Novel Steering and Control Algorithms for Single-Gimbal Control Moment Gyroscopes

Permanent Link: http://ufdc.ufl.edu/UFE0042129/00001

Material Information

Title: Novel Steering and Control Algorithms for Single-Gimbal Control Moment Gyroscopes
Physical Description: 1 online resource (164 p.)
Language: english
Creator: Leve, Frederick
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2010


Subjects / Keywords: attitude, control, elliptic, hyperbolic, singularities, small, steering
Mechanical and Aerospace Engineering -- Dissertations, Academic -- UF
Genre: Aerospace Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation


Abstract: NOVEL STEERING AND CONTROL ALGORITHMS FOR SINGLE-GIMBAL CONTROL MOMENT GYROSCOPES The research presented in this manuscript attempts to first systematically solve the SGCMG steering and control problem. To accomplish this, a better understanding of singularities associated with SGCMGs is required. Next based on a better understanding of singularities, a Hybrid Steering Logic (HSL) is developed and compared to the legacy methods singular direction avoidance(SDA) and local gradient (LG) methods. The HSL is shown analytically and numerically to outperform these legacy methods for a four-SGCMG pyramid arrangement in terms of attitude tracking precision. However, all of the methods are susceptible to gimbal-lock. A method referred to as Orthogonal Torque Compensation (OTC) is developed for singularities with gimbal-lock in SGCMGs, which are known to present a challenge to most steering algorithms. Orthogonal Torque Compensation conditions the attitude control torque by adding torque error orthogonal to the singular direction when at singularity. This method can be combined with any steering algorithm including HSL and is proven analytically to be stable and escape singularities with gimbal-lock. Finally, the problems with scaling systems of SGCMGs are discussed. It is found that scaling produces an increase in the gimbal-wheel assembly inertia of the SGCMGs which in turn increases the effect of the dynamics associated with these inertias. Through analysis and numerical simulations, it is shown that more significant gimbal-flywheel inertia reduces the performance by increasing torque error and reducing torque amplification of SGCMGs. Since most of the legacy algorithms used for singularity avoidance and escape use the gimbal rates for control, the performance is degraded when the dynamics from the gimbal-wheel assembly inertia are increased. This degraded performance is shown to also be dependent on the ratio of gimbal-wheel assembly inertia to nominal SGCMG flywheel angular momentum. The overall result here is that the same legacy steering algorithms that use gimbal rates for control cannot be used for systems of SGCMGs of a reduced scale.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Frederick Leve.
Thesis: Thesis (Ph.D.)--University of Florida, 2010.
Local: Adviser: Fitz-Coy, Norman G.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2010
System ID: UFE0042129:00001

Permanent Link: http://ufdc.ufl.edu/UFE0042129/00001

Material Information

Title: Novel Steering and Control Algorithms for Single-Gimbal Control Moment Gyroscopes
Physical Description: 1 online resource (164 p.)
Language: english
Creator: Leve, Frederick
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2010


Subjects / Keywords: attitude, control, elliptic, hyperbolic, singularities, small, steering
Mechanical and Aerospace Engineering -- Dissertations, Academic -- UF
Genre: Aerospace Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation


Abstract: NOVEL STEERING AND CONTROL ALGORITHMS FOR SINGLE-GIMBAL CONTROL MOMENT GYROSCOPES The research presented in this manuscript attempts to first systematically solve the SGCMG steering and control problem. To accomplish this, a better understanding of singularities associated with SGCMGs is required. Next based on a better understanding of singularities, a Hybrid Steering Logic (HSL) is developed and compared to the legacy methods singular direction avoidance(SDA) and local gradient (LG) methods. The HSL is shown analytically and numerically to outperform these legacy methods for a four-SGCMG pyramid arrangement in terms of attitude tracking precision. However, all of the methods are susceptible to gimbal-lock. A method referred to as Orthogonal Torque Compensation (OTC) is developed for singularities with gimbal-lock in SGCMGs, which are known to present a challenge to most steering algorithms. Orthogonal Torque Compensation conditions the attitude control torque by adding torque error orthogonal to the singular direction when at singularity. This method can be combined with any steering algorithm including HSL and is proven analytically to be stable and escape singularities with gimbal-lock. Finally, the problems with scaling systems of SGCMGs are discussed. It is found that scaling produces an increase in the gimbal-wheel assembly inertia of the SGCMGs which in turn increases the effect of the dynamics associated with these inertias. Through analysis and numerical simulations, it is shown that more significant gimbal-flywheel inertia reduces the performance by increasing torque error and reducing torque amplification of SGCMGs. Since most of the legacy algorithms used for singularity avoidance and escape use the gimbal rates for control, the performance is degraded when the dynamics from the gimbal-wheel assembly inertia are increased. This degraded performance is shown to also be dependent on the ratio of gimbal-wheel assembly inertia to nominal SGCMG flywheel angular momentum. The overall result here is that the same legacy steering algorithms that use gimbal rates for control cannot be used for systems of SGCMGs of a reduced scale.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Frederick Leve.
Thesis: Thesis (Ph.D.)--University of Florida, 2010.
Local: Adviser: Fitz-Coy, Norman G.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2010
System ID: UFE0042129:00001

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2010 Frederick A. Leve

Dedicated to my mother for always being there to support me


I would like to thank first my advisor Dr. Norman Fitz-Coy for providing me with

the guidance and knowledge for this great research I undertook. Second, I would

like to thank my committee members Dr. Warren Dixon, Dr. Anil Rao, Dr. William

Hager from UF, and Dr. Scott Erwin from the Air Force Research Lab Space Vehicles

Directorate. My committee comprises the expertise in the areas of research that would

provide me the best opportunity for my research. Last, but not least, I would like to

thank my colleagues in my research lab who provided input throughout my time as a

graduate student that aided in this research: Dr. Andy Tatsch, Shawn Allgeier, Vivek

Nagabushnan, Josue Munoz, Takashi Hiramatsu, Andrew Waldrum, Sharan Asundi,

Dante Buckley, Jimmy Tzu Yu Lin, Shawn Johnson, Katie Cason, and Dr. William



ACKNOWLEDGMENTS ................... ............... 4

LIST O FTABLES ..................... ................. 8

LIST OF FIGURES .................... ................. 9

ABSTRACT .................... ................... .. 13


1 INTRODUCTION .................... ............... 15

1.1 History and Background ................... ........ 15
1.1.1 Gyroscopic Rate Determination .. .. 15
1.1.2 Spin Stabilized Spacecraft ..... .. .. 15
1.1.3 Spacecraft Attitude Control through Gyrostats .... 15
1.1.4 3-axis Attitude Control of Spacecraft ... 16
1.1.5 Single-Gimbal Control Moment Gyroscopes (SGCMGs) 17
1.1.6 Double-Gimbal Control Moment Gyroscopes (DGCMGs) 17
1.1.7 Variable-Speed Control Moment Gyroscopes (VSCMGs) 18
1.2 Problem Statement .................. ........... 18

2 DYNAMIC MODELS ... .............. ............ .. 20

2.1 Dynamic Formulation ... .............. ......... .. 20
2.2 Singular Surface Equations ........................ 25
2.2.1 Elliptic Singularities .. 26 External singularities . ... 27 Elliptic internal singularities 29
2.2.2 Hyperbolic Singularities ... 29 Non-degenerate hyperbolic singularities ... 30 Degenerate hyperbolic singularities ... 30
2.2.3 Gimbal-Lock.. ........................... 30
2.3 Singularities for SGCMGs Mathematically Defined .... 31


3.1 Common SGCMG Arrangements ... 34
3.1.1 R ooftop . .. 34
3.1.2 Box ... . 35
3.1.3 3 B ox . ... 44
3.1.4 Scissor Pair . 45
3.1.5 Pyram id . .. 46
3.2 Choice of Arrangement ............................ 48
3.3 S im ulation . . 50

4 SURVEY OF STEERING ALGORITHMS ................ ... 54

4.1 Moore-Penrose Pseudo-Inverse ................ ....... 55
4.2 Singularity Avoidance Algorithms ... ..... 55
4.2.1 Constrained Steering Algorithms ..... 56
4.2.2 Null Motion Algorithms ......................... 56 Local gradient (LG) ................. .. 56 Global avoidance/Preferred trajectory tracking 57 Generalized Inverse Steering Law (GISL) ... 58
4.3 Singularity Escape Algorithms ... 59 Singularity Robust (SR) inverse ... 59 Generalized Singularity Robust (GSR) inverse 60 Singular Direction Avoidance (SDA) ... 60 Feedback Steering Law (FSL) .. 62 Singularity Penetration with Unit-Delay (SPUD) 63
4.4 Singularity Avoidance and Escape Algorithms .. 64 Preferred gimbal angles . ... 64 Optimal steering law (OSL) ... 64
4.5 Other Steering Algorithms ................ ......... 66
4.6 Steering Algorithm Computation Comparison .. 66


5.1 Hybrid Steering Logic ................. .......... 68
5.1.1 Internal Singularity Metrics ... 68
5.1.2 Hybrid Steering Logic Formulation .. 69
5.2 Lyapunov Stability Analysis .......................... 71
5.3 Numerical Simulation ................. ...... .. .. 76
5.3.1 Case 1: At Zero Momentum Configuration 6 = [0 0 0 0]' deg 79 Local gradient simulation results ... 80 Singular Direction Avoidance simulation results 82 Hybrid Steering Logic simulation results ... 85
5.3.2 Case 2: Near Elliptic External Singularity 6 = [105 105 105 105]T
d e g . . 8 5 Local gradient simulation results ... 88 Singular Direction Avoidance simulation results 91 Hybrid Steering Logic simulation results ... 94
5.3.3 Case 3: Near Hyperbolic Internal Singularities J = [15 105 195
75]T deg . .. 96 Local gradient simulation results ... 96 Singular Direction Avoidance simulation results 99 Hybrid Steering Logic simulation results ... 102
5.4 Hybrid Steering Logic Summary ... ... ... 104


6.1 Attitude Controller with OTC ................ .......... 106
6.2 Lyapunov Stability Analysis ................ .......... 107
6.3 Num erical Sim ulation .. .. .. .. .. .. .. 110
6.3.1 Case : 60 = [0 0 0 0]T deg ....................... 111
6.3.2 Case II a: 6o = [90 90 90 90]T deg ... 116
6.3.3 Case II b (HSL/OTC): = [90 90 90 90]T deg 123
6.4 Orthogonal Torque Compensation Summary ... 127

7 SCALABILITY ISSUES FOR SGCMGS ........... ........... 129

7.1 Scalability Problems with SGCMG Hardware ..... 129
7.2 Effect of Igw on Torque Error ........... ... .. .......... 130
7.3 Num erical Sim ulation ............................. 132
7.3.1 Case l: Kg, = 0 . 134
7.3.2 Case Il: Kg = 2 ................. ........ 138
7.4 Effect of g,, on Torque Torque Amplification .... 142
7.5 Sum m ary . . 143

8 CO NCLUSIO N . . 144


ACTUATORS (SGCMG/VSCMG) ..........................147

A .1 Assum options . . 147
A .2 Dynam ics . . 147

B MOMENTUM ENVELOPE CODE ......................... 151


REFERENC ES . . .. 157

BIOGRAPHICAL SKETCH ................................ 164


Table page

3-1 M odel Param eters . .. 50

4-1 Algorithm Flops m = row(A) and n = column(A). 66

5-1 M odel Param eters . .. 78

5-2 Performance Comparisons for Case I: Zero Momentum ... 85

5-3 Performance Comparisons for Case II: Elliptic Singularity ... 96

5-4 Performance Comparisons for Case III: Hyperbolic Singularity 104

6-1 M odel Param eters . . 110

6-2 Hybrid Steering Logic Parameters ..... .. ...... 123

6-3 Performance Comparisons ............................. 128

7-1 M odel Param eters .. .. .. .. .. .. .. .. 133

7-2 M odel Param eters .. .. .. .. .. .. .. .. 142

C-1 Off-the-Shelf CMG Specifications ..... .. .... .. 156

Figure page

2-1 Rigid body with a constant c.m ............................ 21

2-2 Gimbal frame Tg) of IMPAC SGCMG (Patent Pending) ... 22

2-3 Singularity shown when CMG torque vectors lie in a plane (IMPAC SGCMGs
Patent Pending) ............. ...... ................ 25

2-4 Singularities for SGCMGs .............................. 27

2-5 External singular surfaces for a four-SGCMG pyramid ... 28

2-6 Internal singular surfaces for a four-SGCMG pyramid ..... 30

3-1 Four-SGCMG rooftop arrangement ......... ....... ........ 34

3-2 Four-SGCMG box arrangement ......... ....... .......... 35

3-3 Angular momentum envelope for a four-SGCMG box arrangement. ...... ..36

3-4 Planes of torque for a four-CMG rooftop arrangement ... 37

3-5 Torque planes traced out for a four-SGCMG rooftop arrangement ... 39

3-6 Angular momentum envelope with plotted angular momentum combinations
for the four-SGCMG box arrangement . .. 42

3-7 Degenerate hyperbolic singularities for the four-SGCMG box arrangement 43

3-8 Singular surfaces showing 1 ho singularity free region ..... 45

3-9 3 Orthogonal scissor pairs of SGCMGs . 46

3-10 Planes of angular momentum and torque for a four-SGCMG pyramid 47

3-11 Four-SGCMG pyramid arrangement ......... ........ ........ 47

3-12 Optimization process block diagram ..... ..... 49

3-13 Singular surfaces for the optimized arrangement at the Euler angles 0* =
[170.2 13.6 85.5 168.0]T deg and = [17.7 167.0 304.3 92.5]T deg 51

3-14 Gimbal rates for the optimized and pyramid arrangements ... 51

3-15 Torque error for the optimized and pyramid arrangements ... 52

3-16 Singularity measure for the optimized and pyramid arrangements ... 52

3-17 Optimization cost for the optimized and pyramid arrangements ... 53

4-1 Outer and inner loops of GNC system

4-2 Steering algorithms ........

5-1 Zero-momentum configuration of a

5-2 Simulation results for LG with aco =
momentum .............

5-3 Simulation results for LG with aco =
momentum (contd.) ........

5-4 Simulation results for SDA with co
zero momentum ..........

5-5 Simulation results for SDA with co
zero momentum (contd.) ......

5-6 Simulation results for HSL with aco
1 at zero momentum ........

5-7 Simulation results for HSL with aco
1 at zero momentum (contd.) .

5-8 Simulation results for LG with aco =
elliptic singularities .........

5-9 Simulation results for LG with aco =
elliptic singularities (contd.) .

5-10 Simulation results for SDA with aco
near elliptic singularities ......

5-11 Simulation results for SDA with aco
near elliptic singularities (contd.)

5-12 Simulation results for HSL with aco
1 near elliptic singularities .....

5-13 Simulation results for HSL with aco
1 near elliptic singularities (contd.)

5-14 Simulation results for LG with aco =
hyperbolic singularities .......

5-15 Simulation results for LG with aco =
hyperbolic singularities (contd.) .

5-16 Simulation results for SDA with ao
near hyperbolic singularities .

four-SGCMG pyramid arrangement ..

a = b = 1

a = b =

= 0.01,3o =

= 0.01,3o =

=0.01,3o = 2

=0.01,3o = 2

a = b =

a = b =

= 0.01,/3o =

= 0.01,/3o =

= 0.01, o =2

= 0.01,3o = 2

a = b =

a = b =

= 0.01,/3o =

= 0and 2 =

= 0and 2 =

S=b = 12 =

S=b = 12 =


, a= 1, b =

= 0 and 12

= 0 and 12

a =b =2 =

a= b= /2

_,a =l,b=

, a =, b=

= 0 and -12

= 0 and -12

0, a = 0, b

3o 1 at zero

3o = 1 at zero

0, and /1 = 1 at

0, and /1 = 1 at

3, and 11 = 12 =

3, and 11 = 12 =

= o = 1 near

= o = 1 near

0 and p = 1

S0, and 1 = 1

3, and p1 = 12 =

\, and 11 = 12 =

= o = 1 near

= o = 1 near

0, and p = 1




















5-17 Simulation results for SDA with co = 0.01, /o = 0, a = 0, b = 0, and = = 1
near hyperbolic singularities (contd.) ... 101

5-18 Simulation results for HSL with ao = 0.01, /o = 2, a = 1, b = 3, and p1 = 12 =
1 near hyperbolic singularities ..... ... 103

5-19 Simulation results for HSL with co = 0.01, /o = 2, a = 1, b = 3, and /i = 12 =
1 near hyperbolic singularities (contd.) ... 104

6-1 Satellite attitude control system block diagram ... 106

6-2 G im bal rates . ...... 112

6-3 O utput torque . ...... 113

6-4 Vector elements of the error quaternion ... 114

6-5 Singularity m measure . . 115

6-6 Singularity parameter (OTC) ............................ 116

6-7 Quaternion error difference: (A) eGSR eSDA (B) eSDA/OTC eSDA .. 116

6-8 G im bal rates . . ..... 118

6-9 O utput torque . ...... 119

6-10 Vector elements of the error quaternion ... 120

6-11 Singularity m measure . . 121

6-12 Gim bal-lock m measure ................................ 122

6-13 Singularity parameter (OTC) ............................ 123

6-14 G im bal rates . ...... 124

6-15 O utput torque .. .. .. .. .. .. .. .. ...... 125

6-16 Vector elements of the error quaternion ... 126

6-17 Singularity m measure . . 126

6-18 Gim bal-lock m measure ................................ 127

6-19 Singularity parameter (OTC) ............................ 127

7-1 Off-the-shelf CM G s . . 129

7-2 Gim bal rates for Kg, = 0 ...............................135

7-3 Gimbal accelerations for Kg = 0 .... .. .... .. 136


Torque error for Kg, =0 .. .

Singularity measure for Kg,

Gimbal rates for Kg, =2 ...

Gimbal accelerations for Kg,

Torque error for Kg = 2 .. .

Singularity measure for Kgw

. . 1 3 7

. . 1 3 8

. . 1 3 9

. . 1 4 0

. . 1 4 1

. 14 1







Abstract of Dissertation Presented to the Graduate School
of the University of Florida in Partial Fulfillment of the
Requirements for the Degree of Doctor of Philosophy


Frederick A. Leve

August 2010

Chair: Norman Fitz-Coy
Major: Aerospace Engineering

The research presented in this manuscript attempts to first systematically solve the

SGCMG steering and control problem. To accomplish this, a better understanding of

singularities associated with SGCMGs is required.

Next based on a better understanding of singularities, a Hybrid Steering Logic

(HSL) is developed and compared to the legacy methods singular direction avoidance

(SDA) and local gradient (LG) methods. The HSL is shown analytically and numerically

to outperform these legacy methods for a four-SGCMG pyramid arrangement in terms of

attitude tracking precision. However, all of the methods are susceptible to gimbal-lock.

A method referred to as Orthogonal Torque Compensation (OTC) is developed

for singularities with gimbal-lock in SGCMGs, which are known to present a challenge

to most steering algorithms. Orthogonal Torque Compensation conditions the attitude

control torque by adding torque error orthogonal to the singular direction when at

singularity. This method can be combined with any steering algorithm including HSL and

is proven analytically to be stable and escape singularities with gimbal-lock.

Finally, the problems with scaling systems of SGCMGs are discussed. It is

found that scaling produces an increase in the gimbal-wheel assembly inertia of the

SGCMGs which in turn increases the effect of the dynamics associated with these

inertias. Through analysis and numerical simulations, it is shown that more significant

gimbal-flywheel inertia reduces the performance by increasing torque error and

reducing torque amplification of SGCMGs. Since most of the legacy algorithms used

for singularity avoidance and escape use the gimbal rates for control, the performance

is degraded when the dynamics from the gimbal-wheel assembly inertia are increased.

This degraded performance is shown to also be dependent on the ratio of gimbal-wheel

assembly inertia to nominal SGCMG flywheel angular momentum. The overall result

here is that the same legacy steering algorithms that use gimbal rates for control cannot

be used for systems of SGCMGs of a reduced scale.


1.1 History and Background

1.1.1 Gyroscopic Rate Determination

The behavior of gyroscopic systems comes from the principle of conservation of

angular momentum. Both rate determination and rate generation are possible through

this gyroscopic behavior. For example, determination of attitude rates can be found

by mechanical gyroscopes. From these devices, spacecraft angular rates are inferred

from their reaction onto the gimbals of the mechanical gyroscope. The first known

gyroscopes were passive and the first of these was developed by Johann Bohnenberger

in 1817 [1, 2]. They were later developed for the navy as the naval gyrocompass [3].

There are now passive gyroscopes used for determining angular rates developed via

micro-electrical and mechanical systems (MEMS) that are smaller than the human eye

can detect.

1.1.2 Spin Stabilized Spacecraft

Early spacecraft did not have active attitude control but where spin stabilized about

their major axis of inertia by initiating a spin after launch. The major-axis rule is required

for directional stability (i.e., the spacecraft must spin about its major axis) [4]. This

provided gyroscopic stability or directional stability to the spacecraft which was defined

as stability provided the spacecraft does not have energy dissipation capability (i.e., a

rigid body). However, a spacecraft is truly not a rigid body (e.g., flexible booms, solar

arrays, internal movable parts, and outgassing) which may dissipate energy and become

asymptotically stable.

1.1.3 Spacecraft Attitude Control through Gyrostats

A gyrostat is any rigid body that has attached to it a wheel that through the

conservation of angular momentum provides either rotational stability or control. To

provide attitude stability, constant speed flywheels known as momentum wheels (MWs)

were added internal to the spacecraft to provide gyroscopic stiffness which in turn

supplies attitude stability [5]. This system is an example of a gyrostat. The gyrostat was

the first fundamental dynamical system that considered a spinning flywheel within or

attached to a rigid body. In this type of system the flywheel imparts angular momentum

stiffness to the body internally through the principle of the conservation of angular

momentum. The use of momentum wheels for gyroscopic stability has found its use for

other vehicles than spacecraft (e.g., boats, trains, buses [6]).

1.1.4 3-axis Attitude Control of Spacecraft

When active attitude control was needed, momentum wheels were exchanged

for reaction wheels (RWs) which provide a reaction torque on the spacecraft through

flywheel accelerations. These devices were then used in combination with thrusters to

provide attitude control to spacecraft [7]. Further innovation came to spacecraft attitude

control through the introduction of gimbaled MWs known as control moment gyroscopes

(CMGs). They have been used in satellite attitude control for decades due to their

high precision and property of torque amplification (i.e., larger torque output onto the

spacecraft than the input torque from the gimbal motors). Typically, they have been used

for large satellites that require high agility while maintaining pointing precision and have

even found their uses onboard the international space station (ISS) [8]. Control moment

gyroscopes produce a gyroscopic torque through rotation of angular momentum

about one or two gimbal axes. Momentum wheels, RWs and CMGs all are known as

momentum exchange devices because their torque is produced through redistribution

of angular momentum from the CMGs to the spacecraft. Control moment gyroscopes

come in two classes: 1) those with a single controllable degree of freedom and 2) those

with multiple controllable degrees of freedom.

The choice of a CMG or attitude actuator in general depends on the needs

of the mission. All CMGs have specific challenges associated with their use. The

challenges of CMGs are as follows: single DOF CMGs known as single-gimbal control

moment gyroscopes (SGCMGs) suffer from instantaneous internal singularities (i.e.,

situations within the performance envelope where a given torque cannot be produced);

multiple DOF CMGs known double-gimbal control moment gyroscopes (DGCMGs)

are mechanically complex and have singularities known as gimbal-lock when their

gimbals align thus eliminating their extra DOF; other multiple DOF CMGs known as

variable-speed control moment gyroscopes (VSCMGs) are difficult to mitigate induced

vibration (i.e., due to the variation of flywheel speeds) and require more complicated

control laws, motor driver circuitry, and larger flywheel motors.

1.1.5 Single-Gimbal Control Moment Gyroscopes (SGCMGs)

Single-gimbal control moment gyroscopes have a long heritage of flight on larger

satellites and the ISS [9-15]. They are known to have the highest torque amplification of

all CMGs, are less mechanically complex than DGCMGs and have less mathematically

complex dynamics than VSCMGs. These actuators suffer from internal singularities that

must be handled on-the-fly, where torque cannot be generated in a specific direction.

There is no single method that has been proven to avoid all internal singularities while

tracking an arbitrary torque perfectly (i.e., without the use of torque error or constraining

the torque). Thus, there is merit in finding alternate solutions to control of SGCMGs for

attitude control.

1.1.6 Double-Gimbal Control Moment Gyroscopes (DGCMGs)

Double-gimbal control moment gyroscopes contain two controllable DOFs through

their two gimbal axes. They are the most mechanically complex of CMGs although, the

redundancy in the additional gimbal may lead to less than three DGCMGs required for

3-axis control. The benefit of this redundancy is lost for DGCMGs when they encounter

a gimbal-lock singularity. Gimbal-lock is encountered when the gimbal axes are aligned

and are no longer linearly independent. As a consequence, the extra controllable DOF

in this case is lost. Effective methods already exist for avoiding gimbal-lock singularities

associated with DGCMGs [16-20].

1.1.7 Variable-Speed Control Moment Gyroscopes (VSCMGs)

Variable-speed control moment gyroscopes utilize an extra controllable DOF

through flywheel accelerations. As a consequence of the extra DOF, the flywheel motors

must be larger and it is more troublesome to isolate unwanted induced vibration. In

addition, this extra degree of freedom makes a system of two or more non-collinear

VSCMGs free from singularities through the extra degree of freedom, (i.e., at CMG

internal singularities, the needed torque is provided by flywheel accelerations). Several

algorithms have been developed that are effective in reducing the amount of flywheel

accelerations used and thus providing better torque amplification [21-30]. In addition,

methods have been developed to use the VSCMG's extra controllable DOF to spin down

the flywheels and store their kinetic energy. These methods are known as the integrated

power and attitude control system (IPACS) and the flywheel attitude control and energy

transmission system (FACETS) in literature [31-34].

1.2 Problem Statement

A new paradigm that requires highly agile small spacecraft is ongoing through

efforts by government agencies and labs such as Operationally Responsive Space

(ORS), Air Force Research Laboratory (AFRL), the National Reconnaissance Organization

(NRO), and National Aeronautics and Space Administration (NASA) to perform such

missions as intelligence, surveillance and reconnaissance (ISR), space situational

awareness (SSA), and space science missions (e.g., the imaging of gamma ray bursts)

[35-37]. Many of these missions are in LEO and require higher agility and attitude

precision to track targets on earth. Attitude control systems (ACSs) based on reaction

control devices (e.g., thrusters) can achieve great agility but cannot meet the pointing

requirements and space needed for propellant storage on small satellites [38].

Single-gimbal control moment gyroscopes are being considered the actuator of

choice to provide higher agility to smaller satellites based on their perceived torque

amplification. However, many problems exist that must be solved prior to using

SGCMGs for small satellite attitude control. Traditionally, SGCMGs have been oversized

for their mission and the angular momentum envelope constrained to avoid internal

singularities on larger satellites. For smaller satellite systems, however, the extra volume

and mass needed for oversized SGCMGs may be unacceptable. Therefore, small

satellite SGCMGs should utilize more of the entire angular momentum envelope where

singularities may be encountered in the momentum space (both internal and external).

Thus, legacy steering algorithms from larger satellite applications may not provide the

same performance for systems of a smaller scale requiring a new approach for steering

and control of SGCMGs.

For the succeeding chapters: Chapter 2 discusses and reviews the fundamentals

of CMG dynamics and describes the different forms of singularities associated with

SGCMGs; Chapter 3 describes the possible arrangements for systems of SGCMGs

and their desirable and undesirable qualities; Chapter 4 provides the background on

previously published methods of steering algorithms for implementation of SGCMGs;

Chapter 5 discusses the development of the Hybrid Steering Logic (HSL) for SGCMGs;

Chapter 6 discusses the development of Orthogonal-Torque-Compensation (OTC) for

gimbal-lock escape of SGCMGs; Chapter 7 discusses the performance degradation

encountered when scaling systems of SGCMGs; and Chapter 8 provides conclusions of

the research.


2.1 Dynamic Formulation

The dynamic formulation presented in this section addresses momentum exchange

devices for attitude control systems where the angular momentum of the spacecraft

system (i.e., the spacecraft and system of CMGs) is assumed constant. This is in

contrast to reaction control devices (e.g., thrusters) and/or energy dissipation devices

(e.g., magnet torquers) which change the angular momentum or energy of the system.

It is assumed that the center of mass (c.m.) of each CMG lies along the gimbal axis.

This is equivalent to stating that the rotation of the gimbal-flywheel system about the

gimbal axis does not move the position of the c.m. of the system. It is also assumed

that the spacecraft-CMG system is a rigid body; and this system is absent of friction and

external torques. For a coordinateless derivation of the dynamics see Appendix A. The

first assumption can be visualized by treating the CMGs as cylinders and having their

spin axis along their c.m. as shown in Figure 2-1.

The total centroidal angular momentum of the spacecraft-CMG system in the

spacecraft body frame is

H = Jw + h (2-1)

which is composed of the spacecraft centroidal angular momentum Jc, composed of

the spacecraft centroidal inertia J, and angular velocity w and the angular momentum h

contributed from the CMGs.

Considering the spacecraft modeled as a rigid body, its centroidal inertia is

composed of both constant and time-varying inertias and is expressed as

J, = J + mi(rril rir,) + CBG BG (2-2)


Figure 2-1. Rigid body with a constant c.m.

where J represents the constant spacecraft-CMG system inertia in referencing position
about the spacecraft's c.m; mi(rTril rrT) are the parallel axis terms, with mass mi
of the ith gimbal-flywheel assembly and its c.m., position ri with respect to the c.m.
of the spacecraft, and CBG IgwCG, are the time varying inertias from rotation of the
gimbal-flywheel system inertia Igw. The angular momentum contributed from the ith
CMG in a gimbal frame .g) is expressed as

hi= 0 (2-3)

which consists of angular momentum from the flywheel (I/Q) and that from the
gimbal-wheel system (IgwB). It should be noted that the CMG angular momentum
expression in Eq.(2-3) is based on an SGCMG or VSCMG only and the following
development of angular momentum for CMGs will be for a multiple gimbal CMG. The

resultant angular momentum from the CMG system in the body frame is found through

the summation of the contributions of angular momentum from all CMGs rotated from

their respective gimbal frames into the spacecraft body frame; i.e.,

h = CBGhi


where n is the number of SGCMGs, CBG, is the direction cosine matrix (DCM) from the

gimbal frame Fg) shown in Figure 2-2 to spacecraft body frame F3.


Figure 2-2. Gimbal frame T) of IMPAC SGCMG (Patent Pending)

With the assumption of no external torques and frictionless devices, the total

angular momentum is constant, thus the inertial time derivative of Eq.(2-4) shows the

redistribution of the systems's angular momentum (i.e., on the mechanism by which the

torque is produced by CMGs). Differentiation of Eq.(2-1) yields

dH d(Jcj + h)
dt dt

X (Jxc + h) = 0


The time derivative of the spacecraft angular momentum yields

-C = Jc + (2-6)
The spacecraft inertia is assumed to only vary by the gimbal angles of the CMGs thus

making the second term on the right hand side of Eq.(2-6)

S a(Jw) d6 k
JW 96ASj (2-7)
j=1 j=1
where A cE R3Xn is the Jacobian matrix resulting from the coupling of the spacecraft and

CMG kinematics from the jth gimbal of a multiple gimbal CMG and n is the number of


The angular momentum of the CMG system is a function of the flywheel angular

velocities Q e Rx I gimbal angles 6j e Rnxl, and gimbal rates 6j e Rnxl, respectively.

The time derivative of the CMGs angular momentum can be expressed as

dh Oh d6 9h d6 j h dl
= h = 6 +) + (2-8)
dt Y. dt 6 dt 96 dt 2dt
where the individual Jacobian matrices are,

A eI Rs n (2-9)

BJ = E R3X" (2-10)

C= E R 3xn (2-11)

Assuming a constant flywheel speed f = 0 and a single gimbal (k= 1) configuration,

then Eq.(2-5) can be rewritten as

S= Jc + wxJc + wh +DX = 0 (2-12)


DX= [(A + A2) B] =h +J = T (2-13)

The general equations for the SGCMG output torque in terms of a given internal

control torque 7 is expressed as,

DX = -- wX>h = T (2-14)

where T is the total torque output from the system of SGCMGs. It should be noted that 6

and 6 are kinematically coupled, thus it is not possible to find both states simultaneously

when mapping D e R3s"s3 onto T (i.e., only one gimbal state can be chosen as a control

variable). For SGCMG systems that contain significant flywheel angular momentum

and gyroscopic torque, the dynamics of the gimbal-wheel assembly inertias can be

considered insignificant (i.e., A1 z 0 and B z 0). For such systems, it is customary

to neglect the inertia variations due to the gimbal motion (i.e., Jc = 0) resulting in the

composite D reducing to A2 e Ra3n. Therefore, the Jacobian D is simply A2 and the

solution of the output torque from the SGCMGs is contributed solely from the gimbal

rates as

h = hoA26 = ho[il, 2, ...rn]6 (2-15)

where A2 = hoA and -, is the torque vector direction of the ith CMG as shown in

Figure 2-2. The coefficient matrix in Eq.(2-15) is 3 x n and when n > 3 the system

is over-actuated. When this matrix becomes rank deficient, the system is said to be

singular. Physically, when these singularities occur, the torque vector directions of each

SGCMG in the body frame lie in a plane as shown in Figure 2-3.

For convenience, from this point until Chapter 7, it is assumed A2 = A.

Figure 2-3. Singularity shown when CMG torque vectors lie in a plane (IMPAC SGCMGs
Patent Pending)

2.2 Singular Surface Equations

It is customary to define an orthonormal basis {hi, ,, } as shown in Figure 2-2
where ii is the spin axis of the flywheel, f, is the SGCMG torque direction, and 6, is the
gimbal axis direction.
Therefore, the singular direction s c R3X1 is defined from

{s e R3 : sTi = 0} (2-16)

This constraint constitutes a maximum (or minimum) projection of fi onto s. There is a
fundamental assumption that ho, is equal to ho (i.e., the magnitude of nominal angular
momentum is the same for all SGCMGs in the system). For a given singular direction
s # 6, (i.e., which only occurs for DGCMGs and for rooftop arrangements), the
conditions for singularity are

sT-i = 0 and sThi 0 (2-17)

If we define A sTfi, then the torque and spin axis directions can be expressed as

I x s
T = s f 6i, i = 1,...,n (2-18)
113, x s||

h X (8, X s )8i, =1, ..., n (2-19)
II6, x s|l
Combining Eqs.(2-18) and (2-19), the total normalized angular momentum from the
SGCMGs is expressed as

n n x S) X S
S s x ,, = (2-20)
1i=6 i=x s
It is important to note that when s = 6i Eqs.(2-18)-(2-20) are indeterminate.
The locus of total normalized angular momentum h from Eq.(2-20) for all s e RR3

and all e, 0 (i.e., s not collinear to 6,) produces the external singular surface known
as the angular momentum envelope shown in Figure 2-5 for a four-SGCMG pyramid
arrangement. Similarly, each of the four internal singular surfaces shown in Figure 2-6
for a four-SGCMG pyramid arrangement are found by setting one of the ce < 0. Matlab

code for both of these surfaces can be seen in Appendix C. Singularities for SGCMGs
can be classified into the groups/subgroups as shown in Figure 2-4.
2.2.1 Elliptic Singularities

Elliptic singularities are those in which null solutions to the gimbal angles do not
exist for a specific point of CMG angular momentum space. Null motion is a continuous
set of null solutions for gimbal angles (i.e., there is a continuous transfer from one null
solution to the next) that does not change the CMG's angular momentum and thus, does
not produce any motion to the spacecraft. Since elliptic singularities do not have null
solutions, the angular momentum must be perturbed thus inducing torque error to the
system to escape from these singularities. Elliptic singularities are not limited to internal
singularities; (e.g., all external singularities are elliptic and hence cannot be avoided or
escaped through null motion).

I External Internal Internal

Gimbal-Lock Gimbal-Lock I )>n>Iil. Non-degenerate

Gimbal-Lock Gimbal-Lock

Figure 2-4. Singularities for SGCMGs External singularities

External singularities also known as saturation singularities are associated

with the maximum projection of CMG angular momentum in any direction. These

singularities cannot be avoided by null motion and therefore by definition are elliptic.

These singularities occur on the surface of the angular momentum envelope and an
example of this surface for a four-SGCMG pyramid is showed in Figure 2-5. When

these singularities are encountered, the CMGs are unable to produce any more angular
momentum in the saturated direction. External singularities are addressed a priori in the

design process through sizing of the CMG actuators.

1., -- ,-' '- ,,. .. "" |

20 4

Y(ho) X(h-)

Figure 2-5. External singular surfaces for a four-SGCMG pyramid

Consider a four-SGCMG pyramid arrangement, Eq.(2-21) can be used to express

the angular momentum as

-c(0)s(1) c(2) + c(O)s(6) + c(4)
h = ho c(1) c(0)s(62) c(63)+ c()s(64) (2-21)

s(O)(s() + s(2) +s(63) +s(6))

where 0 is the skew angle and 6, is the ith gimbal angle. Further consider the set of

gimbal angles 5es = [90 90 90 90] deg, then the momentum vector becomes

0 0
h(6es)= 0 0 (2-22)

hos(O)(s(61) + s(2) + s() + s(64)) 4hos()

It is clear that there is only one set of gimbal angles 6es = [90 90 90 90] deg that will

give the angular momentum in Eq.(2-22). Therefore, null solutions do not exist, and this

angular momentum vector corresponds to the elliptic saturation singularity along the

z-axis. Elliptic internal singularities

Elliptic singularities which lie on the internal singular surfaces such as that shown

for the four-SGCMG pyramid arrangement in Figure 2-6 are referred to as elliptic internal

singularities. Unlike external singularities, these singularities cannot be accounted for

in the design process; furthermore, since they occur instantaneously, they cannot be

generally avoided.

2.2.2 Hyperbolic Singularities

Hyberbolic singularities are those in which null motion is possible. Thus, all

hyperbolic singularities are therefore internal (i.e., these singularities occur on the

internal singular surfaces). The points on the internal singular surface corresponding

to a hyperbolic singularity have null solutions of gimbal angles, corresponding to the

null space of the Jacobian matrix. The null solutions are typically chosen to avoid

the singular configurations of the system. Shown in Figure 2-6, is an example of this

surface for a four-SGCMG pyramid. Singularities occur only when the point on this

surface corresponds to a singular Jacobian matrix (i.e., there may be nonsingular sets

of gimbal angles at this point on the surface). When these singularities are encountered,

the SGCMG torque vector directions lie in a plane and as a consequence there is

no torque available out of the plane. These singularities, like elliptic singularities, are

instantaneous and must be handled on the fly. For a four-SGCMG pyramid arrangement

with angular momentum in Eq.(2-21), a set of gimbal angles 6hs = [180 90 0 90]T deg is

an hyperbolic singularity that has the following momentum vector,

h(hs) = -2ho (2-23)


It is clear that there multiple solutions (i.e., null solutions) to Eq.(2-23). A null solution of

the gimbal angles satisfying 61 = 63 = 90 o and s(64) = -(62) = substituted into

Eq.(2-21) will also satisfy Eq.(2-23).


0 -.
,- 0 -
-1 -- "
-2 -
-3- -



-5 -2

Figure 2-6. Internal singular surfaces for a four-SGCMG pyramid Non-degenerate hyperbolic singularities

Non-Degenerate hyperbolic internal singularities are those in which null motion

is possible and some of the null solutions are nonsingular providing the possibility of

singularity avoidance. Degenerate hyperbolic singularities

Degenerate hyperbolic internal singularities occur when the null solutions to gimbal

angles correspond to singular sets of gimbal angles leaving no room for avoidance or

escape. These singularities are also considered impassable and therefore are handled

in a similar manner to elliptic singularities when approached.

2.2.3 Gimbal-Lock

Gimbal-lock for SGCMGs occurs at singularity when the mapped output torque

vector is in the singular direction. When this occurs, the system becomes trapped in this

; ~ar

singular configuration with only a few methods that are capable of escape from it. One

such method is known as the Generalized Singularity Robust (GSR) Inverse [39, 40].

This method has been shown numerically to escape gimbal-lock of SGCMGs but not

analytically and there is no formal proof to suggest that it is always successful.
2.3 Singularities for SGCMGs Mathematically Defined

To quantifiy the effectiveness of avoiding internal singularities through null motion,

we must define their forms (i.e., hyperbolic and elliptic) mathematically. Typically,

topology and differential geometry are used to represent hyperbolic and elliptic
internal singularities as surfaces or manifolds [41, 42]. The behavior of these internal

singularities can also be explained through the use of linear algebra [43]. To accomplish

this, a Taylor series expansion of the SGCMG angular momentum about a singular
configuration gives

h(6) -h(65) = [ A 6 ,2 A+ H. O.T.1 (2-24)
i= 1
where h(65) is the angular momentum at a singular set of gimbal angles 6s, A6j

6, 6s, and n is the number of SGCMGs in the system.

The first term on the right-hand side (RHS) of Eq.(2-24) contains the ith column

of the Jacobian matrix f, = associated with the ith SGCMG's torque direction.

The second term on the RHS contains the partial derivative of the Jacobian matrix's

ith column with respect to the ith gimbal angle ,i s. Furthermore, from Eq.(2-15), the
RHS of Eq.(2-24) can be transformed through the realization of the following operations:

a2hi -?91
-h-;=i = -hi (2-25)

where () denotes a unit vector. Next, Eq.(2-25) is substituted into Eq.(2-24) and the

inner product of the result with the singular direction s obtained from null(AT) yields

sT[h() h(6s)] 1= hs Ab (2-26)
The first term on the RHS of Eq.(2-24) has zero contribution because of the definition of

the singular direction (i.e, ATs = 0). Equation (2-26) can be written more compactly as

sT[h(6) h(6s)] = -ATPA6 (2-27)
where P is the singularity projection matrix defined as P = diag(h's).

By definition, null motion does not affect the total system angular momentum and

which equates to h(6) = h(6s). Consequently, the left-hand side (LHS) of Eq.(2-27)

is zero (i.e., A6TPA6 = 0). Null motion is expressed in terms of the basis N = null(A)

concatenated in matrix form as follows

A6= Aivi = NA (2-28)
where A is a column matrix of the scaling components of the null space basis vectors vi

and N e Rnx(n-r(A)) is the dimension of the null space basis for any system of SGCMGs

with r(A) = rank(A). Substituting Eq.(2-28) into Eq.(2-27) observing the null motion

constraint yields

ATQA = 0 (2-29)

As a result of this analysis, a matrix Q is defined as

Q = NTPN (2-30)

Therefore, when away from singularity Q e RIi; when at a rank 2 singularity, Q e R2x2;

and when at a rank 1 singularity, Q e RS3x. The eigenvalues of the Q matrix determine

whether a singularity is hyperbolic or elliptic. If Q is definite (i.e., has all positive or

negative eigenvalues), it does not contain a null space since a nonzero null vector A

does not exist that satisfies Eq.(2-29) [44]. Therefore, situations where the matrix Q is
definite constitute elliptic singularities.

