode pressure. It should be noted that it is

preferable for the bulk plasma voltage drop to dominate the total voltage drop

because it is the bulk plasma voltage drop that determines the power golng

into propellant heating in an arc Jet application. The cathode and anode fall

voltages on the other hand are proportional to the power going into heating

the cathode and anode.

The capacity of a hollow cathode to sustain lower cathode interior pressures at prescribed interelectrode pressures by using reverse rather than normal flow is demonstrated by the data of Fig. 49. For example, at an

interelectrode pressure of 50 Torr, a 4 A discharge current and a 0.51 mm dla

orifice the data of Fig. 49 show the interior pressure in the reverse flow

case is 13 Tort compared with SO Torr for normal flow. These date were obtained using a cathode body and orifice plate made of boron nitride. In both the

normal and reverse flow situations, operation had to be terminated because the

orifice enlarged due to what appeared to be melting. In fact, the nonlinearity of the reverse flow data above about 60 Torr suggests that in the

reverse flow case, the orifice may have begun to enlarge at an interelectrode

pressure near 60 Torr. Had orifice enlargement not occurred it is believed

tne cathode interior pressure would have only been about 20 Torr at an

interelectrode pressure or 90 Torr rather than 30 Torr as the data show.

Above 90 Torr in the reverse flow case dramatic enlargement of the boron

nltrlde cathode orifice occurred, the cathode interior pressure increased

rapidly and the test had to be stopped. Note from the data of Fig. 49,

-94-
I I I I

0 0 0 0 0 0 0

0

0

(J,o.I.) 3_11"1SS3_1d _IOI_I3.LNI 3(30HJ.Y3

t_

-95-
however, that higher interelectrode pressures were realized before orifice

failure with reverse flow than with normal flow (90 Torr for reverse flow

versus 65 Torr for normal flow).

An objective of the study was to achieve operation on ammonia at

interelectrode pressure approaching 1,000 Torr. Both analytical and experimental results obtained suggest that the cathode interior pressure should

be around 10 Torr or less to sustain a diffuse emission from the cathode

insert. In order to estimate of the orifice size required to achieve this,

the data of Fig. 50 were collected while ammonia was flowing through the

cathode orifice in the reverse direction but no discharge was occurring.

While it was recognized that the actual relationship between interior and

interelectrode pressures would be a function of discharge current, these data

did provide an indication that the orifice diameter would have to be less than

-0.5 mm to realize a sufficiently low cathode interior pressures to assure

diffuse emission from the insert. These data were also used to estimate

whether a cathode orifice was enlarging during a test by simply comparing the

interior and Interelectrode pressures; thls was generally done by turning off

the discharge and making a comparison. Tests conducted with various orifice

sizes at a 4 A discharge current suggested that a change in cathode orifice

diameter affected the cathode interior pressure but did not influence its

interelectrode pressure vs. anode voltage characteristic. This result, shown

by the data of Fig. 51, indicates essentially the same anode voltage versus

interelectrode pressure plot for orifice diameters ranging from 0.5 to 1.8 ram.

For all of the tests conducted with the boron nltride cathode orifice plates,

the orifice plate tended to fall because of orlfice enlargement due to what

appeared to be melting. In one attempt to overcome this problem, copper cooling coils were cemented to the boron nltrlde but they did not facilitate

-96-
0

0

0

0

0

co

>

.-r- hi 0

Z E Z

0

0

o

(JJOl) 3_lnss3_ld _IOI_I31NI 3OOHIVO

0

0

0

,.,,-4

c_ 0

o

0

0 0

0 "_

o

0

-97-
hi

<] bJ I,I I,I

(J (J (J

(A) 39V170A 3(]ONV

o

o

o=

_:_

o_ |

o

0 r_

-98-
operation without melting at higher pressures.

The data presented in Figs. 48 through 51 were all collected using the

horizontal orientation of the cathode axis and a cathode body and orifice

plate made of HBR grade boron nltrlde. It was postulated that the experiments

might be extended to higher pressures if the melting temperature of the boron

nitrlde orifice plate could be increased and if flow changes could be introduced to encourage an electron current path along the centerllne of the

cathode at the orifice. In order to accomplish this, the vertically oriented

tes_ apparatus utilizing vortex stabilized flow, shown in Fig. 43, was used.

In addition, the boron nitrlde grade was improved from the 1200°C melting

point HBR grade to the 2000°C melting temperature HBC grade. These changes

were Insufflclent to prevent boron nltrlde orifice plate melting, although

they did facilitate some increase in the interelectrode pressure that could be

sustained before orifice plate failure occurred.

Using the apparatus of FiE. 4S it was possible to demonstrate that

increases in discharge current facilitated a larger pressure drop across the

cathode orifice in the reverse flow conflguratlon. Typical data demonstrating

this are presented in Fig. $2 where the changes in interelectrode pressure

are plotted as a function of discharge current for a cathode with a relatively large (0.94 mm dia) orifice in which expellant flows were controlled so

cathode interior pressure remained constant. The data of Fig. $2 were actually obtained by increasing the discharge current while a fixed expellant flow

rate through the vortex tubes and a fixed pumping environment on the cathode

interior were maintained. Anode voltage decreased slightly as current was

increased through the range indicated. Increasing the expellant flow induced

increases in both inter-electrode and cathode interior pressures and produced the

-99-
t_

O

I--

I.iJ

0::

U)

I.=J

n,-

O.

7O

6O

5O

4O

3O

2O

I0

0

CATHODE INTERIOR PRESSURE-_

O ,0 _ 0

HBC GRADE BORON NITRIDE HOLLOW CATHODE AND ORIFICE

0.94 mm DIA ORIFICE PLATE

R-500 TREATED To FOIL INSERT

20 mm DIA WATER-COOLED ANODE

NH 3 EXPELLANT

REVERSE FLOW

I I I I I I 1 I I I I I I I

) 2 4 6 8 I0 12 14

DISCHARGE CURRENT (A)

Flg. 52 Effect of Discharge Current on Operating

Pressures at Constant Reverse Flow

-100-
voltage versus inter-electrode pressure curves shown in Fig. 53. In cases reported

in Figs. 52 and 53 the higher current results in a lower anode voltage due to a

9

negative resistance phenomenon that has been recognized and described previously.

This occurs because plasma resistivity is proportional to the ratio of electron

collision frequency to electron density and increasing the discharge current

causes the electron density to increase more rapidly than the collision frequency.

As the data of Fig. S3 suggest the discharge extinguished repeatedly at a 50 Tort

interelectrode pressure but at pressures near 220 Tort it began to fail as a result

of boron nitride melting which led to orifice erosion followed by rapid degradation

of the insert. This degradation occurred when the orifice had enlarged to the

point where cathode interior pressures were sufficiently large so the emission spot

on the insert was small and localized overheating occurred.

Because the objective of the research effort was to demonstrate hollow cathode

operation at high interelectrode pressures repeated attempts were made to operate

at these pressures using a variety of different cathode designs. On one occasion

for example operation of an HBC grade boron nltride cathode with a 1.2 mm dia

orifice in the normal flow configuration was extended to an interelectrode pressure

of 600 Torr. Between pressures of I and 300 Torr the discharge appeared to be

stable, and at pressures above 300 Torr, the discharge was somewhat unstable and

the anode voltage fluctuated with a 20 V amplitude and a fraction of a second

period. The mean anode voltage required to sustain the discharge as shown in Fig.

54 for this test was relatively constant in the range of 100 to 130 V at pressures

from I00 to 600 Torr. These voltages were lower than the 200 V values reached at

pressures near 200 Torr is previous tests (see Fig. 53 for example) presumably

because of the normal flow environment and this apparently facilitated operation to

the higher interelectrode pressures. With the normal flow, however, the interior

cathode pressure became very high and this caused severe melting and erosion of the

boron nitride orifice plate and rapid destruction of the insert. In fact post-

-101-
I I I I

0 0 0 0

0 _0 0

(A) 3OVI-IOA 3(]ONV

0

oo

o

=i

@

!

i °°

o_ _

e_ e_

Z _'_

e_ e_

0 o.