When Q is semi-definite (i.e., it has at least one zero eigenvalue), then a null space

exists since there exists a A 7 0 that satisfies Eq.(2-29). Therefore, null motion is

possible near singularity and the possibility of singularity avoidance may hold (i.e.,

does not for degenerate-hyperbolic singularities) [43]. If the matrix Q is indefinite (i.e.,

the eigenvalues are positive and negative), the result of Eq.(2-29) has the possibility

of being equal to zero. Therefore, both of these situations constitute an hyperbolic


The tools developed in this chapter for describing the existence of elliptic singularities

and hyperbolic internal singularities in a system of SGCMGs are used in the next

chapter to more specifically discuss which of these internal singularities exist in common

arrangements of SGCMGs. In addition, Chapter 5 introduces a novel steering algorithm

known as the Hybrid Steering Logic (HSL) which uses these tools in its derivation.


3.1 Common SGCMG Arrangements

Several common SGCMG arrangements have been studied. Typically, the factors

that determine the choice of a specific SGCMG arrangement are: (i) available volume (ii)

desirable angular momentum envelope, and (iii) associated singularities. In this chapter,

we examine the common arrangements and use the tools developed in Chapter 2 to

characterize their singularities.

3.1.1 Rooftop

The rooftop arrangement shown in Figure 3-1 has two sets of parallel SGCMGs,

each with parallel gimbal axes where 0 is the skew angle relating the planes of torque. A

four-SGCMG rooftop arrangement is shown in Figure 3-1.

X B y

Figure 3-1. Four-SGCMG rooftop arrangement

Since these arrangements are free from elliptic internal singularities, they have a

significant flight heritage on satellites and thus, their control is well understood [45].

However, degenerate hyperbolic singularities which are also impassable still exist and

like elliptic singularities cannot be addressed through the use of null motion. In addition,

there are degenerate cases of hyperbolic singularities for the rooftop arrangement when

the Jacobian matrix is rank 1 which may provide difficulty to singularity escape laws that
regulate the smallest singular value.
3.1.2 Box

The box arrangement is a subset of the rooftop arrangement and has two parallel
sets of two SGCMGs with an angle of 90 deg between the two planes of torque as
shown in Figure 3-2. This arrangement is given its name because the planes of angular
momentum can form a box [46]. Like the rooftop array, there are situations where this
arrangement may have a rank 1 Jacobian.



Figure 3-2. Four-SGCMG box arrangement

The angular momentum envelope for the four-SGCMG box arrangement is shown
in Figure 3-3. The angular momentum envelope for all rooftop arrangements is an
ellipsoidial surface and thus, there is not equal momentum saturation in all directions.
Rooftop arrangements are chosen for their compactness and the fact that they are
free from elliptic internal singularities. Analysis provided in the literature proves that
though these arrangements are free from elliptic internal singularities. However, they are
not free from the impassable degenerate hyperbolic singularities [47].

-1 -
-2,, : ..- '.== L,- : '

-2-2 2
% 1 0*
Y(ho) -4 -2 1

Figure 3-3. Angular momentum envelope for a four-SGCMG box arrangement.

In Chapter 2, it was shown that the definiteness of the matrix Q determines if a

system of SGCMGs is at an elliptic singularity. For a system of four SGCMGs, the

largest Q can be is RR2x2 excluding the case when the Jacobian goes rank 1 which will

be discussed later. Therefore, excluding a rank 1 Jacobian, and if the determinant of

Q is strictly positive, then the system is at an elliptic singularity (e.g., AIA2 > 0 where

A1 2 are an eigenvalues of Q). Consequently, when det(Q) < 0, the system is at an

hyperbolic singularity. There are a few general cases where singularities may occur for a

four-SGCMG rooftop arrangement. The first case occurs when the torque vectors lying

in the same plane of torque are either parallel or anti-parallel as shown in Figures 3-4 A

and B where 6 62 + 1800, 63 64 1800 and r is the axis intersecting the two planes of


/l W / W
A Torque planes with parallel torque B Torque planes with one parallel and
vectors anti-parallel torque vectors
Figure 3-4. Planes of torque for a four-CMG rooftop arrangement

A four CMG rooftop system in the configuration of Figure 3-4 A has a Jacobian

c(O)c(61) c(O)c(6G) -c(O)c(3) -c(O)c(63)
A = s(61) s(61) -s(63) -s(63) (3-1)

s(O)c((6) s(O)c(61) s(0)c(63) s(0)c(63)
with angular momentum represented in the spacecraft-body frame (i.e., where the
gimbals are enumerated counter-clockwise beginning at the spacecraft body x-axis)

c(0)s(61) -c(0)s(63)
h = h + h2+ h3 h4= 2ho -c(61) 2ho c(63) (3-2)

s(0)s(61) s(0)s(63)
The singular direction for this case is found by cross product of ^1 and ^3

s()s(01 + 63)
= -2s(0)c(0)c(61)c(63) (3-3)

c(0)s(1 63)
with the resultant projection matrix

c(61) 0
0 c(6i)
P = 2hos(O)c(O) 0 )
0 0
0 0
and the null-space of the Jacobian concatenated in

-1 0
1 0

0 0
0 0

-c(63) 0
0 -c(3)
matrix form is



The configuration in Figure 3-4 has the det(Q) = -16s2(O)c2(O)c(61)c(3). It should

be noted that when P and Q are definite (i.e., e = [+ + ++] or c = [- --] where

ci = sgn(hfj s)), the system is at a external singularity. It was discussed in Chapter 2 that
external (or saturation) singularities are elliptic. Therefore, if the matrix P is definite, then

the matrix Q is also definite. The det(Q) = -16s2(0)c2(0)c( )c((3) will only be positive

when sgn(c(l6)) z sgn(c(O3)) (i.e., at saturation singularity). The saturation singularity

is not an internal singularity and therefore neither the rooftop and box arrangements

contain elliptic internal singularities for this case. It can be verified that the variations

of these cases such as (f1 = -t2), (f-2 = -4), (f1 = -2) and (f2 = 4) all have

det(Q) = 0 and therefore are hyperbolic internal singularities.

The other case when the Jacobian of a four-CMG rooftop is singular occurs when

the torque vectors of the two parallel SGCMGs lie in the direction of the intersecting

torque planes r shown in Figures 3-5.

r r

A Torque planes with parallel torque B Torque planes with anti-parallel torque
vectors along r vectors along r

Figure 3-5. Torque planes traced out for a four-SGCMG rooftop arrangement

For Figure 3-5 A, the Jacobian is

A = -S(63)

with angular momentum vector

-(0)C(64) 0
-s(64) 1

s(0)C(64) 0

c(0) -c(0)s(63)
h = hi h2 h3 h4= 2ho 0 ho c(63)

s()_ s(O)s(63)

assuming that the intersection of the planes of torque r = [0

shown in Figure 3-5. For this situation, the singular direction

intersection of the planes and is found to be


Sho c(64) (3-7)


1 O]T for this arrangement

is orthogonal to the

s= 0


with the resulting projection matrix


1 0

0 1
P = 2hos(0)c(0)c(63)
0 0

0 0
and the null space basis for the Jacobian concatenated

N = null(A)

0 0

0 0

0 0

0 0
in matrix form is


1 s(64 3)/C(3)


) -(C(64)/C63))


For this case, det(Q) = 2hosin2(20) > 0 which is an elliptic singularity but is not known
yet to be one that is internal or external. For the anti-parallel case when 63 = 180 + 64 =
2700 deg, the det(Q) = 2hosin2(20) < 0 and the singularity is hyperbolic. Notice that the

diagonal entries of P that are zero correspond to the gimbal axis of that roof-side being
along the singular direction s of the system.

Recall, from Chapter 2 that s = Si is a special case that happens only for DGCMGs
and rooftop arrangements. For example, consider the case when 63 = 180 + 64 = 2700

and an angular momentum of

h= h2 (3-11)


is desired along with, 61 = -62. The result of Eq.(3-11) is c(61) = c(2) = 2. There

are two solutions for any possible value of h2 inside the momentum envelope due to

symmetry and thus there is one null solution. Both of these solutions are singular and
abide by 63 = 1800 + 64 = 2700 and therefore the null solutions that exist do not help in

escape from the singularity. Thus, this is a case of a degenerate hyperbolic singularity at
that specific point on the momentum space. In addition, the value of d = 0 (i.e., where

m = det(AAT)) for both 63 = 1800 + 4 = 2700 and for 61 and 62 free, and thus there is

no set of gimbal angles that will provide a change in m (i.e., no null solutions to escape

To determine if the other cases of singularity when 63 = 64 = 900 deg are elliptic,
we check if null motion exists orthogonal to the singularity (i.e., d 6 0). The result of

this case when 63 = 64 = 900 and 61 and 62 are free is consistent with Eq.(3-6). It is
found that for any choice of 61 and 62 gives d = 0 and therefore this is a family of elliptic

singularities because not only does det(Q) > 0 but also d = 0. To visualize where the

external singularities occur for this case, the SGCMG angular momentum is plotted for

all variations of 61 and 62 of a four-SGCMG box arrangement shown in Figure 3-6.


.-- '
-- ,----

-3 '


-2- 3

Figure 3-6. Angular momentum envelope with plotted angular momentum combinations
for the four-SGCMG box arrangement

In Figure 3-6, all possible combinations of angular momentum are plotted onto

the angular momentum envelope in black for 61 and 62 and when 63 = 64 = 90

In this figure, every combination of this situation is an external singularity. It can be

shown from symmetry that all permutations of this case have the same result in that

they are external singularities. Therefore, rooftop arrangements do not contain elliptic

internal singularities. In addition, the case when the Jacobian for a four-SGCMG rooftop

arrangement approaches rank 1 is a subset of this family. This family of configurations is

defined as the cases where at least two torque vectors are parallel along the intersection

of the two torque planes r and the outcome of the det(Q) is not dependent on the gimbal

angles. The rank 1 Jacobian case is a degenerate hyperbolic singularity for when the

torque vectors are all anti-parallel (i.e., 0 or 2ho) and a external singularity when all the

angular momentum vectors are parallel (i.e., 4 ho).

These results confirm those obtained via topology and differential geometry by

Kurokawa [41]. In reference [41], it was stated that any rooftop arrangements with

no less than six units are free from internal impassable surfaces (i.e., elliptic internal

singularities not including external singularities). Kurokawa concluded that there are

impassible internal surfaces in the four-SGCMG rooftop arrangements that correspond

to the singular direction s not contained in the plane spanned by the two gimbal axes 81

and 62. This is exactly the degenerate case shown in Eqs.(3-6)-(3-9) where s = 31. The

degenerate hyperbolic singularities for these arrangements lie on two circles with radius

2ho when 61 = 62 and 63 = 1800 + 64 = 2700 and at zero momentum Oho when 61 = -62

and 63 = 1800 + 4 = 2700 shown as an example for the box arrangement in Figure 3-7




0 2

X(ho) 2 2Y(ho)

Figure 3-7. Degenerate hyperbolic singularities for the four-SGCMG box arrangement

The curves and point in Figure 3-7 are compartmentalized and not spread

throughout the entire angular momentum envelope unlike elliptic singularities and thus,

constrained steering algorithms exist to avoid these regions while providing singularity

avoidance using null motion.

3.1.3 3 Box

The 3 box arrangement is a subset of the box arrangement in which one of the

SGCMGs is not used and left as a spare. This arrangement has the longest heritage

of flight due to its conservative nature. For this arrangement, a pseudo-inverse is not

required to obtain a solution to the gimbal rates since the gimbal rates are found directly

from the inverse of a 3 x 3 Jacobian matrix [48]. The combination of this arrangement

and the constrained angular momentum steering law limiting controllable SGCMG

angular momentum to a 1 ho radius of the angular momentum envelope, ensure the

safest control of a system of SGCMGs [49]. This design although safe, may not be

practical for small satellite applications because the SGCMGs must be oversized

to provide the desired performance (i.e., they do not take advantage of the entire

momentum envelope).



2 3 -1 5

Y(ho) -3 -3 X(h0) -2 -1
Y(ho) X(ho)
A External singular surface for 3 box B Internal singular surface 3 box

Figure 3-8. Singular surfaces showing 1 ho singularity free region

In this arrangement, there is a 1 ho radius of the angular momentum envelope

that is guaranteed to be singularity free [48] which is shown by the red circle drawn

on the external and internal singular surfaces of Figure 3-8. It should be mentioned

that because there is no longer a null space (i.e., using only 3 of the 4 SGCMGs),

any angular momentum point on the singular surfaces correspond to the location of a

elliptic singularity. This makes the constraining to the 1 ho sphere of angular momentum


3.1.4 Scissor Pair

The scissor-pair arrangement has three sets of collinear pairs of two SGCMGs

orthogonal to each other. This arrangement is constrained to have 61 = -62 at all times

for both CMGs to avoid internal singularities. With this arrangement shown in Figure 3-9,

full three axis control is possible with a full rank Jacobian matrix as long as it does not

extend past the maximum angular momentum of the system. As a consequence of the

constraint for these pairs, only one-third of the entire angular momentum envelope is

utilized which will be troublesome for use on small satellites (i.e., six SGCMGs needed

for 2ho of angular momentum). Use of these arrangements was found to conserve power

when a single gimbal motor is used for each scissor pair [50]. Also, analysis has shown

that scissor pairs may be beneficial for space robotics application since their torque is

unidirectional [51].


Figure 3-9. 3 Orthogonal scissor pairs of SGCMGs

Due to the gimbal angle constraint associated with this arrangement, internal

singularities do not exist here. Also because of the gimbal angle constraint 61 = -62,

three orthogonal scissor-pairs contain only external singularities that occur when

one or more of the pairs has an overall zero torque vector (i.e., undefined torque for

scissor pair). When this occurs, the Jacobian matrix contains a column of zeros for the

associated pair.

3.1.5 Pyramid

Pyramid arrangements of SGCMGs have independent planes of angular momentum

and torque which form a pyramid. As a consequence of these independent planes of

torque, these arrangements will never have the Jacobian matrix with rank less than 2.

This is shown in Figure 3-10 for a four-SGCMG pyramid cluster.

In terms of small satellite constraints and when utilizing the entire momentum

envelope, the four SGCMG arrangement seems practical among all previously

discussed arrangements for platforms requiring high torque and slew rates with near

equal momentum saturation in three directions .

Figure 3-10. Planes of angular momentum and torque for a four-SGCMG pyramid

Figure 3-11. Four-SGCMG pyramid arrangement

Control of these arrangements is more complicated than rooftop and box arrangements

due to the presence of elliptic internal singularities because null motion solutions do

not exist. In addition, elliptic singularities do not have continuous gimbal trajectories

associated the corresponding continuous angular momentum trajectories [47]. These

arrangements are studied for their desirable momentum envelope (i.e., it is possible to

get a near spherical angular momentum envelope with a skew angle of 0 = 54.74 deg)

[52]. If high agility is what is needed and there are more relaxed pointing requirements,

the pyramid may provide benefits over the other arrangements. Even if this is not the

case, if this arrangement is hosted on a small satellite and the attitude error induced

from the torque error provided by the singularity escape of elliptic singularities is on the

same order of the attitude determination sensors and/or methods, then the torque error

used for singularity escape will be inconsequential.

3.2 Choice of Arrangement

Beyond the common arrangements of SGCMGs previously discussed, it is difficult

to choose the arrangement of SGCMGs through shaping of the angular momentum

envelope of the system. This is due to the locations of where internal singularities lie

within the angular momentum envelope denoted by the internal singular surfaces (e.g.,

see Figure 2-6). These singularities are dispersed and may cover the entire angular

momentum envelope leaving only very small singularity-free areas. Formulating

the problem as a parameter optimization as in [53] can only provide the optimal

arrangement for a given set of slews and initial gimbal angles which makes the problem

more constrained than useful. For example, we can express the gimbal axes relation to

the spacecraft body frame in terms of the Euler angles, two of which are the optimized

constants inclination angle 0i, spacing angle 0i, and the third is the gimbal angle 6,. The

DCM that is used to transform from the body to the gimbal frame is

CGB = C3 i)C2(0i)C3(i) (3-12)

The angular momentum of the SGCMGs is transformed from the gimbal TFg to the

spacecraft body frame Fb through this DCM as

h= CBGhi (3-13)
which is consistent with Eq.(2-4). Therefore, holding the spacing and inclination angles

constant, the resultant angular momentum of the CMG system is an instantaneous

function of only the gimbal angles for SGCMGs. Considering this and the truncated

dynamic model of SGCMGs from Chapter 2, the cost function

M = f(-m2 + aTeTTe b6 T) dt


where a and b are scalars making the cost function unitless and Tre = hoA6, can

optimize the system with respect to minimal torque error through the choice of the

Euler angles for a given slew, slew time, and initial conditions of the gimbal angles. This

procedure for the parameter optimization is shown in Figure 3-12.

Dynamics Solved First


Figure 3-12. Optimization process block diagram

3.3 Simulation

An example simulation of a rest-to-rest attitude maneuver has the parameters in

Table 3-1. This simulation will shown the benefit of different arrangements on performing

this maneuver (i.e., tracking the torque from the controller). It should be noted that the

initial conditions of the gimbal angles although are the same for every arrangement, they

produce a different initial CMG angular momentum.

Table 3-1. Model Parameters
Variable Value
/100 -2.0 1.5
S-2.0 900 -60!
1.5 -60 1000
60 [0 0 0 0]T
eo [0.04355 -0.08710 0.04355 0.99430]T
eo [0 0 0]T
ho 128
k 0.05
c 0.15
a 1
b 1
At 0.02
e., 0.0001




1/N2 m2s2

The results were simulated using the following eigen-axis control logic [54]

7- = -2kJe cJw + wxJw


combined with a fourth-order Runga-Kutta integrator at a timestep At until the

steady-state error tolerance of the error quaternion eigen-angle ess was reached. The

simulation compares the optimized solution to the four-SGCMG pyramid arrangement at

a skew angle 0 = 54.74 deg.

The results for this example at initial conditions 60, eo, and wo have the solution

for the system's singular surfaces with calculated arrangement Euler angles shown in

Figure 3-13.

2- 1

0 0
N i N -

A External singular surface B Internal singular surface

Figure 3-13. Singular surfaces for the optimized arrangement at the Euler angles
0* = [170.2 13.6 85.5 168.0]T deg and 0* = [17.7 167.0 304.3 92.5]T deg

The gimbal rates for the optimized arrangement in Figure 3-14 A are approximately

the same magnitude than that for the pyramid arrangement, although they have a

smoother transient response.

400 _d61/dt 200
300 -_d82/dt 100
d3 /dt
200 d 3dt 0 .d1d
g Id64/dt d68/dt
S100 -100 _d82,/dt

-200 d83/dt
/10 -300 4
0 20 40 60 0 20 40 60
Times(s) Times(s)
A Gimbal rates for optimized arrangement B Gimbal rates for pyramid arrangement

Figure 3-14. Gimbal rates for the optimized and pyramid arrangements

The torque error shown in Figure 3-15 A for the optimized case is smaller

magnitude than that for the pyramid arrangement due to the area under the curves

thus, more torque error is added during the maneuver for the pyramid arrangement.

z 0.04

0 20 40 60
A Torque error for optimized arrangement

0 20 40 60
B Torque error for pyramid arrangement

Figure 3-15. Torque error for the optimized and pyramid arrangements

For the optimized method, the singularity measure is far from singularity initially

and does not encounter it as shown in Figure 3-16 A. This is in contrast to that for

the pyramid arrangement shown in Figure 3-16 B, which starts out initially far from

singularity and then encounters singularity several times during the maneuver. The

negative quadratic term present in the cost function of Eq.(3-14) for this singularity does

not weight distance from singularity as high as torque error which can be seen when

comparing Figures 3-16 A and B to 3-15 A and B.



0 20 40 60 0 20 40 60
Times(s) Times(s)
Singularity measure for optimized arrange- B Singularity measure for pyramid arrangement

Figure 3-16. Singularity measure for the optimized and pyramid arrangements

Finally, the cost function of the optimized arrangement in Figure 3-17 A is less than

that for the pyramid in B due to the are under the curves.


2 30 30
20 20

10o 10
o 0
0 20 40 60 0 20 40 60
Times(s) Times(s)
A Cost for optimized arrangement B Cost for pyramid arrangement

Figure 3-17. Optimization cost for the optimized and pyramid arrangements

These simulations support the idea that if it where mechanically possible to

reconfigure the gimbal-axis arrangements in a timely manner, and the initial gimbal

angles and maneuver of interest were known, a solution to the optimal CMG arrangement

can be found. In addition, these simulations prove that you cannot simply choose

an optimal arrangement because the problem is not only dependent on the attitude

maneuver, but also dependent on the initial conditions of the gimbal angles.

If a desired arrangement is known while on-orbit and there was a mechanical way to

reconfigure the SGCMG gimbal axes, such as in the Honeywell patent [55], then there

would be merit in finding an algorithm that was successful in reorienting the gimbal axes

of the CMG arrangement. Although, no algorithm exists to reorient the SGCMG gimbal

axes while keeping spacecraft unperturbed. Also, such an algorithm would still require

angular momentum offloading due to the nature of SGCMGs.


A guidance, navigation, and control (GNC) system is composed of the loops shown

in Figure 4-1

Figure 4-1. Outer and inner loops of GNC system

The outer most loop of a spacecraft GNC system concerns the navigation (i.e.,

provides the state knowledge) and is usually the minimum loop needed for any mission.

The second most outer loop is concerns the guidance of the system (i.e., provides the

desired trajectories) (e.g., trajectories avoiding pointing a star camera towards the sun).

A loop inner to the guidance loop concerns the control of the system (i.e., generates an

error of the state knowledge from the navigation loop with the desired trajectories from

the guidance loop to be minimized). The inner most loop concerns the distribution of the

desired control to the systems actuators (e.g., what thrusters need to fire, what reaction

wheels or CMGs need to move). Steering algorithms are concerned with the inner most

loop of Figure 4-1 when the differential equation relating the control to the actuators is

singular. When this equation is singular, the steering law realizes a solution.

4.1 Moore-Penrose Pseudo-Inverse

An early method used to map the gimbal rates from the required output torque

uses the minimum two-norm least squares solution also known as the Moore-Penrose

pseudo-inverse. The solution of the gimbal rates using this pseudo-inverse mapping has

the form

6- A+h= AT(AAT)-lh (4-1)
ho ho
where A+ is the Moore-Penrose pseudo-inverse, ho is the magnitude of SGCMG

angular momentum, h is the SGCMG output torque, and 6 are the gimbal rates. The

Moore-Penrose pseudo-inverse, however, is singular when the Jacobian matrix A

has rank i 3 [56]. It might seem intuitive that the addition of more SGCMG actuators

increases the possibility of having full rank, but the performance is not equally increased

for all of CMG arrangements. This is because there are 2" singular configurations

for any given singular direction of a system containing n SGCMGs [41]. Also, the

Moore-Penrose pseudo-inverse and variations of it cause the system to move toward

singular states when performing discrete time control [47].

To handle cases when singularities may be encountered, steering algorithms are

used. Steering algorithms can be broken down into the following groups as shown in

Figure 4-2

4.2 Singularity Avoidance Algorithms

Singularity avoidance algorithms, are methods which steer the gimbals of the

SGCMGs away from internal singularities. These methods either constrain the angular

momentum envelope and/or gimbal angles, or apply null motion to avoid singularity

encounters. As discussed in Chapter 2, a method that uses only null motion cannot

avoid or escape elliptic internal singularities [44, 47].

Figure 4-2. Steering algorithms

4.2.1 Constrained Steering Algorithms

Constrained steering algorithms either constrain the gimbal angles and/or

useable angular momentum to avoid singularities. These steering laws are a form of

singularity avoidance that takes into account the locations of singularities a priori. As

a consequence of not using the entire angular momentum space, these steering logics

are typically more effective for systems where the SGCMGs are oversized. Honeywell

has patented methods that do not explicitly use null space but that implicitly do so by

creating constraints that keep the gimbals away from singularity without needing to

recognize their presence explicitly [57-59]. A simple example is the steering logic for

scissored pairs in Chapter 3, where mere constraints are used to keep the array out

of trouble. This method is able to guarantee singularity avoidance and an a available

torque but reduces the available workspace of the system by requiring it to be singularity

free [52, 60].

4.2.2 Null Motion Algorithms Local gradient (LG)

Singularity avoidance algorithms known as local gradient (LG) methods use null

motion to keep the Jacobian matrix from becoming singular. This is accomplished

through choice of the null vector d to maximize an objective function that relates the

distance from singularity such as the Jacobian matrix condition number, smallest
singular value [61, 62], or the singularity measure m which is expressed as

m = /det(AAT) (4-2)

An example of the null vector calculation is

9f (8m) -1 (Om) T
d = V6f (- m (4-3)
am ( 6 m2 (86)

where the objective function f = 1/m [42, 63, 64]. Minimization of this objective function

maximizes the distance from singularity by maximizing the singularity measure. The LG

methods, however, cannot avoid or escape elliptic internal singularities because they

apply only null motion [44, 64]. The null vector d can be arbitrary, although the projection
matrix which maps it onto the null space is constrained. Global avoidance/Preferred trajectory tracking

A way of choosing the null motion vector to steer gimbals to alternate nonsingular

configurations before maneuvering is known as preferred trajectory tracking [63, 65-67].

Preferred trajectory tracking is a global method that calculates nonsingular gimbal

trajectories offline. The gimbals converge to these trajectories using null motion to

minimize an error (6 6*). The gimbal rates using this method are

6 = A+h + [1 A+A](6 6*) (4-4)
where 6* are the preferred trajectories and A is the singularity parameter defined by

A = Xoexp-pm2 (4-5)

with constants A0 and p. Since this method calculates the preferred trajectories offline, it

is not real-time implementable. Also, preferred tracking relies entirely on null motion and

thus will be unable to escape elliptic internal singularities. Generalized Inverse Steering Law (GISL)

The Generalized Inverse Steering Law (GISL) provides a pseudo-inverse which is a
variation of the Moore-Penrose pseudo-inverse. This method defines another Jacobian
matrix B which has each of its columns orthogonal to the associated column of the
original Jacobian matrix A (i.e., aiLbi, not necessarily aZbj where A = [al a2 a3 a4] and
B = [b, b2 b3 b4]) [68]. Therefore, as an example for a four-CMG pyramid arrangement,
the matrix A and B have the following form

-cOcos(61) sin(62) cOcos(63) -sin(64)
A = -sin(61) -cOcos(62) sin(63) cOcos(64) (4-6)

s0cos(61) sOcos(62) s0cos(63) sOcos(64)

-cOsin(6) cos(62) csin(63) cos(64)
B = cos(61) -cOsin(62) -cos(63) cOsin(64) (4-7)

sOsin(61) sOsin(62) sOsin(63) sOsin(64)
where cO, sO are the cosine and sine of the pyramid skew angle 0 and 6, are the gimbal
angles, respectively. The pseudo-inverse of this steering law with the discussed matrices

AGISL = (A + B)T(A(A +B) )-1 (4-8)

It is important to note that this pseudo-inverse does not eliminate the problem of internal
singularities. The GISL adds null motion from the addition of B and therefore couples
the null and forced solution into a single inverse and thus, it is not able to avoid elliptic
internal singularities.
Claim: The GISL provides only null motion through B

6 = AGISL = (A + B)T(A(A + B)T)-l

The torque error is

h AJ = A(A + B)T(A(A + B)T)-h i h = h h = 0

If the matrix B = aA


AGISL = ((1 + )A)T((1 + )AAT)-1 = A

Therefore, the matrix B whose components are orthogonal to A must only provide

null motion and those that are along A vanish. Because, the GISL or any generalized

inverse used for singularity avoidance only adds null motion, it is unable to avoid elliptic

internal singularities.

4.3 Singularity Escape Algorithms

Singularity escape methods, known as pseudo-inverse solutions, add torque error to

pass through or escape internal singularity [39, 40, 69, 70]. These methods do not take

into consideration the type of internal singularity that is being approached when adding

torque error. Singularity Robust (SR) inverse

The Singularity Robust (SR) inverse is a variation of the Moore-Penrose pseudo-inverse

[69] where, a positive definite matrix Al composed of an identity matrix scaled by the

singularity parameter in Eq.(4-5) is added to the positive semi-definite matrix AAT. The

pseudo-inverse of this method has the form

ASR = AT(AAT +Al)-1 (4-9)

The SR inverse is able to escape both hyperbolic and elliptic singularities [44], although,

is ineffective in gimbal-lock escape. To overcome this situation, a modified SR inverse

known as the Generalized Singularity Robust (GSR) pseudo-inverse was developed

[39, 40]. Generalized Singularity Robust (GSR) inverse

The GSR inverse approach replaces the constant diagonal positive definite matrix

Al with a time-varying positive definite symmetric matrix AE

1 C1 2
E= c1 1 3 ci = cosin(wit + i,) (4-10)

C2 C3 1
where the off-diagonal terms of E are time dependent trigonometric functions with

frequency uw and phase shift 0i. The GSR inverse provides a means of escape

of the gimbal-lock configuration associated with a system of SGCMGs. The GSR

pseudo-inverse has the form

AGSR = AT(AAT + E)-1 (4-11)

and like the SR inverse, is guaranteed to avoid both hyperbolic and elliptic internal

singularities. Singular Direction Avoidance (SDA)

Another modification of the SR inverse known as the Singular Direction Avoidance

(SDA) only applies torque error in the singular direction and therefore reduces the

amount of torque error needed for singularity escape. The SDA method decomposes the

Jacobian matrix using a singular value decomposition (SVD) to determine its singular

values. The matrix of singular values is regulated with the addition of error to the

smallest singular value o-3 so that the pseudo-inverse is defined. The pseudo-inverse

using SDA has the form

0 0

0 0
ASDA = V 2= VYSDAUT (4-12)
3 +A

0 0

where o, are the singular values. Regulating only the smallest singular value, reduces

the amount of torque error added and creates smoother gimbal rate trajectories when

compared to the SR and GSR inverses [70]. This is obvious when the SR inverse

decomposed using SVD as

0 "2 0
ASR = V 22A UT = ASR = VYSRUT (4-13)

where all the singular values are regulated and hence there is torque error in all

directions. It is clear from Eqs.(4-12) and (4-13) that SR inverse and SDA are

susceptible to gimbal-lock because when the output torque is along the singular

direction h oc s = u3 then it is in the null(ISDAUT) and null(lSRUT) thus encountering

gimbal-lock as no consequence to the size of the torque error added from A. Without the

perturbations to the Jacobian matrix that are not gimbal state dependent at gimbal-lock

the system remains locked in a singular configuration.

Recall from Chapter 3, that for a four CMG pyramid arrangement, the rank is never

less than two and therefore it is acceptable to regulate only the smallest singular value.

However, if the skew angle is made close to 0, 90, 180, or 270 deg (i.e., box or planar

arrangement), the Jacobian matrix for these arrangements will have at least two small

singular values when near singularity and regulation of the smallest singular value may

be ineffective. Feedback Steering Law (FSL)

The Feedback Steering Law (FSL) provides a solution to the gimbal rates without

using an inverse. This method is derived from a minimization of the torque error which

is similar to how the SR inverse is derived. The optimization for defining FSL has the

following structure

1 Te KI 0 Te
min- (4-14)
,6R4 2 0 K

where K1 and K2 are positive definite gain matrices, and Te = h A6. This minimization

reduces to the SR inverse when K1 = 1 and K2 = Al and where A = Aoexp-pm2) from


The FSL method has K2 = 1 and K1 = K(s) as a compensator. The compensator is

derived from an H, minimization

w (s)[1 +AK(s)]-1
mmin (4-15)
K(s)GR3x3 W2K(s)[l AK(s)]-1/

where wi(s) and w2 are weighting matrices. The wi(s) matrix is defined below

wi(s) = : ] (4-16)
CK 0

where AK, BK, and CK are matrices associated with state-space model of the system.

The w2 matrix is constant and is

W2 = 14x4 (4-17)
where the constant w bounds the gimbal rates. The state-space model of the system

has the form

^ = AKX + BiKTe
(4-1 8)
6 = CKX

The output matrix CK is an explicit function of the CMG gimbal angles expressed as

CK = ATb2P (4-19)

with b as a positive scalar associated with the bandwidth and P is the steady-state

solution to the Riccati equation of the state-space system in Eq.(4-18). Using the

feedback of the system, Eq(4-18) will provide a solution to the gimbal rates that does

not require a pseudo-inverse. It should be noted that the system may start out stable,

however, the observability of the system may be lost resulting in instability, due to CK'S

explicit dependence on the gimbal angle (i.e., H(s) = CK(sl AK)-1BK where CK

constant). For further information on the development of this method, please see [71].

This steering algorithm was shown to go unstable for certain values of CK corresponding

to specific gimbal angles sets (see [72]). Singularity Penetration with Unit-Delay (SPUD)

The Singularity Penetration with Unit Delay (SPUD) algorithm escapes singularity

through reuse of the previous gimbal rate command when at a certain threshold of

singularity [73]. The previous command is saved through a zeroth-order hold to the

system. Escape of a singularity is always possible unless the system is initially at

the threshold of singularity, then there is no previous command to use for singularity

avoidance. Also, SPUD is not intended for attitude tracking maneuvers. The SPUD

algorithm accumulates attitude tracking error while escaping singularity and there are

no guarantees on how long it will take to escape singularity and how large the torque

disturbance will be on the spacecraft as its performance is directly associated with the

system and the choice of singularity threshold.

4.4 Singularity Avoidance and Escape Algorithms

Singularity avoidance and escape algorithms avoid singularities through null motion

whenever possible and use torque error for escape when they are not. Preferred gimbal angles

Preferred gimbal angles are a set of initial gimbal angles for SGCMGs that can

be reached by null motion. These angles are preferred since maneuvers originating

from them avoid a singular configuration [74]. This set of angles is found by backwards

integration of the Eq.(4-9) and the attitude equations of motion. It has been shown that

this method cannot avoid singularities if the initial set of gimbal angles is 60 = [45 -45 45

-45]T deg [74]. Since the null space projection matrix is undefined at singularity, the SR

inverse is used in place of the Moore-Penrose pseudo-inverse of A in ,n as

6n = [1 ASRA]d (4-20)

As a result, this causes the system to add torque error when at singularity. In practice,

this method acts as an offline optimization which determines the initial set(s) of the

gimbal angles that will give singularity free maneuver(s). However, it is not possible to go

from one to any point in gimbal space through null motion itself because there will never

be n dimensions of null space. Optimal steering law (OSL)

The Inner-Product Index (IPI) combined with the optimal steering law (OSL) is

used to determine a steering algorithm that produces minimum torque error while both

avoiding and escaping internal singularities [75]. The singularity index is added to the


min [-cV(6+6At) 16TW +eTR-le] (4-21)
6ER4,e6R3 2

where At is the one-step time delay, V(6 + 6At) is the IPI, re = h A6 (i.e., torque

error), c is a positive scalar, and W and R-1 are positive definite weighting matrices. The

IPI is approximated by a Taylor series expansion up to the 2nd order as

OVT 1 6 1 2V
V(6 +65 t) V + TAt +6TV &6t2 (4-22)
a +6 0 2 962 6
where the IPI V is expressed as a sum of square of inner products of the column vectors

of the Jacobian.

V=- (a aj)2 (4-23)
The result of the minimization in Eq.(4-21) using this approximation of V is

6 = H-AT(AH-AT + R)-l + [H-AT(AH-1AT + R)-AH-1 H-1]g (4-24)

where the Hessian matrix H is defined as

H = cAt2ggT + W (4-25)

with gradient g = a The weighting matrix R shown previously in the minimization of

Eq.(4-21), is expressed as

00 0
R =U 0 0 0 U (4-26)

0 0Aoexp- "
where o3 is the smallest of the singular values of the Jacobian matrix, A0 and p are

positive scalars and U is the unitary matrix made up of the left singular vectors from

the singular value decomposition of the Jacobian matrix A. This addition of torque error

into the gimbal rate state equation is analogous to the SDA method except that it is also

added to the free response solution [70]. It should be noted that this steering algorithm

does not consider the form of internal singularities and therefore, does not truly minimize

the amount of torque error for singularity escape. This is because non-degenerate

hyperbolic singularities are avoidable through null motion without the use of torque error.

At a non-degenerate hyperbolic singularity R is nonzero and thus torque error is still

added (see [72]).

4.5 Other Steering Algorithms

Other published steering algorithms that have not been discussed can be found in

the references [44, 61, 76-81]. These methods include mathematical techniques such

as neural networks, optimization, and game-theory.

4.6 Steering Algorithm Computation Comparison

An analysis comparing the computation for the implementation of the mentioned

steering algorithms is difficult due to lack of information on how some were coded in

literature. For example, some of these algorithms are offline and may require a large

number of memory calls and stored memory but not as many flops. It is however,

useful to quantifying some of the previously discussed steering algorithms in terms of

floating point operations that are not calculated offiline. These are shown in Table 4-1 for

algorithms where flops make a good comparison. In this table, the metric of comparison

is an approximate number of flops per time step.

Table 4-1. Algorithm Flops m = row(A) and n = column(A)
Variable Value
MP O(m4)
LG O(m4)
GISL O(m4)
SR O(m4)
GSR O(m4)
SDA 6(nm3)
FSL O(mn2)
SPUD O(m4)
Optimal Steering O(m4) 3+ (nm3)

It should be mentioned that many of the steering algorithms discussed have the

same order of magnitude of flops (e.g., MP, LG, GISL, SR, GSR, and SPUD) due to

the approximate number of flops for a Gauss-Jordan matrix inverse. It is assumed that

the calculation of the gradient vector for LG, and OSL is in memory and that the flops

associated with then are on a lower order of magnitude that an SVD or Gauss-Jordan

matrix inverse. The OSL has to do both SVD for calculation of the R matrix and the

Gauss-Jordan matrix inverse and therefore has O(m4) + O(nm3) flops. The addition of

O(m4) + 6(nm3) is inserted for the approximate flops of OSL because depending on

the amount of SGCMGs this algorithm is working for O(m4) < O(nm3). The FSL has

the lowest number of flops because it does not require and SVD or inverse, although it is

also not an exact mapping as previously discussed.


5.1 Hybrid Steering Logic

Existing steering logics (see Chapter 4) do not explicitly consider the type of

singularity that is being encountered and thus, do not completely address attitude

tracking performance of SGCMG attitude control systems. A proposed method known as

the Hybrid Steering Logic (HSL) which utilizes the knowledge of the type of singularity

encountered (i.e., elliptic or hyperbolic singularities) to improve the attitude tracking

performance of the SGCMG attitude control system, is developed for a four-SGCMG

pyramid arrangement at a skew angle 0 = 54.74 deg. By using a hybrid approach, HSL

acts as an LG method (i.e., null motion for singularity avoidance) at hyperbolic singularity

and an SDA method (i.e., pseudo-inverse solutions for singularity escape) at elliptic

singularity. Also, because HSL is developed for a four-SGCMG pyramid arrangement,

there is no existence of degenerate-hyperbolic singularities [41]. The challenge is to

develop the appropriate singularity metrics such that the LG and SDA components of the

hybrid strategy do not counteract each other during operation.

5.1.1 Internal Singularity Metrics

The singularity metrics developed are of similar form as the singularity parameter in

Eq.(4-5) with the addition of terms relating to the form of the actual singularity.

a = oaexp-aexp-pm (5-1)

/ = Po exp-bexp-2m (5-2)

where a, b, p/, p2, ao and /o are positive scalar constants and m is the singularity

measure as defined in Eq.(4-2). Away from singularity, a four-SGCMG pyramid

arrangement at a skew angle 0 = 54.75 deg, has the matrix Q e R. At singularity

this SGCMG arrangement has Q e R2x2 (see Chapter ??) and therefore, the det(Q)

will be zero or negative (i.e., Q is negative semi-definite or indefinite) for hyperbolic

singularities and positive (i.e., Q is definite) for elliptic singularities. Taking this into

account, parameters a and p are defined as

a = Qo det(Q)| (5-3)

1 1
= (5-4)
SQo det(Q) a
where Qo is a scalar value chosen on the same order of magnitude of det(Q) but

greater to scale the response of a and /. It is difficult to analytically define Qo since it

depends on the maximum value of det(Q) (i.e., det(Q) varies with gimbal angle and

therefore the maximum must span all combinations of the gimbal angles) which is of

high dimensionality and highly nonlinear. However, through simulation of a four-SGCMG

pyramid arrangement at a skew angle 0 = 54.75 deg, it was found that Idet(Q)| < 1

and therefore we define Qo 1. In addition, it is important to note that the constant

parameters a, b, p1, PJ, ao, and /o are used to morph the HSL steering logic into the

respective LG and SDA methods when appropriate: (e.g., if the parameters a = b =

ao = 0 and /o a 0 then the HSL method is the LG method). Therefore, the choice of

metrics a and / in this way ensures that null motion will be added when approaching

a hyperbolic singularity and torque error with less null motion will be added when

approaching an elliptic singularity. It should be noted that when using HSL det(Q) is

normalized by the nominal angular momentum ho.

5.1.2 Hybrid Steering Logic Formulation

The proposed steering logic is defined as

6 = ASDAa + /3[1 A+A]d (5-5)
where ASDA a is

0 0
0 0
ASDAa = V (2 UT (5-6)
00 0
(T32 2+
0 0
If it is assumed that the analytic function for the gradient vector d is derived offline and

the calculation of it at each time-step is less than that for SVD, this algorithm has the

same number of flops on order as SDA from Table 4-1 of 6(nm3) from the SVD. The

difference between the conventional ASDA and ASDA,a is the parameter that regulates

o3. In ASDA, the regulation parameter is 7 (i.e., different from A in Chapter 4 by using m
instead of m2) which is

7 = 70exp-pm (5-7)

with positive constants 70 and p, but with ASDAa the singularity parameter is a defined

in Eq.(5-1) which regulates the amount of induced torque error in the vicinity of elliptic

singularities. Through a SVD decomposition of A, Eq.(5-5) can be written as

1 [1
6 = ASDAah + /[1 V V]d (5-8)
h0 OT 0

Here, the null motion projection matrix is expressed as a function of nonsingular
matrices V. Also, very robust numerical algorithms exist for computing the SVD, so

its computational risk in a real-time implementation is not particularly high.

The scalar that regulates the magnitude of the null motion is 3. The null vector d is

in the direction of the gradient of f = -det(AAT) = -m2 and maximizes the distance

from singularity.

This choice of this objective function reduces the computation needed for the

gradient (i.e., the derivative of (-det(AAT)) is less computationally intensive than the

derivative /(det(AAT)) and ensures that the addition of null motion will not approach
infinity at the region of singularity for cases such as f = 1 and then t = It
m W65 m2 5 "
should be mentioned that the null vector is a nonlinear function of the gimbal angles and

is simplified due to the symmetry of the four CMG pyramid arrangement. To prove the

feasibility of HSL, a stability analysis is conducted.

5.2 Lyapunov Stability Analysis

The candidate Lyapunov function

V = w eTK-J + eTe +(1 e4)2 (5-9)
is chosen for this analysis and can be rewritten as

V = zTMz (5-10)

where z = [WT eT (1 e4)]T and M = diag(~K-'J, 1,1). Consequently the Lyapunov

function is bounded as

Ami, z12 < V < Amax lz12 (5-11)

where Ami, and Amax denote the minimum and maximum eigenvalues of M. This bound

will become useful later in the analysis.

A rest-to-rest quaternion regulator controller is given by Eq.(5-12) for the internal

control torque 7, is chosen for its flown heritage and the fact that it yields an global

asymptotic stable control solution proven through LaSalle's Invariant Theorem [63].

7 = -Ke Cw + XJ (5-12)

Gain matrices K = 2kJ and C = cJ of Eq.(5-12) are positive definite and symmetric.

Assuming rigid body dynamics, the spacecraft's angular momentum is given by

H =Jw +h

The rotational equations of motion come from taking the inertia time derivative of

Eq.(5-13) as

Lj = -l[Tact oXJ] (5-14)

with SGCMG output torque

h=- -xh = hoA (5-15)

where w is the spacecraft angular velocity, J is the spacecraft centroidal inertia, H is

the total system angular momentum, and h is the angular momentum from the CMGs.

The spacecraft's angular velocity and the CMG angular moment are governed by

Eqs.(5-14) and (5-15) respectively, where Tact is the actual control torque (i.e., may

differ due to induced torque error for singularity escape). It is assumed here that the

contribution to the dynamics from the gimbal-flywheel assembly inertias is negligible and

therefore J is constant.

The actual control torque Tact based on the mapping of the gimbal rates is

Tact= -hoA6 -xh= -A(ASDAah +3[1- V 1 0 VT]d)-wh (5-16)
0T 0

and needs to be considered in the Lyapunov analysis for stability of the attitude

controller/steering algorithm combination. When simplified, Eq.(5-16) becomes

00 0

Tact =U 0 0 0 UT[- -- wXh] + (5-17)

0 0 -
where the stability of the system is affected by the torque perturbation matrix CHSL from

ASDAa defined as


00 0

EHSL = U 0 0 0 UT (5-18)

0 0 a+
o +a
The spacecraft attitude error kinematics is governed by

1 1
e= --wxe + Ie4 (5-19)
2 2

e4 Te (5-20)

where e is the quaternion error vector elements and e4 is its scalar element. The time

derivative of the Lyapunov function is

1 1
V = wK- [-rat WJ] + 2eT[-2 Xe+ Iwe4] + 2(1 e4) w e (5-21)
2 2 2

Equation(5-21) can be reduced by substituting in the expression for ,act from

Eq.(5-17) and the desired control torque vector 7 from Eq.(5-12). The time derivative of

the Lyapunov function now yields

V = W-cTK- [C- EHSL(C+ HX )]w K- 6HSLKe (5-22)

or more compactly

V = 2kTW + TJ -HSLtJ +C1 (5-23)
2k 2k

e1 = w TJ- HSLHX + WTJ-IEHSLJe (5-24)

Since T2a < 1, Eq.(5-18) can be used to rewrite Eq.(5-23) as

10 0
V < -C TjU 0 1 0 UTJW + e1 = -_(TR + e (5-25)
0 0 A1

and can be further bounded as

V < -Allz112 + (5-26)

where A1 is the minimum eigenvalue of the positive semi-definite matrix R, (i.e., A0 = 0

at singularity), ( = 2, and ||z|| is defined in Eq.(5-10). Substituting Eq.(5-11) into

Eq.(5-26) yields
V < -AV + (5-27)
A max
The solution to the differential equation in Eq.(5-27) in a Volterra integral form is

V < V(0)exp(- t) + ,(t c1 ,T d(5-28)

The error can be bounded from Eq.(5-28) as

12 < V(0) X t) 1 t(t-7))(7)d
|z 12 < exp +, 1 t exp maxt (5-29)
Amin Amin Jto
At singularity when A = 0, the error is

|z||2 < V m c1(-)dT (5-30)

where Vs is the error at time the occurrence of singularity at time ts.

Stability cannot be proven from Eq.(5-30). It is assumed that the system will not

remain locked in singularity except for the special case of gimbal-lock. If components

of the torque needed for stability are actually in the singular direction, periods of

instability may occur at singularity. The duration of this instability is dependent on

the selection and sizing of the singularity parameters a and / which provide torque

error for singularity escape and/or null motion for singularity avoidance. From a practical

perspective, stability cannot be proven for this Lyapunov function at singularity, since at

singularity, there is no torque available in the singular direction.

For the special case of a singularity with gimbal-lock, the angular velocity of

the spacecraft is constant assuming the absence of friction and external torques in

the system. In this case, the contribution from c1 to the error is bounded and even

sometimes zero. This can be shown by evaluating the expression for the angular

acceleration at gimbal-lock which is

S= 0 = J-1[0 + wH] (5-31)

It is clear that Eq.(5-31) is satisfied only when the product w" H = 0 which is only

true when w is parallel to H, w = 0, or H = 0. When w is parallel to H or H = 0,

1i = wTJ-EHSLJe which is a bounded sinusoid whose integral is also a bounded

sinusoid. Therefore, for these two cases, the error is bounded at gimbal-lock. When

w = 0, c0 = 0 and the error is simply locked at Vs. It should be mentioned that at

times away from singularity, the error monotonically decreases because the contribution

from cl to the error becomes negligible. Care needs to be taken in the design of the

singularity parameter so that the minimum steady-state error is achieved while meeting

the constraints of the actuators.

The steady-state error assuming that the systems has a singularity free period

towards the end of the maneuver (i.e., does not end at singularity) is

iiz(0)l12 < 1 r,
-z(-oo) -exp -ax) exp)mx ())d? (5-32)
Amin ts
This expression is indeterminate so application of LHopital's rule to Eq.(5-33) yields

d 1 t ( A 1 -)
lim dt Am, J exp x (00o) (5-33)
t*oo dxp t) Amin i
dt x

which suggests that a sufficiently large value of c (i.e., larger () will lower the amount

of steady-state error giving you a uniformly-ultimately bounded (UUB) result away from

singularity. When the maneuver is finished, the effect of e1 on the error will become

a constant assuming the maneuver ends at rest. It should be noted that away from

singularity the size of e1 exponentially decreases due to the behavior of EHSL. The

difference in impact of HSL rather than SR inverse on stability can be observed from the

magnitude of the positive semi-definite matrix EHSL in Eq.(5-18) compared to the matrix

shown in Eq.(5-34). The ESR matrix has a larger norm and therefore has a worse UUB

even for sufficiently large values of (. From comparing Eqs.(5-6) and (5-18), the SDA

method has a similar amount of torque error added when compared to HSL, although

it will add this torque error whenever the singularity approached not taking into account

the form.

0 0

SR = U 0 0 (5-34)

0 0 2
U3+ .
The above results are only for the attitude controller/steering algorithm combination.

For example, an attitude controller whose torque trajectory was chosen to avoid

the occurrence of singularities may not have the periods of possible instabilities at

singularity and thus may provide better stability performance. However, no real-time

controller of this form exists (i.e., one that ensures singularity avoidance) and thus was

not considered in the following simulations.

5.3 Numerical Simulation

To evaluate the performance of the proposed HSL against heritage steering

logics (i.e., LG and SDA), simulations were performed using a four-SGCMG pyramidal

arrangement with a skew angle of 0 = 54.74 deg. To ensure a fair comparison, the

control logic and satellite model were identical for all simulations. For each steering

algorithm, three different scenarios were simulated: (1) starting in a zero-momentum
configuration 6 = [0 0 0 0]T deg (i.e., far from singularity); (2) starting near an elliptic
external singularity 6 = [105 105 105 105]T deg; and (3) starting near an hyperbolic
singularity 6 = [15 105 195 -75]T deg. The singularity conditions were verified for each
case by observing the singularity measure defined in Eq.(4-2). For these simulations,
the following performance measures were compared: (i) the transient response of the
error quaternion, (ii) the amount and duration of singularity encounter, (iii) the magnitude
of gimbal rate, (iv) the amount of torque error (i.e., h hoA6) for singularity escape, and
(v) null motion contribution. Additionally, a, 3, and det(Q) are also considered.
The Jacobian associated with this pyramidal configuration is

-c(0)c(61) S(2) c(0)C(63) -S(6
A = -s(61) -c(O)c(2) s(3) c(0)c(4) (5-35)

s(0)c((1) s(O)c(62) s(O)c(63) s(O)c(64)
and the associated angular momentum vector is

-c(0)s(1) C(62) + C(O)S(3) c(4)
h = ho c(6) c(0)s(62) c(63) c()s(64) (5-36)

s(O)(s(6) +S(62) + (3) +S(4))
All simulations are performed using a fourth-order fixed time step Runga Kutta with the
parameters shown in Table 5-1. The actuator parameters chosen for this simulation are
based on the Honeywell M95 SGCMGs, which are sized for the satellite system chosen
for simulation [82].

Table 5-1. Model Parameters

Variable Value Units
/100 -2 1.5 \
J -2 900 -60 kgm
\1.5 -60 1000/
0 54.74 deg
eo [0.04355 0.08710 0.04355 0.99430]T
w0 [0 0 0]T deg/s
ho 128 Nms
k 0.05 1/s2
c 0.15 1/s
mo 0.5
ess 0.001 deg
At 0.02 sec

It should be noted that care must be taken when numerically defining the singular
direction since s = 0 when the system has a full rank Jacobian. Because the rank
is numerically determined, a tolerance should be set on the singularity measure to
determine what is considered full rank. For the results presented in this paper, rank
deficiency for the HSL was defined as m < mo where for this simulation mo = 0.5. The
simulations terminate when the steady state error ess defined in Eq.(5-37) is achieved.

es = min[2sin-l(llel ), 27 2sin-l(l||e|)] (5-37)

The magnitude of ess given in Table 5-1 is based off reference [38].
5.3.1 Case 1: At Zero Momentum Configuration 6 = [0 0 0 0]T deg
The first set of simulations has initial gimbal angles at 6 = [0 0 0 0]T deg
which represents a scenario starting far away from singularities. Figure 5-1 shows this
configuration which is also a typical startup configuration for a four-SGCMG pyramid

h i:h

x z li I


Figure 5-1. Zero-momentum configuration of a four-SGCMG pyramid arrangement
 Local gradient simulation results

The parameters for the LG simulation are: ao = a = b = p/ = 0 and P2 = o = 1.

Figures 5-2 A and B show that the LG method was able to perform the maneuver to the

given error tolerance ess without inducing torque error. The absence of torque error in

Figure 5-2 B is due to the zero value of singularity metric a in Figure 5-2 C (i.e., LG is

an exact mapping). The null motion shown in Figure 5-3 B is small but significant when

compared to the total output gimbal rates in Figure 5-2 A. This is a consequence of the

singularity metric 3 in Figure 5-3 D. Figures 5-3 C and D show that the maneuver was

completed without singularity encounter.


0.02 _e2
0 -e3


0 20 40 60 8(
A Quaternion error vector elements



s 0



x 10-14





0 20 40 60 80
B Torque Error






0 20 40 60 80 0 20 40 60
Times(s) Times(s)
C Alpha D Beta

Figure 5-2. Simulation results for LG with ao = a = b = / = 0 and P2 = Ao

1 at zero


*^ ^ -- d ,/d
............... At

0 20 40 60
A CMG gimbal rates



E 1.0888



0 20 40
B Null motion



0 20 40 60
C Singularity measure

20 40
D det(Q)

60 80

60 80

Figure 5-3. Simulation results for LG with ao = a = b
momentum (contd.)

p = 0 and 2 = 3 = 1 at zero





0.05 \

0.05 Singular Direction Avoidance simulation results

The parameters for the SDA simulation are: ao = 0.01, / = a = b = 2 = 0,

and p/ = 1. This method shows similar results in the transient response of the error

states in Figure 5-4 A to that for LG in Figure 5-2 A with the exception of nonzero torque

error seen in Figure 5-4 B. Also, this method had a slower rate of convergence to the

steady-state error ess than LG as evident from the time in simulation in Figure 5-2 A.

This is due to the small nonzero value of the singularity metric a, shown in Figure 5-4 C.

Figure 5-5 B shows a zero null motion contribution to the gimbal rates in Figure 5-5 A for

SDA. Figures 5-5 C and D and Figures 5-3 C and D are almost equivalent because the

system started far away from singularity.

0.02\ _e

0 e,

< -0.02



0 20 40 60 80
A Quaternion error vector elements





x 10-3

20 40 60

20 40 60
C Alpha


0 20 40 60 80
B Torque Error






20 40 60
D Beta

Figure 5-4. Simulation results for SDA with ao = 0.01, /o
at zero momentum

= a = b = 2 = 0, and /p = 1


_ d82/dt

S d 4/dt



0 20 40 60
A CMG gimbal rates

20 40 60
B Null motion






1.0885 V
0 20 40 60
C Singularity measure




Figure 5-5. Simulation results for SDA with co = 0.01, /o
at zero momentum (contd.)

20 40 60
D det(Q)

=a = b = 2 =

0, and p = 1





H, Hybrid Steering Logic simulation results

The parameters for the HSL simulation are: ao = 0.01, /o = 2, a = 1, b = 3,

and pi = p2 = 1. In Figure 5-6 A, the HSL method shows similar results to that of

the SDA shown in Figure 5-4 A, with exception to the faster rate of convergence of

the transient error response. However, the torque error in Figure 5-6 B added into the

system is smaller than that of SDA in Figures 5-4 B and null motion in Figure 5-7 B is

smaller than that of the LG method in 5-2 B. This is due to the nonzero value for both

singularity metrics a and 3 in Figures 5-6 C and D. Singularity was not encountered in

this simulations as is shown by a value m > 0.5 in Figure 5-7 C and a zero value of

det(Q) in 5-7 D.

For Case I at zero-momentum, Table 5-2, compares the root-mean squared (RMS)

gimbal rates (deg/s) and tracking performance in terms of RMS torque error (Nm) for

LG, SDA, and HSL. In this table it is shown that all three methods have approximately

the same performance which is expected for a four-SGCMG pyramid arrangement at

zero-momentum, far from singularity. The steady-state error for LG or any of the other

methods is nonzero as a consequence of the controller's performance is captured here.

Table 5-2. Performance Comparisons for Case I: Zero Momentum
Steering Algorithm sRMS TeRMS
LG 5.7366 2.2437e-06
SDA 5.7279 3.5902
HSL (mo = 0.5) 5.7317 2.2573

5.3.2 Case 2: Near Elliptic External Singularity 6 = [105 105 105 105]T deg

The second set of simulations starts at initial gimbal angles 6 = [105 105 105 105]T

deg, which represents a scenario near an elliptic external singularity at (i.e., 15 deg for

each SGCMG away from the external singularity 6 = [90 90 90 90]T deg).



a -0.02




E 8

0 20 40 60
A Quaternion error vector element

0 20 40 60 80
C Alpha

80 0 20 40 60 80
s B Torque Error






Figure 5-6. Simulation results for HSL with ro = 0.01, /o
P1 = p2 = 1 at zero momentum

20 40 60 80
D Beta

2, a= l, b= 3, and

_d8 /dt

* d84/dt

0 20 40 60
A CMG gimbal rates

0 20 40
B Null motion



0 20 40 60
C Singularity measure

Figure 5-7. Simulation results for HSL with co = 0.01, /o
l = P 2 = 1 at zero momentum (contd.)

20 40
D det(Q)

=2, a= 1, b

60 80

60 80

3, and


0.05 Local gradient simulation results

The parameters for the LG simulation are: o = a = b = p = 0 and 2 = /3 = 1.

The plots in Figure 5-8 A show that the LG method appears to have successfully

performed the maneuver as shown in Figures 5-8 A and B. However, this is misleading

since non-implementable gimbal rates and accelerations are required to do so as

shown in Figure 5-9 A. The singularity metrics a 0 as expected for this method and

3 = 1 at the singularity encounter. Even though = 1 at singularity, null motion at the

exact time of singularity encounter is zero as shown in Figure 5-9 because the gradient

vector d for LG is zero at elliptic singularities (i.e., no gradient vector exists that is in

the direction away from singularity). Also, the singularity, verified to be elliptic from the

positive value of det(Q) in Figure 5-9 D, was escaped immediately with the help of the

non-implementable gimbal rates and accelerations, shown by Figure 5-9 C.

x 10-12

0.02\ e

O_ e

a -0.02



0 20 40 60 8
A Quaternion error vector elements

) 20 40 60 80
B Torque Error






20 40
C Alpha

60 80 0

Figure 5-8. Simulation results for LG with ao = a = b
elliptic singularities

20 40
D Beta

/i1 = 0 and p2

60 80

:/3 =1 near




d. 4/dt

0 20 40 60 81
A CMG gimbal rates



0 _- Y

0 20 40 60 80
B Null motion







0 20 40 60
C Singularity measure

Figure 5-9. Simulation results for LG with
elliptic singularities (contd.)



- 0.2

_ 0.15




ao = a = b

20 40
D det(Q)

/1 = 0 and P2 ~

60 80

S=/3 1 near Singular Direction Avoidance simulation results

The parameters for the SDA simulation are: ao = 0.01, /o = a = b = /2 = 0, and

1 = 1. The transient response of the error for the SDA method shown in Figure 5-10 A

is comparable to that of the LG method in Figure 5-8 A, but with implementable gimbal

rates and accelerations as shown in Figure 5-11 A. The SDA method escapes the elliptic

external singularity as shown in Figure 5-11 C at the expense of significant torque error

shown in Figure 5-10 B. The torque error scaled by the singularity metric a shown in

Figure 5-10 C decreases away from singularity as shown in Figure 5-11 C. As expected

for SDA, the singularity metric / in Figure 5-10 C is zero resulting in zero null motion

as shown in Figure 5-11 B. In contrast to the LG method, for SDA, the system lingers in

singularity for around 15 seconds before escaping as shown in Figure 5-11 C. Elliptic

singularity for this simulation is verified by the positive value of det(Q) in Figure 5-11 D.

0.02\ e.
0 --e,

a -0.02



0 50 100
A Quaternion error vector elements

50 100
C Alpha

0 50 100
B Torque Error




D Beta

Figure 5-10. Simulation results for SDA with ao = 0.01, /o = a = b = 2 = 0, and / = 1
near elliptic singularities


- da1/dt
d6 4/dt

50 11
A CMG gimbal rates

50 10
C Singularity measure






B Null motion

D det(Q)

Figure 5-11.

Simulation results for SDA with co = 0.01, /o = a = b = 2 = 0, and / = 1
near elliptic singularities (contd.) Hybrid Steering Logic simulation results

The parameters for the HSL simulation are: ao = 0.01, /o = 2, a = 1, b = 3,
and /i = /p2 = 1. The results for HSL shown in Figures 5-12 and 5-13 are almost
identical to the corresponding results of SDA for this simulation. The only difference

between HSL and SDA simulated results, lies in the nonzero singularity metrics a and
/ in Figures 5-12 C and D. Due to the choice of the HSL parameters a, b, / ,, pL2 I, ao,

/3, the threshold for singularity m < 0.5, and Qo, the HSL acts as the SDA at an elliptic

0.02\A _e,
O .0- e,
a -0.02
0 50 100
A Quaternion error vector elements

x 10-3

0 20 40 60 80 100
B Torque Error

0 20 40 60 80 100 0 20 40 60 80 100
Times(s) Times(s)
C Alpha D Beta

-12. Simulation results for HSL with ao = 0.01, /o = 2, a = 1, b = 3, and
1 = P2 = 1 near elliptic singularities


Figure 5





20 40 60 80 100
A CMG gimbal rates

0 20 40 60 80 100
C Singularity measure




a 5


-5 ..'

0 20 40 60 80 100
B Null motion



20 40 60 80 100
D det(Q)

Figure 5-13.

Simulation results for HSL with co = 0.01,/3o
1 = P2 = 1 near elliptic singularities (contd.)

2, a = 1, b = 3, and

For Case II near an elliptic singularity, Table 5-3 compares the RMS gimbal rates

(deg/s) and tracking performance in terms of RMS torque error (Nm) for LG, SDA,

and HSL. In this table the LG method is said to have an infinite RMS gimbal rate to

point out that it failed for elliptic singularity. Also, it is shown that SDA and HSL were

successful in completing the maneuver while escaping singularity. Both SDA and HSL

had approximately the same performance for elliptic singularity with the exception of

slightly better tracking performance for HSL.



Table 5-3. Performance Comparisons for Case II: Elliptic Singularity
Steering Algorithm sRMS TeRMS
LG c0 7.7159e-06
SDA 8.2564 29.8989
HSL (mo = 0.5) 8.1366 26.6946

5.3.3 Case 3: Near Hyperbolic Internal Singularities J = [15 105 195 -75]T deg

The final set of simulations starts at initial gimbal angles 6 = [15 105 195 -75]T

deg which represents a scenario near an hyperbolic singularity at (i.e., a distance 15 deg

from each CMG away from the singularity at 6 = [0 90 180 -90]T deg). Local gradient simulation results

The parameters for the LG simulation are: ao = a = b = p = 0 and /2 = / = 1.

The transient response of the error for the LG method in Figure 5-14 A is identical to

that for the other two cases. This is because the LG method is an exact mapping evident

from a = 0 in Figure 5-14 C and has no torque error associated with it in theory as

shown in Figure 5-14 B. The null motion in Figure 5-15 B makes up almost the entire

contribution of the gimbal rates in Figure 5-15 A due to the nonzero value of 3 in Figure

5-14 D. The LG method by itself is able to avoid the hyperbolic singularity swiftly and

remain away as shown in Figure 5-14 C and D.

0.02\/ e

0 -e,

< -0.02



0 20 40 60 8
A Quaternion error vector elements






20 40
C Alpha

60 80

x 10-13





0 20 40 60 80
B Torque Error

0 20 40
D Beta

60 80

Figure 5-14.

Simulation results for LG with ao = a =
hyperbolic singularities

b = 1 = 0 and 2 = Po = 1 near


d6 4/dt








20 40 60 80 0 20 40 60 80
Times(s) Times(s)
A CMG gimbal rates B Null motion

0 20 40 60 80
C Singularity measure

Figure 5-15.








Simulation results for LG with ao = a = b
hyperbolic singularities (contd.)

20 40
D det(Q)

/i = 0 and P2

60 80

= /o = 1 near Singular Direction Avoidance simulation results

The parameters for the SDA simulation are: /o = a = b = p2 = 0,/1 = 1 and

co = 0.01. The transient response of the error for the SDA method in Figure 5-16 A

is different in the rate of convergence to ess, but on the same order of magnitude to

that for the LG method. However, the gimbal rates for SDA shown in Figure 5-17 B are

an order of magnitude smaller than that for the LG method. This method escapes the

hyperbolic singularity successfully with torque error as shown in Figures 5-16 B as a

consequence of the nonzero value of a in Figure 5-16 C. The singularity metric P in

5-16 D is zero because SDA does not use null motion. Added torque error for singularity

escape versus null motion for singularity avoidance is the trade off between SDA and

LG. The singularity in this simulation is verified to be hyperbolic from the negative result

shown in Figure 5-17 D. Also, the SDA method did not escape by what is considered

singularity in Figure 5-17 C by the threshold m < 0.5. However, this did not affect the

decaying of the errors transient response. This is due to the fact that the torque error is

scaled by the needed output torque being mapped and therefore, is not seen to have a

significant effect towards the end of the maneuver.

0.02\X e

0__ e,

a -0.02



0 50 100
A Quaternion error vector elements

0 50 100
C Alpha

0 50 100
B Torque Error




D Beta

Figure 5-16. Simulation results for SDA with ao = 0.01, /o = 0, a = 0, b = 0, and / = 1
near hyperbolic singularities




d6 4/dt




A CMG gimbal rates



-o -0.03

50 1
C Singularity measure



B Null motion

50 100
D det(Q)

Figure 5-17.

Simulation results for SDA with co = 0.01, /o
near hyperbolic singularities (contd.)

= 0, a = 0, b = 0, and p = 1


0 Hybrid Steering Logic simulation results

The parameters for the HSL simulation are: ao = 0.01, /o = 2, a = 1, b = 3, and

p 1 = 12 = 1. The transient response of the error in Figure 5-18 A is almost identical to
the LG method for this case and has a faster rate of convergence to ess than SDA. This

is attributed to the nonzero values of the singularity metrics a and / in Figures 5-18 C

and D which provide an order of magnitude less null motion for singularity avoidance

than LG and orders of magnitude less torque error than SDA for this case shown in

Figure 5-19 B and 5-18 A when avoiding the hyperbolic singularity verified in Figures

5-19 C and D. Unlike SDA, HSL escaped and then avoided the singularity which is

due to the addition of null motion for this method (see Figures 5-16 C and 5-17 C).

Therefore, HSL relies more on null motion for singularity avoidance rather than soley

trying to pass through the hyperbolic singularities as SDA, SR, and GSR do. Precision in

attitude tracking with the threat of hyperbolic singularities while still being able to escape

elliptic singularities is the strength of HSL.


0.02 e
0 e e3
< -0.02
0 20 40 60 80
A Quaternion error vector elements

E 0.04
P 0.03

0 20 40 60 80
B Torque Error




0 20 40 60 80 0 20 40 60 80
Times(s) Times(s)
C Alpha D Beta

Figure 5-18.

Simulation results for HSL with co = 0.01,/o =
11 = 1/2 = 1 near hyperbolic singularities

For Case III near a hyperbolic singularity, Table 5-3 compares the RMS gimbal

rates (deg/s) and tracking performance in terms of RMS torque error (Nm) for LG, SDA,

and HSL. In this table the LG method has the largest RMS gimbal rate among the three

methods, which is needed for singularity avoidance. Also, LG performed the method with

the best tracking performance among the three methods which is an expected result

for an exact method. The HSL had better tracking performance in terms of RMS torque

error than SDA as a consequence of the larger gimbal rates needed for null motion

singularity avoidance. This is an expected strength of HSL at hyperbolic singularity.


2, a=l, b

3, and



S1 -d6 dt 1.
: 0, 4 Foa 1 44


-2 -0.5

0 20 40 60 80 0 20 40 60 80
Times(s) Times(s)
A CMG gimbal rates B Null motion


1 -0.02

E 0.8 -0.04

0.6 -0.06

0 20 40 60 80 0 20 40 60 80
Times(s) Times(s)
C Singularity measure D det(Q)

Figure 5-19. Simulation results for HSL with co = 0.01, /0 = 2, a = 1, b = 3, and
P1 = P2 = 1 near hyperbolic singularities (contd.)

Table 5-4. Performance Comparisons for Case III: Hyperbolic Singularity
Steering Algorithm sRMS TeRMS
LG 10.3905 1.4742e-05
SDA 6.3611 14.4937
HSL (mo = 0.5) 9.9330 4.7925

5.4 Hybrid Steering Logic Summary

The HSL was found numerically to preserve attitude tracking precision in the

presence of hyperbolic singularities, act comparably to SDA in the presence of elliptic

singularities, and perform better than SDA away from singularity. The performance

of this algorithm is attributed to the new singularity metrics, which allow smooth

transition between singularity avoidance using LG and singularity escape using SDA.

By reducing the times where torque error is induced for singularity escape, this method

provides improved attitude tracking performance. Analytic and simulated results show

that HSL has many benefits over the two other methods for singularity avoidance

and escape. These benefits are: it can be implemented real-time; although SVD

may be computationally intensive, it removes the need for an inverse and provides

all the information needed for HSL; numerically robust algorithms exist for SVD;

HSL induces less torque error than SDA by itself; and finally, the HSL provides a

nonsingular expression that can start at singularity. The HSL is not successful in

avoiding gimbal-lock because null motion is nonexistent at elliptic singularities and SDA

fails at gimbal-lock (see Chapter 4).



6.1 Attitude Controller with OTC

Traditionally the control law and steering algorithm are separated for attitude control

systems using SGCMGs as shown in Figure 6-1. This is done to facilitate understanding

of the attitude control system and actuator dynamics separately. However, considering

the steering algorithm separate from the control law may reduce the possibility of an

increase in the performance in the system.

W"s Separated!
(d 7 /^ l qe 7 WWWWWWW h,,/

q, q ---------


Figure 6-1. Satellite attitude control system block diagram

Many steering logics by themselves are incapable of avoiding gimbal lock.

Gimbal-lock occurs when the required torque for an attitude maneuver is along the

singular direction. This produces a local minimum condition where the gimbal rate

solution is zero while the required torque is still not met. Open-loop methods that

provide a gimbal trajectory free of this condition exist; examples of such methods are

forward propagation from preferred gimbal angles, global steering, and optimal control

[65, 66, 74, 83]. These methods are time consuming and cannot guarantee a solution

exists for the constraints provided.

Real-time solutions to gimbal-lock avoidance exist such as the GSR inverse which

uses off-diagonal dither components in its perturbation matrix to escape gimbal-lock

(see Chapter 4). There is no formal proof that these methods will always be successful

in avoiding gimbal-lock. Through the use of nonlinear control, an orthogonal torque


compensation (OTC) methodology can be augmented with a suitable steering and

control algorithm to also avoid or escape gimbal-lock. Through this nonlinear control

framework, stability can be proven and the steering algorithm can be chosen separately

in contrast to GSR which relies entirely on handling gimbal-lock avoidance/escape

through the steering algorithm.

Open-loop methods such as optimal control for gimbal-lock escape or avoidance

may not find a feasible solution or a solution at all depending on how the cost function

and constraints are formulated.

It is possible that combination of an optimal control maneuver with a pseudo-inverse

method (e.g., SR inverse) will drive the system toward the vicinity of singularity as

the maneuver is completed. This may occur since the required gimbal rates are not

only scaled from the distance to singularity, but also by the needed output torque from

the SGCMG system. As the next rest-to-rest maneuver is needed the torque may be

required about the singular direction. When this occurs, the maneuver could cause the

local minimum previously discussed.

6.2 Lyapunov Stability Analysis

For the cases considered, OTC will be a modification to the quaternion regulator

control logic from reference [54] shown in Eq.(5-12). It should be noted that this

modification could, in theory, work with any control algorithm which in turn can be

combined with any steering algorithm for SGCMGs. Therefore, it is not restricted to any

steering algorithm or the quaternion regulator control law if the proper stability analysis

is carried out. The quaternion regulator control logic assumes perfect information and

has the following nominal form

h = Ke + Cw + H (6-1)

where K = 2kJ and C = cJ are positive-definite symmetric gain matrices based on the

spacecraft's centroidal inertia J, e is the vector elements of the quaternion error vector,


w is the spacecraft angular velocity, and H is the total spacecraft system centroidal

angular momentum from Eq.(2-1).

Recall from Eq.(4-12) that the Jacobian's left singular vectors U is an orthonormal

basis for the output torque h. This basis is composed of a unit vector in the direction

of the singular direction ull when at singularity, and two unit vectors orthogonal to the

singular direction, u_ and u, (i.e., even when A is nonsingular, the basis from U still

exists). Utilizing this basis in the formation of the output torque yields

h* = Pull + u + Cu, h (6-2)

with coefficients

p= hTll
h = h Tu + ag(m) (6-3)

S= iun bg(m)

The quantity g(m) is a augmentation to the orthogonal to the singular direction

components of torque that is an explicit function of the singularity measure. It will

be referenced as the OTC singularity parameter and a and b are switching elements

defined by

1 if h u > 0

-1 if nun < 0

Substituting Eq.(6-2) into Eq.(5-23) and bounding yields,


10 0
V =< TJU O 1 0 UTJW + 2 = -_ TR2W + 2 (6-4)
0 0 A1

where C2 = C1 + g(m)wTK-l(au1 + bun) and R2 = R1 in Eq.(5-25) with the singularity

parameter A from Eq.(4-5) in place of a for the HSL in Eq.(5-1). Similar to the Lyapunov

analysis in Chapter 5 for HSL, the error z is bounded with a Volterra integral expression


| |z| 2 < V(0) exp( x ( 1 exp(_ (t) r) d (6-5)
Amin Amin to
Since the SGCMG output torque will always have components orthogonal to the singular

direction when near singularity, it is assumed that a system using OTC will never

encounter gimbal-lock up to a specific size of ||lel and A from Eq.(4-5) where

g(m) = Allel| (6-6)

Therefore, situations of singularity other than those with gimbal-lock are of concern.

When at singularity, the expression for the error is

|I|z|12 < + e r)d (6-7)

because the error z is based off the transient term of the Lyapunov equation Vs from

Eq.(7-2) and the dynamic term containing the effect of the torque error added for

gimbal-lock escape ft C2(r)dr. Recall from Chapter 5, while using HSL, that when

singularity occurs with the exception of gimbal-lock, there may be a period of instability

and it is assumed that the maneuver does not end at singularity. With this in mind, the

steady-state error of the of a system away from singularity using SDA combined with
OTC is bounded as


iz(oo)l12 < m 2(00) (6-8)

through the use of L'Hospital's rule as in Eq.(5-33).

The result of OTC is UUB for sufficiently large the choices of c rather than (. This is

true because short periods of instability may arise, but the negative semi-definite term

of Eq.(6-4) becomes negative definite away from singularity and will become dominate

for sufficiently large values of c. With the choice of g(m) in Eq.(6-6), whenever there is

an attitude error and the system is in proximity to a singularity, there will be torque error

added orthogonal to the singular direction, and thus gimbal-lock will be escaped.

6.3 Numerical Simulation

For the steering algorithms of SDA, GSR, and SDA with OTC (SDA/OTC)

augmented to the attitude controller, two cases were simulated for a four-SGCMG

pyramidal cluster at 8 = 54.74 deg and the model parameters in Table 6-1: (1) a z-axis

maneuver starting at initially at the zero momentum configuration from Chapter 5 (i.e.,

6 = [0 0 0 O]T deg) and (2) a z-axis maneuver starting at gimbal-lock configuration (i.e.,
6 = [90 90 90 90]T deg). Both cases use the same control gains applied to a pyramidal

arrangement of four-SGCMGs. Also, the simulation was propagated with a discrete

fourth-order Runga-Kutta at a time-step of At = 0.02 sec.

Table 6-1. Model Parameters
Variable Value Units
J 13x3 kgm2
0 54.74 deg
eo [0 0 0.3]
wo [0 0 0]T deg/s
ho 1 Nms
k 2 1/s2
c 10 1/s
ei 0.1 rand(1)
Ao 0.1
At 0.02 sec


6.3.1 Case 1: 60 = [0 0 0 0]T deg

For a z-axis maneuver originating from an initial zero-momentum configuration,

Figures 6-2 show that SDA, GSR, and the SDA/OTC appear identical. This is what

is expected for a maneuver far from singularity. The results for the torque Figure 6-3

confirm this because the transient response for this case (i.e., away from singularity) is

short. The transient response of the output torque shown in Figure 6-3 D of SDA/OTC

has significant jitter but with a small magnitude. This jitter has negligible effect on the

gimbal rates shown in Figure 6-2 D which is due to the mapping of the output torque

onto the gimbal rates. The difference in quaternion error and singularity measure in

Figures 6-4 and 6-5 are small. This should not be surprising since for SDA, GSR, and

SDA/OTC, the contributions of torque error are designed to be significant only when the

system is close to a singularity.

The OTC singularity parameter shown in Fig. 6-6, while initially nonzero for this

case, converges to zero rapidly. The fact that this parameter is nonzero initially and

there is no significant differences in the quaternion error responses as shown in Figure

6-7, might suggest that the torque error from the SDA method itself was dominant. In

addition, it should be noted that the difference in quaternion error responses while small

(10-8), is not on the order of machine precision (10-16) or (10-32).
















80 100


1 2 3 4
B SDA (transient response)



20 40 60

80 100


1 2 3 4
D GSR (transient response)



20 40 60

80 100

0 1 2 3 4
F SDA/OTC (transient response)

Figure 6-2. Gimbal rates


20 40 60











0 20 40 60 80 100





0 20 40 60 80 100




. 0.8
z 0.6
c 0.4


0 20 40 60 80 100

B SDA (transient response)


0 0


1 2 3 4 5
D GSR (transient response)

F SDA/OTC (transient response)

Figure 6-3. Output torque





a 0.15



0 20 40 60 80 100




a< 0.15


0 20 40 60 80 100




a 0.15



0 20 40 60 80 100

Figure 6-4. Vector elements of the error quaternion


E 1.09



0 50 100



E 1.09



0 50 100



E 1.09


0 50 100

Figure 6-5. Singularity measure


0 10 20

Figure 6-6. Singularity parameter (OTC)

x 10-9




0 10 2

0 10 20

30 40

0 10 20 30 40

Figure 6-7. Quaternion error difference: (A) eGSR eSDA (B) eSDA/OTC eSDA

6.3.2 Case II a: 60 = [90 90 90 90]T deg

For this case, the gimbals are initially oriented such that the system is in a

gimbal-lock configuration. Figure 6-8 A shows that the gimbal rates of the SDA method

are unchanged throughout the simulation since the system starts in a gimbal-lock

configuration and SDA cannot generate the necessary commands to escape. The

gimbal rates for GSR and SDA/OTC (Fig. 6-8 B and D), however, are nonzero because

the addition of the torque error has provided the system with the ability to escape

gimbal-lock. In addition, the controller approaches the original quaternion regulator

controller as the system moves away from singularity.


30 40

The transient response of the gimbal rates for the GSR and SDA/OTC in Figure

6-8 C and E are both oscillatory with GSR having the higher amplitude and duration.

This is attributed to the fact that unlike OTC, the functions adding torque error in GSR

for gimbal-lock escape are not clearly visualized (i.e., depend on the combination of

sinusoids with possible different frequencies and phases for dither) when mapped to the

gimbal rates.

Figures 6-9 and 6-10 show, respectively, the required torque and attitude error. In

both cases, the results show a similar trend as the gimbal rates for GSR, and SDA/OTC

approach zero;the GSR and SDA/OTC were able to generate the torque required to

drive the attitude error to zero.

An examination of the singularity measures shown in Figure 6-11 reaffirms the

responses shown in Figures 6-8 through 6-10 where the SDA remains at singularity

unlike GSR and the SDA/OTC which escape singularity but transition back to it

as the maneuver is completed. This transition back to singularity is common for all

pseudo-inverse steering algorithms, which work by approaching a singular configuration

and then making a rapid transition for escape [47]. Recall previously from 6.1, that it

was stated that it is possible to end in the vicinity of a singularity when the maneuver

was completed; this is an example of such a case shown in Figure 6-10 and 6-11.

The measure of how far the system of SGCMGs is from gimbal lock can be found

as the norm IIATh 0. Both the GSR and the SDA/OTC were successful in escaping

gimbal-lock as shown in Figure 6-12. It should be noted that because this measure is a

function of h, it goes to zero as the maneuver is completed.

The OTC singularity parameter is shown in Figure 6-13. It has a nonzero initial

value and converges rapidly to zero which makes it effective for helping in singularity





-0.1 50 100

0 -0.5




0 50 100 0 1 2 3 4 5
times) times)

B GSR C GSR (transient response)

o -1
1 -2 I


20 2 40 60 80 100 0 1 2 3 4 5
times) times)
D SDA/OTC E SDA/OTC (transient response)

Figure 6-8. Gimbal rates




0 20 40 60 80 100

0 20 40 60 80 100



z 0.6


0 20 40 60 80 100

C GSR (transient response)

0 1 2 3 4
E SDA/OTC (transient response)

Figure 6-9. Output torque






W 0.15



0 20 40 60 80 100




a 0.15



0 20 40 60 80 100

0 20 40 60 80 100

Figure 6-10. Vector elements of the error quaternion






00 50 100



E 0.4


0 20 40 60 80 100


E 0.4


0 20 40 60 80 100

Figure 6-11. Singularity measure










n 10-20

10 20

0 10 20





0 1 20,

0 10 20

30 40

30 40

30 40

Figure 6-12. Gimbal-lock measure


E 0.015

0 10 20 30 40
Figure 6-13. Singularity parameter (OTC)

6.3.3 Case II b (HSL/OTC): 60 = [90 90 90 90]T deg
Recall, from 6.1 that OTC can be used in combination with any steering algorithm
for gimbal-lock avoidance/escape. This case verifies through simulation that this is
indeed true by comparing HSL/OTC to GSR starting at gimbal-lock (60 = [90 90 90 90]T
deg). The HSL parameters are shown in Table 6-2.
Table 6-2. Hybrid
Variable Value
ao 0.01
0o 2
Pil 1
P2 1
a 1
b 3
mo 0.5

With the exception of the initial transient, the gimbal rates for GSR and those of
HSL/OTC in Figure 6-14, are approximately the same magnitude.



0 U.0



V -2 V \
o "o I
0 50 100 0 1 2 3 4 5
times) times)
A GSR B GSR (transient response)

3 0.5



0 0

0 20 40 60 80 100 0 1 2 3 4 5
times) times)
C HSL/OTC D HSL/OTC (transient response)

Figure 6-14. Gimbal rates

The transient response of the gimbal rates for the GSR in Figure 6-14 B is highly

oscillatory and not as smooth as that for HSL/OTC (compare with Figure 6-14 D). This

is due to oscillatory behavior of the dither-used for gimbal-lock escape that may be of

any duration depending on the frequencies and phases of the off-diagonal components

of the GSR perturbation matrix E. The HSL method acts as a SDA method, but when

combined with OTC will avoid/escape a singularity at the speed of the parameters

chosen for A in Eq.(6-6) in the which the duration will be understood for all singularities

and their combinations to the norm of quaternion error. Figures 6-15 and 6-16 show,

respectively, the required torque and attitude error. In both cases, the results show that

both methods (GSR and HSL/OTC) were successful in meeting the required torque and

completing the attitude maneuver.

20 40 60



c 0.4


0 20 40 60

0 20 40 60

80 100

B GSR (transient response)

80 100

0 1 2 3 4
D HSL/OTC (transient response)

Figure 6-15. Output torque

The singularity parameters for both methods in Figure 6-17, escape singularity

although they transition back to it as the maneuver is completed. Recall, it was

mentioned previously that a maneuver can be completed (i.e., e 0) while the

gimbal angles settle into a singular configuration; Figure 6-11 and 6-17 show this trend

(compare with Figure 6-11).

Figure 6-18 shows instantaneous escape from gimbal-lock for both methods.