_ °

0 _

_ 0

m

0 _

-102-
I I I I I

0 0 0 0 0

_ _ 0

(A) 39V±_OA 300NV

8

o_

o

o_

_5

o_:

z

m

0

_..

0

o

a.a

o]

.e.l

[..

o

0] _j

0_

_°

-103-
operation inspection of the cathode interior showed the insert had completely

disintegrated. A boron nitrlde-tantalum mass from which electron emission was

apparently occurring appeared to remain in its place. The same test was attempted

using reverse flow, but the boron nltride near the orifice melted repeatedly making

it impossible to maintain low interior cathode pressures. When high interior

cathode pressures do develop because of normal flow or orifice enlargement when

reverse flow is being used, it appears the insert emission region is a small

overheated spot and typically emission cannot be sustained from such a spot. If an

insert material like the boron nltrlde/tantalum mass that was produced during the

collection of the data of Fig. 54 could be made to operate for long periods of

time, the data of Fig. 54 suggest it might be possible to operate at high

interelectrode pressures in the normal flow mode.

Because it was considered more likely that successful hollow cathode operation

at high interelectrode pressures could be achieved if the cathode interior pressure

could be kept sufficiently low so the electron emission reglon on the insert could

be kept larEe, no search for an improved insert material was started. Bather the

focus of the research was directed at maintaining the integrity of the cathode

orifice so low cathode interior pressures could be sustained in the reverse flow

mode of operation.

Because boron nitrlde orifice plates melted under the high temperatures

encountered during operation, it was decided that a tungsten orifice plate

should De substituted for the boron r_Itrlde one. When this cathode was

operated in the reverse flow conflgura_ion, it was possible to extend operation up to interelectrode pressures of 1,000 Torr. Cathode interior pressures

measured as a function of interelectrode pressure for these tests are plotted

in Fig. $5. This figure shows that the interior cathode pressure remained

below i0 Torr over the entire interelectrode pressure range. The

-104-
0

0

0

1

0

-0

0

0

0

cO

I I I I I I I I I I

0 _ co _. _ _ _ _o _ -- 0

(J.Jol) 3_lrlSs3_d 3GOH.LVO _01_131NI

"c

o

• S

e.

0

•_ 0

m c.

o

O,.-i

I_ l-i

-105-
interior cathode pressures observed with the 1 mm dla orifice used in the test

are much lower than expected. In addition the data show quite a bit of

scatter. After the test had been completed and the cathode disassembled, it

was observed that a small object, presumably boron nitride, was partially

blocking the orifice in the tungsten plate and apparently causing of both the

low interior cathode pressures and the data scatter. The anode voltage during

this same test is shown plotted as a function of interelectrode pressure in

Fig. 56. During the tests that the discharge operated stably at both the 12 A

discharge current pertaining to Figs. 55 and 56 and at lower current levels.

The discharge could not be observed visually during the tests, but it is

believed based on subsequent tests where the discharge could be seen that

the current was flowing in the form of a thin moving filament from the

orifice radially outward toward the anode. After about 30 minutes of testing

this current filament apparently caused the ceramic adhesive holding the

tungsten orifice plate in place to loosen and erode" This eventually resulted

in the opening of a flow passage between the Interelectrode and cathode interior regions which allowed the cathode interior pressure to increase and the

insert to degrade rapidly.

In an effort to overcome the tendency of the discharge to melt a hole

between the tungsten orifice plate and the boron nltrlde body a configuration

which used a tungsten orifice plate electron beam welded to a tantalum cathode

body in the manner shown in Fig. 46 was tested. These attempts to operate the

tantalum body cathode shown in Fig. 46 were completely unsuccessful. Immediately after startup it appeared that the quartz tube, which insulated the

insert from the cathode tube melted. This allowed the insert to come in contact with the cathode tube and as a result the discharge apparently

transferred from inside the cathode to a point on the edge of the tungsten

orifice plate. An attempt to use a boron nltride insulator in place of the

-106-
0 0

0

0

0

0

0 0

0

0

(A) 39Vi'lOA 300NV

0

0

m

0

-0

6o

¢;

t.

L

0

0

f..,

c_

I--I

d

0

o_

-107-
quartz tube insulator shown in Fig. 46 was also unsuccessful.

It was felt that the difficulties encountered with the tungsten orifice

plate electrode beam welded to the tantalum cathode body could be overcome if

the diameter of the assembly could be increased beyond the 0.7 cm value that

was belr_ used. Larger tantalum and tungsten assemblies, however, could not

be located and so it was decided to attempt to make relatively large diameter

(-1.5 cm) graphite caps. Three graphite orifice cap configurations were

tested in conjunction with the boron nltrlde cathode body; anode voltage

versus interelectrode pressure characteristics li_e those shown in Fig. 57

were observed. These characteristics are very similar to those observed with

the tungsten orifice plate. _hen the cylindrical orifice cap having the centered orifice was removed and examined after the tests, hemispherical pits

were observed at the outside edge of the cap near the anode. Their appearance

suggested the discharge may have been coming from the cap surface rather than

through the orifice. To prevent this, the seml-conlcal cap configuration was

tested but again the discharge pitting suggestive of emission sites on the

outside surface of the cap were apparent near the outer edge of the cap. Even

when the orifice was offset and data represented by the squares in Fig. 57

were obtained, this pitting effect was still observed. It should be noted

that in all cases described the insert was isolated from the orifice cap by

the boron nitrlde so if emission was occurring from the outside surfaces of

the orifice caps the discharge current had to be flowing from the insert to

the inside of the orifice cap, through the orifice cap and then from the exterior of the orifice cap to the anode.

In all of the tests that utilized boron nitride or graphite at the orifice

boundary, evidence of orifice enlargement was found after the associated tests had

been completed. This enlargement as well as pitting observed on the outside of the

-108-
0 0 0 0 0 0

0

0

(A) 39V.L-IOA 3CIONV

0

0

0

rJ

q;

0 q;

O_

-109-
graphite cap must have been caused by a combination of discharge-induced heating,

sputtering and chemical attack between the carbon or boron nitrlde and ammonia.

After one 30 minute period of operation a new graphite cap was observed to have

been eroded completely through the side of the orifice cap. In this case it

appeared that the discharge had attached itself to the exterior of the cap at a

location near the anode and dlscharge-induced erosion had eventually resulted in

hole completely through the cap wall. Tungsten orifice plates on the other hand

did not show any evidence of orifice enlargement during the conduct of tests that

utilized them. It also appeared that insert degradation frequently accompanied

cathode operation as evidenced by increases in anode voltage at a particular

operating condition. During the tests conducted wltb the graphite and tungsten

orifice regions the discharge could not be observed directly so it is not known

when or if the discharges actually began to pass through the cap wall rather than

through the orifice.

Conclusions

Hollow cathodes can operate on ammonia at interelectrode pressures up to the

1,000 Tort range where arcJets operate, but it is not known at this time if they

operate in the conventional manner, i.e. so the discharge current passes through

the cathode orifice under all conditions. If the cathode orifice plate is made

from an electrical conductor the current may flow from the insert through the

cathode interior plasma and then through the conducting orifice plate to couple to

the interelectrode plasma under some conditions. In view of this, the following

criteria have been identified as necessary but not sufficient conditions for high

pressure hollow cathode operation.

1. The cathode interior pressure should be maintained at a low value (_10 Tort).

-110-
This can be realized by using reverse flow through an orifice with a

sufficiently small diameter so the 10 Tort pressure is sustained within the

cathode interior. The low cathode interior pressure is required to prevent

emission from a tiny spot which eventually results in localized overheating

and damage to typical insert materials.

m The insert must be electrically isolated from the orifice plate. This

condition ensures electron emission spot migration from the insert to the

cathode exterior does not occur at high pressures. Under conditions where

the orifice plate is made of an InsulatlnE materlal, it can be stated

that the discharge is comlr_ from the insert through the orifice. If

the orifice plate is made of an electrically conducting material this

statement cannot be made. In this case the discharge may pass from

the insert through the interior plasma to the interior surface of the

orifice, through the orifice plate, from a second emission site on the

exterior of the orifice plate and finally through the interelectrode

plasma to the anode.

. The orifice plate must be made of a material that can tolerate a high

temperature and is chemically non-reactlve with the ammonia expellant

if long-term operation is to be realized.

For these tests it was necessa_ to water-cool the anode at high

interelectrode pressures to prevent the anode from overheating due to the

large voltage drop across the anode sheath.

-IIi-
VERIFICATIONOF RING CUSP DISCHARGE CHAMBER PERFORMANCE

Jason Vaughn

IDtroduct$on

The ring cusp discharge chamber having a cathode adjacent to the screen

grid has been shown by Hiatt 12 to exhibit optimum performance when the anode

is positioned downstream of the magnetic ring cusp at the optimum "virtual

anode" location. This optimum performance was also determined by the location of the virtual cathode which was in turn determined by the position of

the ring magnet and the size of the filament cathode used to supply electrons

to the discharge. A study intended to provide further insight into the

discharge chamber optimization process was initiated on the same 8 cm ring

Cusp discharge chamber used by Hiatt. The objective of the first phase of

this work was to verify his best performance results. In order to do this it

was necessary to first adjust the discharge chamber geometrical parameters to

the values used by Hiatt and then to measure the discharge chamber performance. Experlmental results obtained from thls study will be analyzed in

terms of the discharge chamber model developed by Brophy $ and performance will

be characterized in terms of the plasma ion energy cost (Sp) and the extracted

ion fraction (fb) parameters which form the basis of this model

Apparatu_ and Procedure

The 8 cm dla ring cusp discharge chamber used in this verification study

is shown schematically in Fig. 58. All of the symbols used to define the

discharge chamber dimensions that can be varied are defined in th%s figure.

The magnetic field associated with this discharge chamber is constructed by

placing one magnet on the centerline at the rear of the chamber and

-112-
-113-
arranging a ring of magnets end-to-end around the circumference of the wall.

The 1.9 cm by 1.3 cm by 0.6 cm samarium cobalt magnets used to construct

the field have a flux density of 0.27 Tesla at their surface. They are

arranged so the polarity of the outward facing surface of the magnets comprising the ring is opposite to that of the single magnet at the rear of the

chamber.

The cathode used in the discharge chamber is a 0.25 mm dia tungsten wire

formed into a loop of diameter 3.1 cm. This loop is located 1.4 cm upstream

of the screen grid. Under the action of the lead screw and yoke the 5.2 cm

dla copper anode could be moved through the region between the axial location

of the ring cusp and the axial location of the cathode loop. The grids had a

cold spacing of 0.86 mm, the screen grid holes had diameters of 1.9 mm, the

accelerator (accel) grid holes generally had diameters of 1.5 mm although a

few near the centerline had enlarged to 1.7 mm as a result of sputter erosion,

and the hole-to-hole spacing in both grids was 2.2 mm. The screen grid

potential was 750 V positive and the accel grid was maintained at -250 V

throughout the study. The argon propellent used to conduct the research was

fed into the discharge chamber through the sidewalls in the manner shown in

Fig. 58. The ball Jar used was 46 cm in dia. and it was diffusion pumped to

an ultimate pressure in the low 10 -6 Tort range and operating pressures that

were typically in the mid to low 10 -5 Torr range. Neutral densities in the

discharge chamber and propellant utilization data associated with its

operation were corrected for neutral back flow into the discharge chamber

from the bell Jar. An electrical schematic diagram showing where power

supplies are connected to the ion thruster and the points at which

measurements are made is shown in Fig. 59.

-114-
,',,- _ a::: IWa.

ILl W r--._ N _._)'-N_:l z_l I o I I

7 7,.,,1, _, I.l_--._ I. I

_: 0 ._,ll-I:Lr _l_ I'l--_.ll_" It I

,- _",-00 [ _Ul -- o_ I

::3 .,J (,3

,,, _ ,,, I '7 _--I

Z . Z I "------"

>

II

rjl-- + _> -r _..j I

_ IF I1 I .

I I / / ell-- _ :::)l T T o1_=_1 r

/ I + i.. o---,., ,.,.. z_ _ = ,,, _ P,

-- _,., _ _ z o o o_ _,= _ >

/ z __.-z z "3_-,,,,--,--_ J ___

/ • _ 0 --u_ I'-_ _ :_ -.% LIJ 0 _

I _ _ I- p (,.-,n _ _

(l)_.l + _- o ,,, ,._ i- z z z < ,- -,r := 1 .. ' ,,l_n I <__.- :_ _ v_ =-- <[ -- W_ ___ "' 17 .> n* ,-, ,-,

(/) ,L. m 3:: I.,': (,,) (-) (,.) "r. 1-- 3: (,.) I,,1.1 r,n U I,- -- 4{I

Z .., Ul ,_ ,., Z Z Z <[ Ul ,., _ ,.,

/ _: 9 3 _ _ E= o o o _ _ T= E _

I "'

I"" I I I I I I I I I I I I I

@ " " " " I

, ,., ®_, o . = _-., _®o=°>

m

o

.la

o

g)

,--I

g)

g.

.=

E-,

0

Q. o

-115-
The basic test procedure used involved first establishing a flowrate (m)

in the range between 100 and 500 mA eq., then heating the cathode, starting

the discharge and allowing it to stabilize. Typically this stabilization took

place at a discharge voltage (VD) of 50 V and a moderate discharge current

ranging from 200 to 300 mA. After stabilization the discharge current was

varied incrementally from 100 to 1,500 mA and the beam current (JB) and ion

production expressed as a current (Jp) were measured. The ion production

current was measured by biasing the chamber walls and the screen grids 30 V

negative of the cathode and measuring the ion currents arriving at these surfaces. The plasma ion energy cost (ep) and the extracted ion fraction (fb)

were calculated using these measured currents. The procedure was repeated

over a range of flowrates from 100 to 500 mA eq and then the loop anode was

moved downstream (_a in Fig. 58 was increased incrementally from its initial

locatlon at the ring cusp). At each incremental location the discharge

current was again varied through the 100 to 1500 mA range and the currents and

voltages of interest were measured. This process continued until the anode

was sufficiently far downstream so the discharge was on the verge of extinction. This corresponded to the optimum performance configuration in agreement

with Hiatt's results. The beam diameter d b was kept constant at 8 cm

throughout this test, however, some adjustments in the ring cusp location _r

and the cathode diameter dc were effected in an effort to reproduce Hiatt's

test configuration exactly.

Results

A typical plot of plasma ion energy cost versus neutral density parameter

generated using data collected in the manner Just described is shown in Fig.

60. The trial-and-error procedure used to fit the curve to the data involves

selecting values for the primary electron containment factor (C o ) and the

-116-
I

0

0

N.o-®_._'o o

II II 0

0 0

O0

0

0

0

0

0

0

0

0

I I I

0 0 0

IF) 0 Lo

(u0! owsolCllA_) [ d)] 1SO0 AOa3N3 NOI VlAISV7cl

cO

0

<_..

o -

v

0

.j _

< _

w _

W _

Z "_

0 °

e,L

O

0

-117-
S

baseline plasma ion energy cost (ep). This plot shows typical data scatter

observed during a typical experiment. To avoid the clutter associated with

many data points on curves that are located close together the actual data

points will be left off of subsequent plots.