- 0.4


40 60 80 100

40 60 80 100


20 40 60

80 100

Figure 6-16. Vector elements of the error quaternion

40 60 80 100 0 20 40 60
times) times)

Figure 6-17. Singularity measure

The OTC singularity parameter shown in Figure 6-19 has an initial nonzero value
and converges rapidly to zero similarly to Figure 6-13, which makes it effective for
helping in singularity escape.
The Table 6-3 compares the root-mean squared (RMS) gimbal rates (rad/s),
tracking performance in terms of RMS torque error (Nm), and pointing performance
in terms of the norm of the steady-state error quaternion for GSR, SDA/OTC, and
HSL/OTC. From Table 6-3 it can be shown that the choice of the singularity threshold
mo has an effect on the tracking and pointing performance of the HSL method combined


a 0.15

0 20



0.15 0.15
0.1 0.1

0.05 0.05

0 10 20 30 40 0 10 20 30 40
times) times)
Figure 6-18. Gimbal-lock measure

E 0.015

0 10 20 30 40
Figure 6-19. Singularity parameter (OTC)

with OTC. In fact, when this value is mo = 0.5 for this model and with the set of control
gains that differ from the model in Chapter 5, the tracking and pointing performance of
HSL/OTC is actually worse. This is expected as shown by the Lyapunov analysis in 6.2
where the steady-state error of SDA is dependent on the torque error added into the
system; and a larger threshold value of mo will increase the steady-state error.
6.4 Orthogonal Torque Compensation Summary
Orthogonal torque compensation (OTC) methodology was developed to ensure
escape from all singularities, particularly scenarios involving gimbal-lock configurations.
The compensation methodology can be incorporated with any attitude controller/steering



Table 6-3. Performance Comparisons
Steering Algorithm 6RMS TeRMS I essII
GSR 7.7901 10.1864 0.0024
SDA/OTC 7.3146 10.4080 0.0020
HSL/OTC (mo = 0.5) 7.1693 6.1819 0.0038
HSL/OTC (mo = 0.05) 7.0041 3.0416 2.1474e-09
HSL/OTC (mo = 0.005) 6.9166 2.9852 1.8843e-09

logic combination and was shown through analysis to ensure stability with sufficiently

large choice of the controller gain c. Since the compensator was designed to work with

any attitude controller, then it is compatible with any steering algorithms. This could

prove very beneficial for steering algorithms like HSL which reduce the amount of torque

error at hyperbolic singularities (see Chapter 5). The OTC was also demonstrated

through numerical simulation where it was shown to be effective in escaping gimbal-lock

with near zero steady-state attitude error. These simulations were based on a

four-SGCMG pyramidal arrangement using an quaternion regulator controller combined

with the steering algorithms SDA and HSL and compared with GSR.


7.1 Scalability Problems with SGCMG Hardware
Currently available CMG actuators are shown in Figure 7-1 with specifications from
Table C-1 in Appendix C do not meet the power, mass, and volume requirements for
satellites smaller than the micro-sat class. Currently, development of CMG hardware
underway will meet some of the constraints for these smaller classes of satellites. New
steering algorithms to complement these newly developed CMG hardware is not being
emphasized and will have a major effect on how systems of miniature CMGs perform.
This chapter highlights the effect of scaling on the performance of miniature CMGs.


Power and Mass

Figure 7-1. Off-the-shelf CMGs


Torque "0 to 50 Nm

CM s
Be^^ ing^^^

Power ~ 0 to 20 W
Mass 0 to 15 kg

7.2 Effect of g,, on Torque Error

The gimbal accelerations are kinematically dependent on the choice of the gimbal

rates and as a consequence, only one of them can be used as a control variable.

Therefore, the gimbal rates are considered as measurable quantities and the gimbal

accelerations are the control. The solution to the gimbal accelerations as a control is

defined as

= BT(BBT)-1[T A26] (7-1)

where A = A1 + A2 from Eqs.(2-7) and (A-26) and T = h + A16 equivalent to

Eq.(2-13). This solution is considered an exact solution but for some cases may be

highly oscillatory and/or unstable for the gimbal rates and accelerations. A Lyapunov

analysis is presented below.

It was stated previously that the direct solution in Eq.(7-1) may be unstable. To

prove this we start with the given candidate Lyapunov function

V1 = -w'K-1Jc + eTe + (1 e4)2 (7-2)
Taking the time derivative, yields

V = TK-1[-T wx H] + Te (7-3)

For the system to be globally asymptotically stable (i.e., h, e, and w -> 0 as t oo)

T = A + B6 = Ke + Cw w H (7-4)

Next, a second candidate Lyapunov function is required to analyze the behavior of the

gimbal rates and accelerations as time approaches infinity.

2 (7
V2 = _T6 (7-5)


Taking the time derivative utilizing Eq.(7-1), we obtain

V2= 6 6 = TBT(BB T)- [T A6] (7-6)

From the previous Lyapunov analysis where h, e, and w 0 as t oo it can be

assumed that

V2= -TBT(BBT)- -1A (h )6TS (7-7)
where B = IgwB, A = hoA, and S = -BT(BBT)-A. Matrix S is semi-indefinite and

therefore the gimbal rates can be unstable. Furthermore, a Lyapunov analysis of V,+ V2
shows that the ratio ( ') plays a key role in the stability of the whole system.

Next, we consider the use of the SR inverse where the gimbal rates are found from

6sR = ASR(T Igw^) (7-8)

ASR AT(AAT + A)-1 (7-9)

where A is the singularity parameter defined in Eq.(4-5).

Assuming the SR inverse is used to apply the gimbal rates as a control variable we

find that the torque error is expressed as

Te = Tct T = hoA6sR + IgwB T (7-10)


AASR = [ + A(AAT)-]-1 (7-11)

Away from singularity, a series expansion of Eq.(7-11) with only the linear terms yields


This series expansion is convergent if away from singularity because the term

I(AAT)-11 < 1. Substituting Eq.(7-12) into the torque error, Eq.(7-10) we have

Te I/gw,(AAT)-1 A(AAT)-1T (7-13)

It can be seen that the torque error may be amplified by the magnitude of the gimbal-flywheel

inertia Igw. Furthermore, if 6 or Ig, is considered negligible then the torque error is only

affected by the singularity parameter A, the distance from singularity which is related to

the determinant of (AAT)-1, and T. It should be noted that an increase in g,, is followed

by a decrease in AAT, but its effectiveness in lowering the torque error requires a large

ratio of 1- (i.e., effective when >> 1 which could be thought of as being a system of


The eigen-axis control logic from Eq.(5-12) is used to define the torque needed for

a given maneuver to be mapped onto the gimbal states. The SGCMG system proposed

in this analysis assumes that it is self-contained and therefore the metric of the hosted

algorithm performance is independent of the control logic chosen as long as it meets the

constraints of the SGCMG actuators. Therefore, no generality is lost for the choice of the

control logic in the analysis.

7.3 Numerical Simulation

The cases compared here are the SR Inverse and a filtered gimbal-acceleration

control law based on Oh and Vadali [84]. The filtered gimbal acceleration control law has

the following form

6= K6(6sR 6) 6R (7-14)



where K6 is the gain matrix that sizes the amount of gimbal acceleration utilized for

control and 6SR and 6SR are the gimbal rates and accelerations from the SR inverse and

the time derivative of that rate. The effect of the gimbal-flywheel inertia is scaled in the

simulation by the gain Kg, (i.e., gw, = Kg1wgw) where Kg, = 0 signifies that their is no

torque or angular momentum contributed from the gimbal dynamics.

Simulations of these two steering algorithms were compared by scaling g,, through

two different values of Kg,: i) Kg, = 0 and ii) Kg, = 2. The model parameters for the

nominal satellite inertia J and gimbal-flywheel inertia g,, are based on a four-SGCMG

pyramidal arrangement sized for a 1 U CubeSat. Both simulations were for a maneuver

of 1800 about the z-axis. The initial gimbal angles for all simulations are 60 = [-90
-90 -90 -90]T deg corresponding to a elliptic saturation singularity about the z-axis.

This set of initial gimbal angles along with the required maneuver will force the system

to enter gimbal-lock (i.e., ATh = 0) and accumulate a steady state attitude error. This

situation was chosen to test the system to its performance limit. The parameters that

were used for all of the results are shown in Table 7-1.

Table 7-1. Model Parameters
Variable Value Units
533.8 0 0
Js 0 533.8 0 x 10-6 kgm2
0 0 895.6
0 52 deg
eo [0 0 1 0]T
Wo [0 0 0]T deg/s
ho 4.486 x 10-4 Nms
7gw 5.154 x 10-6 kgm2
k 10 1/s2
c 50 1/s
K6 10 14x4
Ao 0.5
/P 10
60 [0 0 0 0]T deg/s


7.3.1 Case I: Kg,= 0

For this case, Figures 7-2 and 7-3 show that the gimbal rates and accelerations are

quite similar for both the SR inverse and filtered acceleration control law, except at the

very beginning. Since the filtered acceleration control law also uses an SR inverse in

its formulation, it is reasonable to assume that the differences can be attributed to the

filter (see Eq.(7-14) ). The torque errors in Figure 7-4 are initially larger for the filtered

acceleration algorithm as compared to the SR inverse by itself. Again, this is most likely

due to the lower initial gimbal rates and accelerations attributed to the filter. Also, the

torque error for both cases has a steady state offset where the system encounters




A SR Inverse

-5 1



C Filtered gimbal acceleration



0 2 4 6 8
B SR Inverse (transient response)

0 2 4 6 8 10
D Filtered gimbal acceleration (transient re-

Figure 7-2. Gimbal rates for Kg, = 0



A SR Inverse


0 20 40 6(
C Filtered gimbal acceleration


0 2 4 6 8
B SR Inverse (transient response)



0 2 4 6 8


Filtered gimbal acceleration (transient re-

Figure 7-3. Gimbal accelerations for Kgw




E -1


A SR Inverse

x 10-3

E -4 '



0 20 40
C Filtered gimbal acceleration

Figure 7-4. Torque error for Kg, = 0

B SR Inverse (transient response)

x 10-3

0 5 10
D Filtered gimbal acceleration (transient re-


The singularity measure m shown in Figure 7-5 is identical for both methods in

these plots. The only discrepancy between the gimbal rates and accelerations of the

two methods for this cases was at the very beginning of the maneuver. Therefore, the

differences in m would not be obvious in these plots. The value of m here is shown

to transit away from but return to singularity in Figure 7-5 for both methods. This is

common to steering algorithms of the SR inverse type.

Time(s) Time(s)
A SR Inverse B Filtered gimbal acceleration

Figure 7-5. Singularity measure for Kg, = 0

7.3.2 Case II: Kg = 2

The initial gimbal rates and accelerations are less for the filtered acceleration

algorithm than for SR inverse itself. This is the shown in Figures 7-6 and 7-7.



0 0

-V -0.01

0 20 40 60 0 20 40
Time(s) Time(s)
A SR Inverse B SR Inverse (transient response)

Sx 10-3



-u -1

0 20 40 60 0 10 20 30 40 50
Time(s) Time(s)
C Filtered gimbal acceleration D Filtered gimbal acceleration (transient re-

Figure 7-6. Gimbal rates for Kg, = 2

A closer look at the transient response for the gimbal rates and accelerations

of the two methods is shown in Figures 7-6 B and D and 7-7 B and D. It is gathered

from Figure 7-6 B and 7-7 B that the steady-state response for the gimbal rates

and accelerations of the SR and filtered acceleration law inverse is nonzero. The

steady-state response for the gimbal rates and accelerations of the filtered acceleration

law although nonzero, is considerably smaller than that for the SR inverse which in turn

prevents the torque error from diverging.

However, Figure 7-8 shows the torque error for the SR inverse appears to diverge

where as the torque error for the filtered acceleration algorithm remains bounded.




r 0

_ -100

A SR Inverse

10 20 30 40 50
B SR Inverse (transient response)


0 20 40 60
C Filtered gimbal acceleration

Figure 7-7. Gimbal accelerations for Kg,

0 20 40
D Filtered gimbal acceleration (transient re-

Furthermore, it appears as though the precision is improved with larger Ig, for the

filtered acceleration algorithm. This may be due to the fact that it relies less on the SR

inverse which would be the source of the torque error in this example.



0 20 40
A SR Inverse

Figure 7-8. Torque error for Kg,

0 20 40
B Filtered gimbal acceleration

The singularity measures shown in Figure 7-9 are identical except that the

singularity is approached quicker for SR inverse. It should be noted that as the value of

Ig, is increased, the singularity measure approaches singularity later in the simulation

for the filtered steering algorithm (i.e., larger gimbal rates lead to larger gimbal angle

excursions which in turn, make the system approach singularity quicker).

x 10-10

x 10-10

Time(s) Time(s)
A SR Inverse B Filtered gimbal acceleration

Figure 7-9. Singularity measure for Kgw = 2

The gimbal-flywheel assembly inertia will also have a degrading effect on the actual

torque amplification of SGCMG actuators. This will be explained in the next section.

7.4 Effect of Igw on Torque Torque Amplification

The torque amplification of a single SGCMG can be described by its output torque

divided by the input torque as

,Tout| ll ||lh x6 + gw1,, 6
h (7-15)
1ini, I Ilw X h + lgwJ ll
From Eq.(7-15) it is seen that as the gimbal-flywheel inertia Igw oo, the other terms

in the equation become less dominant and the torque amplification converges to one.

This is undesirable for SGCMGs because at the point that the torque amplification

converges to one, the system essentially becomes a reaction-wheel system and the

benefits of using SGCMGs are lost. Fortunately, a system of SGCMGs of this scale

does not exist.

The scaling of SGCMGs does reduce the torque amplification. To show actually

how much the torque amplification is degraded by scaling, the value of torque amplification

is calculated for the IMPAC SGCMGs in Figure 2-2 with the parameters in Table 7-2.

Table 7-2. Model Parameters
Variable Value Units
533.8 0 0
Js 0 533.8 0 x 10-6 kgm2
0 0 895.6
0 52 deg
ho 4.486 x 10-4 Nms
Igw 5.154 x 10-6 kgm2
wmax 3 deg/s
6max 1 rad/s
6max 1 rad/s2

For this example,

II-otl Ia +holn + /w max
'out|| hJomax 'gwmax 43.4 (7-1 6)
117n 1 Wmaxho + Igw6max
Therefore, there is a significant value of torque amplification even when scaling as seen

by the result in Eq.(7-16).


7.5 Summary

Singularities from systems utilizing SGCMGs cannot be easily scaled when

describing the algorithms for their control. Just as the performance on a hardware

level for SGCMGs will eventually flatline with scaling, so will the use of current steering

algorithms for singularity avoidance. This chapter showed that some current steering

algorithms may have difficulty when the gimbal-wheel assembly inertia becomes

significant. In addition, this chapter also showed that as a consequence of significant

values for gimbal-wheel assembly inertia, the performance of the SGCMG system is

affected by the torque amplification approaching 1 as g,, oo.



Control of spacecraft attitude with single-gimbal control moment gyroscopes

(SGCMGs) is difficult and becomes more so with the scaling of these actuators to

small satellites. The research presented in this manuscript began with a discussion of

the dynamic model for systems containing CMGs and their singularities. For SGCMG

systems, singularities were classified and tools were developed to quantify the form

of the singularity. These tools provided insights into these singularities (i.e., singular

surfaces) and were used to quantify them mathematically.

The singularities associated with SGCMGs were discussed in detail and classified

by the tools developed. Through this discussion, it was shown a complete explanation

of SGCMG singularities is absent from the literature. For example, it was found that

the special case of where the singular direction s is along a gimbal axis 6 can occur for

rooftop arrangements when the rank of the Jacobian is 2. This was further shown to be

a degenerate case which could lead to degenerate hyperbolic singularities that were

previously neglected in the literature for systems of SGCMGs. Using linear algebra,

it was proven that rooftop arrangements are free of elliptic internal singularities but

still contained elliptic external singularities (i.e., all elliptic singularities do not have null

motion are are thus impassable by null motion) associated with angular momentum

saturation. Furthermore, degenerate hyperbolic singularities were shown to exist for

rooftop arrangements (i.e., degenerate hyperbolic singularities contain only singular null

solutions exist and are also impassable by null motion).

It was shown that selecting an arrangement of SGCMGS through choice of a

desirable angular momentum envelope is difficult. Thus, a method of offline optimization

was suggested in a very constrained case that will provide the best SGCMG arrangement

in terms of Euler angles. However, this method is not that applicable for real spacecraft

design, although, it suggested that the current common configurations do not necessarily

have the best performance.

Next, it was shown that legacy steering algorithms, which can be categorized into

the three families of singularity avoidance, singularity escape, and singularity avoidance

and escape, did not consider the form of internal singularity (i.e., hyperbolic or elliptic).

This was shown to be problematic when precise attitude tracking is required because

the same amount of torque error was used for both hyperbolic and elliptic singularities.

A Hybrid Steering Logic (HSL) was developed that takes into account the form of

singularity. This algorithm uses this knowledge to apply null motion from a local gradient

(LG) method for singularity avoidance when near a hyperbolic singularity and torque

error from Singular Direction Avoidance (SDA) when near a elliptic singularities. Through

analytic derivations and numerical simulations, HSL was shown to perform better (i.e.,

lower torque error at hyperbolic singularities than pseudo-inverse methods and the

ability to escape both elliptic and hyperbolic singularities unlike LG methods) than legacy

methods for precise attitude tracking when using a four-SGCMG pyramid arrangement

of SGCMGs. Also, HSL was shown to have computation of flops on the same order as

many legacy methods.

Gimbal-lock was shown to be a special case of singularity when the output torque

lies in the singular direction of the Jacobian. With the exception of the Generalized-Singularity

Robust (GSR) inverse, legacy steering algorithms are known to be ineffective in

escaping gimbal-lock. To provide other steering algorithms with the same benefit as

GSR, a attitude controller augmentation defined as Orthogonal Torque Compensation

(OTC) was developed. This method was shown to be successful in escaping gimbal-lock

by adding orthogonal components of torque error when at singularity. This method was

combined with two separate steering algorithms, simulated, and compared to GSR

where it was shown numerically to have a much smoother transient response for the

gimbal rates.


Finally, the problems with scaling SGCMGs were discussed. It was shown that

the performance of SGCMGs is degraded (i.e., a lower torque amplification) and

same legacy algorithms previously used on larger SGCMGs could be ineffective for

scaled SGCMGs. A mathematical proof was used to show that with the increase in the

gimbal-flywheel assembly inertia Ig, compared to the flywheel angular momentum ho

causes this degradation in performance and the ineffectiveness of SGCMG control with

use of the Singularity Robust (SR) inverse. The utility of scaled SGCMGs is still viable

because the approximate SGCMG torque amplification for a single acutator was shown

to be on the order of 50 which is far more than the less than one-to-one ratio for systems

of reaction wheels.



A.1 Assumptions

The dynamic formulation for single gimbal and variable speed control moment

gyroscope (CMG) actuators assumes the absence of friction and external torque

in the system (spacecraft including CMGs). In addition, it is also assumed that the

center of mass (cm) of each CMG is along its gimbal axis and therefore does not affect

the position of the overall cm of the system. These assumptions are valid for current

state-of-the-art CMGs.

A.2 Dynamics

The centroidal angular momentum of the system consisting of that from the

spacecraft and a single CMG is

H = hw + hG + hs/c (A-1)

with contributions from the flywheel hw, gimbal hG, and the spacecraft hs/c. The

flywheel and gimbal angular momentum are expressed as

hw = Iwih6 (A-2)


hG = g66 (A-3)

where the gimbal frame basis [6, -, 6] is related to the spacecraft body-fixed basis

through a 3-2-1 rotation through the angles [6, 0, b] by

h = (s6s c c6CcO)ebl (S6Cb C6SCO)eb2 (CSO)eb3 = h bl + h2eb2 + h3b3 (A-4)


i = c, s',e, + s' s,-e, + c06b3 = tlebl + t26b2 + t36b3

6 = -(c6s s6c ceO)bl (c6c s6s~ceO)b2+ (s6sO)eb3 = d1bl +d2b2 +d3b3 (A-6)

where c(.) = cos(.) and s(.) = sin(.) and [ebl, 6b2, 6b3] is the basis for the spacecraft

body frame. Therefore, equivalent vector components for these angular moment shown

in the spacecraft body-fixed basis are

hw = Iwj(hlbl + h2b2 + h3b3) (A-7)


hG = Ig3(dleb d2b2 d3eb3) (A-8)

where Iw and Ig3 are the first and third components of the flywheel and gimbal inertias.

The angular momentum from the spacecraft is expressed as the tensor product of the

spacecraft centroidal inertia dyadic Jc with the inertial spacecraft angular velocity w.

hs/c = Jc. (A-9)

The spacecraft centroidal inertia dyadic is

Jc = IG + JO mGW (rC ricl r 0 rc) (A-10)

where re is the position of the cm of a CMG's cm from the cm of the system expressed

rc = rlebl + rc2eb2 + rc3eb3 (A-1 1)

and the static spacecraft inertia dyadic Jo is made up of constant inertias (i.e.,

assuming that the cm of the CMGs lies along the gimbal axis) and the inertias due



to the gimbal-wheel assembly IG are time varying due to the rotation about the gimbal

axis. The expression of the static spacecraft inertia dyadic is

3 3
Jo = #bi eb (A-12)
i=1 j 1
where (ebi ebj) ebi = 0 and (ebi ebj) ebj ey= bi. It is assumed that the gimbal-wheel

assembly inertia is aligned with the principle axes and can be expressed as

IGW =Igh h + lg2T 0 + g36 0 6 (A-13)

3 3
6 6 = hihJb, 0 ib (A-1 4)
i=1 j 1

3 3
S0+ = titibi 0 bj (A-15)
i=1 j=1

3 3
6 = Y d d, dj bi bJ (A-16)
i=1 j 1
The equations of motion (EOM) assuming torque free motion (i.e., no external torques)

are found through taking the inertial time derivative of the total system centroidal angular

momentum in Eq.(A-1) as

dH, o
-dH ,i M H +l w x Hc =0 (A-17)


He = [(al + a21)6 b i + c Q] Jc W (A-18)

The final EOM for a single CMG that has a single gimbal is

[(all + a21)6 + b1j + c] +Jc -+W + w x H= 0 (A-1 9)

The Jacobian matrices all, a21, bl, and cl are



a21 -

bi =




c =- A (A-23)

where the CMG angular momentum h = hw + hG.

For a system of CMGs with a single gimbal, the EOM concatenated into matrix

which is a consequence of Eq.(A-19), is expressed as

[(A1 + A2)6 + B6 + CQ] + JcW + 0x He = 0


where for n CMGs the Jacobian matrices are represented as

A = [all, a12, a13, ...ain]

A2 = [a21, a22, a23, ...a2n]

B = [b b ,, b3,...b]

C = [c, C2, c3 ...cn]





This concludes the development of the EOM for a rigid body spacecraft system of n

CMGs which contain a single gimbal.



function [hx,hy,hz]



% This code is generate the singularity surfaces for a

% any general SGCMG cluster with skew angle theta or

% inclination angle phi(i) and spacing angle si(i)

% where i = num_CMG.


% The angles th(i) and si(i) are the Euler angles relating the

% spin axis of each CMG to the body frames X-axis


%Frederick Leve

%Last updated: 07/08/08

%_______________________________________--- --- --- --


% This function simulated the CMG algorithms

% _- - -


% hO = nominal SGCMG angular momentum (could be vector if each

% CMG does not have the same angular momentum


% th = vector of inclination angles

% si = vector of spacing angles


% hx = angular momentum of envelope in the x-direction

% hy = angular momentum of envelope in the x-direction


% hz = angular momentum of envelope in the x-direction

%_____________________________________________________________ %

%epsilon parameter vector for surface generation

%to show internal singular surface make one epsilon -1 instead of 1

num_CMG = length(si);

if length(hO)

if int_ext

% external singular surface

eps = ones(num_CMG,1);

elseif intext

% internal singular surface

% eps = [ones(num_CMG-l,1);-1];

% eps = [1 1 -1 1];

eps = [1 1 -1 1];


display('intext must be either 0 or 1...

for internal or external singular surface')



min_h0 = min(h0);

if intext == 0

% external singular surface

for i



elseif int

% internal


= hO(i)/min(hO);

.ext ==1

singular surface


for i = l:num_CMG-1

eps(i) = h0(i)/min(h0);


eps(num_CMG) = -hO(num_CMG)/min(hO);


display('intext must be either 0 or 1...

for internal or external singular surface')



for 1 = l:num_CMG

% The transformation C1 is about the inclination angle phi(i)

C1(:,:,l) = [cos(th(l)+3*pi/2) 0 -sin(th(1)+3*pi/2);

0 1 0;

sin(th(l)+3*pi/2) 0 cos(th(1)+3*pi/2)];

C2 (:, :,1)

[cos(si(l)) sin(si(l)) 0;

-sin(si(l)) cos(si(l)) 0;

g(:,l) = transpose(C1(:,:,l)*C2(:,:,1))*[1;0;0];


87 %total angular momentum at the singular states corresponding to singular

%direction u

H = zeros(3,1);

n = 100;



%number of grid point for unit sphere

%generate the unit sphere (domain of u)

redlight = 5;

trafficlight = zeros( n+1 n+1 );


for i = l:n+l

for j = l:n+l

u = [ x(i,j) ; y(i,j) ; z(i,j) ];

for k = l:num_CMG

%this is the cosine of angle


%both of

vectors since

unit norm

u_dotgk = abs(u'*g(:,k));

if ( udot_gk > 0.95 )




for i = l:n+l

for j = l:n+l

u = [ x(i,j) ; y(i,j) ; z(i,j) ]; %compose the

%singularity vector u

for k = l:num_CMG


H = H + eps(k)/norm(cross( g(:,k), u ) )...

*cross( cross( g(:,k), u ) g(:,k) );



hx(i,j) = H(1);

hy(i,j) = H(2);

hz(i,j) = H(3);

H = zeros(3,1);

%parse out the components of the

%momentum vector for later

%surface or mesh plotting.







Table C-1. Off-the-Shelf CMG Specifications
CMG Output Torque (Nm) Mass (kg)) Power (W)
M50 74.6 33.1 75.0
M95 128.8 38.56 129.0
M160 216.9 44.0 217.0
M225 305.1 54.0 305.0
M325 441.0 61.2 441.0
M325D 441.0 61.3 441.0
M600 813.5 81.6 814
M715 969.4 89.8 949.0
M1300 1762.6 125.2 1716
M1400 1898.1 131.5 1899.0
EHCMG 2304.9 146.4482 2306
MicroWheel-10S-E 0.01 1.1 5.0
CMG 15-45S 45.0 18.4 25.0



[1] Johann Bohennberger. Description of a machine for the explanation of the laws of
motion of earth around an axis, and the change in orientation of the latter. Tubinger
Blatter Fur Naturwissenschaften und Arzneikunde, 3:72-83, 1817.

[2] Simeon-Denis Poisson. Memoir on a special case of rotational movement of
massive bodies. Journal de I'Ecole Polytechnique, 9:247-262, 1813.

[3] Donald MacKenzie. Inventing accuracy: A historical sociology of nuclear missile
guidance. MIT Press, pages 31-40, 1990.

[4] P.C. Hughes and P. Carlisle. Spacecraft attitude dynamics. Number p. 156-184. J.
Wiley, 1986.

[5] H. Krishnan, N.H. McClamroch, and M. Reyhanoglu. Attitude stabilization of a rigid
spacecraft using two momentum wheel actuators. In In its Efficient Reorientation
Maneuvers for Spacecraft with Multiple Articulated Payloads 20 p (SEE N93-29988
11-18), 1993.

[6] J.D. Adams and S.W. McKenney. Gyroscopic roll stabilizer for boats, June 4 2003.
US Patent App. 10/454,905.

[7] II Skelton and C. Eugene. Mixed Control Moment Gyro and Momentum Wheel
Attitude Control Strategies. PhD thesis, Virginia Polytechnic Institute and State
University, 2003.

[8] A Pothiawala and MA Dahleh. Hoo Optimal Control for the Attitude Control and
Momentum Management of the Space Station. MIT Press.

[9] F Wu. Fixed-Structure Robust CMG Momentum Manager Design for the
International Space Station. In AIAA Guidance, Navigation, and Control Con-
ference and Exhibit, Denver, CO, Aug. 14-17, 2000.

[10] S.R. Vadali and H.S. Oh. Space Station Attitude Control and Momentum
Management- A Nonlinear Look. Journal of Guidance, Control, and Dynamics,
15(3):577-586, 1992.

[11] L.R. Bishop, R.H. Bishop, and K.L. Lindsay. Proposed CMG Momentum
Management Scheme for Space Station. In IN: AIAA Guidance, Navigation
and Control Conference, Monterey, CA, Aug. 17-19, Technical Papers, volume 2,

[12] P. Hattis. Predictive Momentum Management for the Space Station. Journal of
Guidance, Control, and Dynamics, 9(4):454-461, 1986.

[13] B. Wie and et al. New Approach to Attitude/Momentum Control for the Space
Station. Journal of Guidance, Control, and Dynamics, 12(5):714-721, 1989.


[14] R.H. Bishop, S.J. Paynter, and J.W. Sunkel. Adaptive Control of Space Station with
Control Moment Gyros. IEEE Control Systems Magazine, 12(5):23-28, 1992.

[15] R.H. Bishop, S.J. Paynter, and J.W. Sunkel. Adaptive Control of Space Station
During Nominal Operations with CMGs. In Proceedings of IEEE Conference on
Decision and Control, 30th, Brighton, United Kingdom, Dec. 11-13, volume 3, 1991.

[16] J. Zhou and D. Zhou. Spacecraft Attitude Control with Double-Gimbaled Control
Moment Gyroscopes. In Robotics and Biomimetics, 2007. ROBIO 2007. IEEE
International Conference on, pages 1557-1562, 2007.

[17] T. Yoshikawa. A Steering Law for Double Gimbal Control Moment Gyro System.
NASA TM-X-64926, 1975.

[18] B.K. Powell, G.E. Lang, S.I. Lieberman, and S.C. Rybak. Synthesis of Double
Gimbal Control Moment Gyro Systems for Spacecraft Attitude Control. In AIAA
Guidance, Navigation and Control Conference, Snowmass, CO, August 19-21,
1985, pages 71s-937.

[19] H.F Kennel. A Control Law for Double-Gimballed Control Moment Gyros Used for
Space Vehicle Attitude Control. Technical report, NASA, 1970.

[20] J. Ahmed and D.S. Bernstein. Adaptive Control of Double-Gimbal Control-Moment
Gyro with Unbalanced Rotor. Journal of Guidance, Control, and Dynamics,
25(1):105-115, 2002.

[21] D.J. Richie, V.J. Lappas, and G Prassinos. A Practical Small Satellite
Variable-Speed Control Moment Gyroscope For Combined Energy Storage and
Attitude Control. In AIAA/AAS Astrodynamics Specialist Conference and Exhibit
Aug 18-21, Honolulu, HI, 2008.

[22] D.J. Richie, V.J. Lappas, and B. Wie. A Practical Variable-Speed Control Moment
Gyroscope Steering Law for Small Satellite Energy Storage and Attitude Control. In
AIAA/AAS Astrodynamics Specialist Conference and Exhibit Aug 18-21, Honolulu,
HI, 2008.

[23] H. Yoon and P. Tsiotras. Spacecraft Line-of-Sight Control Using a Single
Variable-Speed Control Moment Gyro. Journal of Guidance Control and Dynamics,
29(6):1295, 2006.

[24] H. Yoon. Spacecraft Attitude and Power Control Using Variable Speed Control
Moment Gyros. PhD thesis, Georgia Institute of Technology, 2004.

[25] H. Yoon and P. Tsiotras. Spacecraft Angular Velocity and Line-of-Sight Control
Using A Single-Gimbal Variable-Speed Control Moment Gyro. In 2005 AIAA
Guidance, Navigation, and Control Conference and Exhibit, pages 1-22, 2005.


[26] H Schaub and JL Junkins. Singularity Avoidance Using Null Motion and
Variable-Speed Control Moment Gyros. Journal of Guidance, Control, and Dy-
namics, 23(1):11-16, 2000.

[27] H. Lee, I.H. Lee, and H. Bang. Optimal Steering Laws for Variable Speed Control
Moment Gyros. In 2005 AIAA Guidance, Navigation, and Control Conference and
Exhibit, pages 1-11,2005.

[28] H. Schaub, S.R. Vadali, and J.L. Junkins. Feedback Control Law for Variable Speed
Control Moment Gyros. Journal of the Astronautical Sciences, 46(3):307-328,

[29] H. Schaub and J.L. Junkins. CMG Singularity Avoidance Using VSCMG Null
Motion (Variable Speed Control Moment Gyroscope). In AIAA/AAS Astrodynamics
Specialist Conference and Exhibit, Boston, MA, pages 213-220, 1998.

[30] J.Y. Shin, KB Lim, and DD Moerder. Attitude Control for an Aero-Vehicle Using
Vector Thrusting and Variable Speed Control Moment Gyros. In AIAA Guidance,
Navigation, and Control Conference and Exhibit, pages 1-10, 2005.

[31] D. DeVon, R. Fuentes, and J. Fausz. Closed-Loop Power Tracking for an Integrated
Power and Attitude Control System Using Variable-Speed Control Moment
Gyroscopes. In AIAA Guidance, Navigation, and Control Conference and Ex-
hibit, AIAA, Aug, 2004.

[32] H. Yoon and P. Tsiotras. Spacecraft Adaptive Attitude and Power Tracking with
Variable Speed Control Moment Gyroscopes. Journal of Guidance Control and
Dynamics, 25(6):1081-1090, 2002.

[33] D.A. DeVon and R.J. Fuentes. Adaptive Attitude Control and Closed-Loop Power
Tracking for an Integrated Power and Attitude Control System using Variable
Speed Control Moment Gyroscopes. In AIAA Guidance, Navigation, and Control
Conference and Exhibit, Aug 15-18, San Francisco, CA., pages 1-23, 2005.

[34] J.L. Fausz and D.J. Richie. Flywheel Simultaneous Attitude Control and Energy
Storage Using a VSCMG Configuration. In Proceedings of the 2000 IEEE Interna-
tional Conference on Control Applications, 2000., pages 991-995, 2000.

[35] D.A. Bearden. Small-satellite costs. Crosslink, 2(1):32-44, 2001.

[36] E.B. Tomme. The Paradigm Shift to Effects-Based Space: Near-Space as a
Combat Space Effects Enabler, 2005.

[37] K. Schilling. Distributed small satellites systems in Earth observation and

[38] W.J. Larson and J.R. Wertz. Space Mission Analysis and Design. Microcrosm, 3rd
edition, 1999.


[39] B. Wie, D. Bailey, and C. Heiberg. Singularity Robust Steering Logic for Redundant
Single-Gimbal Control Moment Gyros. Journal of Guidance, Control, and Dynam-
ics, 24(5):865-872, 2001.

[40] B. Wie, D. Bailey, and C. Heiberg. Rapid Multitarget Acquisition and Pointing
Control of Agile Spacecraft. Journal of Guidance Control and Dynamics,
25(1):96-104, 2002.

[41] H. Kurokawa. A Geometric Study of Single Gimbal Control Moment Gyros. PhD
thesis, University of Tokyo, 1998.

[42] G. Margulies and JN Aubrun. Geometric Theory of Single-Gimbal Control Moment
Gyro Systems. Journal of the Astronautical Sciences, 26(2):159-191, 1978.

[43] N.S. Bedrossian, J. Paradise, E.V. Bergmann, and D. Rowell. Steering Law Design
for Redundant Single-Gimbal Control Moment Gyroscopes. Journal of Guidance,
Control, and Dynamics, 13(6):1083-1089, 1990.

[44] B Wie. Singularity Escape/Avoidance Steering Logic for Control Moment Gyro
Systems, August 2003.

[45] B. Hamilton and B. Underhill. Modern Momentum Systems for Spacecraft Attitude
Control. Advances in the Astronautical Sciences, 125:57, 2006.

[46] FB Abbott, B. Hamilton, T. Kreider, P. Di Leonardo, and D. Smith. MCS Revolution.
Advances in the Astronautical Sciences, 125:99, 2006.

[47] H. Kurokawa. Survey of Theory and Steering Laws of Single-Gimbal Control
Moment Gyros. Journal of Guidance Control and Dynamics, 30(5), 2007.

[48] T.A. Sands, J.J. Kim, and B. Agrawal. 2H Singularity-Free Momentum Generation
with Non-Redundant Single Gimbaled Control Moment Gyroscopes. In 45th IEEE
Conference on Decision and Control, pages 1551-1556, 2006.

[49] B. Underhill and B. Hamilton. Momentum Control System and Line-of-Sight
Testbed. Advances in the Astronautical Sciences, 125:543, 2006.

[50] D. Brown and M.A. Peck. Scissored-Pair Control-Moment Gyros: A Mechanical
Constraint Saves Power. Journal OF Guidance, Control, AND Dynamics, 31 (6),

[51] D. Brown. Control Moment Gyros as Space-Robotics Actuators. AIAA/AAS
Astrodynamics Specialist Conference and Exhibit 18 21 August 2008, Honolulu,
HI, 2008.

[52] H. Kurokawa. Constrained Steering Law of Pyramid-Type Control Moment Gyros
and Ground Tests. Journal of Guidance, Control, and Dynamics, 20(3):445-449,


[53] F Leve, G Boyarko, and N Fitz-Coy. Optimization in choosing gimbal axis
orientations of optimization in choosing gimbal axis orientations of a cmg attitude
control system. Infotech@Aerospace Conference April 6-10, Seattle WA, 2009.

[54] B. Wie and J. Lu. Feedback Control Logic for Spacecraft Eigenaxis Rotations Under
Slew Rate and Control Constraints. Journal of Guidance, Control, and Dynamics,
18(6):1372-1379, 1995.

[55] Honeywell. Dynamic cmg array and method. Patent 20070029447, 2009.

[56] J. Baillieul, J. Hollerbach, and R. Brockett. Programming and Control of
Kinematically Redundant Manipulators. In Decision and Control, 1984. The
23rd IEEE Conference on, volume 23, 1984.

[57] M.A. Peck, B.J. Hamilton, and B. Underhill. Method and System for Optimizing
Torque in a CMG Array, July 30 2004. US Patent App. 10/903,774.

[58] D.A. Bailey, C.J. Heiberg, and B. Wie. Continuous Attitude Control that Avoids CMG
Array Singularities, October 10 2000. US Patent 6,131,056.

[59] M.R. Elgersma, D.P. Johnson, M.A. Peck, B.K. Underhill, G. Stein, B.G. Morton, and
B.J. Hamilton. Method and System for Controlling Sets of Collinear Control Moment
Gyroscopes, November 30 2005. US Patent App. 11/291,706.

[60] L. Jones and M. Peck. A generalized framework for linearly-constrained
singularity-free control moment gyro steering laws. In AIAA Guidance, Naviga-
tion and Control Conference, 209.

[61] D. E. Cornick. Singularity Avoidance Control Laws for Single Gimbal Control
Moment Gyros. AIAA, pages 20-33, 1979.

[62] R. D. Hefner and C. H. McKenzie. A Technique for Maximizing Torque Capability of
Control Moment Gyro Systems. Astrodynamics, 54(83-387):905-920, 1983.

[63] B. Wie. Space Vehicle Dynamics and Control. AIAA, 1998.

[64] D. Jung and P. Tsiotras. An Experimental Comparison of CMG Steering Control
Laws. In Proceedings of the AIAA Astrodynamics Specialist Conference, 2004.

[65] M.D. Kuhns and A.A. Rodriguez. A Preferred Trajectory Tracking Steering Law
for Spacecraft with Redundant CMGs. In Proceedings of the American Control
Conference, 1995., volume 5, 1995.

[66] J.A. Paradise. Global steering of single gimballed control moment gyroscopes using
a directed search. Journal of Guidance, Control, and Dynamics, 15(5):1236-1244,

[67] J.A. Paradise. A Search-Based Approach to Steering Single Gimballed CMGs.
Technical report, NASA Johnson Space Center (JSC), Houston, TX, CSDL, 1991.

[68] S. Asghar, PL Palmer, and M. Roberts. Exact steering law for pyramid-type four
control moment gyro systems. In AIAA/AAS Astrodynamics Specialist Conference
and Exhibit, Keystone, CO, 2006.

[69] Y. Nakamura and H. Hanafusa. Inverse Kinematic Solutions with Singularity
Robustness for Robot Manipulator Control. ASME, Transactions, Journal of
Dynamic Systems, Measurement, and Control, 108:163-171, 1986.

[70] K.A. Ford and C.D. Hall. Singular Direction Avoidance Steering for Control-Moment
Gyros. Journal of Guidance Control and Dynamics, 23(4):648-656, 2000.

[71] A.N. Pechev. Feedback-based steering law for control moment gyros. Journal of
Guidance Control and Dynamics, 30(3):848, 2007.

[72] F Leve, G Boyarko, J Munoz, and N Fitz-Coy. Comparison of state-of-the-art
steering logics for single-gimbal control moment gyroscopes. In AAS Astrodynam-
ics Conference, Pittsburgh, PA, 2009.

[73] T.A. Sands. Fine Pointing of Military Spacecraft. PhD thesis, Naval Postgraduate
School, March 2007.

[74] S.R. Vadali, S.R. Walker, and H.S. Oh. Preferred gimbal angles for single
gimbal control moment gyros. Journal of Guidance, Control, and Dynamics,
13(6):1090-1095, 1990.

[75] H. Leeghim, H. Bang, and J.O. Park. Singularity avoidance of control moment gyros
by one-step ahead singularity index. Acta Astronautica, 2008.

[76] D.N. Nenchev and Y. Tsumaki. Singularity-Consistent Attitude Motion Analysis
and Control Based on Euler Angle Parameterization. In SICE Annual Conference,
volume 1, 2003.

[77] M.D. Kuhns and A.A. Rodriquez. Singularity Avoidance Control Laws For a Multiple
CMG Spacecraft Attitude Control System. In American Control Conference,
volume 3, 1994.

[78] J.S. Lee, H. Bang, and H. Lee. Singularity Avoidance by Game Theory for Control
Moment Gyros. In AIAA Guidance, Navigation, and Control Conference and
Exhibit, pages 1-20, 2005.

[79] S.R. Vadali. Variable-Structure Control of Spacecraft Large-Angle Maneuvers.
Journal of Guidance, Control, and Dynamics, 9(2):235-239, 1986.

[80] W MacKunis, F Leve, A Waldrum, N Fitz-Coy, and W Dixon. Adaptive Neural
Network Satellite Attitude Control with Experimental Validation. IEEE Transactions
on Control Systems Technology, 2009.


[81] W. MacKunis, K. Dupree, N. Fitz-Coy, and W. Dixon. Adaptive Satellite Attitude
Control in the Presence of Inertia and CMG Gimbal Friction Uncertainties. In
American Control Conference, 2007.

[82] F Leve. Development of the spacecraft orientation buoyancy experimental kiosk.
Master's thesis, University of Florida, 2008.

[83] A. Fleming and I.M. Ross. Singularity-Free Optimal Steering Of Control Moment
Gyros. Advances in the Astronautical Sciences, 123, 2006.