Figure 61 shows how varying ring cusp location (_r) affects the performance of the thruster operating wlth a 3.1 cm dla cathode and a 5.2 cm dla

l

anode. Thls figure indicates that there is an optimum ring cusp location at _r

of 4.7 cm at which the value of the primary electron containment factor is 12

(A eq) -1 and the baseline plasma Ion energy cost Is 48 eV/lon. To describe

the performance of the Ion source completely using the model developed by Brophy $ one must calculate the extracted ion fraction. The effect of rlng cusp

location on thls fraction is shown in Flg. 62. The data points in this figure

are actually mean values of the extracted ion fraction recorded during data

collection of data over the range of neutral density conditions investigated

at each rlng cusp location. These data show the extracted Ion fraction reaching a very mlld peak of -0.42 at a ring cusp location of 4.7 cm. Because a

minimum plasma Ion energy cost and a maximum extracted ion fraction were both

observed at a 4.7 cm rlng cusp location, the beam ion energy cost versus propellant utilization curve should also be optimal at thls condition. Figure 63

shows how ring cusp location affects ion thruster performance when the anode is

positioned downstream of the ring cusp at the point where the discharge is on the

verge of extinction (optimized location). This figure indicates thruster performance is optimal when the ring cusp is at an _r of 4.7 cm for the ion source

described by the parameters listed in the legend. It should be noted that the

anode was positioned to yield optimum performance at each rlng cusp location

associated with the curves of Fig. 63. It is believed that the optimum performance shown in these data corresponds to a rlng cusp location that causes

-118-
0

._ E w 0 l

_ -- I •

_ _,_-_ I

m _ _ -" n _ II

II II II II 0 II Jl

" .- II

.9

u < _ IB

Ill

T ® .o " " " I II

•_, ,¢_ _ _- __ "_0. I II

& _ E .-- _ _ r._ w III

® _ o ._ _ III

E < > . • _o o_ . _ o__ \///

m co ,, ,, ,, /x,,/l

CO _" ,- O,ll._. 1 fl

,, ,, ,, .o . //i

t I I I

0 0 0 0

0 _O 0 Io

1 1

(uo! oWsold/Ae) [ d_ ] 1SO0 Ag_I':IN3 NOI VlAISV7,:I

(D

c;

u_

d ,

,,=,

8

-(_ ,.-10 _

_ o

_ o

0 "_

-119-
Z

_o

I--

(.3

<¢

IJ.

Z

0

1

Q

l.iJ

I-.-

C)

n..

l--

X

LIJ

0.50 -

0.45

0,40

0

db = 8 cm

dc = 3.1 cm

do = 5.2 cm

2d = 11.8 cm

go OPTIMIZED

d_ = 200-500 mA eq. (Ar)

I I I I

4.5 4.6 4.7 4.B

RING CUSP LOCATION

Fig. 62 Effect of Ring Cusp Location on Extracted

Ion Fraction

-120-
5OO

4OO

m

w

_--_ 500

01

0

¢.)

)-

(9

0_ 200

Z

UJ

Z

0

:E I00

m

db=8 cm

dc = 3.1 cm

da = 5.2 cm

_d = 11.8 cm

£0 OPTIMIZED , _,

rh=270mA eq. (Ar) _r = 4.8 cm

I I

0 0.2 0.4

_r = 4.7 cm

PROPELLANT

I

0.6

UTILIZATION

I

0.8

I

1.0

Fig. 63 Effect of Ring Cusp Location on Beam Ion

Energy Cost

-121-
the virtual anode magnetic field line to intercept the outer most ring of

holes in the screen grid. At this condition optimum use is being made of the

screen grid. Movement of the virtual anode field llne in either direction

would degrade performance. Moving the virtual anode field line further

out would cause significant numbers of the ions being produced to be

directed beyond the outer-most holes in the screen grid onto the region of

the grid where there are no extraction holes and onto the discharge

chamber wall. Ions that strike these surfaces pick up an electron and

they therefore come off of the surfaces as atoms and they must be

reionized thereby using additional energy before they can be extracted as ions.

On the other hand moving the virtual anode field llne at the screen grid

further in would reduce the ion current through the outer-most holes in the

screen grid while neutral atoms would still be able to escape the discharge

chamber through these holes. This also induces a degradation in discharge

chamber performance. It is the balancing of these two competing effects that

determines the optimum virtual anode field line location.

By operatlng the thruster at the optimum rlr_ cusp location it was possible to make a comparison of the optimized data obtained in the present tests

with those obtained previously by Hiatt. 12 Figure 64 shows such a comparison

at a relatively high argon flowrace (380 mA eq) and these data show good

agreement. At a reduced flowrate of 270 mA eq., however, the data of Fig. 65

were obtained and they do not show such good agreement. The reason why the

data agree at high flowrates and not at low flowrates is because the primary

electron containment factor (Co) as determined in the present tests was -12 (A

eq) -1 while it was 14 (A eq) -1 in Hiatt's tests. The baseline plasma ion

energy cost and the extracted ion fraction parameters determined in past and

-122-
500 -

c

0

400

>

0

m

w

300-

(n

0

0

>-

-" 200-

hi

Z

i,i

Z

0

I00 -

w

0

d b =8 cm

de =.3.1 cm

da =5.2 cm

Qd =11.8 cm

_r = 4.7 cm

_a OPTIMIZED

rh =380 mA eq. (Ar)

PREVIOUS

OPTIMIZED

CONDITION _ ..,,-z_

g.,_Fw" '_ -"-OPTIMIZED DATA

FOR COMPARISON

I I I I

o._ o.., o., 0.8

P,O,'_LL,,TUT,,.,Z,,,O, [ _=]

I

1.0

Fig. 64 Performanoe Comparison at High Flowrate

-123-
500

db= 8 cm

dc = ,.'3.1cm

da= 5.2cm

2d= 11.8 cm

2r = 4.7 cm

OPTIMIZED

rh =270mAeq.(Ar)

OPTIMIZE D DATA _ j_s"

FOR COMPARISON _ ..,,-'"-

' _,,0._.. ,,,"_SO

"PREVIOUS

A OPTIMIZED

CONDITION

(REF. 12)

0

J.. / / .j.

0.2 0.4 0.6 0.8

P_OPEL_NT OT,,,ZAT,ON [_o]

I

1.0

FiE. 65 Performanoe Comparison at Moderate

Flowrate

-124-
present tests both agree well and the effects of these two parameters dominate

and cause the good agreement observed in Figs. 64 and 65 at low propellant

utilizations.

There are several reasons why the primary electron contalnmenc factor

could be lower in the present set of experiments than in Hiatt's tests. The

measurement of the neutral flow into the discharge chamber could be in error

in either of the experiments. The two sets of data would agree if a neutral

flow loss due to a leak or a neutral flow measurement error of 60 mA eq could

be identified in the current experiments. The difference could also be

explained if somehow the magnetic field were not contalnlr_ primary electrons

as effectively In the present experiments as it did in ELatt's tests. To

date, it has not been possible to Idenclfy errors of thls type in the present

experiments or to find an indication of errors in the opposite sense that

occurred In the previous set of experiments.

Verification of experimental results obtained previously by Hiatt has

been achieved at flowrates above about $80 mA eq. At lower flowrates an error

has been identified that could be explained by a loss of 60 mA eq of neutral

atoms in the present experiments. In general the trends and tendencies

observed by Hiatt were confirmed in the present study. By varying the ring

cusp location for a particular discharge chamber geometry a location was found

that yielded optimum performance. With the ring cusp at this location, it is

believed that the virtual anode magnetic field line intercepts the outer most

ring of holes in the screen grid. This results in optimum utilization of the

screen grid by facilitating ion extraction over the complete open region of

the screen grid without allowing substantial numbers of ions to pass beyond

the outer-most holes in the screen grid to a region of the discharge chamber

where they recombine and are lost.