[84] H.S. Oh and SR Vadali. Feedback control and steering laws for spacecraft using
single gimbal control moment gyros. Master's thesis, Texas A& M University,



Frederick Aaron Leve was born in Hollywood, Florida, in 1981. In August 2000 he

was accepted into the University of Floridas Department of Aerospace Engineering

in the College of Engineering where he pursued his bachelors degrees in Mechanical

and Aerospace Engineering. After completing his bachelors degrees in May 2005, he

was accepted into the masters program in aerospace engineering at the University of

Florida. While in the masters program, he received two awards in academia. In January

2007, he received the American Institute of Aeronautics and Astronautic's Abe Zarem

Award for Distinguished Achievement in Astronautics. For this award he was invited

to Valencia, Spain, where he competed in the International Astronautical Federations

International Astronautical Congress Student Competition. Here he received the silver

Herman Oberth medal in the graduate category. He completed the masters program in

May 2008 and continued on to his PhD. In May 2006, he was accepted to the Air Force

Research Lab (AFRL) Space Scholars Program, where spent his summer conducting

space research. After space scholars, he was employed as a student temporary

employee at AFRL where he received the Civilian Quarterly Award for all of AFRL in his

category. Currently he works in the Guidance, Navigation, and Control group at AFRL

Space Vehicles Directorate. His interests include, applied math, satellite attitude control,

satellite pursuit evasion, astrodynamics, and orbit relative motion.




2010FrederickA.Leve 2


Dedicatedtomymotherforalwaysbeingtheretosupportme 3


ACKNOWLEDGMENTS IwouldliketothankrstmyadvisorDr.NormanFitz-CoyforprovidingmewiththeguidanceandknowledgeforthisgreatresearchIundertook.Second,IwouldliketothankmycommitteemembersDr.WarrenDixon,Dr.AnilRao,Dr.WilliamHagerfromUF,andDr.ScottErwinfromtheAirForceResearchLabSpaceVehiclesDirectorate.Mycommitteecomprisestheexpertiseintheareasofresearchthatwouldprovidemethebestopportunityformyresearch.Last,butnotleast,Iwouldliketothankmycolleaguesinmyresearchlabwhoprovidedinputthroughoutmytimeasagraduatestudentthataidedinthisresearch:Dr.AndyTatsch,ShawnAllgeier,VivekNagabushnan,JosueMunoz,TakashiHiramatsu,AndrewWaldrum,SharanAsundi,DanteBuckley,JimmyTzuYuLin,ShawnJohnson,KatieCason,andDr.WilliamMackunis. 4


TABLEOFCONTENTS page ACKNOWLEDGMENTS .................................. 4 LISTOFTABLES ...................................... 8 LISTOFFIGURES ..................................... 9 ABSTRACT ......................................... 13 CHAPTER 1INTRODUCTION ................................... 15 1.1HistoryandBackground ............................ 15 1.1.1GyroscopicRateDetermination .................... 15 1.1.2SpinStabilizedSpacecraft ....................... 15 1.1.3SpacecraftAttitudeControlthroughGyrostats ............ 15 1.1.43-axisAttitudeControlofSpacecraft ................. 16 1.1.5Single-GimbalControlMomentGyroscopes(SGCMGs) ...... 17 1.1.6Double-GimbalControlMomentGyroscopes(DGCMGs) ...... 17 1.1.7Variable-SpeedControlMomentGyroscopes(VSCMGs) ...... 18 1.2ProblemStatement ............................... 18 2DYNAMICMODELS ................................. 20 2.1DynamicFormulation ............................. 20 2.2SingularSurfaceEquations .......................... 25 2.2.1EllipticSingularities ........................... 26 .................... 27 ................. 29 2.2.2HyperbolicSingularities ........................ 29 .......... 30 ............ 30 2.2.3Gimbal-Lock ............................... 30 2.3SingularitiesforSGCMGsMathematicallyDened ............. 31 3CONTROLMOMENTGYROSCOPEARRANGEMENTS ............ 34 3.1CommonSGCMGArrangements ....................... 34 3.1.1Rooftop ................................. 34 3.1.2Box .................................... 35 3.1.33 4Box .................................. 44 3.1.4ScissorPair ............................... 45 3.1.5Pyramid ................................. 46 3.2ChoiceofArrangement ............................ 48 3.3Simulation .................................... 50 5


4SURVEYOFSTEERINGALGORITHMS ..................... 54 4.1Moore-PenrosePseudo-Inverse ....................... 55 4.2SingularityAvoidanceAlgorithms ....................... 55 4.2.1ConstrainedSteeringAlgorithms ................... 56 4.2.2NullMotionAlgorithms ......................... 56 ..................... 56 ...... 57 ......... 58 4.3SingularityEscapeAlgorithms ........................ 59 .............. 59 ...... 60 ............ 60 ............... 62 ...... 63 4.4SingularityAvoidanceandEscapeAlgorithms ................ 64 ................... 64 ................. 64 4.5OtherSteeringAlgorithms ........................... 66 4.6SteeringAlgorithmComputationComparison ................ 66 5STEERINGALGORITHM-HYBRIDSTEERINGLOGIC ............. 68 5.1HybridSteeringLogic ............................. 68 5.1.1InternalSingularityMetrics ....................... 68 5.1.2HybridSteeringLogicFormulation .................. 69 5.2LyapunovStabilityAnalysis .......................... 71 5.3NumericalSimulation ............................. 76 5.3.1Case1:AtZeroMomentumConguration=[0000]Tdeg ... 79 .............. 80 ...... 82 .......... 85 5.3.2Case2:NearEllipticExternalSingularity=[105105105105]Tdeg .................................... 85 .............. 88 ...... 91 .......... 94 5.3.3Case3:NearHyperbolicInternalSingularities=[15105195)]TJ /F4 11.955 Tf 9.3 0 Td[(75]Tdeg ................................ 96 .............. 96 ...... 99 .......... 102 5.4HybridSteeringLogicSummary ....................... 104 6


6CONTROLALGORITHM-ORTHOGONALTORQUECOMPENSATION .... 106 6.1AttitudeControllerwithOTC .......................... 106 6.2LyapunovStabilityAnalysis .......................... 107 6.3NumericalSimulation ............................. 110 6.3.1CaseI:0=[0000]Tdeg ....................... 111 6.3.2CaseIIa:0=[90909090]Tdeg .................. 116 6.3.3CaseIIb(HSL/OTC):0=[90909090]Tdeg ............ 123 6.4OrthogonalTorqueCompensationSummary ................ 127 7SCALABILITYISSUESFORSGCMGS ...................... 129 7.1ScalabilityProblemswithSGCMGHardware ................ 129 7.2EffectofIgwonTorqueError .......................... 130 7.3NumericalSimulation ............................. 132 7.3.1CaseI:Kgw=0 ............................. 134 7.3.2CaseII:Kgw=2 ............................ 138 7.4EffectofIgwonTorqueTorqueAmplication ................. 142 7.5Summary .................................... 143 8CONCLUSION .................................... 144 APPENDIX ARIGIDBODYDYNAMICSFORMULATIONFORCONTROLMOMENTGYROSCOPEACTUATORS(SGCMG/VSCMG) .......................... 147 A.1Assumptions .................................. 147 A.2Dynamics .................................... 147 BMOMENTUMENVELOPECODE ......................... 151 CCONTROLMOMENTGYROSCOPEACTUATORSPECIFICATIONS ...... 156 REFERENCES ....................................... 157 BIOGRAPHICALSKETCH ................................ 164 7


LISTOFTABLES Table page 3-1ModelParameters .................................. 50 4-1AlgorithmFlopsm=row(A)andn=column(A) ................. 66 5-1ModelParameters .................................. 78 5-2PerformanceComparisonsforCaseI:ZeroMomentum ............. 85 5-3PerformanceComparisonsforCaseII:EllipticSingularity ............ 96 5-4PerformanceComparisonsforCaseIII:HyperbolicSingularity .......... 104 6-1ModelParameters .................................. 110 6-2HybridSteeringLogicParameters ......................... 123 6-3PerformanceComparisons ............................. 128 7-1ModelParameters .................................. 133 7-2ModelParameters .................................. 142 C-1Off-the-ShelfCMGSpecications .......................... 156 8


LISTOFFIGURES Figure page 2-1Rigidbodywithaconstantc.m. ........................... 21 2-2GimbalframeFGiofIMPACSGCMG(PatentPending) .............. 22 2-3SingularityshownwhenCMGtorquevectorslieinaplane(IMPACSGCMGsPatentPending) ................................... 25 2-4SingularitiesforSGCMGs .............................. 27 2-5Externalsingularsurfacesforafour-SGCMGpyramid .............. 28 2-6Internalsingularsurfacesforafour-SGCMGpyramid ............... 30 3-1Four-SGCMGrooftoparrangement ......................... 34 3-2Four-SGCMGboxarrangement ........................... 35 3-3Angularmomentumenvelopeforafour-SGCMGboxarrangement. ....... 36 3-4Planesoftorqueforafour-CMGrooftoparrangement .............. 37 3-5Torqueplanestracedoutforafour-SGCMGrooftoparrangement ........ 39 3-6Angularmomentumenvelopewithplottedangularmomentumcombinationsforthefour-SGCMGboxarrangement ....................... 42 3-7Degeneratehyperbolicsingularitiesforthefour-SGCMGboxarrangement ... 43 3-8Singularsurfacesshowing1h0singularityfreeregion ............... 45 3-93OrthogonalscissorpairsofSGCMGs ...................... 46 3-10Planesofangularmomentumandtorqueforafour-SGCMGpyramid ...... 47 3-11Four-SGCMGpyramidarrangement ........................ 47 3-12Optimizationprocessblockdiagram ........................ 49 3-13SingularsurfacesfortheoptimizedarrangementattheEulerangles=[170.213.685.5168.0]Tdegand=[17.7167.0304.392.5]Tdeg ...... 51 3-14Gimbalratesfortheoptimizedandpyramidarrangements ............ 51 3-15Torqueerrorfortheoptimizedandpyramidarrangements ............ 52 3-16Singularitymeasurefortheoptimizedandpyramidarrangements ........ 52 3-17Optimizationcostfortheoptimizedandpyramidarrangements ......... 53 9


4-1OuterandinnerloopsofGNCsystem ....................... 54 4-2Steeringalgorithms ................................. 56 5-1Zero-momentumcongurationofafour-SGCMGpyramidarrangement ..... 79 5-2SimulationresultsforLGwith0=a=b=1=0and2=0=1atzeromomentum ...................................... 80 5-3SimulationresultsforLGwith0=a=b=1=0and2=0=1atzeromomentum(contd.) ................................. 81 5-4SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1atzeromomentum ................................... 83 5-5SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1atzeromomentum(contd.) ............................... 84 5-6SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1atzeromomentum ................................. 86 5-7SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1atzeromomentum(contd.) ............................ 87 5-8SimulationresultsforLGwith0=a=b=1=0and2=0=1nearellipticsingularities .................................. 89 5-9SimulationresultsforLGwith0=a=b=1=0and2=0=1nearellipticsingularities(contd.) ............................. 90 5-10SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1nearellipticsingularities ............................... 92 5-11SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1nearellipticsingularities(contd.) .......................... 93 5-12SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearellipticsingularities .............................. 94 5-13SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearellipticsingularities(contd.) ......................... 95 5-14SimulationresultsforLGwith0=a=b=1=0and2=0=1nearhyperbolicsingularities ................................ 97 5-15SimulationresultsforLGwith0=a=b=1=0and2=0=1nearhyperbolicsingularities(contd.) ........................... 98 5-16SimulationresultsforSDAwith0=0.01,0=0,a=0,b=0,and=1nearhyperbolicsingularities ............................. 100 10


5-17SimulationresultsforSDAwith0=0.01,0=0,a=0,b=0,and=1nearhyperbolicsingularities(contd.) ........................ 101 5-18SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearhyperbolicsingularities ............................ 103 5-19SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearhyperbolicsingularities(contd.) ....................... 104 6-1Satelliteattitudecontrolsystemblockdiagram .................. 106 6-2Gimbalrates ..................................... 112 6-3Outputtorque ..................................... 113 6-4Vectorelementsoftheerrorquaternion ...................... 114 6-5Singularitymeasure ................................. 115 6-6Singularityparameter(OTC) ............................ 116 6-7Quaternionerrordifference:(A)eGSR)]TJ /F7 11.955 Tf 11.96 0 Td[(eSDA(B)eSDA=OTC)]TJ /F7 11.955 Tf 11.96 0 Td[(eSDA ........ 116 6-8Gimbalrates ..................................... 118 6-9Outputtorque ..................................... 119 6-10Vectorelementsoftheerrorquaternion ...................... 120 6-11Singularitymeasure ................................. 121 6-12Gimbal-lockmeasure ................................ 122 6-13Singularityparameter(OTC) ............................ 123 6-14Gimbalrates ..................................... 124 6-15Outputtorque ..................................... 125 6-16Vectorelementsoftheerrorquaternion ...................... 126 6-17Singularitymeasure ................................. 126 6-18Gimbal-lockmeasure ................................ 127 6-19Singularityparameter(OTC) ............................ 127 7-1Off-the-shelfCMGs ................................. 129 7-2GimbalratesforKgw=0 ............................... 135 7-3GimbalaccelerationsforKgw=0 .......................... 136 11


7-4TorqueerrorforKgw=0 ............................... 137 7-5SingularitymeasureforKgw=0 .......................... 138 7-6GimbalratesforKgw=2 ............................... 139 7-7GimbalaccelerationsforKgw=2 .......................... 140 7-8TorqueerrorforKgw=2 ............................... 141 7-9SingularitymeasureforKgw=2 .......................... 141 12


AbstractofDissertationPresentedtotheGraduateSchooloftheUniversityofFloridainPartialFulllmentoftheRequirementsfortheDegreeofDoctorofPhilosophyNOVELSTEERINGANDCONTROLALGORITHMSFORSINGLE-GIMBALCONTROLMOMENTGYROSCOPESByFrederickA.LeveAugust2010Chair:NormanFitz-CoyMajor:AerospaceEngineering TheresearchpresentedinthismanuscriptattemptstorstsystematicallysolvetheSGCMGsteeringandcontrolproblem.Toaccomplishthis,abetterunderstandingofsingularitiesassociatedwithSGCMGsisrequired. Nextbasedonabetterunderstandingofsingularities,aHybridSteeringLogic(HSL)isdevelopedandcomparedtothelegacymethodssingulardirectionavoidance(SDA)andlocalgradient(LG)methods.TheHSLisshownanalyticallyandnumericallytooutperformtheselegacymethodsforafour-SGCMGpyramidarrangementintermsofattitudetrackingprecision.However,allofthemethodsaresusceptibletogimbal-lock. AmethodreferredtoasOrthogonalTorqueCompensation(OTC)isdevelopedforsingularitieswithgimbal-lockinSGCMGs,whichareknowntopresentachallengetomoststeeringalgorithms.OrthogonalTorqueCompensationconditionstheattitudecontroltorquebyaddingtorqueerrororthogonaltothesingulardirectionwhenatsingularity.ThismethodcanbecombinedwithanysteeringalgorithmincludingHSLandisprovenanalyticallytobestableandescapesingularitieswithgimbal-lock. Finally,theproblemswithscalingsystemsofSGCMGsarediscussed.Itisfoundthatscalingproducesanincreaseinthegimbal-wheelassemblyinertiaoftheSGCMGswhichinturnincreasestheeffectofthedynamicsassociatedwiththeseinertias.Throughanalysisandnumericalsimulations,itisshownthatmoresignicantgimbal-ywheelinertiareducestheperformancebyincreasingtorqueerrorand 13


reducingtorqueamplicationofSGCMGs.Sincemostofthelegacyalgorithmsusedforsingularityavoidanceandescapeusethegimbalratesforcontrol,theperformanceisdegradedwhenthedynamicsfromthegimbal-wheelassemblyinertiaareincreased.Thisdegradedperformanceisshowntoalsobedependentontheratioofgimbal-wheelassemblyinertiatonominalSGCMGywheelangularmomentum.TheoverallresulthereisthatthesamelegacysteeringalgorithmsthatusegimbalratesforcontrolcannotbeusedforsystemsofSGCMGsofareducedscale. 14


CHAPTER1INTRODUCTION1.1HistoryandBackground 1.1.1GyroscopicRateDetermination Thebehaviorofgyroscopicsystemscomesfromtheprincipleofconservationofangularmomentum.Bothratedeterminationandrategenerationarepossiblethroughthisgyroscopicbehavior.Forexample,determinationofattituderatescanbefoundbymechanicalgyroscopes.Fromthesedevices,spacecraftangularratesareinferredfromtheirreactionontothegimbalsofthemechanicalgyroscope.TherstknowngyroscopeswerepassiveandtherstofthesewasdevelopedbyJohannBohnenbergerin1817[ 1 2 ].Theywerelaterdevelopedforthenavyasthenavalgyrocompass[ 3 ].Therearenowpassivegyroscopesusedfordeterminingangularratesdevelopedviamicro-electricalandmechanicalsystems(MEMS)thataresmallerthanthehumaneyecandetect.1.1.2SpinStabilizedSpacecraft Earlyspacecraftdidnothaveactiveattitudecontrolbutwherespinstabilizedabouttheirmajoraxisofinertiabyinitiatingaspinafterlaunch.Themajor-axisruleisrequiredfordirectionalstability(i.e.,thespacecraftmustspinaboutitsmajoraxis)[ 4 ].Thisprovidedgyroscopicstabilityordirectionalstabilitytothespacecraftwhichwasdenedasstabilityprovidedthespacecraftdoesnothaveenergydissipationcapability(i.e.,arigidbody).However,aspacecraftistrulynotarigidbody(e.g.,exiblebooms,solararrays,internalmovableparts,andoutgassing)whichmaydissipateenergyandbecomeasymptoticallystable.1.1.3SpacecraftAttitudeControlthroughGyrostats Agyrostatisanyrigidbodythathasattachedtoitawheelthatthroughtheconservationofangularmomentumprovideseitherrotationalstabilityorcontrol.Toprovideattitudestability,constantspeedywheelsknownasmomentumwheels(MWs) 15


wereaddedinternaltothespacecrafttoprovidegyroscopicstiffnesswhichinturnsuppliesattitudestability[ 5 ].Thissystemisanexampleofagyrostat.Thegyrostatwastherstfundamentaldynamicalsystemthatconsideredaspinningywheelwithinorattachedtoarigidbody.Inthistypeofsystemtheywheelimpartsangularmomentumstiffnesstothebodyinternallythroughtheprincipleoftheconservationofangularmomentum.Theuseofmomentumwheelsforgyroscopicstabilityhasfounditsuseforothervehiclesthanspacecraft(e.g.,boats,trains,buses[ 6 ]).1.1.43-axisAttitudeControlofSpacecraft Whenactiveattitudecontrolwasneeded,momentumwheelswereexchangedforreactionwheels(RWs)whichprovideareactiontorqueonthespacecraftthroughywheelaccelerations.Thesedeviceswerethenusedincombinationwiththrusterstoprovideattitudecontroltospacecraft[ 7 ].FurtherinnovationcametospacecraftattitudecontrolthroughtheintroductionofgimbaledMWsknownascontrolmomentgyroscopes(CMGs).Theyhavebeenusedinsatelliteattitudecontrolfordecadesduetotheirhighprecisionandpropertyoftorqueamplication(i.e.,largertorqueoutputontothespacecraftthantheinputtorquefromthegimbalmotors).Typically,theyhavebeenusedforlargesatellitesthatrequirehighagilitywhilemaintainingpointingprecisionandhaveevenfoundtheirusesonboardtheinternationalspacestation(ISS)[ 8 ].Controlmomentgyroscopesproduceagyroscopictorquethroughrotationofangularmomentumaboutoneortwogimbalaxes.Momentumwheels,RWsandCMGsallareknownasmomentumexchangedevicesbecausetheirtorqueisproducedthroughredistributionofangularmomentumfromtheCMGstothespacecraft.Controlmomentgyroscopescomeintwoclasses:1)thosewithasinglecontrollabledegreeoffreedomand2)thosewithmultiplecontrollabledegreesoffreedom. ThechoiceofaCMGorattitudeactuatoringeneraldependsontheneedsofthemission.AllCMGshavespecicchallengesassociatedwiththeiruse.ThechallengesofCMGsareasfollows:singleDOFCMGsknownassingle-gimbalcontrol 16


momentgyroscopes(SGCMGs)sufferfrominstantaneousinternalsingularities(i.e.,situationswithintheperformanceenvelopewhereagiventorquecannotbeproduced);multipleDOFCMGsknowndouble-gimbalcontrolmomentgyroscopes(DGCMGs)aremechanicallycomplexandhavesingularitiesknownasgimbal-lockwhentheirgimbalsalignthuseliminatingtheirextraDOF;othermultipleDOFCMGsknownasvariable-speedcontrolmomentgyroscopes(VSCMGs)aredifculttomitigateinducedvibration(i.e.,duetothevariationofywheelspeeds)andrequiremorecomplicatedcontrollaws,motordrivercircuitry,andlargerywheelmotors.1.1.5Single-GimbalControlMomentGyroscopes(SGCMGs) Single-gimbalcontrolmomentgyroscopeshavealongheritageofightonlargersatellitesandtheISS[ 9 15 ].TheyareknowntohavethehighesttorqueamplicationofallCMGs,arelessmechanicallycomplexthanDGCMGsandhavelessmathematicallycomplexdynamicsthanVSCMGs.Theseactuatorssufferfrominternalsingularitiesthatmustbehandledon-the-y,wheretorquecannotbegeneratedinaspecicdirection.Thereisnosinglemethodthathasbeenproventoavoidallinternalsingularitieswhiletrackinganarbitrarytorqueperfectly(i.e.,withouttheuseoftorqueerrororconstrainingthetorque).Thus,thereismeritinndingalternatesolutionstocontrolofSGCMGsforattitudecontrol.1.1.6Double-GimbalControlMomentGyroscopes(DGCMGs) Double-gimbalcontrolmomentgyroscopescontaintwocontrollableDOFsthroughtheirtwogimbalaxes.TheyarethemostmechanicallycomplexofCMGsalthough,theredundancyintheadditionalgimbalmayleadtolessthanthreeDGCMGsrequiredfor3-axiscontrol.ThebenetofthisredundancyislostforDGCMGswhentheyencounteragimbal-locksingularity.Gimbal-lockisencounteredwhenthegimbalaxesarealignedandarenolongerlinearlyindependent.Asaconsequence,theextracontrollableDOFinthiscaseislost.Effectivemethodsalreadyexistforavoidinggimbal-locksingularitiesassociatedwithDGCMGs[ 16 20 ]. 17


1.1.7Variable-SpeedControlMomentGyroscopes(VSCMGs) Variable-speedcontrolmomentgyroscopesutilizeanextracontrollableDOFthroughywheelaccelerations.AsaconsequenceoftheextraDOF,theywheelmotorsmustbelargeranditismoretroublesometoisolateunwantedinducedvibration.Inaddition,thisextradegreeoffreedommakesasystemoftwoormorenon-collinearVSCMGsfreefromsingularitiesthroughtheextradegreeoffreedom,(i.e.,atCMGinternalsingularities,theneededtorqueisprovidedbyywheelaccelerations).Severalalgorithmshavebeendevelopedthatareeffectiveinreducingtheamountofywheelaccelerationsusedandthusprovidingbettertorqueamplication[ 21 30 ].Inaddition,methodshavebeendevelopedtousetheVSCMG'sextracontrollableDOFtospindowntheywheelsandstoretheirkineticenergy.Thesemethodsareknownastheintegratedpowerandattitudecontrolsystem(IPACS)andtheywheelattitudecontrolandenergytransmissionsystem(FACETS)inliterature[ 31 34 ].1.2ProblemStatement AnewparadigmthatrequireshighlyagilesmallspacecraftisongoingthrougheffortsbygovernmentagenciesandlabssuchasOperationallyResponsiveSpace(ORS),AirForceResearchLaboratory(AFRL),theNationalReconnaissanceOrganization(NRO),andNationalAeronauticsandSpaceAdministration(NASA)toperformsuchmissionsasintelligence,surveillanceandreconnaissance(ISR),spacesituationalawareness(SSA),andspacesciencemissions(e.g.,theimagingofgammaraybursts)[ 35 37 ].ManyofthesemissionsareinLEOandrequirehigheragilityandattitudeprecisiontotracktargetsonearth.Attitudecontrolsystems(ACSs)basedonreactioncontroldevices(e.g.,thrusters)canachievegreatagilitybutcannotmeetthepointingrequirementsandspaceneededforpropellantstorageonsmallsatellites[ 38 ]. Single-gimbalcontrolmomentgyroscopesarebeingconsideredtheactuatorofchoicetoprovidehigheragilitytosmallersatellitesbasedontheirperceivedtorqueamplication.However,manyproblemsexistthatmustbesolvedpriortousing 18


SGCMGsforsmallsatelliteattitudecontrol.Traditionally,SGCMGshavebeenoversizedfortheirmissionandtheangularmomentumenvelopeconstrainedtoavoidinternalsingularitiesonlargersatellites.Forsmallersatellitesystems,however,theextravolumeandmassneededforoversizedSGCMGsmaybeunacceptable.Therefore,smallsatelliteSGCMGsshouldutilizemoreoftheentireangularmomentumenvelopewheresingularitiesmaybeencounteredinthemomentumspace(bothinternalandexternal).Thus,legacysteeringalgorithmsfromlargersatelliteapplicationsmaynotprovidethesameperformanceforsystemsofasmallerscalerequiringanewapproachforsteeringandcontrolofSGCMGs. Forthesucceedingchapters:Chapter 2 discussesandreviewsthefundamentalsofCMGdynamicsanddescribesthedifferentformsofsingularitiesassociatedwithSGCMGs;Chapter 3 describesthepossiblearrangementsforsystemsofSGCMGsandtheirdesirableandundesirablequalities;Chapter 4 providesthebackgroundonpreviouslypublishedmethodsofsteeringalgorithmsforimplementationofSGCMGs;Chapter 5 discussesthedevelopmentoftheHybridSteeringLogic(HSL)forSGCMGs;Chapter 6 discussesthedevelopmentofOrthogonal-Torque-Compensation(OTC)forgimbal-lockescapeofSGCMGs;Chapter 7 discussestheperformancedegradationencounteredwhenscalingsystemsofSGCMGs;andChapter 8 providesconclusionsoftheresearch. 19


CHAPTER2DYNAMICMODELS2.1DynamicFormulation Thedynamicformulationpresentedinthissectionaddressesmomentumexchangedevicesforattitudecontrolsystemswheretheangularmomentumofthespacecraftsystem(i.e.,thespacecraftandsystemofCMGs)isassumedconstant.Thisisincontrasttoreactioncontroldevices(e.g.,thrusters)and/orenergydissipationdevices(e.g.,magnettorquers)whichchangetheangularmomentumorenergyofthesystem. Itisassumedthatthecenterofmass(c.m.)ofeachCMGliesalongthegimbalaxis.Thisisequivalenttostatingthattherotationofthegimbal-ywheelsystemaboutthegimbalaxisdoesnotmovethepositionofthec.m.ofthesystem.Itisalsoassumedthatthespacecraft-CMGsystemisarigidbody;andthissystemisabsentoffrictionandexternaltorques.ForacoordinatelessderivationofthedynamicsseeAppendix A .TherstassumptioncanbevisualizedbytreatingtheCMGsascylindersandhavingtheirspinaxisalongtheirc.m.asshowninFigure 2-1 Thetotalcentroidalangularmomentumofthespacecraft-CMGsysteminthespacecraftbodyframeis H=Jc!+h(2) whichiscomposedofthespacecraftcentroidalangularmomentumJc!composedofthespacecraftcentroidalinertiaJcandangularvelocity!andtheangularmomentumhcontributedfromtheCMGs. Consideringthespacecraftmodeledasarigidbody,itscentroidalinertiaiscomposedofbothconstantandtime-varyinginertiasandisexpressedas Jc=J+mi(rTiri1)]TJ /F8 11.955 Tf 11.95 0 Td[(rirTi)+nXi=1CBGiIgwCTBGi(2) 20


Figure2-1. Rigidbodywithaconstantc.m. whereJrepresentstheconstantspacecraft-CMGsysteminertiainreferencingpositionaboutthespacecraft'sc.m;mi(rTiri1)]TJ /F8 11.955 Tf 12.56 0 Td[(rirTi)aretheparallelaxisterms,withmassmioftheithgimbal-ywheelassemblyanditsc.m.,positionriwithrespecttothec.m.ofthespacecraft,andCBGiIgwCTBGiarethetimevaryinginertiasfromrotationofthegimbal-ywheelsysteminertiaIgw.TheangularmomentumcontributedfromtheithCMGinagimbalframeFGiisexpressedas hi=266664Iw0Igw_i377775(2) whichconsistsofangularmomentumfromtheywheel(Iw)andthatfromthegimbal-wheelsystem(Igw_i).ItshouldbenotedthattheCMGangularmomentumexpressioninEq.( 2 )isbasedonanSGCMGorVSCMGonlyandthefollowingdevelopmentofangularmomentumforCMGswillbeforamultiplegimbalCMG.The 21


resultantangularmomentumfromtheCMGsysteminthebodyframeisfoundthroughthesummationofthecontributionsofangularmomentumfromallCMGsrotatedfromtheirrespectivegimbalframesintothespacecraftbodyframe;i.e., h=nXi=1CBGihi(2) wherenisthenumberofSGCMGs,CBGiisthedirectioncosinematrix(DCM)fromthegimbalframeFGishowninFigure 2-2 tospacecraftbodyframeFB. Figure2-2. GimbalframeFGiofIMPACSGCMG(PatentPending) Withtheassumptionofnoexternaltorquesandfrictionlessdevices,thetotalangularmomentumisconstant,thustheinertialtimederivativeofEq.( 2 )showstheredistributionofthesystems'sangularmomentum(i.e.,onthemechanismbywhichthetorqueisproducedbyCMGs).DifferentiationofEq.( 2 )yields dH dt=d(Jc_!+h) dt+!(Jc_!+h)=0(2) 22


Thetimederivativeofthespacecraftangularmomentumyields d(Jc!) dt=Jc_!+_Jc!(2) ThespacecraftinertiaisassumedtoonlyvarybythegimbalanglesoftheCMGsthusmakingthesecondtermontherighthandsideofEq.( 2 ) _Jc!kXj=1@(Jc!) @jdj dt=kXj=1Aj1_j(2) whereAj12R3nistheJacobianmatrixresultingfromthecouplingofthespacecraftandCMGkinematicsfromthejthgimbalofamultiplegimbalCMGandnisthenumberofCMGs. TheangularmomentumoftheCMGsystemisafunctionoftheywheelangularvelocities2Rn1,gimbalanglesj2Rn1,andgimbalrates_j2Rn1,respectively.ThetimederivativeoftheCMGsangularmomentumcanbeexpressedas dh dt=_h=kXj=1@h @jdj dt+@h @_jdj dt+@h @d dt(2) wheretheindividualJacobianmatricesare, Aj2=@h @j2R3n(2) Bj=@h @_j2R3n(2) C=@h @2R3n(2) Assumingaconstantywheelspeed_=0andasinglegimbal(k=1)conguration,thenEq.( 2 )canberewrittenasdH dt=Jc_!+!Jc!+!h+D_X=0(2) 23


where, D_X=[(A1+A2)B]264_375=_h+_Jc!=T(2) ThegeneralequationsfortheSGCMGoutputtorqueintermsofagiveninternalcontroltorqueisexpressedas, D_X=)]TJ /F10 11.955 Tf 9.29 0 Td[()]TJ /F10 11.955 Tf 11.95 0 Td[(!h=T(2) whereTisthetotaltorqueoutputfromthesystemofSGCMGs.Itshouldbenotedthat_andarekinematicallycoupled,thusitisnotpossibletondbothstatessimultaneouslywhenmappingD2R33nontoT(i.e.,onlyonegimbalstatecanbechosenasacontrolvariable).ForSGCMGsystemsthatcontainsignicantywheelangularmomentumandgyroscopictorque,thedynamicsofthegimbal-wheelassemblyinertiascanbeconsideredinsignicant(i.e.,A10andB0).Forsuchsystems,itiscustomarytoneglecttheinertiavariationsduetothegimbalmotion(i.e.,_Jc=0)resultinginthecompositeDreducingtoA22R3n.Therefore,theJacobianDissimplyA2andthesolutionoftheoutputtorquefromtheSGCMGsiscontributedsolelyfromthegimbalratesas _h=h0^A2_=h0[^1,^2,...^n]_(2) whereA2=h0^A2and^iisthetorquevectordirectionoftheithCMGasshowninFigure 2-2 .ThecoefcientmatrixinEq.( 2 )is3nandwhenn>3thesystemisover-actuated.Whenthismatrixbecomesrankdecient,thesystemissaidtobesingular.Physically,whenthesesingularitiesoccur,thetorquevectordirectionsofeachSGCMGinthebodyframelieinaplaneasshowninFigure 2-3 Forconvenience,fromthispointuntilChapter 7 ,itisassumed^A2=A. 24


Figure2-3. SingularityshownwhenCMGtorquevectorslieinaplane(IMPACSGCMGsPatentPending) 2.2SingularSurfaceEquations Itiscustomarytodeneanorthonormalbasisf^hi,^i,^igasshowninFigure 2-2 where^hiisthespinaxisoftheywheel,^iistheSGCMGtorquedirection,and^iisthegimbalaxisdirection. Therefore,thesingulardirections2R31isdenedfrom fs2R3:sT^i=0g(2) Thisconstraintconstitutesamaximum(orminimum)projectionof^hiontos.Thereisafundamentalassumptionthath0iisequaltoh0(i.e.,themagnitudeofnominalangularmomentumisthesameforallSGCMGsinthesystem).Foragivensingulardirections6=^i(i.e.,whichonlyoccursforDGCMGsandforrooftoparrangements),theconditionsforsingularityare sT^i=0andsT^hi6=0(2) Ifwedenei,sT^hi,thenthetorqueandspinaxisdirectionscanbeexpressedas ^i=i^is jj^isjj,s6=^i,i=1,...,n(2) 25


^hi=^i^i=i(^is)^i jj^isjj,s6=^i,i=1,...,n(2) CombiningEqs.( 2 )and( 2 ),thetotalnormalizedangularmomentumfromtheSGCMGsisexpressedas ^h=nXi=1^hi=nXi=1^i^i=i(^is)^i jj^isjj(2) Itisimportanttonotethatwhens=^iEqs.( 2 )-( 2 )areindeterminate. Thelocusoftotalnormalizedangularmomentum^hfromEq.( 2 )foralls2R3andalli6=0(i.e.,snotcollinearto^i)producestheexternalsingularsurfaceknownastheangularmomentumenvelopeshowninFigure 2-5 forafour-SGCMGpyramidarrangement.Similarly,eachofthefourinternalsingularsurfacesshowninFigure 2-6 forafour-SGCMGpyramidarrangementarefoundbysettingoneofthei<0.MatlabcodeforbothofthesesurfacescanbeseeninAppendix C .SingularitiesforSGCMGscanbeclassiedintothegroups/subgroupsasshowninFigure 2-4 .2.2.1EllipticSingularities EllipticsingularitiesarethoseinwhichnullsolutionstothegimbalanglesdonotexistforaspecicpointofCMGangularmomentumspace.Nullmotionisacontinuoussetofnullsolutionsforgimbalangles(i.e.,thereisacontinuoustransferfromonenullsolutiontothenext)thatdoesnotchangetheCMG'sangularmomentumandthus,doesnotproduceanymotiontothespacecraft.Sinceellipticsingularitiesdonothavenullsolutions,theangularmomentummustbeperturbedthusinducingtorqueerrortothesystemtoescapefromthesesingularities.Ellipticsingularitiesarenotlimitedtointernalsingularities;(e.g.,allexternalsingularitiesareellipticandhencecannotbeavoidedorescapedthroughnullmotion). 26


Figure2-4. SingularitiesforSGCMGs ExternalsingularitiesalsoknownassaturationsingularitiesareassociatedwiththemaximumprojectionofCMGangularmomentuminanydirection.Thesesingularitiescannotbeavoidedbynullmotionandthereforebydenitionareelliptic.Thesesingularitiesoccuronthesurfaceoftheangularmomentumenvelopeandanexampleofthissurfaceforafour-SGCMGpyramidisshowedinFigure 2-5 .Whenthesesingularitiesareencountered,theCMGsareunabletoproduceanymoreangularmomentuminthesaturateddirection.ExternalsingularitiesareaddressedaprioriinthedesignprocessthroughsizingoftheCMGactuators. 27


Figure2-5. Externalsingularsurfacesforafour-SGCMGpyramid Considerafour-SGCMGpyramidarrangement,Eq.( 2 )canbeusedtoexpresstheangularmomentumash=h0266664)]TJ /F7 11.955 Tf 9.3 0 Td[(c()s(1))]TJ /F7 11.955 Tf 11.96 0 Td[(c(2)+c()s(3)+c(4)c(1))]TJ /F7 11.955 Tf 11.95 0 Td[(c()s(2))]TJ /F7 11.955 Tf 11.95 0 Td[(c(3)+c()s(4)s()(s(1)+s(2)+s(3)+s(4))377775(2) whereistheskewangleandiistheithgimbalangle.Furtherconsiderthesetofgimbalangleses=[90909090]deg,thenthemomentumvectorbecomes h(es)=26666400h0s()(s(1)+s(2)+s(3)+s(4))377775=266664004h0s()377775(2) Itisclearthatthereisonlyonesetofgimbalangleses=[90909090]TdegthatwillgivetheangularmomentuminEq.( 2 ).Therefore,nullsolutionsdonotexist,andthis 28


angularmomentumvectorcorrespondstotheellipticsaturationsingularityalongthez-axis. Ellipticsingularitieswhichlieontheinternalsingularsurfacessuchasthatshownforthefour-SGCMGpyramidarrangementinFigure 2-6 arereferredtoasellipticinternalsingularities.Unlikeexternalsingularities,thesesingularitiescannotbeaccountedforinthedesignprocess;furthermore,sincetheyoccurinstantaneously,theycannotbegenerallyavoided.2.2.2HyperbolicSingularities Hyberbolicsingularitiesarethoseinwhichnullmotionispossible.Thus,allhyperbolicsingularitiesarethereforeinternal(i.e.,thesesingularitiesoccurontheinternalsingularsurfaces).Thepointsontheinternalsingularsurfacecorrespondingtoahyperbolicsingularityhavenullsolutionsofgimbalangles,correspondingtothenullspaceoftheJacobianmatrix.Thenullsolutionsaretypicallychosentoavoidthesingularcongurationsofthesystem.ShowninFigure 2-6 ,isanexampleofthissurfaceforafour-SGCMGpyramid.SingularitiesoccuronlywhenthepointonthissurfacecorrespondstoasingularJacobianmatrix(i.e.,theremaybenonsingularsetsofgimbalanglesatthispointonthesurface).Whenthesesingularitiesareencountered,theSGCMGtorquevectordirectionslieinaplaneandasaconsequencethereisnotorqueavailableoutoftheplane.Thesesingularities,likeellipticsingularities,areinstantaneousandmustbehandledonthey.Forafour-SGCMGpyramidarrangementwithangularmomentuminEq.( 2 ),asetofgimbalangleshs=[18090090]Tdegisanhyperbolicsingularitythathasthefollowingmomentumvector, h(hs)=2666640)]TJ /F4 11.955 Tf 9.29 0 Td[(2h02h0s()377775(2) 29


Itisclearthattheremultiplesolutions(i.e.,nullsolutions)toEq.( 2 ).Anullsolutionofthegimbalanglessatisfying1=3=90ands(4)=)]TJ /F7 11.955 Tf 9.3 0 Td[(s(2)=1 c()substitutedintoEq.( 2 )willalsosatisfyEq.( 2 ). Figure2-6. Internalsingularsurfacesforafour-SGCMGpyramid Non-Degeneratehyperbolicinternalsingularitiesarethoseinwhichnullmotionispossibleandsomeofthenullsolutionsarenonsingularprovidingthepossibilityofsingularityavoidance. Degeneratehyperbolicinternalsingularitiesoccurwhenthenullsolutionstogimbalanglescorrespondtosingularsetsofgimbalanglesleavingnoroomforavoidanceorescape.Thesesingularitiesarealsoconsideredimpassableandthereforearehandledinasimilarmannertoellipticsingularitieswhenapproached.2.2.3Gimbal-Lock Gimbal-lockforSGCMGsoccursatsingularitywhenthemappedoutputtorquevectorisinthesingulardirection.Whenthisoccurs,thesystembecomestrappedinthis 30


singularcongurationwithonlyafewmethodsthatarecapableofescapefromit.OnesuchmethodisknownastheGeneralizedSingularityRobust(GSR)Inverse[ 39 40 ].Thismethodhasbeenshownnumericallytoescapegimbal-lockofSGCMGsbutnotanalyticallyandthereisnoformalprooftosuggestthatitisalwayssuccessful.2.3SingularitiesforSGCMGsMathematicallyDened Toquantiytheeffectivenessofavoidinginternalsingularitiesthroughnullmotion,wemustdenetheirforms(i.e.,hyperbolicandelliptic)mathematically.Typically,topologyanddifferentialgeometryareusedtorepresenthyperbolicandellipticinternalsingularitiesassurfacesormanifolds[ 41 42 ].Thebehavioroftheseinternalsingularitiescanalsobeexplainedthroughtheuseoflinearalgebra[ 43 ].Toaccomplishthis,aTaylorseriesexpansionoftheSGCMGangularmomentumaboutasingularcongurationgives h())]TJ /F8 11.955 Tf 11.95 0 Td[(h(S)=nXi=1@hi @iSii+1 2@2hi @2iSi2i+H.O.T.(2) whereh(S)istheangularmomentumatasingularsetofgimbalanglesS,i=i)]TJ /F3 11.955 Tf 11.95 0 Td[(Si,andnisthenumberofSGCMGsinthesystem. Thersttermontheright-handside(RHS)ofEq.( 2 )containstheithcolumnoftheJacobianmatrix^i=@hi @ijSi,associatedwiththeithSGCMG'storquedirection.ThesecondtermontheRHScontainsthepartialderivativeoftheJacobianmatrix'sithcolumnwithrespecttotheithgimbalangle@2hi @2ijSi.Furthermore,fromEq.( 2 ),theRHSofEq.( 2 )canbetransformedthroughtherealizationofthefollowingoperations: @2hi @2i=@^i @i=)]TJ /F7 11.955 Tf 9.3 0 Td[(hi^hi=)]TJ /F8 11.955 Tf 9.29 0 Td[(hi(2) where^()denotesaunitvector.Next,Eq.( 2 )issubstitutedintoEq.( 2 )andtheinnerproductoftheresultwiththesingulardirectionsobtainedfromnull(AT)yields 31


sT[h())]TJ /F8 11.955 Tf 11.96 0 Td[(h(S)]=)]TJ /F4 11.955 Tf 10.5 8.08 Td[(1 2nXi=1hTis2i(2) ThersttermontheRHSofEq.( 2 )haszerocontributionbecauseofthedenitionofthesingulardirection(i.e,ATs=0).Equation( 2 )canbewrittenmorecompactlyas sT[h())]TJ /F8 11.955 Tf 11.96 0 Td[(h(S)]=)]TJ /F4 11.955 Tf 10.5 8.09 Td[(1 2TP(2) wherePisthesingularityprojectionmatrixdenedasP=diag(hTis). Bydenition,nullmotiondoesnotaffectthetotalsystemangularmomentumandwhichequatestoh()=h(S).Consequently,theleft-handside(LHS)ofEq.( 2 )iszero(i.e.,TP=0).NullmotionisexpressedintermsofthebasisN=null(A)concatenatedinmatrixformasfollows =(n)]TJ /F5 7.97 Tf 6.59 0 Td[(r(A))Xi=1ii=N(2) whereisacolumnmatrixofthescalingcomponentsofthenullspacebasisvectorsiandN2Rn(n)]TJ /F5 7.97 Tf 6.59 0 Td[(r(A))isthedimensionofthenullspacebasisforanysystemofSGCMGswithr(A)=rank(A).SubstitutingEq.( 2 )intoEq.( 2 )observingthenullmotionconstraintyields TQ=0(2) Asaresultofthisanalysis,amatrixQisdenedasQ=NTPN(2) Therefore,whenawayfromsingularityQ2R11;whenatarank2singularity,Q2R22;andwhenatarank1singularity,Q2R33.TheeigenvaluesoftheQmatrixdeterminewhetherasingularityishyperbolicorelliptic.IfQisdenite(i.e.,hasallpositiveornegativeeigenvalues),itdoesnotcontainanullspacesinceanonzeronullvector 32


doesnotexistthatsatisesEq.( 2 )[ 44 ].Therefore,situationswherethematrixQisdeniteconstituteellipticsingularities. WhenQissemi-denite(i.e.,ithasatleastonezeroeigenvalue),thenanullspaceexistssincethereexistsa6=0thatsatisesEq.( 2 ).Therefore,nullmotionispossiblenearsingularityandthepossibilityofsingularityavoidancemayhold(i.e.,doesnotfordegenerate-hyperbolicsingularities)[ 43 ].IfthematrixQisindenite(i.e.,theeigenvaluesarepositiveandnegative),theresultofEq.( 2 )hasthepossibilityofbeingequaltozero.Therefore,bothofthesesituationsconstituteanhyperbolicsingularity. ThetoolsdevelopedinthischapterfordescribingtheexistenceofellipticsingularitiesandhyperbolicinternalsingularitiesinasystemofSGCMGsareusedinthenextchaptertomorespecicallydiscusswhichoftheseinternalsingularitiesexistincommonarrangementsofSGCMGs.Inaddition,Chapter 5 introducesanovelsteeringalgorithmknownastheHybridSteeringLogic(HSL)whichusesthesetoolsinitsderivation. 33