-12S-
HIGH CURRENT HOLLOW CATHODE RF.SEARCH

There is an anticipated need for a high current (-50 A) xenon hollow

cathode in eleatrodynamlc tether plasma contaotor and hlgh-power ion thruster

applications. Currently available hollow cathodes have been tested to 10 and

20 A levels and it might be possible to extend their operating envelope to

this current. However, it is considered desirable to analyze present hollow

catnode designs to determine if 2 to S-fold increases in operating current

levels might cause operational problems. In order to do this the hollow

cathode model developed by Siegfrled 11 was programmed and exercised to determine the operatlng conditions that might be expected at the $0 A emission

current level. The key design parameters associated with this model which are

used in the analysls are illustrated in Fi_ 66.

The hollow cathode shown in Fi_ 66 consists of a cathode tube and orifice plate enclosing a low work function insert from whloh electrons are

extracted. The basis of the Siegfrled model is that electrons, emitted from

the downstream segment of the insert having a length Re are drawn into the

region enclosed by the dotted llne in FiE. 66. It is argued that within this

reglon the plasma potential, plasma density and electron temperature are relatively constant and that primary electrons emitted from the insert undergo

ionlzlng collisions with the neutral gas atoms which are also present within

this reglon. This neutral gas is maintained at the pressure P in the reglon

by expellant gas which is fed in at a flowrate m. Electrons in the region,

which are both drawn from the insert and produced as a consequence of ionization, are extracted through an orifice of diameter d to an anode located

o

downstream of the orifice plate. Ions produced in the dlscharge are on the

other hand, extracted across the dotted line boundary to the insert, orifice

-126-
ILl

Cbl,l

,..;.._

ew j

On

a.

,E

ILl

_o

-- _ o

J

z

,.,o_, _°_

_z o ,,,w _ _ _.

II II n II II II II II

•-- 0 Q .*- :J_ Q.

_ _m "E O. ,C_ I--

-127-
plate or cathode tube. Those ions crossing the boundary adjacent to the

insert strike the insert thereby heating it to sustain its temperature and

facilitate electron emission through the field-enhanced thermionlc emission

mechanism. The Siegfried model applies the equations of conservation of

energy at the insert surface, conservation of energy within the ionizing

volume, conservation of ion current lost from the ionizing volume at the Bohm

velocity and electric field enhanced electron emission from the insert to

determine the relationship between the variables indicated in the legend of

Fig. 66. In order to apply the model it is necessary to assume only an electron temperature fop the plasma within the dotted boundary and a value for the

ratio of emission region length (le) to primary electron scattering mean free

path since other design variables are defined. The present analysis assumes

0.7 eV and unity respectively for the parameters on the basis of xenon hollow

cathode experiments conducted by Siegfried on xenon at low currents 13.

Simultaneous solution of the equations Just described yields results such

as those shown in Fig. 67 for a 50 A xenon cathode with an insert diameter of

0.5 cm and an insert work function of 2.1 eV. The upper plot of Fig. 67 shows

insert temperature variation as a function of emission region length _e with the

heat removal rate from the insert Qthas a parameter varying from 200 W to 1000 W.

This heat removal rate from the insert is determined in a piece of hardware by

such parameters as the thermal contact resistance between the insert and the

cathode body, the radiative heat transfer environment of the cathode and the

thermal conductance of the cathode tube between the insert and some structure

tending to remain at a fixed temperature. It should be noted that the heat

removal rates indicated of Fig. 67 are one or two orders of magnitude greater than

those computed for previously tested low-current cathodes 11 but they are

considered appropriate for operation at the hlgh current involved.

-128-
0

0

13..

:_ 1200

i,i

I--

In.,"

i,i

(n I000

Z

m

50A XENON CATHODE

1400- di =0.5cm

HEAT REMOVAL RATE _ =2. I V

0

_ __.._...200 W f400 W

m

I I I I I I I I

2 4 6 8 I0 12 14

EMISSION LENGTH (ram)

>

_1

I--

Z

i,i

I--

0

0..

(f)

..J

in

6O

5O

40

3O

20

I0

0

HEAT REMOVAL RATE

I000 W

6O0 W

8OO W

200 W 400 W

I I I I I I I

2 4 6 8 I0 12 14

EMISSION LENGTH (ram)

o

I.'-

i,i

e:

:3

(/')

LIJ

n."

in

13::

0

nbJ

Z

Fig. 67

12

I0

8

6

4

2

W 400 W

0 2 4 6

HEAT REMOVAL RATE

6OO W )OW

I000 W

8 I0 12 14

Effect of Heat Removal Rate and Emission

Region Length on Cathode Performance

Parameters

-129-

EMISSION LENGTH (ram)
The data in the upper plot of Fig. 67 show that insert temperatures in

excess of 1200°C will be realized for the full range of emission lengths associated with a 200 W heat removal rate and that the insert emission length can

be expected to extend only a little beyond 2 mm in this case. As the heat

removal rate is increased, the curves of Fig. 67 show the insert temperature

should drop and the emission length should increase to the point where excessively high emission current densities will not be demanded. It should be

noted in considering the data that an insert temperature less than 1100°C is

generally preferred to prevent migration and loss of the low work function

material loaded into the insert.

The middle graph in Fig. 67 shows how plasma potential within the cathode

varies as a function of heat removal rate and emission length. It indicates it is

desirable to operate with a long emission length in order to keep the interior

plasma potential and therefore the keeper voltage low. Hence increased emission length induces both of the desirable effects associated with the top two

graphs (low insert temperature and low keeper voltage). However, the desirable effect for the upper graph (lower temperature) is realized by increasing

heat removal rate while the desirable effect for the lower one (lower potentlals) is realized by decreasing it. On the other hand mlnlmlzlng the insert

heat removal rate, an operating loss, maximizes the efficiency of at,7 hollow

cathode device so in this sense it is also desirable. In view of these considerations then it is probably desirable to operate with a large emission

length and small heat removal rate. The bottom plot of Fig. 67 shows the

desired large emission length is achieved by operating with a low cathode

interior pressure. Unfortunately the plot also shows emission length limits

are encountered at low heat removal rates (a little over 2 mm for 200 W heat

rejection rate and about 7 mm for the 400 W rate) where interior cathode

-130-
pressure drop to near-zero values. What is occurring here can be understood by

recognizing that increases in emission length result from increases in the primary

electron mean free path. This path is increases by reducing the densities of

either Maxwelllan electron electrons or neutral atoms. In fact, however, reducing

the neutral density to zero results in a near-zero pressure. Under this condition

only the Maxwelllan electrons and the ions needed to conduct the current being

demanded are left in the cathode and there is no mechanism to reduce the mean free

path below the limit imposed by this near-zero pressure condition. It is the

imposition of this limit that causes the 200 W and 400 W curves in the upper and

central plots of Fig. 67 to terminate suddenly.

An overview of the data of Fig. 67 suggests the preferred operating mode

involves reducing the heat removal rate (to say 200 W). Figure 68 shows plots

of interior cathode pressure as a function of emission length again for the 50

A xenon cathode and a 2.1 V work function insert but for inserts having diameters that vary from 0.$ cm to 1.5 cm. These curves show that one can indeed

increase the emission length at which the interior cathode pressure drops to

zero. This is accomplished by increasing the insert diameter. In fact, for

the 200 W heat removal rate case, increasing the insert diameter from 0.5 to

1.0 cm increases the emission length corresponding to zero interior cathode

pressure from 2 to over 6 mm.

Figure 69 shows a comparison of insert temperature, plasma potential, and

xenon flux through the hollow cathode (proportional to flowrate per unit area)

as a function of insert length for the SO A xenon cathode at a heat rejection

rate of 200 W and for two insert diameters (0.5 and 1 cm). These data suggest

that the 1 cm dia insert (solid llne) facilitates operation with greater emission

lengths, lower insert temperatures and acceptably low plasma potentials if the

xenon flux level is kept low. Since these operating conditions are preferred, the

-131-
Fig. 68

12 / 50A XENON CATHODE

=o _ _ =2.=v

/N

\ k__ HEAT REMOVAL RATE

\ \______ f 600w _eoow

, , ; i

0 2 4 6 8 I0 12 14

EMISSION LENGTH (mm)

12-

i,i

o_ I0-

n

g6

FZ

W

Q

0

-r. 0

_J

d i = 1.0 cm

HEAT REMOVAL RATE

200 W 400 W 600 TO I000 W

I I I I

2 4 6 8 I0 12 14

EMISSION LENGTH (ram)

12-

I0

8

6

4

2

d i =1.5 ¢m

HEAT REMOVAL RATE

_, /200w

_ ,_ /4oow

L I J l

0 2 4 6 8 I0 12 14

EMISSION LENGTH (ram)

Effect of Insert Diameter on Cathode

Interior Pressure

-132-
1400 -

1200

In,"

bJ

(/3

z I000:

0

INSERT DI A

I

2

50A XENON CATHODE

 -2iv

-(_th : 200 W

I I I I

4 6 8 I0

EMISSION LENGTH ( mm )

I I

12 14

60

v

_1

i

_'- 40

Z

UJ

I--

0

n

_¢ 20

o_

c¢

..J

n

0

600

• 4O0

x E

D

_1 o"

g aoo

X

0

Fig. 69

,_I cm

I I I I I

2 4 6 8 I0

EMISSION LENGTH (ram)

I I

12 14

- I

I

! m

!t_/_.- O. 5 cm

1 cm

%%I _ I I I

2 4 6 8 I0

EMISSION LENGTH (ram)

Effect of Insert Diameter on Cathode

Performance Parameters

-133-

I I

12 14
data suggest the larger insert diameter is preferred at high current levels.

should be noted that xenon flux, which is the mass flowrate of xenon per unit

orifice diameter squared, is directly related to cathode internal pressure.

It

If one were to assume the flow through the cathode orifice were

isentropic, choked flow, one might expect a linear relationship between

cathode interior pressure and cathode expellant flux. In the present case,

however, the discharge current also affects the pressure that would be

measured at a fixed flowrate as a results of the collisions that undoubtedly

occur in the orifice region. Since this collision process has not been

11

described analytically, Siegfried's approach was followed and an

empirical relationship between the discharge current and the pressure per unit

xenon flux was obtained. This was done by measuring the cathode internal

pressure and flowrate associated with a cathode having a 0.71 mm dia orifice

as a function of discharge current. The results are sho_n in Fig. 70. Since

the experlmental apparatus allowed meast_ements to be made only to a discharge

current of 15 A an extrapolation to 50 A was needed to project the xenon flux

requirements associated with the lowest curve in Fig. 69. This extrapolation

was not difficult because the curves of Fig. 70 seem to level out at currents

beyond about 15 A. The xenon flux level specified in the lowest plot of Fig. 69

should probably be considered to be accurate to within about 100_ because of

the uncertainty associated with this assumption. It should be noted regarding

to the overall analysis that there are uncertainties associated with the work

function associated with a particular insert, the collision cross sections

which have been incorporated into the model to compute primary electron mean

free paths and the pressure /flux correlation of Fig. VO. This suggests that

precise values of the insert diameter required to sustain operation will be

uncertain. Still the trends suggested by Fig. 69 are considered correct and

they indicate that the occurrence of overheating in an insert can be mitigated

-134-
g 0 _ <

X _

0 <

\

(Vw/aww- JJO..L)

0

m

q

0

3_nss3_d

0

O<t

Z

ILl

0

W

0

r_

0

0

4-_ 4,a

,,--I

L r_,

g,, :_

0 rj

C_

"

•8.1 e_

e.

-13_-
by increasing the insert diameter and reducing the xenon flowrate to the point

where the emission wlll be occurring over a substantial length of the insert.

It can be stated that the model results suggest a 1 cm insert diameter is

preferred over a 0.5 cm one.

-136-
CONSTRAINED SHEATH OPTICS RESEARCH

The high specific impulse capabilities of ion thrusters have lone been

recognized as well suited to high characteristic velocity missions such as

planatary and deep space probes. Over the past decade, however, the program

emphasis within NASA has been directed more toward near-earth missions where

ion thrusters must perform at relatively low specific impulses in order to

minimize trip times under the constrained of fixed power. Further, it is

desirable to use a small number of high thrust-denslty ion thrusters operating

at these low specific impulses to produce the required thrust. This could be

aocompllshed by operating at a low net accelerating voltage to achieve low

specific impulse operation and a high total accelerating voltage to achieve

the desired high thrust density. Both experimental and theoretical data have

shown, however, that conventional two-grld ion optics systems tend to exhibit

high ion beamlet divergence losses and high direct ion impingement currents

when they are operated at the low ratios of the net-to-total accelerating voltage (8<0.5) required. 14'15 Operation can be reallzed at low net-to-total

accelerating voltage ratios by using mechanically more complex three-grid ion

optics systems but the associated reduction in R to about 0.2 is still

15,16

relatively modest. The effort described here was undertaken to examine

the possibility of constraining the shape of the sheath through which ions are

extracted from the discharge plasma to facilitate operation at high current

densities and low values of R without encountering the high beamlet divergence

behavior that limits the performance of free-sheath optics systems.

In order to understand the mechanisms by which ions are accelerated and

diverted from axial trajectories in Ion-optlcs systems, it is instructive to

consider the diagram shown in Fig. 71. This diagram shows equlpotential contour lines associated with the case where ions are extracted from a flat plane

-137-
>

0

0

io

I

> >

0 0

0 0

0 to

I I

</ > >

> 0 0 0

0 > 0 0 0

Io 0 I0

0 , "T V

0

o

o

E_

0

I-4

"0

:3

0

0

C_

c_

0

e._ e_

-138-
at 500 V potential, accelerated through a 460 V screen grid aperture and a

-2500 V accelerator grid aperture to a zero potential, flat plane at which

neutralization of the ions is presumed to occur. The equipo_entlal lines

shown In the figure are the solid ones and the dotted llne shows a conceptualized trajectory for an ion comlng from the ion source plane and passing

through the system at a substantial distance from the centerllne of the apertures. The arrows emanating from the dots along the trajectory suggest the

direction of electric field vectors along which forces act on the ions as they

pass through the grid system along th_s trajectory. The arrow at position A

suggests initial electric field forces acting on the ion have both a downstream axial component and an inward radial component. This inward radial

force component causes the ion to be directed closer to the aperture centerllne as it accelerates between the grids. Thls process continues until point

B where the electric field contour is observed to be flat and essentially perpendlular to the centerllne. At this point, only the axial component of electric field acts. Beyond point B the electric field vectors begin to have a

component of electric field directed radially outward as shown at point C. At

point D the ion begins to decelerate due to a component of electric field that

is directed axially upstream. By the time the ion reaches point E the component of electric field as indicated by the arrow again has a significant

outwardly directed radial component in addition to the upstream directed axial

component. Figure 71 taken in total shows that the ion sees an outwardly

directed radial force that causes it to diverge from point B to the neutralization plane. As the accelerator grid is biased more negative (i.e. the R

value is reduced) the Ion sees a component of outward directed electric field

for a greater and greater fraction of the time associated with its passage

through the optics system.