CHAPTER3CONTROLMOMENTGYROSCOPEARRANGEMENTS3.1CommonSGCMGArrangements SeveralcommonSGCMGarrangementshavebeenstudied.Typically,thefactorsthatdeterminethechoiceofaspecicSGCMGarrangementare:(i)availablevolume(ii)desirableangularmomentumenvelope,and(iii)associatedsingularities.Inthischapter,weexaminethecommonarrangementsandusethetoolsdevelopedinChapter 2 tocharacterizetheirsingularities.3.1.1Rooftop TherooftoparrangementshowninFigure 3-1 hastwosetsofparallelSGCMGs,eachwithparallelgimbalaxeswhereistheskewanglerelatingtheplanesoftorque.Afour-SGCMGrooftoparrangementisshowninFigure 3-1 Figure3-1. Four-SGCMGrooftoparrangement Sincethesearrangementsarefreefromellipticinternalsingularities,theyhaveasignicantightheritageonsatellitesandthus,theircontroliswellunderstood[ 45 ].However,degeneratehyperbolicsingularitieswhicharealsoimpassablestillexistandlikeellipticsingularitiescannotbeaddressedthroughtheuseofnullmotion.Inaddition,therearedegeneratecasesofhyperbolicsingularitiesfortherooftoparrangementwhen 34


theJacobianmatrixisrank1whichmayprovidedifcultytosingularityescapelawsthatregulatethesmallestsingularvalue.3.1.2Box TheboxarrangementisasubsetoftherooftoparrangementandhastwoparallelsetsoftwoSGCMGswithanangleof90degbetweenthetwoplanesoftorqueasshowninFigure 3-2 .Thisarrangementisgivenitsnamebecausetheplanesofangularmomentumcanformabox[ 46 ].Liketherooftoparray,therearesituationswherethisarrangementmayhavearank1Jacobian. Figure3-2. Four-SGCMGboxarrangement Theangularmomentumenvelopeforthefour-SGCMGboxarrangementisshowninFigure 3-3 .Theangularmomentumenvelopeforallrooftoparrangementsisanellipsoidialsurfaceandthus,thereisnotequalmomentumsaturationinalldirections. Rooftoparrangementsarechosenfortheircompactnessandthefactthattheyarefreefromellipticinternalsingularities.Analysisprovidedintheliteratureprovesthatthoughthesearrangementsarefreefromellipticinternalsingularities.However,theyarenotfreefromtheimpassabledegeneratehyperbolicsingularities[ 47 ]. 35


Figure3-3. Angularmomentumenvelopeforafour-SGCMGboxarrangement. InChapter 2 ,itwasshownthatthedenitenessofthematrixQdeterminesifasystemofSGCMGsisatanellipticsingularity.ForasystemoffourSGCMGs,thelargestQcanbeisR22excludingthecasewhentheJacobiangoesrank1whichwillbediscussedlater.Therefore,excludingarank1Jacobian,andifthedeterminantofQisstrictlypositive,thenthesystemisatanellipticsingularity(e.g.,12>0where1,2areaneigenvaluesofQ).Consequently,whendet(Q)0,thesystemisatanhyperbolicsingularity.Thereareafewgeneralcaseswheresingularitiesmayoccurforafour-SGCMGrooftoparrangement.Therstcaseoccurswhenthetorquevectorslyinginthesameplaneoftorqueareeitherparalleloranti-parallelasshowninFigures 3-4 AandBwhere12+180,34+180andristheaxisintersectingthetwoplanesoftorque. 36


ATorqueplaneswithparalleltorquevectors BTorqueplaneswithoneparallelandanti-paralleltorquevectors Figure3-4. Planesoftorqueforafour-CMGrooftoparrangement 37


AfourCMGrooftopsysteminthecongurationofFigure 3-4 AhasaJacobian A=266664c()c(1)c()c(1))]TJ /F7 11.955 Tf 9.3 0 Td[(c()c(3))]TJ /F7 11.955 Tf 9.3 0 Td[(c()c(3)s(1)s(1))]TJ /F7 11.955 Tf 9.3 0 Td[(s(3))]TJ /F7 11.955 Tf 9.3 0 Td[(s(3)s()c(1)s()c(1)s()c(3)s()c(3)377775(3) withangularmomentumrepresentedinthespacecraft-bodyframe(i.e.,wherethegimbalsareenumeratedcounter-clockwisebeginningatthespacecraftbodyx-axis) h=h1+h2+h3+h4=2h0266664c()s(1))]TJ /F7 11.955 Tf 9.3 0 Td[(c(1)s()s(1)377775+2h0266664)]TJ /F7 11.955 Tf 9.29 0 Td[(c()s(3)c(3)s()s(3)377775(3) Thesingulardirectionforthiscaseisfoundbycrossproductof^1and^3 s=266664s()s(1+3))]TJ /F4 11.955 Tf 9.3 0 Td[(2s()c()c(1)c(3)c()s(1)]TJ /F3 11.955 Tf 11.96 0 Td[(3)377775(3) withtheresultantprojectionmatrix P=2h0s()c()266666664c(1)0000c(1)0000)]TJ /F7 11.955 Tf 9.3 0 Td[(c(3)0000)]TJ /F7 11.955 Tf 9.3 0 Td[(c(3)377777775(3) andthenull-spaceoftheJacobianconcatenatedinmatrixformis N=266666664)]TJ /F4 11.955 Tf 9.3 0 Td[(10100)]TJ /F4 11.955 Tf 9.3 0 Td[(101377777775(3) 38


ThecongurationinFigure 3-4 hasthedet(Q)=)]TJ /F4 11.955 Tf 9.3 0 Td[(16s2()c2()c(1)c(3).ItshouldbenotedthatwhenPandQaredenite(i.e.,=[++++]or=[)-276()-276(\000]wherei=sgn(^his)),thesystemisataexternalsingularity.ItwasdiscussedinChapter 2 thatexternal(orsaturation)singularitiesareelliptic.Therefore,ifthematrixPisdenite,thenthematrixQisalsodenite.Thedet(Q)=)]TJ /F4 11.955 Tf 9.3 0 Td[(16s2()c2()c(1)c(3)willonlybepositivewhensgn(c(1))6=sgn(c(3))(i.e.,atsaturationsingularity).Thesaturationsingularityisnotaninternalsingularityandthereforeneithertherooftopandboxarrangementscontainellipticinternalsingularitiesforthiscase.Itcanbeveriedthatthevariationsofthesecasessuchas(^1=)]TJ /F8 11.955 Tf 10.31 0 Td[(^2),(^2=)]TJ /F8 11.955 Tf 10.31 0 Td[(^4),(^1=)]TJ /F8 11.955 Tf 10.31 0 Td[(^2)and(^2=^4)allhavedet(Q)=0andthereforearehyperbolicinternalsingularities. TheothercasewhentheJacobianofafour-CMGrooftopissingularoccurswhenthetorquevectorsofthetwoparallelSGCMGslieinthedirectionoftheintersectingtorqueplanesrshowninFigures 3-5 ATorqueplaneswithparalleltorquevectorsalongr BTorqueplaneswithanti-paralleltorquevectorsalongr Figure3-5. Torqueplanestracedoutforafour-SGCMGrooftoparrangement ForFigure 3-5 A,theJacobianis 39


A=266664)]TJ /F7 11.955 Tf 9.3 0 Td[(c()c(3))]TJ /F7 11.955 Tf 9.29 0 Td[(c()c(4)00)]TJ /F7 11.955 Tf 9.3 0 Td[(s(3))]TJ /F7 11.955 Tf 9.29 0 Td[(s(4)11s()c(3)s()c(4)00377775(3) withangularmomentumvector h=h1+h2+h3+h4=2h0266664c()0s()377775+h0266664)]TJ /F7 11.955 Tf 9.29 0 Td[(c()s(3)c(3)s()s(3)377775+h0266664)]TJ /F7 11.955 Tf 9.3 0 Td[(c()s(4)c(4)s()s(4)377775(3) assumingthattheintersectionoftheplanesoftorquer=[010]TforthisarrangementshowninFigure 3-5 .Forthissituation,thesingulardirectionisorthogonaltotheintersectionoftheplanesandisfoundtobe s=266664s()0c()377775(3) withtheresultingprojectionmatrix P=2h0s()c()c(3)2666666641000010000000000377777775(3) andthenullspacebasisfortheJacobianconcatenatedinmatrixformis N=null(A)=266666664)]TJ /F4 11.955 Tf 9.29 0 Td[(1s(4)]TJ /F3 11.955 Tf 11.96 0 Td[(3)=c(3)100)]TJ /F4 11.955 Tf 9.3 0 Td[((c(4)=c(3))01377777775(3) 40


Forthiscase,det(Q)=2h0sin2(2)>0whichisanellipticsingularitybutisnotknownyettobeonethatisinternalorexternal.Fortheanti-parallelcasewhen3=180+4=270deg,thedet(Q)=2h0sin2(2)<0andthesingularityishyperbolic.NoticethatthediagonalentriesofPthatarezerocorrespondtothegimbalaxisofthatroof-sidebeingalongthesingulardirectionsofthesystem. Recall,fromChapter 2 thats=^iisaspecialcasethathappensonlyforDGCMGsandrooftoparrangements.Forexample,considerthecasewhen3=180+4=270andanangularmomentumof h=2666640h20377775(3) isdesiredalongwith,1=)]TJ /F3 11.955 Tf 9.3 0 Td[(2.TheresultofEq.( 3 )isc(1)=c(2)=)]TJ /F5 7.97 Tf 6.59 0 Td[(h2 2.Therearetwosolutionsforanypossiblevalueofh2insidethemomentumenvelopeduetosymmetryandthusthereisonenullsolution.Bothofthesesolutionsaresingularandabideby3=180+4=270andthereforethenullsolutionsthatexistdonothelpinescapefromthesingularity.Thus,thisisacaseofadegeneratehyperbolicsingularityatthatspecicpointonthemomentumspace.Inaddition,thevalueofdm d=0(i.e.,wherem=p det(AAT))forboth3=180+4=270andfor1and2free,andthusthereisnosetofgimbalanglesthatwillprovideachangeinm(i.e.,nonullsolutionstoescapesingularity). Todetermineiftheothercasesofsingularitywhen3=4=90degareelliptic,wecheckifnullmotionexistsorthogonaltothesingularity(i.e.,dm d6=0).Theresultofthiscasewhen3=4=90and1and2arefreeisconsistentwithEq.( 3 ).Itisfoundthatforanychoiceof1and2givesdm d=0andthereforethisisafamilyofellipticsingularitiesbecausenotonlydoesdet(Q)>0butalsodm d=0.Tovisualizewherethe 41


externalsingularitiesoccurforthiscase,theSGCMGangularmomentumisplottedforallvariationsof1and2ofafour-SGCMGboxarrangementshowninFigure 3-6 Figure3-6. Angularmomentumenvelopewithplottedangularmomentumcombinationsforthefour-SGCMGboxarrangement InFigure 3-6 ,allpossiblecombinationsofangularmomentumareplottedontotheangularmomentumenvelopeinblackfor1and2andwhen3=4=90.Inthisgure,everycombinationofthissituationisanexternalsingularity.Itcanbeshownfromsymmetrythatallpermutationsofthiscasehavethesameresultinthattheyareexternalsingularities.Therefore,rooftoparrangementsdonotcontainellipticinternalsingularities.Inaddition,thecasewhentheJacobianforafour-SGCMGrooftoparrangementapproachesrank1isasubsetofthisfamily.Thisfamilyofcongurationsisdenedasthecaseswhereatleasttwotorquevectorsareparallelalongtheintersectionofthetwotorqueplanesrandtheoutcomeofthedet(Q)isnotdependentonthegimbalangles.Therank1Jacobiancaseisadegeneratehyperbolicsingularityforwhenthe 42


torquevectorsareallanti-parallel(i.e.,0or2h0)andaexternalsingularitywhenalltheangularmomentumvectorsareparallel(i.e.,4h0). TheseresultsconrmthoseobtainedviatopologyanddifferentialgeometrybyKurokawa[ 41 ].Inreference[ 41 ],itwasstatedthatanyrooftoparrangementswithnolessthansixunitsarefreefrominternalimpassablesurfaces(i.e.,ellipticinternalsingularitiesnotincludingexternalsingularities).Kurokawaconcludedthatthereareimpassibleinternalsurfacesinthefour-SGCMGrooftoparrangementsthatcorrespondtothesingulardirectionsnotcontainedintheplanespannedbythetwogimbalaxes^1and^2.ThisisexactlythedegeneratecaseshowninEqs.( 3 )-( 3 )wheres=^1.Thedegeneratehyperbolicsingularitiesforthesearrangementslieontwocircleswithradius2h0when1=2and3=180+4=270andatzeromomentum0h0when1=)]TJ /F3 11.955 Tf 9.3 0 Td[(2and3=180+4=270shownasanexamplefortheboxarrangementinFigure 3-7 [ 41 ]. Figure3-7. Degeneratehyperbolicsingularitiesforthefour-SGCMGboxarrangement ThecurvesandpointinFigure 3-7 arecompartmentalizedandnotspreadthroughouttheentireangularmomentumenvelopeunlikeellipticsingularitiesandthus, 43


constrainedsteeringalgorithmsexisttoavoidtheseregionswhileprovidingsingularityavoidanceusingnullmotion.3.1.33 4Box The3 4boxarrangementisasubsetoftheboxarrangementinwhichoneoftheSGCMGsisnotusedandleftasaspare.Thisarrangementhasthelongestheritageofightduetoitsconservativenature.Forthisarrrangement,apseudo-inverseisnotrequiredtoobtainasolutiontothegimbalratessincethegimbalratesarefounddirectlyfromtheinverseofa33Jacobianmatrix[ 48 ].ThecombinationofthisarrangementandtheconstrainedangularmomentumsteeringlawlimitingcontrollableSGCMGangularmomentumtoa1h0radiusoftheangularmomentumenvelope,ensurethesafestcontrolofasystemofSGCMGs[ 49 ].Thisdesignalthoughsafe,maynotbepracticalforsmallsatelliteapplicationsbecausetheSGCMGsmustbeoversizedtoprovidethedesiredperformance(i.e.,theydonottakeadvantageoftheentiremomentumenvelope). 44


AExternalsingularsurfacefor3 4box BInternalsingularsurface3 4box Figure3-8. Singularsurfacesshowing1h0singularityfreeregion Inthisarrangement,thereisa1h0radiusoftheangularmomentumenvelopethatisguaranteedtobesingularityfree[ 48 ]whichisshownbytheredcircledrawnontheexternalandinternalsingularsurfacesofFigure 3-8 .Itshouldbementionedthatbecausethereisnolongeranullspace(i.e.,usingonly3ofthe4SGCMGs),anyangularmomentumpointonthesingularsurfacescorrespondtothelocationofaellipticsingularity.Thismakestheconstrainingtothe1h0sphereofangularmomentumimperative.3.1.4ScissorPair Thescissor-pairarrangementhasthreesetsofcollinearpairsoftwoSGCMGsorthogonaltoeachother.Thisarrangementisconstrainedtohave1=)]TJ /F3 11.955 Tf 9.3 0 Td[(2atalltimesforbothCMGstoavoidinternalsingularities.WiththisarrangementshowninFigure 3-9 ,fullthreeaxiscontrolispossiblewithafullrankJacobianmatrixaslongasitdoesnotextendpastthemaximumangularmomentumofthesystem.Asaconsequenceoftheconstraintforthesepairs,onlyone-thirdoftheentireangularmomentumenvelopeisutilizedwhichwillbetroublesomeforuseonsmallsatellites(i.e.,sixSGCMGsneededfor2h0ofangularmomentum).Useofthesearrangementswasfoundtoconservepowerwhenasinglegimbalmotorisusedforeachscissorpair[ 50 ].Also,analysishasshown 45


thatscissorpairsmaybebenecialforspaceroboticsapplicationsincetheirtorqueisunidirectional[ 51 ]. Figure3-9. 3OrthogonalscissorpairsofSGCMGs Duetothegimbalangleconstraintassociatedwiththisarrangement,internalsingularitiesdonotexisthere.Alsobecauseofthegimbalangleconstraint1=)]TJ /F3 11.955 Tf 9.3 0 Td[(2,threeorthogonalscissor-pairscontainonlyexternalsingularitiesthatoccurwhenoneormoreofthepairshasanoverallzerotorquevector(i.e.,undenedtorqueforscissorpair).Whenthisoccurs,theJacobianmatrixcontainsacolumnofzerosfortheassociatedpair.3.1.5Pyramid PyramidarrangementsofSGCMGshaveindependentplanesofangularmomentumandtorquewhichformapyramid.Asaconsequenceoftheseindependentplanesoftorque,thesearrangementswillneverhavetheJacobianmatrixwithranklessthan2.ThisisshowninFigure 3-10 forafour-SGCMGpyramidcluster. Intermsofsmallsatelliteconstraintsandwhenutilizingtheentiremomentumenvelope,thefourSGCMGarrangementseemspracticalamongallpreviouslydiscussedarrangementsforplatformsrequiringhightorqueandslewrateswithnearequalmomentumsaturationinthreedirections. 46


Figure3-10. Planesofangularmomentumandtorqueforafour-SGCMGpyramid Figure3-11. Four-SGCMGpyramidarrangement Controlofthesearrangementsismorecomplicatedthanrooftopandboxarrangementsduetothepresenceofellipticinternalsingularitiesbecausenullmotionsolutionsdonotexist.Inaddition,ellipticsingularitiesdonothavecontinuousgimbaltrajectoriesassociatedthecorrespondingcontinuousangularmomentumtrajectories[ 47 ].Thesearrangementsarestudiedfortheirdesirablemomentumenvelope(i.e.,itispossibletogetanearsphericalangularmomentumenvelopewithaskewangleof=54.74deg)[ 52 ].Ifhighagilityiswhatisneededandtherearemorerelaxedpointingrequirements,thepyramidmayprovidebenetsovertheotherarrangements.Evenifthisisnotthe 47


case,ifthisarrangementishostedonasmallsatelliteandtheattitudeerrorinducedfromthetorqueerrorprovidedbythesingularityescapeofellipticsingularitiesisonthesameorderoftheattitudedeterminationsensorsand/ormethods,thenthetorqueerrorusedforsingularityescapewillbeinconsequential.3.2ChoiceofArrangement BeyondthecommonarrangementsofSGCMGspreviouslydiscussed,itisdifculttochoosethearrangementofSGCMGsthroughshapingoftheangularmomentumenvelopeofthesystem.Thisisduetothelocationsofwhereinternalsingularitiesliewithintheangularmomentumenvelopedenotedbytheinternalsingularsurfaces(e.g.,seeFigure 2-6 ).Thesesingularitiesaredispersedandmaycovertheentireangularmomentumenvelopeleavingonlyverysmallsingularity-freeareas.Formulatingtheproblemasaparameteroptimizationasin[ 53 ]canonlyprovidetheoptimalarrangementforagivensetofslewsandinitialgimbalangleswhichmakestheproblemmoreconstrainedthanuseful.Forexample,wecanexpressthegimbalaxesrelationtothespacecraftbodyframeintermsoftheEulerangles,twoofwhicharetheoptimizedconstantsinclinationanglei,spacinganglei,andthethirdisthegimbalanglei.TheDCMthatisusedtotransformfromthebodytothegimbalframeis CGiB=C3(i)C2(i)C3(i)(3) TheangularmomentumoftheSGCMGsistransformedfromthegimbalFGitothespacecraftbodyframeFBthroughthisDCMas h=nXi=1CBGihi(3) whichisconsistentwithEq.( 2 ).Therefore,holdingthespacingandinclinationanglesconstant,theresultantangularmomentumoftheCMGsystemisaninstantaneous 48


functionofonlythegimbalanglesforSGCMGs.ConsideringthisandthetruncateddynamicmodelofSGCMGsfromChapter 2 ,thecostfunction M=Zti+1ti()]TJ /F7 11.955 Tf 9.3 0 Td[(m2+aeTe+b_T_)dt(3) whereaandbarescalarsmakingthecostfunctionunitlessande=_h)]TJ /F7 11.955 Tf 12.61 0 Td[(h0A_,canoptimizethesystemwithrespecttominimaltorqueerrorthroughthechoiceoftheEuleranglesforagivenslew,slewtime,andinitialconditionsofthegimbalangles.ThisprocedurefortheparameteroptimizationisshowninFigure 3-12 Figure3-12. Optimizationprocessblockdiagram 49


3.3Simulation Anexamplesimulationofarest-to-restattitudemaneuverhastheparametersinTable 3-1 .Thissimulationwillshownthebenetofdifferentarrangementsonperformingthismaneuver(i.e.,trackingthetorquefromthecontroller).Itshouldbenotedthattheinitialconditionsofthegimbalanglesalthougharethesameforeveryarrangement,theyproduceadifferentinitialCMGangularmomentum. Table3-1. ModelParametersVariableValueUnits J0@100)]TJ /F4 11.955 Tf 9.3 0 Td[(2.01.5)]TJ /F4 11.955 Tf 9.3 0 Td[(2.0900)]TJ /F4 11.955 Tf 9.29 0 Td[(601.5)]TJ /F4 11.955 Tf 9.3 0 Td[(6010001Akgm20[0000]Tdege0[0.04355)]TJ /F4 11.955 Tf 11.95 0 Td[(0.087100.043550.99430]T\000!0[000]Tdeg=sh0128Nmsk0.051=s2c0.151=sa11=N2m2s2b11=s2t0.02sess0.0001deg Theresultsweresimulatedusingthefollowingeigen-axiscontrollogic[ 54 ] =)]TJ /F4 11.955 Tf 9.3 0 Td[(2kJe)]TJ /F7 11.955 Tf 11.95 0 Td[(cJ!+!J!(3) combinedwithafourth-orderRunga-Kuttaintegratoratatimesteptuntilthesteady-stateerrortoleranceoftheerrorquaternioneigen-angleesswasreached.Thesimulationcomparestheoptimizedsolutiontothefour-SGCMGpyramidarrangementataskewangle=54.74deg. Theresultsforthisexampleatinitialconditions0,e0,and!0havethesolutionforthesystem'ssingularsurfaceswithcalculatedarrangementEuleranglesshowninFigure 3-13 50


AExternalsingularsurface BInternalsingularsurface Figure3-13. SingularsurfacesfortheoptimizedarrangementattheEulerangles=[170.213.685.5168.0]Tdegand=[17.7167.0304.392.5]Tdeg ThegimbalratesfortheoptimizedarrangementinFigure 3-14 Aareapproximatelythesamemagnitudethanthatforthepyramidarrangement,althoughtheyhaveasmoothertransientresponse. AGimbalratesforoptimizedarrangement BGimbalratesforpyramidarrangement Figure3-14. Gimbalratesfortheoptimizedandpyramidarrangements ThetorqueerrorshowninFigure 3-15 Afortheoptimizedcaseissmallermagnitudethanthatforthepyramidarrangementduetotheareaunderthecurvesthus,moretorqueerrorisaddedduringthemaneuverforthepyramidarrangement. 51


ATorqueerrorforoptimizedarrangement BTorqueerrorforpyramidarrangement Figure3-15. Torqueerrorfortheoptimizedandpyramidarrangements Fortheoptimizedmethod,thesingularitymeasureisfarfromsingularityinitiallyanddoesnotencounteritasshowninFigure 3-16 A.ThisisincontrasttothatforthepyramidarrangementshowninFigure 3-16 B,whichstartsoutinitiallyfarfromsingularityandthenencounterssingularityseveraltimesduringthemaneuver.ThenegativequadratictermpresentinthecostfunctionofEq.( 3 )forthissingularitydoesnotweightdistancefromsingularityashighastorqueerrorwhichcanbeseenwhencomparingFigures 3-16 AandBto 3-15 AandB. ASingularitymeasureforoptimizedarrange-ment BSingularitymeasureforpyramidarrangement Figure3-16. Singularitymeasurefortheoptimizedandpyramidarrangements 52


Finally,thecostfunctionoftheoptimizedarrangementinFigure 3-17 AislessthanthatforthepyramidinBduetotheareunderthecurves. ACostforoptimizedarrangement BCostforpyramidarrangement Figure3-17. Optimizationcostfortheoptimizedandpyramidarrangements Thesesimulationssupporttheideathatifitwheremechanicallypossibletorecongurethegimbal-axisarrangementsinatimelymanner,andtheinitialgimbalanglesandmaneuverofinterestwereknown,asolutiontotheoptimalCMGarrangementcanbefound.Inaddition,thesesimulationsprovethatyoucannotsimplychooseanoptimalarrangementbecausetheproblemisnotonlydependentontheattitudemaneuver,butalsodependentontheinitialconditionsofthegimbalangles. Ifadesiredarrangementisknownwhileon-orbitandtherewasamechanicalwaytoreconguretheSGCMGgimbalaxes,suchasintheHoneywellpatent[ 55 ],thentherewouldbemeritinndinganalgorithmthatwassuccessfulinreorientingthegimbalaxesoftheCMGarrangement.Although,noalgorithmexiststoreorienttheSGCMGgimbalaxeswhilekeepingspacecraftunperturbed.Also,suchanalgorithmwouldstillrequireangularmomentumofoadingduetothenatureofSGCMGs. 53


CHAPTER4SURVEYOFSTEERINGALGORITHMS Aguidance,navigation,andcontrol(GNC)systemiscomposedoftheloopsshowninFigure 4-1 Figure4-1. OuterandinnerloopsofGNCsystem TheoutermostloopofaspacecraftGNCsystemconcernsthenavigation(i.e.,providesthestateknowledge)andisusuallytheminimumloopneededforanymission.Thesecondmostouterloopisconcernstheguidanceofthesystem(i.e.,providesthedesiredtrajectories)(e.g.,trajectoriesavoidingpointingastarcameratowardsthesun).Aloopinnertotheguidanceloopconcernsthecontrolofthesystem(i.e.,generatesanerrorofthestateknowledgefromthenavigationloopwiththedesiredtrajectoriesfromtheguidancelooptobeminimized).Theinnermostloopconcernsthedistributionofthedesiredcontroltothesystemsactuators(e.g.,whatthrustersneedtore,whatreactionwheelsorCMGsneedtomove).Steeringalgorithmsareconcernedwiththeinnermost 54


loopofFigure 4-1 whenthedifferentialequationrelatingthecontroltotheactuatorsissingular.Whenthisequationissingular,thesteeringlawrealizesasolution.4.1Moore-PenrosePseudo-Inverse Anearlymethodusedtomapthegimbalratesfromtherequiredoutputtorqueusestheminimumtwo-normleastsquaressolutionalsoknownastheMoore-Penrosepseudo-inverse.Thesolutionofthegimbalratesusingthispseudo-inversemappinghastheform _=1 h0A+_h=1 h0AT(AAT))]TJ /F2 7.97 Tf 6.59 0 Td[(1_h(4) whereA+istheMoore-Penrosepseudo-inverse,h0isthemagnitudeofSGCMGangularmomentum,_histheSGCMGoutputtorque,and_arethegimbalrates.TheMoore-Penrosepseudo-inverse,however,issingularwhentheJacobianmatrixAhasrank3[ 56 ].ItmightseemintuitivethattheadditionofmoreSGCMGactuatorsincreasesthepossibilityofhavingfullrank,buttheperformanceisnotequallyincreasedforallofCMGarrangements.Thisisbecausethereare2nsingularcongurationsforanygivensingulardirectionofasystemcontainingnSGCMGs[ 41 ].Also,theMoore-Penrosepseudo-inverseandvariationsofitcausethesystemtomovetowardsingularstateswhenperformingdiscretetimecontrol[ 47 ]. Tohandlecaseswhensingularitiesmaybeencountered,steeringalgorithmsareused.SteeringalgorithmscanbebrokendownintothefollowinggroupsasshowninFigure 4-2 4.2SingularityAvoidanceAlgorithms Singularityavoidancealgorithms,aremethodswhichsteerthegimbalsoftheSGCMGsawayfrominternalsingularities.Thesemethodseitherconstraintheangularmomentumenvelopeand/orgimbalangles,orapplynullmotiontoavoidsingularityencounters.AsdiscussedinChapter 2 ,amethodthatusesonlynullmotioncannotavoidorescapeellipticinternalsingularities[ 44 47 ]. 55


Figure4-2. Steeringalgorithms 4.2.1ConstrainedSteeringAlgorithms Constrainedsteeringalgorithmseitherconstrainthegimbalanglesand/oruseableangularmomentumtoavoidsingularities.Thesesteeringlawsareaformofsingularityavoidancethattakesintoaccountthelocationsofsingularitiesapriori.Asaconsequenceofnotusingtheentireangularmomentumspace,thesesteeringlogicsaretypicallymoreeffectiveforsystemswheretheSGCMGsareoversized.Honeywellhaspatentedmethodsthatdonotexplicitlyusenullspacebutthatimplicitlydosobycreatingconstraintsthatkeepthegimbalsawayfromsingularitywithoutneedingtorecognizetheirpresenceexplicitly[ 57 59 ].AsimpleexampleisthesteeringlogicforscissoredpairsinChapter 3 ,wheremereconstraintsareusedtokeepthearrayoutoftrouble.Thismethodisabletoguaranteesingularityavoidanceandanaavailabletorquebutreducestheavailableworkspaceofthesystembyrequiringittobesingularityfree[ 52 60 ].4.2.2NullMotionAlgorithms4.2.2.1Localgradient(LG) Singularityavoidancealgorithmsknownaslocalgradient(LG)methodsusenullmotiontokeeptheJacobianmatrixfrombecomingsingular.Thisisaccomplishedthroughchoiceofthenullvectordtomaximizeanobjectivefunctionthatrelatesthe 56


distancefromsingularitysuchastheJacobianmatrixconditionnumber,smallestsingularvalue[ 61 62 ],orthesingularitymeasuremwhichisexpressedas m=p det(AAT)(4) Anexampleofthenullvectorcalculationisd=rf=@f @m@m @T=)]TJ /F4 11.955 Tf 9.3 0 Td[(1 m2@m @T(4) wheretheobjectivefunctionf=1=m[ 42 63 64 ].Minimizationofthisobjectivefunctionmaximizesthedistancefromsingularitybymaximizingthesingularitymeasure.TheLGmethods,however,cannotavoidorescapeellipticinternalsingularitiesbecausetheyapplyonlynullmotion[ 44 64 ].Thenullvectordcanbearbitrary,althoughtheprojectionmatrixwhichmapsitontothenullspaceisconstrained. Awayofchoosingthenullmotionvectortosteergimbalstoalternatenonsingularcongurationsbeforemaneuveringisknownaspreferredtrajectorytracking[ 63 65 67 ].Preferredtrajectorytrackingisaglobalmethodthatcalculatesnonsingulargimbaltrajectoriesofine.Thegimbalsconvergetothesetrajectoriesusingnullmotiontominimizeanerror()]TJ /F10 11.955 Tf 11.96 0 Td[().Thegimbalratesusingthismethodare _=1 h0A+_h+[1)]TJ /F8 11.955 Tf 11.95 0 Td[(A+A]()]TJ /F10 11.955 Tf 11.96 0 Td[()(4) wherearethepreferredtrajectoriesandisthesingularityparameterdenedby =0exp)]TJ /F11 7.97 Tf 6.59 0 Td[(m2(4) withconstants0and.Sincethismethodcalculatesthepreferredtrajectoriesofine,itisnotreal-timeimplementable.Also,preferredtrackingreliesentirelyonnullmotionandthuswillbeunabletoescapeellipticinternalsingularities. 57

PAGE 58 TheGeneralizedInverseSteeringLaw(GISL)providesapseudo-inversewhichisavariationoftheMoore-Penrosepseudo-inverse.ThismethoddenesanotherJacobianmatrixBwhichhaseachofitscolumnsorthogonaltotheassociatedcolumnoftheoriginalJacobianmatrixA(i.e.,ai?bi,notnecessarilyai)]TJ 1.33 1.32 Td[()]TJ /F6 11.955 Tf -.33 .17 Td[(?bjwhereA=[a1a2a3a4]andB=[b1b2b3b4])[ 68 ].Therefore,asanexampleforafour-CMGpyramidarrangement,thematrixAandBhavethefollowingform A=266664)]TJ /F7 11.955 Tf 9.29 0 Td[(ccos(1)sin(2)ccos(3))]TJ /F7 11.955 Tf 9.3 0 Td[(sin(4))]TJ /F7 11.955 Tf 9.3 0 Td[(sin(1))]TJ /F7 11.955 Tf 9.3 0 Td[(ccos(2)sin(3)ccos(4)scos(1)scos(2)scos(3)scos(4)377775(4) and B=266664)]TJ /F7 11.955 Tf 9.3 0 Td[(csin(1))]TJ /F7 11.955 Tf 9.29 0 Td[(cos(2)csin(3)cos(4)cos(1))]TJ /F7 11.955 Tf 9.3 0 Td[(csin(2))]TJ /F7 11.955 Tf 9.29 0 Td[(cos(3)csin(4)ssin(1)ssin(2)ssin(3)ssin(4)377775(4) wherec,sarethecosineandsineofthepyramidskewangleandiarethegimbalangles,respectively.Thepseudo-inverseofthissteeringlawwiththediscussedmatricesis AGISL=(A+B)T(A(A+B)T))]TJ /F2 7.97 Tf 6.59 0 Td[(1(4) Itisimportanttonotethatthispseudo-inversedoesnoteliminatetheproblemofinternalsingularities.TheGISLaddsnullmotionfromtheadditionofBandthereforecouplesthenullandforcedsolutionintoasingleinverseandthus,itisnotabletoavoidellipticinternalsingularities. Proof: Claim:TheGISLprovidesonlynullmotionthroughB 58


_=AGISL=(A+B)T(A(A+B)T))]TJ /F2 7.97 Tf 6.58 0 Td[(1_h Thetorqueerroris _h)]TJ /F8 11.955 Tf 11.95 0 Td[(A_=A(A+B)T(A(A+B)T))]TJ /F2 7.97 Tf 6.59 0 Td[(1_h)]TJ /F8 11.955 Tf 13.48 2.66 Td[(_h=_h)]TJ /F8 11.955 Tf 13.48 2.66 Td[(_h=0 IfthematrixB=A then AGISL=((1+)A)T((1+)AAT))]TJ /F2 7.97 Tf 6.59 0 Td[(1=A+ Therefore,thematrixBwhosecomponentsareorthogonaltoAmustonlyprovidenullmotionandthosethatarealongAvanish.Because,theGISLoranygeneralizedinverseusedforsingularityavoidanceonlyaddsnullmotion,itisunabletoavoidellipticinternalsingularities.4.3SingularityEscapeAlgorithms Singularityescapemethods,knownaspseudo-inversesolutions,addtorqueerrortopassthroughorescapeinternalsingularity[ 39 40 69 70 ].Thesemethodsdonottakeintoconsiderationthetypeofinternalsingularitythatisbeingapproachedwhenaddingtorqueerror. TheSingularityRobust(SR)inverseisavariationoftheMoore-Penrosepseudo-inverse[ 69 ]where,apositivedenitematrix1composedofanidentitymatrixscaledbythesingularityparameterinEq.( 4 )isaddedtothepositivesemi-denitematrixAAT.Thepseudo-inverseofthismethodhastheform ASR=AT(AAT+1))]TJ /F2 7.97 Tf 6.58 0 Td[(1(4) TheSRinverseisabletoescapebothhyperbolicandellipticsingularities[ 44 ],although,isineffectiveingimbal-lockescape.Toovercomethissituation,amodiedSRinverseknownastheGeneralizedSingularityRobust(GSR)pseudo-inversewasdeveloped[ 39 40 ]. 59

PAGE 60 TheGSRinverseapproachreplacestheconstantdiagonalpositivedenitematrix1withatime-varyingpositivedenitesymmetricmatrixE E=266664112113231377775,i=0sin(!it+i)(4) wheretheoff-diagonaltermsofEaretimedependenttrigonometricfunctionswithfrequency!iandphaseshifti.TheGSRinverseprovidesameansofescapeofthegimbal-lockcongurationassociatedwithasystemofSGCMGs.TheGSRpseudo-inversehastheform AGSR=AT(AAT+E))]TJ /F2 7.97 Tf 6.59 0 Td[(1(4) andliketheSRinverse,isguaranteedtoavoidbothhyperbolicandellipticinternalsingularities. AnothermodicationoftheSRinverseknownastheSingularDirectionAvoidance(SDA)onlyappliestorqueerrorinthesingulardirectionandthereforereducestheamountoftorqueerrorneededforsingularityescape.TheSDAmethoddecomposestheJacobianmatrixusingasingularvaluedecomposition(SVD)todetermineitssingularvalues.Thematrixofsingularvaluesisregulatedwiththeadditionoferrortothesmallestsingularvalue3sothatthepseudo-inverseisdened.Thepseudo-inverseusingSDAhastheform 60


ASDA=V2666666641 10001 20003 32+000377777775=VSDAUT(4) whereiarethesingularvalues.Regulatingonlythesmallestsingularvalue,reducestheamountoftorqueerroraddedandcreatessmoothergimbalratetrajectorieswhencomparedtotheSRandGSRinverses[ 70 ].ThisisobviouswhentheSRinversedecomposedusingSVDas ASR=V2666666641 12+0002 22+0003 32+000377777775UT=ASR=VSRUT(4) whereallthesingularvaluesareregulatedandhencethereistorqueerrorinalldirections.ItisclearfromEqs.( 4 )and( 4 )thatSRinverseandSDAaresusceptibletogimbal-lockbecausewhentheoutputtorqueisalongthesingulardirection_h/s=u3thenitisinthenull(SDAUT)andnull(SRUT)thusencounteringgimbal-lockasnoconsequencetothesizeofthetorqueerroraddedfrom.WithouttheperturbationstotheJacobianmatrixthatarenotgimbalstatedependentatgimbal-lockthesystemremainslockedinasingularconguration. RecallfromChapter 3 ,thatforafourCMGpyramidarrangement,therankisneverlessthantwoandthereforeitisacceptabletoregulateonlythesmallestsingularvalue.However,iftheskewangleismadecloseto0,90,180,or270deg(i.e.,boxorplanararrangement),theJacobianmatrixforthesearrangementswillhaveatleasttwosmallsingularvalueswhennearsingularityandregulationofthesmallestsingularvaluemaybeineffective. 61

PAGE 62 TheFeedbackSteeringLaw(FSL)providesasolutiontothegimbalrateswithoutusinganinverse.ThismethodisderivedfromaminimizationofthetorqueerrorwhichissimilartohowtheSRinverseisderived.TheoptimizationfordeningFSLhasthefollowingstructure min_2R41 2264e_375T264K100K2375264e_375(4) whereK1andK2arepositivedenitegainmatrices,ande=_h)]TJ /F8 11.955 Tf 12.04 0 Td[(A_.ThisminimizationreducestotheSRinversewhenK1=1andK2=1andwhere=0exp)]TJ /F11 7.97 Tf 6.58 0 Td[(m2)fromEq.( 4 ). TheFSLmethodhasK2=1andK1=K(s)asacompensator.ThecompensatorisderivedfromanH1minimization minK(s)2R330B@w1(s)[1+AK(s)])]TJ /F2 7.97 Tf 6.58 0 Td[(1w2K(s)[1+AK(s)])]TJ /F2 7.97 Tf 6.59 0 Td[(11CA1(4) wherew1(s)andw2areweightingmatrices.Thew1(s)matrixisdenedbelow w1(s)=264AKBKCK0375(4) whereAK,BK,andCKarematricesassociatedwithstate-spacemodelofthesystem.Thew2matrixisconstantandis w2=1 w144(4) wheretheconstantwboundsthegimbalrates.Thestate-spacemodelofthesystemhastheform 62


_^x=AK^x+BKe_=CK^x(4) TheoutputmatrixCKisanexplicitfunctionoftheCMGgimbalanglesexpressedas CK=ATb!2 P(4) withbasapositivescalarassociatedwiththebandwidthand Pisthesteady-statesolutiontotheRiccatiequationofthestate-spacesysteminEq.( 4 ).Usingthefeedbackofthesystem,Eq( 4 )willprovideasolutiontothegimbalratesthatdoesnotrequireapseudo-inverse.Itshouldbenotedthatthesystemmaystartoutstable,however,theobservabilityofthesystemmaybelostresultingininstability,duetoCK'sexplicitdependenceonthegimbalangle(i.e.,H(s)=CK(s1)]TJ /F8 11.955 Tf 12.41 0 Td[(AK))]TJ /F2 7.97 Tf 6.59 0 Td[(1BKwhereCK6=constant).Forfurtherinformationonthedevelopmentofthismethod,pleasesee[ 71 ].ThissteeringalgorithmwasshowntogounstableforcertainvaluesofCKcorrespondingtospecicgimbalanglessets(see[ 72 ]). TheSingularityPenetrationwithUnitDelay(SPUD)algorithmescapessingularitythroughreuseofthepreviousgimbalratecommandwhenatacertainthresholdofsingularity[ 73 ].Thepreviouscommandissavedthroughazeroth-orderholdtothesystem.Escapeofasingularityisalwayspossibleunlessthesystemisinitiallyatthethresholdofsingularity,thenthereisnopreviouscommandtouseforsingularityavoidance.Also,SPUDisnotintendedforattitudetrackingmaneuvers.TheSPUDalgorithmaccumulatesattitudetrackingerrorwhileescapingsingularityandtherearenoguaranteesonhowlongitwilltaketoescapesingularityandhowlargethetorquedisturbancewillbeonthespacecraftasitsperformanceisdirectlyassociatedwiththesystemandthechoiceofsingularitythreshold. 63


4.4SingularityAvoidanceandEscapeAlgorithms Singularityavoidanceandescapealgorithmsavoidsingularitiesthroughnullmotionwheneverpossibleandusetorqueerrorforescapewhentheyarenot. PreferredgimbalanglesareasetofinitialgimbalanglesforSGCMGsthatcanbereachedbynullmotion.Theseanglesarepreferredsincemaneuversoriginatingfromthemavoidasingularconguration[ 74 ].ThissetofanglesisfoundbybackwardsintegrationoftheEq.( 4 )andtheattitudeequationsofmotion.Ithasbeenshownthatthismethodcannotavoidsingularitiesiftheinitialsetofgimbalanglesis0=[45)]TJ /F4 11.955 Tf 9.29 0 Td[(4545)]TJ /F4 11.955 Tf 9.3 0 Td[(45]Tdeg[ 74 ].Sincethenullspaceprojectionmatrixisundenedatsingularity,theSRinverseisusedinplaceoftheMoore-Penrosepseudo-inverseofAin_nas _n=[1)]TJ /F8 11.955 Tf 11.96 0 Td[(ASRA]d(4) Asaresult,thiscausesthesystemtoaddtorqueerrorwhenatsingularity.Inpractice,thismethodactsasanofineoptimizationwhichdeterminestheinitialset(s)ofthegimbalanglesthatwillgivesingularityfreemaneuver(s).However,itisnotpossibletogofromonetoanypointingimbalspacethroughnullmotionitselfbecausetherewillneverbendimensionsofnullspace. TheInner-ProductIndex(IPI)combinedwiththeoptimalsteeringlaw(OSL)isusedtodetermineasteeringalgorithmthatproducesminimumtorqueerrorwhilebothavoidingandescapinginternalsingularities[ 75 ].Thesingularityindexisaddedtotheminimization min_2R4,e2R3[)]TJ /F7 11.955 Tf 9.3 0 Td[(cV(+_t)+1 2_TW_+eTR)]TJ /F2 7.97 Tf 6.59 0 Td[(1e](4) 64


wheretistheone-steptimedelay,V(+_t)istheIPI,e=_h)]TJ /F8 11.955 Tf 12.57 0 Td[(A_(i.e.,torqueerror),cisapositivescalar,andWandR)]TJ /F2 7.97 Tf 6.58 0 Td[(1arepositivedeniteweightingmatrices.TheIPIisapproximatedbyaTaylorseriesexpansionuptothe2ndorderas V(+_t)V+@V @T_Tt+1 2_T@2V @2_t2(4) wheretheIPIVisexpressedasasumofsquareofinnerproductsofthecolumnvectorsoftheJacobian. V=1 24Xi=j=1,i6=j(aTiaj)2(4) TheresultoftheminimizationinEq.( 4 )usingthisapproximationofVis _=H)]TJ /F2 7.97 Tf 6.58 0 Td[(1AT(AH)]TJ /F2 7.97 Tf 6.58 0 Td[(1AT+R))]TJ /F2 7.97 Tf 6.59 0 Td[(1_h+[H)]TJ /F2 7.97 Tf 6.59 0 Td[(1AT(AH)]TJ /F2 7.97 Tf 6.59 0 Td[(1AT+R))]TJ /F2 7.97 Tf 6.58 0 Td[(1AH)]TJ /F2 7.97 Tf 6.58 0 Td[(1)]TJ /F8 11.955 Tf 11.96 0 Td[(H)]TJ /F2 7.97 Tf 6.59 0 Td[(1]g(4) wheretheHessianmatrixHisdenedas H=ct2ggT+W(4) withgradientg=@V @T.TheweightingmatrixRshownpreviouslyintheminimizationofEq.( 4 ),isexpressedas R=U266664000000000exp)]TJ /F11 7.97 Tf 6.59 0 Td[(23377775UT(4) where3isthesmallestofthesingularvaluesoftheJacobianmatrix,0andarepositivescalarsandUistheunitarymatrixmadeupoftheleftsingularvectorsfromthesingularvaluedecompositionoftheJacobianmatrixA.ThisadditionoftorqueerrorintothegimbalratestateequationisanalogoustotheSDAmethodexceptthatitisalso 65


addedtothefreeresponsesolution[ 70 ].Itshouldbenotedthatthissteeringalgorithmdoesnotconsidertheformofinternalsingularitiesandtherefore,doesnottrulyminimizetheamountoftorqueerrorforsingularityescape.Thisisbecausenon-degeneratehyperbolicsingularitiesareavoidablethroughnullmotionwithouttheuseoftorqueerror.Atanon-degeneratehyperbolicsingularityRisnonzeroandthustorqueerrorisstilladded(see[ 72 ]).4.5OtherSteeringAlgorithms Otherpublishedsteeringalgorithmsthathavenotbeendiscussedcanbefoundinthereferences[ 44 61 76 81 ].Thesemethodsincludemathematicaltechniquessuchasneuralnetworks,optimization,andgame-theory.4.6SteeringAlgorithmComputationComparison Ananalysiscomparingthecomputationfortheimplementationofthementionedsteeringalgorithmsisdifcultduetolackofinformationonhowsomewerecodedinliterature.Forexample,someofthesealgorithmsareofineandmayrequirealargenumberofmemorycallsandstoredmemorybutnotasmanyops.Itishowever,usefultoquantifyingsomeofthepreviouslydiscussedsteeringalgorithmsintermsofoatingpointoperationsthatarenotcalculatedofline.TheseareshowninTable 4-1 foralgorithmswhereopsmakeagoodcomparison.Inthistable,themetricofcomparisonisanapproximatenumberofopspertimestep. Table4-1. AlgorithmFlopsm=row(A)andn=column(A) VariableValue MPO(m4)LGO(m4)GISLO(m4)SRO(m4)GSRO(m4)SDAO(nm3)FSLO(mn2)SPUDO(m4)OptimalSteeringO(m4)+O(nm3) 66



CHAPTER5STEERINGALGORITHM-HYBRIDSTEERINGLOGIC5.1HybridSteeringLogic Existingsteeringlogics(seeChapter 4 )donotexplicitlyconsiderthetypeofsingularitythatisbeingencounteredandthus,donotcompletelyaddressattitudetrackingperformanceofSGCMGattitudecontrolsystems.AproposedmethodknownastheHybridSteeringLogic(HSL)whichutilizestheknowledgeofthetypeofsingularityencountered(i.e.,ellipticorhyperbolicsingularities)toimprovetheattitudetrackingperformanceoftheSGCMGattitudecontrolsystem,isdevelopedforafour-SGCMGpyramidarrangementataskewangle=54.74deg.Byusingahybridapproach,HSLactsasanLGmethod(i.e.,nullmotionforsingularityavoidance)athyperbolicsingularityandanSDAmethod(i.e.,pseudo-inversesolutionsforsingularityescape)atellipticsingularity.Also,becauseHSLisdevelopedforafour-SGCMGpyramidarrangement,thereisnoexistenceofdegenerate-hyperbolicsingularities[ 41 ].ThechallengeistodeveloptheappropriatesingularitymetricssuchthattheLGandSDAcomponentsofthehybridstrategydonotcounteracteachotherduringoperation.5.1.1InternalSingularityMetrics ThesingularitymetricsdevelopedareofsimilarformasthesingularityparameterinEq.( 4 )withtheadditionoftermsrelatingtotheformoftheactualsingularity. =0exp)]TJ /F5 7.97 Tf 6.58 0 Td[(a exp)]TJ /F11 7.97 Tf 6.59 0 Td[(1m(5) =0exp)]TJ /F5 7.97 Tf 6.59 0 Td[(b exp)]TJ /F11 7.97 Tf 6.59 0 Td[(2m(5) wherea,b,1,2,0and0arepositivescalarconstantsandmisthesingularitymeasureasdenedinEq.( 4 ).Awayfromsingularity,afour-SGCMGpyramidarrangementataskewangle=54.75deg,hasthematrixQ2R.AtsingularitythisSGCMGarrangementhasQ2R22(seeChapter??)andtherefore,thedet(Q) 68


willbezeroornegative(i.e.,Qisnegativesemi-deniteorindenite)forhyperbolicsingularitiesandpositive(i.e.,Qisdenite)forellipticsingularities.Takingthisintoaccount,parameters and aredenedas =jQ0)]TJ /F7 11.955 Tf 11.96 0 Td[(det(Q)j(5) =1 jQ0)]TJ /F7 11.955 Tf 11.96 0 Td[(det(Q)j=1 (5) whereQ0isascalarvaluechosenonthesameorderofmagnitudeofdet(Q)butgreatertoscaletheresponseof and .ItisdifculttoanalyticallydeneQ0sinceitdependsonthemaximumvalueofdet(Q)(i.e.,det(Q)varieswithgimbalangleandthereforethemaximummustspanallcombinationsofthegimbalangles)whichisofhighdimensionalityandhighlynonlinear.However,throughsimulationofafour-SGCMGpyramidarrangementataskewangle=54.75deg,itwasfoundthatjdet(Q)j<1andthereforewedeneQ01.Inaddition,itisimportanttonotethattheconstantparametersa,b,1,2,0,and0areusedtomorphtheHSLsteeringlogicintotherespectiveLGandSDAmethodswhenappropriate:(e.g.,iftheparametersa=b=0=0and06=0thentheHSLmethodistheLGmethod).Therefore,thechoiceofmetricsandinthiswayensuresthatnullmotionwillbeaddedwhenapproachingahyperbolicsingularityandtorqueerrorwithlessnullmotionwillbeaddedwhenapproachinganellipticsingularity.ItshouldbenotedthatwhenusingHSLdet(Q)isnormalizedbythenominalangularmomentumh0.5.1.2HybridSteeringLogicFormulation Theproposedsteeringlogicisdenedas _=1 h0ASDA,_h+[1)]TJ /F8 11.955 Tf 11.96 0 Td[(A+A]d(5) whereASDA,is 69


ASDA,=V2666666641 10001 20003 32+000377777775UT(5) Ifitisassumedthattheanalyticfunctionforthegradientvectordisderivedofineandthecalculationofitateachtime-stepislessthanthatforSVD,thisalgorithmhasthesamenumberofopsonorderasSDAfromTable 4-1 ofO(nm3)fromtheSVD.ThedifferencebetweentheconventionalASDAandASDA,istheparameterthatregulates3.InASDA,theregulationparameteris(i.e.,differentfrominChapter 4 byusingminsteadofm2)whichis =0exp)]TJ /F11 7.97 Tf 6.59 0 Td[(m(5) withpositiveconstants0and,butwithASDA,thesingularityparameterisdenedinEq.( 5 )whichregulatestheamountofinducedtorqueerrorinthevicinityofellipticsingularities.ThroughaSVDdecompositionofA,Eq.( 5 )canbewrittenas _=1 h0ASDA,_h+[1)]TJ /F8 11.955 Tf 11.96 0 Td[(V264100T0375VT]d(5) Here,thenullmotionprojectionmatrixisexpressedasafunctionofnonsingularmatricesV.Also,veryrobustnumericalalgorithmsexistforcomputingtheSVD,soitscomputationalriskinareal-timeimplementationisnotparticularlyhigh. Thescalarthatregulatesthemagnitudeofthenullmotionis.Thenullvectordisinthedirectionofthegradientoff=)]TJ /F7 11.955 Tf 9.3 0 Td[(det(AAT)=)]TJ /F7 11.955 Tf 9.29 0 Td[(m2andmaximizesthedistancefromsingularity. Thischoiceofthisobjectivefunctionreducesthecomputationneededforthegradient(i.e.,thederivativeof()]TJ /F7 11.955 Tf 9.3 0 Td[(det(AAT))islesscomputationallyintensivethanthe 70


derivativep (det(AAT))andensuresthattheadditionofnullmotionwillnotapproachinnityattheregionofsingularityforcasessuchasf=1 mandthen@f @=)]TJ /F2 7.97 Tf 13.7 4.7 Td[(1 m2@m @.ItshouldbementionedthatthenullvectorisanonlinearfunctionofthegimbalanglesandissimpliedduetothesymmetryofthefourCMGpyramidarrangement.ToprovethefeasibilityofHSL,astabilityanalysisisconducted.5.2LyapunovStabilityAnalysis ThecandidateLyapunovfunction V=1 2!TK)]TJ /F2 7.97 Tf 6.59 0 Td[(1J!+eTe+(1)]TJ /F7 11.955 Tf 11.96 0 Td[(e4)2(5) ischosenforthisanalysisandcanberewrittenas V=zTMz(5) wherez=[!TeT(1)]TJ /F7 11.955 Tf 12.37 0 Td[(e4)]TandM=diag(1 2K)]TJ /F2 7.97 Tf 6.58 0 Td[(1J,1,1).ConsequentlytheLyapunovfunctionisboundedas minjjzjj2Vmaxjjzjj2(5) whereminandmaxdenotetheminimumandmaximumeigenvaluesofM.Thisboundwillbecomeusefullaterintheanalysis. Arest-to-restquaternionregulatorcontrollerisgivenbyEq.( 5 )fortheinternalcontroltorque,ischosenforitsownheritageandthefactthatityieldsanglobalasymptoticstablecontrolsolutionproventhroughLaSalle'sInvariantTheorem[ 63 ]. =)]TJ /F8 11.955 Tf 9.3 0 Td[(Ke)]TJ /F8 11.955 Tf 11.95 0 Td[(C!+!J!(5) GainmatricesK=2kJandC=cJofEq.( 5 )arepositivedeniteandsymmetric.Assumingrigidbodydynamics,thespacecraft'sangularmomentumisgivenby 71


H=J!+h(5) TherotationalequationsofmotioncomefromtakingtheinertiatimederivativeofEq.( 5 )as_!=J)]TJ /F2 7.97 Tf 6.59 0 Td[(1[act)]TJ /F10 11.955 Tf 11.95 0 Td[(!J!](5) withSGCMGoutputtorque_h=)]TJ /F10 11.955 Tf 9.3 0 Td[()]TJ /F10 11.955 Tf 11.96 0 Td[(!h=h0A_(5) where!isthespacecraftangularvelocity,Jisthespacecraftcentroidalinertia,Histhetotalsystemangularmomentum,andhistheangularmomentumfromtheCMGs.Thespacecraft'sangularvelocityandtheCMGangularmomentaaregovernedbyEqs.( 5 )and( 5 )respectively,whereactistheactualcontroltorque(i.e.,maydifferduetoinducedtorqueerrorforsingularityescape).Itisassumedherethatthecontributiontothedynamicsfromthegimbal-ywheelassemblyinertiasisnegligibleandthereforeJisconstant. Theactualcontroltorqueactbasedonthemappingofthegimbalratesis act=)]TJ /F7 11.955 Tf 9.29 0 Td[(h0A_)]TJ /F10 11.955 Tf 11.95 0 Td[(!h=)]TJ /F8 11.955 Tf 9.3 0 Td[(A(ASDA,_h+[1)]TJ /F8 11.955 Tf 11.96 0 Td[(V264100T0375VT]d))]TJ /F10 11.955 Tf 11.96 0 Td[(!h(5) andneedstobeconsideredintheLyapunovanalysisforstabilityoftheattitudecontroller/steeringalgorithmcombination.Whensimplied,Eq.( 5 )becomes act=U26666400000000)]TJ /F11 7.97 Tf 6.59 0 Td[( 23+377775UT[)]TJ /F3 11.955 Tf 9.3 0 Td[()]TJ /F10 11.955 Tf 11.95 0 Td[(!h]+(5) wherethestabilityofthesystemisaffectedbythetorqueperturbationmatrixHSLfromASDA,denedas 72


HSL=U26666400000000 23+377775UT(5) Thespacecraftattitudeerrorkinematicsisgovernedby _e=)]TJ /F4 11.955 Tf 10.5 8.08 Td[(1 2!e+1 2!e4(5) _e4=)]TJ /F4 11.955 Tf 10.49 8.09 Td[(1 2!Te(5) whereeisthequaternionerrorvectorelementsande4isitsscalarelement.ThetimederivativeoftheLyapunovfunctionis _V=!TK)]TJ /F2 7.97 Tf 6.59 0 Td[(1[act)]TJ /F10 11.955 Tf 11.96 0 Td[(!J!]+2eT[)]TJ /F4 11.955 Tf 10.49 8.09 Td[(1 2!e+1 2!e4]+2(1)]TJ /F7 11.955 Tf 11.95 0 Td[(e4)1 2!Te(5) Equation( 5 )canbereducedbysubstitutingintheexpressionforactfromEq.( 5 )andthedesiredcontroltorquevectorfromEq.( 5 ).ThetimederivativeoftheLyapunovfunctionnowyields _V=)]TJ /F10 11.955 Tf 9.29 0 Td[(!TK)]TJ /F2 7.97 Tf 6.59 0 Td[(1[C)]TJ /F10 11.955 Tf 11.95 0 Td[(HSL(C+H)]!+!TK)]TJ /F2 7.97 Tf 6.59 0 Td[(1HSLKe(5) ormorecompactly _V=)]TJ /F7 11.955 Tf 13.75 8.09 Td[(c 2k!T!+c 2k!TJ)]TJ /F2 7.97 Tf 6.59 0 Td[(1HSLJ!+1(5) where1=1 2k!TJ)]TJ /F2 7.97 Tf 6.59 0 Td[(1HSLH!+!TJ)]TJ /F2 7.97 Tf 6.59 0 Td[(1HSLJe(5) 73


Since 23+1,Eq.( 5 )canbeusedtorewriteEq.( 5 )as_V)]TJ /F7 11.955 Tf 26.37 8.09 Td[(c 2k!TJ)]TJ /F2 7.97 Tf 6.58 0 Td[(1U266664100010001377775UTJ!+1=)]TJ /F3 11.955 Tf 9.29 0 Td[(!TR1!+1(5) andcanbefurtherboundedas_V)]TJ /F3 11.955 Tf 21.92 0 Td[(1jjzjj2+1(5) where1istheminimumeigenvalueofthepositivesemi-denitematrixR1(i.e.,1=0atsingularity),=c 2k,andjjzjjisdenedinEq.( 5 ).SubstitutingEq.( 5 )intoEq.( 5 )yields_V)]TJ /F3 11.955 Tf 25.72 8.09 Td[(1 maxV+1(5) ThesolutiontothedifferentialequationinEq.( 5 )inaVolterraintegralformisVV(0)exp()]TJ /F28 5.978 Tf 10.1 4.4 Td[(1 maxt)+Ztt0exp()]TJ /F28 5.978 Tf 10.1 4.4 Td[(1 max(t)]TJ /F11 7.97 Tf 6.59 0 Td[())1()d(5) TheerrorcanbeboundedfromEq.( 5 )as jjzjj2V(0) minexp()]TJ /F28 5.978 Tf 10.1 4.4 Td[(1 maxt)+1 minZtt0exp()]TJ /F28 5.978 Tf 10.1 4.4 Td[(1 max(t)]TJ /F11 7.97 Tf 6.58 0 Td[())1()d(5) Atsingularitywhen1=0,theerroris jjzjj2Vs+1 minZtts1()d(5) whereVsistheerrorattimetheoccurrenceofsingularityattimets. StabilitycannotbeprovenfromEq.( 5 ).Itisassumedthatthesystemwillnotremainlockedinsingularityexceptforthespecialcaseofgimbal-lock.Ifcomponentsofthetorqueneededforstabilityareactuallyinthesingulardirection,periodsofinstabilitymayoccuratsingularity.Thedurationofthisinstabilityisdependentontheselectionandsizingofthesingularityparametersandwhichprovidetorque 74


errorforsingularityescapeand/ornullmotionforsingularityavoidance.Fromapracticalperspective,stabilitycannotbeprovenforthisLyapunovfunctionatsingularity,sinceatsingularity,thereisnotorqueavailableinthesingulardirection. Forthespecialcaseofasingularitywithgimbal-lock,theangularvelocityofthespacecraftisconstantassumingtheabsenceoffrictionandexternaltorquesinthesystem.Inthiscase,thecontributionfrom1totheerrorisboundedandevensometimeszero.Thiscanbeshownbyevaluatingtheexpressionfortheangularaccelerationatgimbal-lockwhichis _!=0=J)]TJ /F2 7.97 Tf 6.59 0 Td[(1[0+!H](5) ItisclearthatEq.( 5 )issatisedonlywhentheproduct!H=0whichisonlytruewhen!isparalleltoH,!=0,orH=0.When!isparalleltoHorH=0,1=!TJ)]TJ /F2 7.97 Tf 6.59 0 Td[(1HSLJewhichisaboundedsinusoidwhoseintegralisalsoaboundedsinusoid.Therefore,forthesetwocases,theerrorisboundedatgimbal-lock.When!=0,1=0andtheerrorissimplylockedatVs.Itshouldbementionedthatattimesawayfromsingularity,theerrormonotonicallydecreasesbecausethecontributionfrom1totheerrorbecomesnegligible.Careneedstobetakeninthedesignofthesingularityparametersothattheminimumsteady-stateerrorisachievedwhilemeetingtheconstraintsoftheactuators. Thesteady-stateerrorassumingthatthesystemshasasingularityfreeperiodtowardstheendofthemaneuver(i.e.,doesnotendatsingularity)is jjz(1)jj21 minexp()]TJ /F28 5.978 Tf 10.1 4.39 Td[(1 max1)Z1tsexp(1 max)1()d(5) ThisexpressionisindeterminatesoapplicationofL'Hopital'sruletoEq.( 5 )yields limt)177(!1d dt1 minRttsexp(1 max)1d d dtexp(1 maxt)=)]TJ /F3 11.955 Tf 18.21 8.09 Td[(max min11(1)(5) 75


whichsuggeststhatasufcientlylargevalueofc(i.e.,larger)willlowertheamountofsteady-stateerrorgivingyouauniformly-ultimatelybounded(UUB)resultawayfromsingularity.Whenthemaneuverisnished,theeffectof1ontheerrorwillbecomeaconstantassumingthemaneuverendsatrest.Itshouldbenotedthatawayfromsingularitythesizeof1exponentiallydecreasesduetothebehaviorofHSL.ThedifferenceinimpactofHSLratherthanSRinverseonstabilitycanbeobservedfromthemagnitudeofthepositivesemi-denitematrixHSLinEq.( 5 )comparedtothematrixshowninEq.( 5 ).TheSRmatrixhasalargernormandthereforehasaworseUUBevenforsufcientlylargevaluesof.FromcomparingEqs.( 5 )and( 5 ),theSDAmethodhasasimilaramountoftorqueerroraddedwhencomparedtoHSL,althoughitwilladdthistorqueerrorwheneverthesingularityapproachednottakingintoaccounttheform. SR=U266664 21+000 22+000 23+377775UT(5) Theaboveresultsareonlyfortheattitudecontroller/steeringalgorithmcombination.Forexample,anattitudecontrollerwhosetorquetrajectorywaschosentoavoidtheoccurrenceofsingularitiesmaynothavetheperiodsofpossibleinstabilitiesatsingularityandthusmayprovidebetterstabilityperformance.However,noreal-timecontrollerofthisformexists(i.e.,onethatensuressingularityavoidance)andthuswasnotconsideredinthefollowingsimulations.5.3NumericalSimulation ToevaluatetheperformanceoftheproposedHSLagainstheritagesteeringlogics(i.e.,LGandSDA),simulationswereperformedusingafour-SGCMGpyramidalarrangementwithaskewangleof=54.74deg.Toensureafaircomparison,thecontrollogicandsatellitemodelwereidenticalforallsimulations.Foreachsteering 76


algorithm,threedifferentscenariosweresimulated:(1)startinginazero-momentumconguration=[0000]Tdeg(i.e.,farfromsingularity);(2)startingnearanellipticexternalsingularity=[105105105105]Tdeg;and(3)startingnearanhyperbolicsingularity=[15105195)]TJ /F4 11.955 Tf 9.3 0 Td[(75]Tdeg.ThesingularityconditionswereveriedforeachcasebyobservingthesingularitymeasuredenedinEq.( 4 ).Forthesesimulations,thefollowingperformancemeasureswerecompared:(i)thetransientresponseoftheerrorquaternion,(ii)theamountanddurationofsingularityencounter,(iii)themagnitudeofgimbalrate,(iv)theamountoftorqueerror(i.e.,_h)]TJ /F7 11.955 Tf 12.13 0 Td[(h0A_)forsingularityescape,and(v)nullmotioncontribution.Additionally,,,anddet(Q)arealsoconsidered. TheJacobianassociatedwiththispyramidalcongurationis A=266664)]TJ /F7 11.955 Tf 9.3 0 Td[(c()c(1)s(2)c()c(3))]TJ /F7 11.955 Tf 9.3 0 Td[(s(4))]TJ /F7 11.955 Tf 9.3 0 Td[(s(1))]TJ /F7 11.955 Tf 9.3 0 Td[(c()c(2)s(3)c()c(4)s()c(1)s()c(2)s()c(3)s()c(4)377775,(5) andtheassociatedangularmomentumvectoris h=h0266664)]TJ /F7 11.955 Tf 9.3 0 Td[(c()s(1))]TJ /F7 11.955 Tf 11.96 0 Td[(c(2)+c()s(3)+c(4)c(1))]TJ /F7 11.955 Tf 11.95 0 Td[(c()s(2))]TJ /F7 11.955 Tf 11.95 0 Td[(c(3)+c()s(4)s()(s(1)+s(2)+s(3)+s(4))377775(5) Allsimulationsareperformedusingafourth-orderxedtimestepRungaKuttawiththeparametersshowninTable 5-1 .TheactuatorparameterschosenforthissimulationarebasedontheHoneywellM95SGCMGs,whicharesizedforthesatellitesystemchosenforsimulation[ 82 ]. 77


Table5-1. ModelParametersVariableValueUnits J0@100)]TJ /F4 11.955 Tf 9.3 0 Td[(21.5)]TJ /F4 11.955 Tf 9.3 0 Td[(2900)]TJ /F4 11.955 Tf 9.3 0 Td[(601.5)]TJ /F4 11.955 Tf 9.3 0 Td[(6010001Akgm254.74dege0[0.043550.087100.043550.99430]T\000!0[000]Tdeg=sh0128Nmsk0.051=s2c0.151=sm00.5\000ess0.001degt0.02sec 78


Itshouldbenotedthatcaremustbetakenwhennumericallydeningthesingulardirectionsinces=0whenthesystemhasafullrankJacobian.Becausetherankisnumericallydetermined,atoleranceshouldbesetonthesingularitymeasuretodeterminewhatisconsideredfullrank.Fortheresultspresentedinthispaper,rankdeciencyfortheHSLwasdenedasmm0whereforthissimulationm0=0.5.ThesimulationsterminatewhenthesteadystateerroressdenedinEq.( 5 )isachieved. ess=min[2sin)]TJ /F2 7.97 Tf 6.58 0 Td[(1(jjejj),2)]TJ /F4 11.955 Tf 11.96 0 Td[(2sin)]TJ /F2 7.97 Tf 6.58 0 Td[(1(jjejj)](5) ThemagnitudeofessgiveninTable 5-1 isbasedoffreference[ 38 ].5.3.1Case1:AtZeroMomentumConguration=[0000]Tdeg Therstsetofsimulationshasinitialgimbalanglesat=[0000]Tdegwhichrepresentsascenariostartingfarawayfromsingularities.Figure 5-1 showsthiscongurationwhichisalsoatypicalstartupcongurationforafour-SGCMGpyramidarrangement. Figure5-1. Zero-momentumcongurationofafour-SGCMGpyramidarrangement 79

PAGE 80 TheparametersfortheLGsimulationare:0=a=b=1=0and2=0=1.Figures 5-2 AandBshowthattheLGmethodwasabletoperformthemaneuvertothegivenerrortoleranceesswithoutinducingtorqueerror.TheabsenceoftorqueerrorinFigure 5-2 BisduetothezerovalueofsingularitymetricinFigure 5-2 C(i.e.,LGisanexactmapping).ThenullmotionshowninFigure 5-3 BissmallbutsignicantwhencomparedtothetotaloutputgimbalratesinFigure 5-2 A.ThisisaconsequenceofthesingularitymetricinFigure 5-3 D.Figures 5-3 CandDshowthatthemaneuverwascompletedwithoutsingularityencounter. AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-2. SimulationresultsforLGwith0=a=b=1=0and2=0=1atzeromomentum 80


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-3. SimulationresultsforLGwith0=a=b=1=0and2=0=1atzeromomentum(contd.) 81

PAGE 82 TheparametersfortheSDAsimulationare:0=0.01,0=a=b=2=0,and1=1.ThismethodshowssimilarresultsinthetransientresponseoftheerrorstatesinFigure 5-4 AtothatforLGinFigure 5-2 AwiththeexceptionofnonzerotorqueerrorseeninFigure 5-4 B.Also,thismethodhadaslowerrateofconvergencetothesteady-stateerroressthanLGasevidentfromthetimeinsimulationinFigure 5-2 A.Thisisduetothesmallnonzerovalueofthesingularitymetric,showninFigure 5-4 C.Figure 5-5 BshowsazeronullmotioncontributiontothegimbalratesinFigure 5-5 AforSDA.Figures 5-5 CandDandFigures 5-3 CandDarealmostequivalentbecausethesystemstartedfarawayfromsingularity. 82


AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-4. SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1atzeromomentum 83


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-5. SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1atzeromomentum(contd.) 84

PAGE 85 TheparametersfortheHSLsimulationare:0=0.01,0=2,a=1,b=3,and1=2=1.InFigure 5-6 A,theHSLmethodshowssimilarresultstothatoftheSDAshowninFigure 5-4 A,withexceptiontothefasterrateofconvergenceofthetransienterrorresponse.However,thetorqueerrorinFigure 5-6 BaddedintothesystemissmallerthanthatofSDAinFigures 5-4 BandnullmotioninFigure 5-7 BissmallerthanthatoftheLGmethodin 5-2 B.ThisisduetothenonzerovalueforbothsingularitymetricsandinFigures 5-6 CandD.Singularitywasnotencounteredinthissimulationsasisshownbyavaluem>0.5inFigure 5-7 Candazerovalueofdet(Q)in 5-7 D. ForCaseIatzero-momentum,Table 5-2 ,comparestheroot-meansquared(RMS)gimbalrates(deg/s)andtrackingperformanceintermsofRMStorqueerror(Nm)forLG,SDA,andHSL.Inthistableitisshownthatallthreemethodshaveapproximatelythesameperformancewhichisexpectedforafour-SGCMGpyramidarrangementatzero-momentum,farfromsingularity.Thesteady-stateerrorforLGoranyoftheothermethodsisnonzeroasaconsequenceofthecontroller'sperformanceiscapturedhere. Table5-2. PerformanceComparisonsforCaseI:ZeroMomentum SteeringAlgorithm_RMSeRMS LG5.73662.2437e-06SDA5.72793.5902HSL(m0=0.5)5.73172.2573 5.3.2Case2:NearEllipticExternalSingularity=[105105105105]Tdeg Thesecondsetofsimulationsstartsatinitialgimbalangles=[105105105105]Tdeg,whichrepresentsascenarionearanellipticexternalsingularityat(i.e.,15degforeachSGCMGawayfromtheexternalsingularity=[90909090]Tdeg). 85


AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-6. SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1atzeromomentum 86


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-7. SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1atzeromomentum(contd.) 87

PAGE 88 TheparametersfortheLGsimulationare:0=a=b=1=0and2=0=1.TheplotsinFigure 5-8 AshowthattheLGmethodappearstohavesuccessfullyperformedthemaneuverasshowninFigures 5-8 AandB.However,thisismisleadingsincenon-implementablegimbalratesandaccelerationsarerequiredtodosoasshowninFigure 5-9 A.Thesingularitymetrics=0asexpectedforthismethodand=1atthesingularityencounter.Eventhough=1atsingularity,nullmotionattheexacttimeofsingularityencounteriszeroasshowninFigure 5-9 becausethegradientvectordforLGiszeroatellipticsingularities(i.e.,nogradientvectorexiststhatisinthedirectionawayfromsingularity).Also,thesingularity,veriedtobeellipticfromthepositivevalueofdet(Q)inFigure 5-9 D,wasescapedimmediatelywiththehelpofthenon-implementablegimbalratesandaccelerations,shownbyFigure 5-9 C. 88


AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-8. SimulationresultsforLGwith0=a=b=1=0and2=0=1nearellipticsingularities 89


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-9. SimulationresultsforLGwith0=a=b=1=0and2=0=1nearellipticsingularities(contd.) 90

PAGE 91 TheparametersfortheSDAsimulationare:0=0.01,0=a=b=2=0,and1=1.ThetransientresponseoftheerrorfortheSDAmethodshowninFigure 5-10 AiscomparabletothatoftheLGmethodinFigure 5-8 A,butwithimplementablegimbalratesandaccelerationsasshowninFigure 5-11 A.TheSDAmethodescapestheellipticexternalsingularityasshowninFigure 5-11 CattheexpenseofsignicanttorqueerrorshowninFigure 5-10 B.ThetorqueerrorscaledbythesingularitymetricshowninFigure 5-10 CdecreasesawayfromsingularityasshowninFigure 5-11 C.AsexpectedforSDA,thesingularitymetricinFigure 5-10 CiszeroresultinginzeronullmotionasshowninFigure 5-11 B.IncontrasttotheLGmethod,forSDA,thesystemlingersinsingularityforaround15secondsbeforeescapingasshowninFigure 5-11 C.Ellipticsingularityforthissimulationisveriedbythepositivevalueofdet(Q)inFigure 5-11 D. 91


AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-10. SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1nearellipticsingularities 92


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-11. SimulationresultsforSDAwith0=0.01,0=a=b=2=0,and1=1nearellipticsingularities(contd.) 93

PAGE 94 TheparametersfortheHSLsimulationare:0=0.01,0=2,a=1,b=3,and1=2=1.TheresultsforHSLshowninFigures 5-12 and 5-13 arealmostidenticaltothecorrespondingresultsofSDAforthissimulation.TheonlydifferencebetweenHSLandSDAsimulatedresults,liesinthenonzerosingularitymetricsandinFigures 5-12 CandD.DuetothechoiceoftheHSLparametersa,b,1,2,,0,0,thethresholdforsingularitym0.5,andQ0,theHSLactsastheSDAatanellipticsingularity AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-12. SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearellipticsingularities 94


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-13. SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearellipticsingularities(contd.) ForCaseIInearanellipticsingularity,Table 5-3 ,comparestheRMSgimbalrates(deg/s)andtrackingperformanceintermsofRMStorqueerror(Nm)forLG,SDA,andHSL.InthistabletheLGmethodissaidtohaveaninniteRMSgimbalratetopointoutthatitfailedforellipticsingularity.Also,itisshownthatSDAandHSLweresuccessfulincompletingthemaneuverwhileescapingsingularity.BothSDAandHSLhadapproximatelythesameperformanceforellipticsingularitywiththeexceptionofslightlybettertrackingperformanceforHSL. 95


Table5-3. PerformanceComparisonsforCaseII:EllipticSingularity SteeringAlgorithm_RMSeRMS LG17.7159e-06SDA8.256429.8989HSL(m0=0.5)8.136626.6946 5.3.3Case3:NearHyperbolicInternalSingularities=[15105195)]TJ /F4 11.955 Tf 9.3 0 Td[(75]Tdeg Thenalsetofsimulationsstartsatinitialgimbalangles=[15105195)]TJ /F4 11.955 Tf 9.3 0 Td[(75]Tdegwhichrepresentsascenarionearanhyperbolicsingularityat(i.e.,adistance15degfromeachCMGawayfromthesingularityat=[090180)]TJ /F4 11.955 Tf 9.3 0 Td[(90]Tdeg). TheparametersfortheLGsimulationare:0=a=b=1=0and2=0=1.ThetransientresponseoftheerrorfortheLGmethodinFigure 5-14 Aisidenticaltothatfortheothertwocases.ThisisbecausetheLGmethodisanexactmappingevidentfrom=0inFigure 5-14 CandhasnotorqueerrorassociatedwithitintheoryasshowninFigure 5-14 B.ThenullmotioninFigure 5-15 BmakesupalmosttheentirecontributionofthegimbalratesinFigure 5-15 AduetothenonzerovalueofinFigure 5-14 D.TheLGmethodbyitselfisabletoavoidthehyperbolicsingularityswiftlyandremainawayasshowninFigure 5-14 CandD. 96


AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-14. SimulationresultsforLGwith0=a=b=1=0and2=0=1nearhyperbolicsingularities 97


ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-15. SimulationresultsforLGwith0=a=b=1=0and2=0=1nearhyperbolicsingularities(contd.) 98

PAGE 99 TheparametersfortheSDAsimulationare:0=a=b=2=0,1=1and0=0.01.ThetransientresponseoftheerrorfortheSDAmethodinFigure 5-16 Aisdifferentintherateofconvergencetoess,butonthesameorderofmagnitudetothatfortheLGmethod.However,thegimbalratesforSDAshowninFigure 5-17 BareanorderofmagnitudesmallerthanthatfortheLGmethod.ThismethodescapesthehyperbolicsingularitysuccessfullywithtorqueerrorasshowninFigures 5-16 BasaconsequenceofthenonzerovalueofinFigure 5-16 C.Thesingularitymetricin 5-16 DiszerobecauseSDAdoesnotusenullmotion.AddedtorqueerrorforsingularityescapeversusnullmotionforsingularityavoidanceisthetradeoffbetweenSDAandLG.ThesingularityinthissimulationisveriedtobehyperbolicfromthenegativeresultshowninFigure 5-17 D.Also,theSDAmethoddidnotescapebywhatisconsideredsingularityinFigure 5-17 Cbythethresholdm0.5.However,thisdidnotaffectthedecayingoftheerrorstransientresponse.Thisisduetothefactthatthetorqueerrorisscaledbytheneededoutputtorquebeingmappedandtherefore,isnotseentohaveasignicanteffecttowardstheendofthemaneuver. 99

PAGE 100

AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-16. SimulationresultsforSDAwith0=0.01,0=0,a=0,b=0,and=1nearhyperbolicsingularities 100

PAGE 101

ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-17. SimulationresultsforSDAwith0=0.01,0=0,a=0,b=0,and=1nearhyperbolicsingularities(contd.) 101

PAGE 102 TheparametersfortheHSLsimulationare:0=0.01,0=2,a=1,b=3,and1=2=1.ThetransientresponseoftheerrorinFigure 5-18 AisalmostidenticaltotheLGmethodforthiscaseandhasafasterrateofconvergencetoessthanSDA.ThisisattributedtothenonzerovaluesofthesingularitymetricsandinFigures 5-18 CandDwhichprovideanorderofmagnitudelessnullmotionforsingularityavoidancethanLGandordersofmagnitudelesstorqueerrorthanSDAforthiscaseshowninFigure 5-19 Band 5-18 AwhenavoidingthehyperbolicsingularityveriedinFigures 5-19 CandD.UnlikeSDA,HSLescapedandthenavoidedthesingularitywhichisduetotheadditionofnullmotionforthismethod(seeFigures 5-16 Cand 5-17 C).Therefore,HSLreliesmoreonnullmotionforsingularityavoidanceratherthansoleytryingtopassthroughthehyperbolicsingularitiesasSDA,SR,andGSRdo.PrecisioninattitudetrackingwiththethreatofhyperbolicsingularitieswhilestillbeingabletoescapeellipticsingularitiesisthestrengthofHSL. 102

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AQuaternionerrorvectorelements BTorqueError CAlpha DBeta Figure5-18. SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearhyperbolicsingularities ForCaseIIInearahyperbolicsingularity,Table 5-3 ,comparestheRMSgimbalrates(deg/s)andtrackingperformanceintermsofRMStorqueerror(Nm)forLG,SDA,andHSL.InthistabletheLGmethodhasthelargestRMSgimbalrateamongthethreemethods,whichisneededforsingularityavoidance.Also,LGperformedthemethodwiththebesttrackingperformanceamongthethreemethodswhichisanexpectedresultforanexactmethod.TheHSLhadbettertrackingperformanceintermsofRMStorqueerrorthanSDAasaconsequenceofthelargergimbalratesneededfornullmotionsingularityavoidance.ThisisanexpectedstrengthofHSLathyperbolicsingularity. 103

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ACMGgimbalrates BNullmotion CSingularitymeasure Ddet(Q) Figure5-19. SimulationresultsforHSLwith0=0.01,0=2,a=1,b=3,and1=2=1nearhyperbolicsingularities(contd.) Table5-4. PerformanceComparisonsforCaseIII:HyperbolicSingularity SteeringAlgorithm_RMSeRMS LG10.39051.4742e-05SDA6.361114.4937HSL(m0=0.5)9.93304.7925 5.4HybridSteeringLogicSummary TheHSLwasfoundnumericallytopreserveattitudetrackingprecisioninthepresenceofhyperbolicsingularities,actcomparablytoSDAinthepresenceofellipticsingularities,andperformbetterthanSDAawayfromsingularity.Theperformanceofthisalgorithmisattributedtothenewsingularitymetrics,whichallowsmooth 104

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transitionbetweensingularityavoidanceusingLGandsingularityescapeusingSDA.Byreducingthetimeswheretorqueerrorisinducedforsingularityescape,thismethodprovidesimprovedattitudetrackingperformance.AnalyticandsimulatedresultsshowthatHSLhasmanybenetsoverthetwoothermethodsforsingularityavoidanceandescape.Thesebenetsare:itcanbeimplementedreal-time;althoughSVDmaybecomputationallyintensive,itremovestheneedforaninverseandprovidesalltheinformationneededforHSL;numericallyrobustalgorithmsexistforSVD;HSLinduceslesstorqueerrorthanSDAbyitself;andnally,theHSLprovidesanonsingularexpressionthatcanstartatsingularity.TheHSLisnotsuccessfulinavoidinggimbal-lockbecausenullmotionisnonexistentatellipticsingularitiesandSDAfailsatgimbal-lock(seeChapter 4 ). 105