-139-
An ion moving along the centerline of the aperture, on the other hand,

passes through equlpotentlal surfaces that remain perpendicular to the axis of

the screen and accelerator (accel) holes throughout its time of passaEe through the

system. This ion therefore sees no radial component of electric field and as a

result it would not diverge. One can see from thls discussion that both the

radial location at which an ion begins the acceleration process and the netto-total accelerating voltage ratio influence its divergence angle at the concluslon of the process. One can also see that low divergence operation at a

low net-to-total accelerating voltage ratio could be achieved if the bulk of

the ions could be directed through the grids along trajectories near the aperture

centerllne.

If ions were to reach an ion source plane curved in such a way that the

ions would all be directed toward the aperture centerllne, one might be able

to minimize the radial force components and thereby minimize the divergence of

the beamlet exiting the grid aperture. In conventional two and three-grid

ion optics systems the ion source surface shape, i.e. the sheath shape, is

determined by the relative balance between the rate at which ions are supplied

from the discharge and the rate at which they are extracted through the optics

system. In this case it is not possible to control the ion source surface

shape (free-sheath case). If on the other hand one could apply a properly

shaped fine wire mesh over the screen aperture it would be possible to establish

an immovable ion source surface having any desired shape ("constrained sheath'1).

Previous experimental investigations utilizing the constrained sheath

concept include the work of Byers 17 who discovered that the wire mesh tended

to melt, Relfenschweiler and Frohner 18 who achieved well-focused beams at high

accelerating voltage levels and more recently Aston 19 whose decoupled extrac-

-140-
tion and acceleration concept is quite slmllar to the constrained sheath

optics concept.

A schematic diagram showing the position of the sheath constraining mesh

relative to the discharge plasma and the screen and accelerator grids is shown

in Fig. 72. This figure also shows the various geometrical symbols used in

describing the experiments. For example, the diameters of the screen and

accelerator grid apertures (ds and da) the spacing between the grids (_g) and

the thicknesses of the screen and accelerator grids (t s and ta) are defined.

For all cases considered in this study, the sheath constraining mesh was made

of 0.25 mm dla tantalum wlre mesh contoured to a spherically shaped surface of

radius rc where rc is limited to values between ds and infinity. The

effective acceleration length _e shown in Fig. 72 was used here to

compute perveance. This effective acceleration length can be expressed in

terms of the other parameters defined in the figure by the equation shown

there. Beamlet divergence results will be presented in the form of the divergence half-angle shown in Fig. 72. This angle is defined as the half-angle of

the cone with its apex at the accelerator grid which encloses 95% of the beam

being extracted through the grid system.

Apparatus and Procedure

Tests were conducted using an 8 am diameter ring cusp ion source having a

single 2.9 cm dia screen aperture at its downstream end. All tests were conducted using argon propellant and the ion beam current density profile was

measured about ll am downstream of the accelerator grid using a Faraday probe

that could be swept across the beam. Reference 20 details the ion source and

Faraday probe system. In order to conduct the tests a stable argon discharge

-141-
IfBEAMLET

_g //_ ENVELOPE

DISCHARGE :_._ _ / I _

PLASMA do

I APERTURE PAIR

ICENTERLINE

SHEATH / //////_

/ /// ///

SCREEN ACCEL I

GRID GRID

CONSTRAINING

MESH

I G

(DIVERGENCE

I ANGLE)

I

N_EUTRALIZATION

PLANE

"' ((")

Fig. 72 Constrained Sheath Optics Schematic

Diagram

-142-
was established in the ion source and the source and accelerator grids were

then biased to achieve the desired net and total accelerating voltage condition. The ion beam current extracted was measured directly using a beam

current meter and indirectly by integrating the beam current density profile

obtalned uslnE the Farady probe. The ion source discharge current was then

increased in steps and ion beam profiles were measured at each operating condition.

Figure 73 shows example profiles measured at three difference discharge

and beam current conditions characterized using a normalized perveance associated with the slr4_le aperture belr_ operated by the equation

f ol

p l n •

v312 4

In thls equation JB is the beam current, VT is the total acceleratlng voltage

and _e and ds are defined in Fig. 72. The limiting value of this parameter

determined from the one-dimenslonal Child-Langmuir law is 6.8 x IO-9A/V 3/2 for

the argon propellant being used in the study. The three ion beam profiles shown

in Fig. 73 reveal the beam profile structures typically observed in the tests.

At low perveance values the profiles were gausslan in shape, however at high

perveance values t_ey tended to bifurcate producing two peaks away from the

centerllne of the beamlet. It is suggested that this bifurcation might be due

to excessive repulsion forces that tend to develop when too many ions are

forced along the centerline of the apertures.

In order to characterize the current density profiles, the measured data

were put into a computer program that generated a polynomial curve fit of

-143-
t--9 = t--_-s=0.055

ds ds

{_--_g= O. 22

cls

VNET = 500 V

CURRENT DENSITY

(p.A/cm _)

d s = 2.9 cm

re = 0.66

ds

ds

R =0.048

20 ARGON

x 10 -9 A/V 3/2 PERVEANCE

/ I I _ I I I _ I I L

-2_ -2o -i_ -,o -_ __ _ so ,5 20 25

40 RADIUS (cm)

EANCE

I I _ I | I i i _ _ I

- 5 -20 -15 -I0 -5 _ 5 I0 15 20 2.=

RADIUS (cm)

2 6.8 x I0 -9 A/V zj2

-25 -20 -15 -I0 -5 _ 5 I0 15 2_0 2_

RADIUS (cm)

Fig. 73 Typical Current Density Profiles

-144-
the data. These curve fits were then integrated to determine a beam current

and the divergence half angle. The half angle was found by first determining

the radial distance to which the profile had to be integrated to obtaln a beam

current that was 95_ of the total integrated beam current. The arc tangent of

tnls radius divided by the distance from the accelerator grid to the probing

plane is the desired half-angle. The computer program was also used to determine the ion beam divergence factor which is defined as the ratio of the

thrust produced by the beam to the maximum thrust produced if all of the ion

trajectories leaving the aperture system were parallel to the axis of the

apertures. The computer program determined this factor by integrating the exit

velocity for ions passing through a differential, clrcumferentially shaped band

at radius r times the cosine of the divergence angle for that band of ion

trajectories from the centerline to the edge of the beamlet.

_esults

Typical results obtained experimentally and shown in Fig. 74 are plots

of divergence angle and ion beam divergence factor as a function of

normalized perveance per hole. Impingement current data associated with

results llke those of Fig. 74 will not be presented here because they they

were small (a few percent of the beam current) until substantial direct ion

impingement began to occur near the perveance limit of the system. In the

particular case of the data shown in Fig. 74, the separation distance between

the grids was varied while other parameters were held constant at values

defined in the legend. The general trend suggested is that divergence angle

increases substantially as normalized perveance per hole is increased. In

each case tnls increase in divergence angle is reflected in a reduction in ion

beam divergence factor with values extending as low as about 0.75 for the very

closely spaced grids. As grid separation is increased the extent of diver-

-145-
d s :2.9 cm

r.-,_-c= O. 66

ds

d...g= I

/ _ =o.15_ _ .._.- _ =_ =o.

/ _/%_,., .,-, ._ "g :0 55- VNET: 50C 401- _._, ;-:u.oo..., L_,,-- A _ ds " -_, ....

l y _,s ..). _ R =u.ir

0,6

,,=, IS'"/ I-D MODEL

I01-" u/ THEORETICAL

_L LS MAXIMUM /

/ , ' , , t , ,xlO -9

0 2 4 6 8 I0

NORMALIZED PERVEANCE PER HOLE (A/V 3_z)

i,Zl09- ",_ "'_......_ _ Ttt._

>, T -o.s5

" d_-"" '_. -""" "_

0.7

i i i i i x I0-e

0 2 4 6 8 I0

NORMALIZED PERVEANCE PER HOLE (A/V 3#2)

ILl

Z

0

Fig. 74 Effeat of Grid Separation on Divergence

Charaeterlstles

-146-
gence angle increase drops substantially. It should be noted that the netto-total accelerating voltage ratio for these data are 0.17 and the lowest

value of thls ratio investigated by Aston for conventional, free-sheath

three or two-grld optics systems was about twice thls value. 14'16 When Aston

attempted operation at lower R-values he observed poor divergence

characteristics and hlgh impingement currents. In Aston's case, however, the

adverse effect of perveance on divergence angle was much more subdued and

divergence angles less than 20 ° were generally observed as long as R-values

were above about 0.3.

One startling feature of the data of Fig. 74 is the maximum normalized

perveance at which data were obtained. It was expected that the mesh screen

on the screen grid would collect a portion of the ion current and hence the

maximum normalized perveance at which the grids could be operated would be

su_stantlally below the theoretlcal limit which for argon is 6.8 x 10 -9

A/V 3/2. In the present case however that normalized perveance values

approaching 10 x 10 -9 A/V 312 are achieved before direct ion impingement on

the accelerator grid became excessive and limited the extractable ion beam

current. It could be that the theoretical perveance limit is being

exceeded when the constrained sheath optics sets are used because the

constraining sheath, which remains fixed in space, maintains a large sheath

area exposed to the discharge plasma as the discharge plasma density is

increased. In the free sheath case, on the other hand, the sheath collapses

toward a plane at the screen grid thus reducing its effective area for the

discharge chamber ion collection as perveance is increased. The high perveances realized in the present case could, however, also be related to the

definition of the effective acceleration length _e identified in Fig. 72. This

-147-
length, which is defined to extend from the upstream plane of the accel grid to

the upstream tangent to the sheath constraining mesh_is appropiate for ions

passing on the centerline of the extraction system, but may be too long for

ions extracted from the discharge near the screen grid edge• Finally

regarding the ion beam divergence factors in Fig. 74, it is noted that beamlet

divergence induces some thrust loss, but this loss may be more than compensated by the ability to operate at higher total accelerating voltages where

the current densities are greater than those that could be extracted from freesheath optics systems• More work needs to be done before quantitative responses

to the question suggested by this comment can be made.

The effect of changing the net accelerating voltage while holding the

net-to-total accelerating voltage ratio and the geometrical factors associated

with the grid system constant, is shown in Fig. 75. These data indicate that

changing the net accelerating voltage from 500 down to 250 V induces no substantial change in divergence characteristics. In regard to the Fig. 75 data

it is suggested that the triangular data point at the high perveance associated with 500 V operation is probably erroneous. Figure 76 shows the effect

of changing the net-to-total accelerating voltage ratio on divergence when the

net accelerating voltage and geometrical factors are held constant. These

data show grid performance improves when the higher net-to-total accelerating

voltage ratio is used. At high perveances, however, the data suggest the

divergence characteristics are similar.

It is noted at this point that beam currents determined by integrating

the current density profiles obtained at each operating condition agreed with

the associated measured ion beam current to within about 10_ when the net-to-

-148-
m

¢1)

L

O_

"0

LU

/

CO

Z

l.IJ

0

Z

I.i.I

CO

O_

l.iJ

>

Q

5O

40

3O

2O

I0

0

ET

VNE? 500V

A

I

: 250 V

A

I I I I

2 4 6 8

NORMALIZED PERVEANCE PER HOLE (A/V 3/2)

ds : 2.9 cm

r__c: 0.66

ds

d_go"1

d$

t_aa= t_

d$ ds =0.055

R = 0.091

_--_g: 0.22

ds

ARGON

J xlO-9

I0

1.0"

G

,,<

bU

0

Z

bJ

(9

n_

LU

>

Q

0.9

0.8

0.7

0

I I I

2 4 6

NORMALIZED PERVEANCE PER

VNET=500 V

VNE T = 250 V

I

8

HOLE (A/V 3/2)

J xl0-e

10

Fig. "/5 gffeot of Net Aooeleratlng Voltage on

Divergence Characteristics

-149-
50-

4O

30

=,

2O

Z

LIJ

(D

e: io I,I

>

0

ds = ?__.9cm

r..c.c= 0.66

ds

da

Ls_=l

R =0 091 _ ta ts • _ .__:----;-_'_-'-_ - E- o.o_

- o ,.z_

-,_-_ "_ "" VNET=500 v

t" S_ _-o_

/ // \R-O. 17 d,

f //" .... ARGON

/7

/

/

- X

I I I I _ x I0 -9

?. 4 6 8 I0

NORMALIZED PERVEANCE PER HOLE (A/V sl_')

_-,or _---_,_\.._.f-- R--o., T

,=,,0.9-

0.8 - R =0.091 _

=E

I.IJ

Z

0

0.7-

0

I I I I

2 4 6 8

NORMALIZED PERVEANCE PER HOLE (A/V 312)

J x I0 -9

I0

Fig. ?6 Effect of Net-to-Total Acoeleratlng

Voltage Ratio on Divergence

Charaeterlstlos

-150-
total accelerating voltage ratio was 0.17 and above. As the net-to-total

accelerating voltage ratio was reduced below 0.17, however, the integrated

beam currents determined from current density measurements dropped systematically below the measured beam currents. This suggests that some ions were

escaplr_ detection by the Faraday probe even though it sensed currents leaving

at half-armies of divergence up to about 50°. The fact that integrated beam

currents were less than measured ones indicates that slgnlflcant numbers of

ions may have been escaping on very divergent trajectories at low net-to-total

accelerating voltage ratios. These ions might also be expected to reach the

probe at such large radii that very small current densities could go

unaetected yet contribute substantially to total ion current. These ions

would be expected to pass through the grid aperture away from the aperture

centerllne where they would see a large radial component of the electrlc field

directing them outward.

The effect of chan_ir_ the radius of curvature associated with the screen

constraining mesh is shown in Fig. 77. These data suggest that it is preferable to operate with the smaller radius of curvature mesh which directs ions

nearer to the aperture centerllne where theory sugEests the radial components

of electric field that induce divergence are smaller.

A sheat_ constrainlr_ mesh fabricated from tantalum wire and spot welded

to the screen grid can influence the focusing of ions extracted through twogrid accelerator apertures. Such a mesh need not limit the perveance of the

extraction apertures to values below the theoretical limit for the propellant

used. The divergence of the resulting beamiets can be reduced by increaslr_

the separation distance between the grids, by increasing the net-to-total

-151-
0

r c _'_ dOos = 2.9

/,%/_

i //_ = 0.66

ds _ss = I

/

ts to

, _ =_=o.o_

3o // R =o.17

VNET= 500 V

20 i# _

z ARGON

J a i i i x I0 -9

0 2 4 6 8 I0

NORMALIZED PERVEANCE PER HOLE (A/V 312)

1.0

o_

hi 00.9

Z

W

n,"

LU

>-0.8

a

w

li1

0.7

Z

0

0

I I I I

2 4 6 8

NORMALIZED PERVEANCE PER HOLE (A/V 3/2)

Fig. 77 Effect of Constraining Mesh Radius on

Curvature on Divergence Characteristics

-151-
accelerating voltage ratio and by decreasing the radius of curvature of the

sheath-constrainlng mesh. In this regard, it is noted that the present study

has been limited to hemispherically shaped mesh surfaces and there is no

theory indicating such a shape is optimum. In fact, the preferred shape would

be expected to have a radius of curvature which varles as a function of radius

from the aperture centerllne. The experimental results also indicate that a

variation in the net accelerating voltage range from 250 to 500 V induces no

substantial change in the divergence characteristics of the beamlet.

-153-
REFERENCES

•

•

•

•

.

o

•

Q.

•

Sisson, J. M., "Developmen_ Status of First Tethered Satellite

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-154-
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-155-
19. Aston, G. and W. D. Deinlr_er, "The DE-AC Accelerator System for

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