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CHAPTER6CONTROLALGORITHM-ORTHOGONALTORQUECOMPENSATION6.1AttitudeControllerwithOTC TraditionallythecontrollawandsteeringalgorithmareseparatedforattitudecontrolsystemsusingSGCMGsasshowninFigure 6-1 .Thisisdonetofacilitateunderstandingoftheattitudecontrolsystemandactuatordynamicsseparately.However,consideringthesteeringalgorithmseparatefromthecontrollawmayreducethepossibilityofanincreaseintheperformanceinthesystem. Figure6-1. Satelliteattitudecontrolsystemblockdiagram Manysteeringlogicsbythemselvesareincapableofavoidinggimballock.Gimbal-lockoccurswhentherequiredtorqueforanattitudemaneuverisalongthesingulardirection.Thisproducesalocalminimumconditionwherethegimbalratesolutioniszerowhiletherequiredtorqueisstillnotmet.Open-loopmethodsthatprovideagimbaltrajectoryfreeofthisconditionexist;examplesofsuchmethodsareforwardpropagationfrompreferredgimbalangles,globalsteering,andoptimalcontrol[ 65 66 74 83 ].Thesemethodsaretimeconsumingandcannotguaranteeasolutionexistsfortheconstraintsprovided. Real-timesolutionstogimbal-lockavoidanceexistsuchastheGSRinversewhichusesoff-diagonaldithercomponentsinitsperturbationmatrixtoescapegimbal-lock(seeChapter 4 ).Thereisnoformalproofthatthesemethodswillalwaysbesuccessfulinavoidinggimbal-lock.Throughtheuseofnonlinearcontrol,anorthogonaltorque 106

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compensation(OTC)methodologycanbeaugmentedwithasuitablesteeringandcontrolalgorithmtoalsoavoidorescapegimbal-lock.Throughthisnonlinearcontrolframework,stabilitycanbeprovenandthesteeringalgorithmcanbechosenseparatelyincontrasttoGSRwhichreliesentirelyonhandlinggimbal-lockavoidance/escapethroughthesteeringalgorithm. Open-loopmethodssuchasoptimalcontrolforgimbal-lockescapeoravoidancemaynotndafeasiblesolutionorasolutionatalldependingonhowthecostfunctionandconstraintsareformulated. Itispossiblethatcombinationofanoptimalcontrolmaneuverwithapseudo-inversemethod(e.g.,SRinverse)willdrivethesystemtowardthevicinityofsingularityasthemaneuveriscompleted.Thismayoccursincetherequiredgimbalratesarenotonlyscaledfromthedistancetosingularity,butalsobytheneededoutputtorquefromtheSGCMGsystem.Asthenextrest-to-restmaneuverisneededthetorquemayberequiredaboutthesingulardirection.Whenthisoccurs,themaneuvercouldcausethelocalminimumpreviouslydiscussed.6.2LyapunovStabilityAnalysis Forthecasesconsidered,OTCwillbeamodicationtothequaternionregulatorcontrollogicfromreference[ 54 ]showninEq.( 5 ).Itshouldbenotedthatthismodicationcould,intheory,workwithanycontrolalgorithmwhichinturncanbecombinedwithanysteeringalgorithmforSGCMGs.Therefore,itisnotrestrictedtoanysteeringalgorithmorthequaternionregulatorcontrollawiftheproperstabilityanalysisiscarriedout.Thequaternionregulatorcontrollogicassumesperfectinformationandhasthefollowingnominalform _h=Ke+C!+!H(6) whereK=2kJandC=cJarepositive-denitesymmetricgainmatricesbasedonthespacecraft'scentroidalinertiaJ,eisthevectorelementsofthequaternionerrorvector, 107

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!isthespacecraftangularvelocity,andHisthetotalspacecraftsystemcentroidalangularmomentumfromEq.( 2 ). RecallfromEq.( 4 )thattheJacobian'sleftsingularvectorsUisanorthonormalbasisfortheoutputtorque_h.Thisbasisiscomposedofaunitvectorinthedirectionofthesingulardirectionukwhenatsingularity,andtwounitvectorsorthogonaltothesingulardirection,u?andun(i.e.,evenwhenAisnonsingular,thebasisfromUstillexists).Utilizingthisbasisintheformationoftheoutputtorqueyields _h=uk+ u?+un6=_h(6) withcoefcients =_h Tuk =_h Tu?+ag(m)=_h Tun+bg(m)(6) Thequantityg(m)isaaugmentationtotheorthogonaltothesingulardirectioncomponentsoftorquethatisanexplicitfunctionofthesingularitymeasure.ItwillbereferencedastheOTCsingularityparameterandaandbareswitchingelementsdenedby a=8><>:1if_h Tu?0)]TJ /F4 11.955 Tf 9.3 0 Td[(1if_h Tu?<0 b=8><>:1if_h Tun0)]TJ /F4 11.955 Tf 9.3 0 Td[(1if_h Tun<0 SubstitutingEq.( 6 )intoEq.( 5 )andboundingyields, 108

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_V=)]TJ /F7 11.955 Tf 26.37 8.09 Td[(c 2k!TJ)]TJ /F2 7.97 Tf 6.59 0 Td[(1U266664100010001377775UTJ!+2=)]TJ /F3 11.955 Tf 9.29 0 Td[(!TR2!+2(6) where2=1+g(m)!TK)]TJ /F2 7.97 Tf 6.58 0 Td[(1(au?+bun)andR2=R1inEq.( 5 )withthesingularityparameterfromEq.( 4 )inplaceoffortheHSLinEq.( 5 ).SimilartotheLyapunovanalysisinChapter 5 forHSL,theerrorzisboundedwithaVolterraintegralexpressionas jjzjj2V(0) minexp()]TJ /F28 5.978 Tf 10.1 4.4 Td[(1 maxt)+1 minZtt0exp()]TJ /F28 5.978 Tf 10.1 4.4 Td[(1 max(t)]TJ /F11 7.97 Tf 6.59 0 Td[())2()d(6) SincetheSGCMGoutputtorquewillalwayshavecomponentsorthogonaltothesingulardirectionwhennearsingularity,itisassumedthatasystemusingOTCwillneverencountergimbal-lockuptoaspecicsizeofjjejjandfromEq.( 4 )whereg(m)=jjejj(6) Therefore,situationsofsingularityotherthanthosewithgimbal-lockareofconcern.Whenatsingularity,theexpressionfortheerroris jjzjj2Vs+Ztts2()d(6) becausetheerrorzisbasedoffthetransienttermoftheLyapunovequationVsfromEq.( 7 )andthedynamictermcontainingtheeffectofthetorqueerroraddedforgimbal-lockescapeRtts2()d.RecallfromChapter 5 ,whileusingHSL,thatwhensingularityoccurswiththeexceptionofgimbal-lock,theremaybeaperiodofinstabilityanditisassumedthatthemaneuverdoesnotendatsingularity.Withthisinmind,thesteady-stateerroroftheofasystemawayfromsingularityusingSDAcombinedwithOTCisboundedas 109

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jjz(1)jj2)]TJ /F3 11.955 Tf 30.83 8.08 Td[(max min12(1)(6) throughtheuseofL'Hospital'sruleasinEq.( 5 ). TheresultofOTCisUUBforsufcientlylargethechoicesofcratherthan.Thisistruebecauseshortperiodsofinstabilitymayarise,butthenegativesemi-denitetermofEq.( 6 )becomesnegativedeniteawayfromsingularityandwillbecomedominateforsufcientlylargevaluesofc.Withthechoiceofg(m)inEq.( 6 ),wheneverthereisanattitudeerrorandthesystemisinproximitytoasingularity,therewillbetorqueerroraddedorthogonaltothesingulardirection,andthusgimbal-lockwillbeescaped.6.3NumericalSimulation ForthesteeringalgorithmsofSDA,GSR,andSDAwithOTC(SDA/OTC)augmentedtotheattitudecontroller,twocasesweresimulatedforafour-SGCMGpyramidalclusterat=54.74degandthemodelparametersinTable 6-1 :(1)az-axismaneuverstartingatinitiallyatthezeromomentumcongurationfromChapter 5 (i.e.,=[0000]Tdeg)and(2)az-axismaneuverstartingatgimbal-lockconguration(i.e.,=[90909090]Tdeg).Bothcasesusethesamecontrolgainsappliedtoapyramidalarrangementoffour-SGCMGs.Also,thesimulationwaspropagatedwithadiscretefourth-orderRunga-Kuttaatatime-stepoft=0.02sec. Table6-1. ModelParametersVariableValueUnits J133kgm254.74dege0[000.3]T\000!0[000]Tdeg=sh01Nmsk21=s2c101=si0.1rand(1)\00000.1\000t0.02sec 110

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6.3.1CaseI:0=[0000]Tdeg Foraz-axismaneuveroriginatingfromaninitialzero-momentumconguration,Figures 6-2 showthatSDA,GSR,andtheSDA/OTCappearidentical.Thisiswhatisexpectedforamaneuverfarfromsingularity.TheresultsforthetorqueFigure 6-3 conrmthisbecausethetransientresponseforthiscase(i.e.,awayfromsingularity)isshort.ThetransientresponseoftheoutputtorqueshowninFigure 6-3 DofSDA/OTChassignicantjitterbutwithasmallmagnitude.ThisjitterhasnegligibleeffectonthegimbalratesshowninFigure 6-2 Dwhichisduetothemappingoftheoutputtorqueontothegimbalrates.ThedifferenceinquaternionerrorandsingularitymeasureinFigures 6-4 and 6-5 aresmall.ThisshouldnotbesurprisingsinceforSDA,GSR,andSDA/OTC,thecontributionsoftorqueerroraredesignedtobesignicantonlywhenthesystemisclosetoasingularity. TheOTCsingularityparametershowninFig. 6-6 ,whileinitiallynonzeroforthiscase,convergestozerorapidly.ThefactthatthisparameterisnonzeroinitiallyandthereisnosignicantdifferencesinthequaternionerrorresponsesasshowninFigure 6-7 ,mightsuggestthatthetorqueerrorfromtheSDAmethoditselfwasdominant.Inaddition,itshouldbenotedthatthedifferenceinquaternionerrorresponseswhilesmall(10)]TJ /F2 7.97 Tf 6.59 0 Td[(8),isnotontheorderofmachineprecision(10)]TJ /F2 7.97 Tf 6.58 0 Td[(16)or(10)]TJ /F2 7.97 Tf 6.58 0 Td[(32). 111

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ASDA BSDA(transientresponse) CGSR DGSR(transientresponse) ESDA/OTC FSDA/OTC(transientresponse) Figure6-2. Gimbalrates 112

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ASDA BSDA(transientresponse) CGSR DGSR(transientresponse) ESDA/OTC FSDA/OTC(transientresponse) Figure6-3. Outputtorque 113

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ASDA BGSR CSDA/OTC Figure6-4. Vectorelementsoftheerrorquaternion 114

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ASDA BGSR CSDA/OTC Figure6-5. Singularitymeasure 115

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Figure6-6. Singularityparameter(OTC) AeGSR)]TJ /F29 9.963 Tf 9.96 0 Td[(eSDA BeSDA=OTC)]TJ /F29 9.963 Tf 9.96 0 Td[(eSDA Figure6-7. Quaternionerrordifference:(A)eGSR)]TJ /F7 11.955 Tf 11.96 0 Td[(eSDA(B)eSDA=OTC)]TJ /F7 11.955 Tf 11.96 0 Td[(eSDA 6.3.2CaseIIa:0=[90909090]Tdeg Forthiscase,thegimbalsareinitiallyorientedsuchthatthesystemisinagimbal-lockconguration.Figure 6-8 AshowsthatthegimbalratesoftheSDAmethodareunchangedthroughoutthesimulationsincethesystemstartsinagimbal-lockcongurationandSDAcannotgeneratethenecessarycommandstoescape.ThegimbalratesforGSRandSDA/OTC(Fig. 6-8 BandD),however,arenonzerobecausetheadditionofthetorqueerrorhasprovidedthesystemwiththeabilitytoescapegimbal-lock.Inaddition,thecontrollerapproachestheoriginalquaternionregulatorcontrollerasthesystemmovesawayfromsingularity. 116

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ThetransientresponseofthegimbalratesfortheGSRandSDA/OTCinFigure 6-8 CandEarebothoscillatorywithGSRhavingthehigheramplitudeandduration.ThisisattributedtothefactthatunlikeOTC,thefunctionsaddingtorqueerrorinGSRforgimbal-lockescapearenotclearlyvisualized(i.e.,dependonthecombinationofsinusoidswithpossibledifferentfrequenciesandphasesfordither)whenmappedtothegimbalrates. Figures 6-9 and 6-10 show,respectively,therequiredtorqueandattitudeerror.Inbothcases,theresultsshowasimilartrendasthegimbalratesforGSR,andSDA/OTCapproachzero;theGSRandSDA/OTCwereabletogeneratethetorquerequiredtodrivetheattitudeerrortozero. AnexaminationofthesingularitymeasuresshowninFigure 6-11 reafrmstheresponsesshowninFigures 6-8 through 6-10 wheretheSDAremainsatsingularityunlikeGSRandtheSDA/OTCwhichescapesingularitybuttransitionbacktoitasthemaneuveriscompleted.Thistransitionbacktosingularityiscommonforallpseudo-inversesteeringalgorithms,whichworkbyapproachingasingularcongurationandthenmakingarapidtransitionforescape[ 47 ].Recallpreviouslyfrom 6.1 ,thatitwasstatedthatitispossibletoendinthevicinityofasingularitywhenthemaneuverwascompleted;thisisanexampleofsuchacaseshowninFigure 6-10 and 6-11 ThemeasureofhowfarthesystemofSGCMGsisfromgimballockcanbefoundasthenormjjAT_hjj)166(!0.BoththeGSRandtheSDA/OTCweresuccessfulinescapinggimbal-lockasshowninFigure 6-12 .Itshouldbenotedthatbecausethismeasureisafunctionof_h,itgoestozeroasthemaneuveriscompleted. TheOTCsingularityparameterisshowninFigure 6-13 .Ithasanonzeroinitialvalueandconvergesrapidlytozerowhichmakesiteffectiveforhelpinginsingularityescape. 117

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ASDA BGSR CGSR(transientresponse) DSDA/OTC ESDA/OTC(transientresponse) Figure6-8. Gimbalrates 118

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ASDA BGSR CGSR(transientresponse) DSDA/OTC ESDA/OTC(transientresponse) Figure6-9. Outputtorque 119

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ASDA BGSR CSDA/OTC Figure6-10. Vectorelementsoftheerrorquaternion 120

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ASDA BGSR CSDA/OTC Figure6-11. Singularitymeasure 121

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ASDA BGSR CSDA/OTC Figure6-12. Gimbal-lockmeasure 122

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Figure6-13. Singularityparameter(OTC) 6.3.3CaseIIb(HSL/OTC):0=[90909090]Tdeg Recall,from 6.1 thatOTCcanbeusedincombinationwithanysteeringalgorithmforgimbal-lockavoidance/escape.ThiscaseveriesthroughsimulationthatthisisindeedtruebycomparingHSL/OTCtoGSRstartingatgimbal-lock(0=[90909090]Tdeg).TheHSLparametersareshowninTable 6-2 Table6-2. HybridSteeringLogicParametersVariableValue 00.01021121a1b3m00.5 Withtheexceptionoftheinitialtransient,thegimbalratesforGSRandthoseofHSL/OTCinFigure 6-14 ,areapproximatelythesamemagnitude. 123

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AGSR BGSR(transientresponse) CHSL/OTC DHSL/OTC(transientresponse) Figure6-14. Gimbalrates ThetransientresponseofthegimbalratesfortheGSRinFigure 6-14 BishighlyoscillatoryandnotassmoothasthatforHSL/OTC(comparewithFigure 6-14 D).Thisisduetooscillatorybehaviorofthedither-usedforgimbal-lockescapethatmaybeofanydurationdependingonthefrequenciesandphasesoftheoff-diagonalcomponentsoftheGSRperturbationmatrixE.TheHSLmethodactsasaSDAmethod,butwhencombinedwithOTCwillavoid/escapeasingularityatthespeedoftheparameterschosenforinEq.( 6 )inthewhichthedurationwillbeunderstoodforallsingularitiesandtheircombinationstothenormofquaternionerror.Figures 6-15 and 6-16 show,respectively,therequiredtorqueandattitudeerror.Inbothcases,theresultsshowthat 124

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bothmethods(GSRandHSL/OTC)weresuccessfulinmeetingtherequiredtorqueandcompletingtheattitudemaneuver. AGSR BGSR(transientresponse) CHSL/OTC DHSL/OTC(transientresponse) Figure6-15. Outputtorque ThesingularityparametersforbothmethodsinFigure 6-17 ,escapesingularityalthoughtheytransitionbacktoitasthemaneuveriscompleted.Recall,itwasmentionedpreviouslythatamaneuvercanbecompleted(i.e.,e)166(!0)whilethegimbalanglessettleintoasingularconguration;Figure 6-11 and 6-17 showthistrend(comparewithFigure 6-11 ). Figure 6-18 showsinstantaneousescapefromgimbal-lockforbothmethods. 125

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AGSR BHSL/OTC Figure6-16. Vectorelementsoftheerrorquaternion AGSR BHSL/OTC Figure6-17. Singularitymeasure TheOTCsingularityparametershowninFigure 6-19 hasaninitialnonzerovalueandconvergesrapidlytozerosimilarlytoFigure 6-13 ,whichmakesiteffectiveforhelpinginsingularityescape. TheTable 6-3 comparestheroot-meansquared(RMS)gimbalrates(rad/s),trackingperformanceintermsofRMStorqueerror(Nm),andpointingperformanceintermsofthenormofthesteady-stateerrorquaternionforGSR,SDA/OTC,andHSL/OTC.FromTable 6-3 itcanbeshownthatthechoiceofthesingularitythresholdm0hasaneffectonthetrackingandpointingperformanceoftheHSLmethodcombined 126

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AGSR BHSL/OTC Figure6-18. Gimbal-lockmeasure Figure6-19. Singularityparameter(OTC) withOTC.Infact,whenthisvalueism0=0.5forthismodelandwiththesetofcontrolgainsthatdifferfromthemodelinChapter 5 ,thetrackingandpointingperformanceofHSL/OTCisactuallyworse.ThisisexpectedasshownbytheLyapunovanalysisin 6.2 wherethesteady-stateerrorofSDAisdependentonthetorqueerroraddedintothesystem;andalargerthresholdvalueofm0willincreasethesteady-stateerror.6.4OrthogonalTorqueCompensationSummary Orthogonaltorquecompensation(OTC)methodologywasdevelopedtoensureescapefromallsingularities,particularlyscenariosinvolvinggimbal-lockcongurations.Thecompensationmethodologycanbeincorporatedwithanyattitudecontroller/steering 127

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Table6-3. PerformanceComparisons SteeringAlgorithm_RMSeRMSjjessjj GSR7.790110.18640.0024SDA/OTC7.314610.40800.0020HSL/OTC(m0=0.5)7.16936.18190.0038HSL/OTC(m0=0.05)7.00413.04162.1474e-09HSL/OTC(m0=0.005)6.91662.98521.8843e-09 logiccombinationandwasshownthroughanalysistoensurestabilitywithsufcientlylargechoiceofthecontrollergainc.Sincethecompensatorwasdesignedtoworkwithanyattitudecontroller,thenitiscompatiblewithanysteeringalgorithms.ThiscouldproveverybenecialforsteeringalgorithmslikeHSLwhichreducetheamountoftorqueerrorathyperbolicsingularities(seeChapter 5 ).TheOTCwasalsodemonstratedthroughnumericalsimulationwhereitwasshowntobeeffectiveinescapinggimbal-lockwithnearzerosteady-stateattitudeerror.Thesesimulationswerebasedonafour-SGCMGpyramidalarrangementusinganquaternionregulatorcontrollercombinedwiththesteeringalgorithmsSDAandHSLandcomparedwithGSR. 128

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CHAPTER7SCALABILITYISSUESFORSGCMGS7.1ScalabilityProblemswithSGCMGHardware CurrentlyavailableCMGactuatorsareshowninFigure 7-1 withspecicationsfromTable C-1 inAppendix C donotmeetthepower,mass,andvolumerequirementsforsatellitessmallerthanthemicro-satclass.Currently,developmentofCMGhardwareunderwaywillmeetsomeoftheconstraintsforthesesmallerclassesofsatellites.NewsteeringalgorithmstocomplementthesenewlydevelopedCMGhardwareisnotbeingemphasizedandwillhaveamajoreffectonhowsystemsofminiatureCMGsperform.ThischapterhighlightstheeffectofscalingontheperformanceofminiatureCMGs. Figure7-1. Off-the-shelfCMGs 129

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7.2EffectofIgwonTorqueError Thegimbalaccelerationsarekinematicallydependentonthechoiceofthegimbalratesandasaconsequence,onlyoneofthemcanbeusedasacontrolvariable.Therefore,thegimbalratesareconsideredasmeasurablequantitiesandthegimbalaccelerationsarethecontrol.Thesolutiontothegimbalaccelerationsasacontrolisdenedas =BT(BBT))]TJ /F2 7.97 Tf 6.59 0 Td[(1[T)]TJ /F8 11.955 Tf 11.96 0 Td[(A2_](7) whereA=A1+A2fromEqs.( 2 )and( A )andT=_h+A1_equivalenttoEq.( 2 ).Thissolutionisconsideredanexactsolutionbutforsomecasesmaybehighlyoscillatoryand/orunstableforthegimbalratesandaccelerations.ALyapunovanalysisispresentedbelow. ItwasstatedpreviouslythatthedirectsolutioninEq.( 7 )maybeunstable.ToprovethiswestartwiththegivencandidateLyapunovfunction V1=1 2!TK)]TJ /F2 7.97 Tf 6.58 0 Td[(1Jc!+eTe+(1)]TJ /F7 11.955 Tf 11.95 0 Td[(e4)2(7) Takingthetimederivative,yields _V1=!TK)]TJ /F2 7.97 Tf 6.59 0 Td[(1[)]TJ /F8 11.955 Tf 9.3 0 Td[(T)]TJ /F10 11.955 Tf 11.95 0 Td[(!H]+!Te(7) Forthesystemtobegloballyasymptoticallystable(i.e.,_h,e,and!)166(!0ast)166(!1) T=A_+B=Ke+C!)]TJ /F10 11.955 Tf 11.96 0 Td[(!H(7) Next,asecondcandidateLyapunovfunctionisrequiredtoanalyzethebehaviorofthegimbalratesandaccelerationsastimeapproachesinnity. V2=1 2_T_(7) 130

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TakingthetimederivativeutilizingEq.( 7 ),weobtain _V2=_T=_TBT(BBT))]TJ /F2 7.97 Tf 6.59 0 Td[(1[T)]TJ /F8 11.955 Tf 11.96 0 Td[(A_](7) FromthepreviousLyapunovanalysiswhere_h,e,and!)166(!0ast)166(!1itcanbeassumedthat _V2=)]TJ /F8 11.955 Tf 11.59 2.66 Td[(_TBT(BBT))]TJ /F2 7.97 Tf 6.59 0 Td[(1A_=(h0 Igw)_TS_(7) whereB=Igw^B,A=h0^A,andS=)]TJ /F8 11.955 Tf 10.39 2.66 Td[(^BT(^B^BT))]TJ /F2 7.97 Tf 6.59 0 Td[(1^A.MatrixSissemi-indeniteandthereforethegimbalratescanbeunstable.Furthermore,aLyapunovanalysisofV1+V2showsthattheratio(h0 Igw)playsakeyroleinthestabilityofthewholesystem. Next,weconsidertheuseoftheSRinversewherethegimbalratesarefoundfrom _SR=1 h0ASR(T)]TJ /F7 11.955 Tf 11.96 0 Td[(Igw^B)(7) with ASR=^AT(^A^AT+1))]TJ /F2 7.97 Tf 6.59 0 Td[(1(7) whereisthesingularityparameterdenedinEq.( 4 ). AssumingtheSRinverseisusedtoapplythegimbalratesasacontrolvariablewendthatthetorqueerrorisexpressedas e=Tact)]TJ /F8 11.955 Tf 11.95 0 Td[(T=h0^A_SR+Igw^B)]TJ /F8 11.955 Tf 11.96 0 Td[(T(7) Furthermore, ^AASR=[1+(^A^AT))]TJ /F2 7.97 Tf 6.58 0 Td[(1])]TJ /F2 7.97 Tf 6.58 0 Td[(1(7) Awayfromsingularity,aseriesexpansionofEq.( 7 )withonlythelineartermsyields 131

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^AASR1)]TJ /F3 11.955 Tf 11.95 0 Td[((^A^AT))]TJ /F2 7.97 Tf 6.58 0 Td[(1(7) Thisseriesexpansionisconvergentifawayfromsingularitybecausethetermj(^A^AT))]TJ /F2 7.97 Tf 6.58 0 Td[(1j<1.SubstitutingEq.( 7 )intothetorqueerror,Eq.( 7 )wehave eIgw(^A^AT))]TJ /F2 7.97 Tf 6.59 0 Td[(1^B)]TJ /F3 11.955 Tf 11.95 0 Td[((^A^AT))]TJ /F2 7.97 Tf 6.58 0 Td[(1T(7) Itcanbeseenthatthetorqueerrormaybeampliedbythemagnitudeofthegimbal-ywheelinertiaIgw.Furthermore,iforIgwisconsiderednegligiblethenthetorqueerrorisonlyaffectedbythesingularityparameter,thedistancefromsingularitywhichisrelatedtothedeterminantof(^A^AT))]TJ /F2 7.97 Tf 6.58 0 Td[(1,andT.ItshouldbenotedthatanincreaseinIgwisfollowedbyadecreasein^A^AT,butitseffectivenessinloweringthetorqueerrorrequiresalargeratioofIgw h0(i.e.,effectivewhenIgw h0>>1whichcouldbethoughtofasbeingasystemofRWs). Theeigen-axiscontrollogicfromEq.( 5 )isusedtodenethetorqueneededforagivenmaneuvertobemappedontothegimbalstates.TheSGCMGsystemproposedinthisanalysisassumesthatitisself-containedandthereforethemetricofthehostedalgorithmperformanceisindependentofthecontrollogicchosenaslongasitmeetstheconstraintsoftheSGCMGactuators.Therefore,nogeneralityislostforthechoiceofthecontrollogicintheanalysis.7.3NumericalSimulation ThecasescomparedherearetheSRInverseandalteredgimbal-accelerationcontrollawbasedonOhandVadali[ 84 ].Thelteredgimbalaccelerationcontrollawhasthefollowingform =K(_SR)]TJ /F8 11.955 Tf 14.25 2.66 Td[(_)+SR(7) 132

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whereKisthegainmatrixthatsizestheamountofgimbalaccelerationutilizedforcontroland_SRandSRarethegimbalratesandaccelerationsfromtheSRinverseandthetimederivativeofthatrate.Theeffectofthegimbal-ywheelinertiaisscaledinthesimulationbythegainKgw(i.e.,Igw=Kgw Igw)whereKgw=0signiesthattheirisnotorqueorangularmomentumcontributedfromthegimbaldynamics. SimulationsofthesetwosteeringalgorithmswerecomparedbyscalingIgwthroughtwodifferentvaluesofKgw:i)Kgw=0andii)Kgw=2.ThemodelparametersforthenominalsatelliteinertiaJandgimbal-ywheelinertiaIgwarebasedonafour-SGCMGpyramidalarrangementsizedfora1UCubeSat.Bothsimulationswereforamaneuverof180oaboutthez-axis.Theinitialgimbalanglesforallsimulationsare0=[)]TJ /F4 11.955 Tf 9.3 0 Td[(90)]TJ /F4 11.955 Tf 9.3 0 Td[(90)]TJ /F4 11.955 Tf 9.3 0 Td[(90)]TJ /F4 11.955 Tf 9.3 0 Td[(90]Tdegcorrespondingtoaellipticsaturationsingularityaboutthez-axis.Thissetofinitialgimbalanglesalongwiththerequiredmaneuverwillforcethesystemtoentergimbal-lock(i.e.,AT_h=0)andaccumulateasteadystateattitudeerror.Thissituationwaschosentotestthesystemtoitsperformancelimit.TheparametersthatwereusedforalloftheresultsareshowninTable 7-1 Table7-1. ModelParametersVariableValueUnits Js24533.8000533.8000895.63510)]TJ /F2 7.97 Tf 6.58 0 Td[(6kgm252dege0[0010]T\000!0[000]Tdeg=sh04.48610)]TJ /F2 7.97 Tf 6.59 0 Td[(4Nms Igw5.15410)]TJ /F2 7.97 Tf 6.59 0 Td[(6kgm2k101=s2c501=sK10I44\00000.5\00010\000_0[0000]Tdeg=s 133

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7.3.1CaseI:Kgw=0 Forthiscase,Figures 7-2 and 7-3 showthatthegimbalratesandaccelerationsarequitesimilarforboththeSRinverseandlteredaccelerationcontrollaw,exceptattheverybeginning.SincethelteredaccelerationcontrollawalsousesanSRinverseinitsformulation,itisreasonabletoassumethatthedifferencescanbeattributedtothelter(seeEq.( 7 )).ThetorqueerrorsinFigure 7-4 areinitiallylargerforthelteredaccelerationalgorithmascomparedtotheSRinversebyitself.Again,thisismostlikelyduetothelowerinitialgimbalratesandaccelerationsattributedtothelter.Also,thetorqueerrorforbothcaseshasasteadystateoffsetwherethesystemencountersgimbal-lock. 134

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ASRInverse BSRInverse(transientresponse) CFilteredgimbalacceleration DFilteredgimbalacceleration(transientre-sponse) Figure7-2. GimbalratesforKgw=0 135

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ASRInverse BSRInverse(transientresponse) CFilteredgimbalacceleration DFilteredgimbalacceleration(transientre-sponse) Figure7-3. GimbalaccelerationsforKgw=0 136

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ASRInverse BSRInverse(transientresponse) CFilteredgimbalacceleration DFilteredgimbalacceleration(transientre-sponse) Figure7-4. TorqueerrorforKgw=0 137

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ThesingularitymeasuremshowninFigure 7-5 isidenticalforbothmethodsintheseplots.Theonlydiscrepancybetweenthegimbalratesandaccelerationsofthetwomethodsforthiscaseswasattheverybeginningofthemaneuver.Therefore,thedifferencesinmwouldnotbeobviousintheseplots.ThevalueofmhereisshowntotransitawayfrombutreturntosingularityinFigure 7-5 forbothmethods.ThisiscommontosteeringalgorithmsoftheSRinversetype. ASRInverse BFilteredgimbalacceleration Figure7-5. SingularitymeasureforKgw=0 7.3.2CaseII:Kgw=2 TheinitialgimbalratesandaccelerationsarelessforthelteredaccelerationalgorithmthanforSRinverseitself.ThisistheshowninFigures 7-6 and 7-7 138

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ASRInverse BSRInverse(transientresponse) CFilteredgimbalacceleration DFilteredgimbalacceleration(transientre-sponse) Figure7-6. GimbalratesforKgw=2 AcloserlookatthetransientresponseforthegimbalratesandaccelerationsofthetwomethodsisshowninFigures 7-6 BandDand 7-7 BandD.ItisgatheredfromFigure 7-6 Band 7-7 Bthatthesteady-stateresponseforthegimbalratesandaccelerationsoftheSRandlteredaccelerationlawinverseisnonzero.Thesteady-stateresponseforthegimbalratesandaccelerationsofthelteredaccelerationlawalthoughnonzero,isconsiderablysmallerthanthatfortheSRinversewhichinturnpreventsthetorqueerrorfromdiverging. However,Figure 7-8 showsthetorqueerrorfortheSRinverseappearstodivergewhereasthetorqueerrorforthelteredaccelerationalgorithmremainsbounded. 139

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ASRInverse BSRInverse(transientresponse) CFilteredgimbalacceleration DFilteredgimbalacceleration(transientre-sponse) Figure7-7. GimbalaccelerationsforKgw=2 Furthermore,itappearsasthoughtheprecisionisimprovedwithlargerIgwforthelteredaccelerationalgorithm.ThismaybeduetothefactthatitrelieslessontheSRinversewhichwouldbethesourceofthetorqueerrorinthisexample. 140

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ASRInverse BFilteredgimbalacceleration Figure7-8. TorqueerrorforKgw=2 ThesingularitymeasuresshowninFigure 7-9 areidenticalexceptthatthesingularityisapproachedquickerforSRinverse.ItshouldbenotedthatasthevalueofIgwisincreased,thesingularitymeasureapproachessingularitylaterinthesimulationforthelteredsteeringalgorithm(i.e.,largergimbalratesleadtolargergimbalangleexcursionswhichinturn,makethesystemapproachsingularityquicker). ASRInverse BFilteredgimbalacceleration Figure7-9. SingularitymeasureforKgw=2 Thegimbal-ywheelassemblyinertiawillalsohaveadegradingeffectontheactualtorqueamplicationofSGCMGactuators.Thiswillbeexplainedinthenextsection. 141

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7.4EffectofIgwonTorqueTorqueAmplication ThetorqueamplicationofasingleSGCMGcanbedescribedbyitsoutputtorquedividedbytheinputtorqueas jjoutjj jjinjj=jjh_+Igwjj jj!h+Igwjj(7) FromEq.( 7 )itisseenthatasthegimbal-ywheelinertiaIgw)166(!1,theothertermsintheequationbecomelessdominantandthetorqueamplicationconvergestoone.ThisisundesirableforSGCMGsbecauseatthepointthatthetorqueamplicationconvergestoone,thesystemessentiallybecomesareaction-wheelsystemandthebenetsofusingSGCMGsarelost.Fortunately,asystemofSGCMGsofthisscaledoesnotexist. ThescalingofSGCMGsdoesreducethetorqueamplication.Toshowactuallyhowmuchthetorqueamplicationisdegradedbyscaling,thevalueoftorqueamplicationiscalculatedfortheIMPACSGCMGsinFigure 2-2 withtheparametersinTable 7-2 Table7-2. ModelParametersVariableValueUnits Js24533.8000533.8000895.63510)]TJ /F2 7.97 Tf 6.58 0 Td[(6kgm252degh04.48610)]TJ /F2 7.97 Tf 6.59 0 Td[(4NmsIgw5.15410)]TJ /F2 7.97 Tf 6.59 0 Td[(6kgm2!max3deg=s_max1rad=smax1rad=s2 Forthisexample, jjoutjj jjinjj=h0_max+Igwmax !maxh0+Igwmax43.4(7) Therefore,thereisasignicantvalueoftorqueamplicationevenwhenscalingasseenbytheresultinEq.( 7 ). 142

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7.5Summary SingularitiesfromsystemsutilizingSGCMGscannotbeeasilyscaledwhendescribingthealgorithmsfortheircontrol.JustastheperformanceonahardwarelevelforSGCMGswilleventuallyatlinewithscaling,sowilltheuseofcurrentsteeringalgorithmsforsingularityavoidance.Thischaptershowedthatsomecurrentsteeringalgorithmsmayhavedifcultywhenthegimbal-wheelassemblyinertiabecomessignicant.Inaddition,thischapteralsoshowedthatasaconsequenceofsignicantvaluesforgimbal-wheelassemblyinertia,theperformanceoftheSGCMGsystemisaffectedbythetorqueamplicationapproaching1asIgw)166(!1. 143

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CHAPTER8CONCLUSION Controlofspacecraftattitudewithsingle-gimbalcontrolmomentgyroscopes(SGCMGs)isdifcultandbecomesmoresowiththescalingoftheseactuatorstosmallsatellites.TheresearchpresentedinthismanuscriptbeganwithadiscussionofthedynamicmodelforsystemscontainingCMGsandtheirsingularities.ForSGCMGsystems,singularitieswereclassiedandtoolsweredevelopedtoquantifytheformofthesingularity.Thesetoolsprovidedinsightsintothesesingularities(i.e.,singularsurfaces)andwereusedtoquantifythemmathematically. ThesingularitiesassociatedwithSGCMGswerediscussedindetailandclassiedbythetoolsdeveloped.Throughthisdiscussion,itwasshownacompleteexplanationofSGCMGsingularitiesisabsentfromtheliterature.Forexample,itwasfoundthatthespecialcaseofwherethesingulardirectionsisalongagimbalaxis^canoccurforrooftoparrangementswhentherankoftheJacobianis2.ThiswasfurthershowntobeadegeneratecasewhichcouldleadtodegeneratehyperbolicsingularitiesthatwerepreviouslyneglectedintheliteratureforsystemsofSGCMGs.Usinglinearalgebra,itwasproventhatrooftoparrangementsarefreeofellipticinternalsingularitiesbutstillcontainedellipticexternalsingularities(i.e.,allellipticsingularitiesdonothavenullmotionarearethusimpassablebynullmotion)associatedwithangularmomentumsaturation.Furthermore,degeneratehyperbolicsingularitieswereshowntoexistforrooftoparrangements(i.e.,degeneratehyperbolicsingularitiescontainonlysingularnullsolutionsexistandarealsoimpassablebynullmotion). ItwasshownthatselectinganarrangementofSGCMGSthroughchoiceofadesirableangularmomentumenvelopeisdifcult.Thus,amethodofofineoptimizationwassuggestedinaveryconstrainedcasethatwillprovidethebestSGCMGarrangementintermsofEulerangles.However,thismethodisnotthatapplicableforrealspacecraft 144

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design,although,itsuggestedthatthecurrentcommoncongurationsdonotnecessarilyhavethebestperformance. Next,itwasshownthatlegacysteeringalgorithms,whichcanbecategorizedintothethreefamiliesofsingularityavoidance,singularityescape,andsingularityavoidanceandescape,didnotconsidertheformofinternalsingularity(i.e.,hyperbolicorelliptic).Thiswasshowntobeproblematicwhenpreciseattitudetrackingisrequiredbecausethesameamountoftorqueerrorwasusedforbothhyperbolicandellipticsingularities.AHybridSteeringLogic(HSL)wasdevelopedthattakesintoaccounttheformofsingularity.Thisalgorithmusesthisknowledgetoapplynullmotionfromalocalgradient(LG)methodforsingularityavoidancewhennearahyperbolicsingularityandtorqueerrorfromSingularDirectionAvoidance(SDA)whennearaellipticsingularities.Throughanalyticderivationsandnumericalsimulations,HSLwasshowntoperformbetter(i.e.,lowertorqueerrorathyperbolicsingularitiesthanpseudo-inversemethodsandtheabilitytoescapebothellipticandhyperbolicsingularitiesunlikeLGmethods)thanlegacymethodsforpreciseattitudetrackingwhenusingafour-SGCMGpyramidarrangementofSGCMGs.Also,HSLwasshowntohavecomputationofopsonthesameorderasmanylegacymethods. Gimbal-lockwasshowntobeaspecialcaseofsingularitywhentheoutputtorqueliesinthesingulardirectionoftheJacobian.WiththeexceptionoftheGeneralized-SingularityRobust(GSR)inverse,legacysteeringalgorithmsareknowntobeineffectiveinescapinggimbal-lock.ToprovideothersteeringalgorithmswiththesamebenetasGSR,aattitudecontrolleraugmentationdenedasOrthogonalTorqueCompensation(OTC)wasdeveloped.Thismethodwasshowntobesuccessfulinescapinggimbal-lockbyaddingorthogonalcomponentsoftorqueerrorwhenatsingularity.Thismethodwascombinedwithtwoseparatesteeringalgorithms,simulated,andcomparedtoGSRwhereitwasshownnumericallytohaveamuchsmoothertransientresponseforthegimbalrates. 145

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Finally,theproblemswithscalingSGCMGswerediscussed.ItwasshownthattheperformanceofSGCMGsisdegraded(i.e.,alowertorqueamplication)andsamelegacyalgorithmspreviouslyusedonlargerSGCMGscouldbeineffectiveforscaledSGCMGs.Amathematicalproofwasusedtoshowthatwiththeincreaseinthegimbal-ywheelassemblyinertiaIgwcomparedtotheywheelangularmomentumh0causesthisdegradationinperformanceandtheineffectivenessofSGCMGcontrolwithuseoftheSingularityRobust(SR)inverse.TheutilityofscaledSGCMGsisstillviablebecausetheapproximateSGCMGtorqueamplicationforasingleacutatorwasshowntobeontheorderof50whichisfarmorethanthelessthanone-to-oneratioforsystemsofreactionwheels. 146

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APPENDIXARIGIDBODYDYNAMICSFORMULATIONFORCONTROLMOMENTGYROSCOPEACTUATORS(SGCMG/VSCMG)A.1Assumptions Thedynamicformulationforsinglegimbalandvariablespeedcontrolmomentgyroscope(CMG)actuatorsassumestheabsenceoffrictionandexternaltorqueinthesystem(spacecraftincludingCMGs).Inaddition,itisalsoassumedthatthecenterofmass(cm)ofeachCMGisalongitsgimbalaxisandthereforedoesnotaffectthepositionoftheoverallcmofthesystem.Theseassumptionsarevalidforcurrentstate-of-the-artCMGs.A.2Dynamics ThecentroidalangularmomentumofthesystemconsistingofthatfromthespacecraftandasingleCMGis Hc=hW+hG+hS=C(A) withcontributionsfromtheywheelhW,gimbalhG,andthespacecrafthS=C.Theywheelandgimbalangularmomentumareexpressedas hW=Iw1^h(A) andhG=Ig1_^(A) wherethegimbalframebasis[^h,^,^]isrelatedtothespacecraftbody-xedbasisthrougha3-2-1rotationthroughtheangles[,, ]by ^h=(ss )]TJ /F7 11.955 Tf 11.22 0 Td[(cc c)^eb1)]TJ /F4 11.955 Tf 11.21 0 Td[((sc )]TJ /F7 11.955 Tf 11.22 0 Td[(cs c)^eb2)]TJ /F4 11.955 Tf 11.22 0 Td[((cs)^eb3=h1^eb1+h2^eb2+h3^eb3(A) 147

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^=c s^eb1+s s^eb2+c^eb3=t1^eb1+t2^eb2+t3^eb3(A) ^=)]TJ /F4 11.955 Tf 9.3 0 Td[((cs )]TJ /F7 11.955 Tf 10.51 0 Td[(sc c)^eb1+(cc )]TJ /F7 11.955 Tf 10.51 0 Td[(ss c)^eb2+(ss)^eb3=d1^eb1+d2^eb2+d3^eb3(A) wherec()=cos()ands()=sin()and[^eb1,^eb2,^eb3]isthebasisforthespacecraftbodyframe.Therefore,equivalentvectorcomponentsfortheseangularmomentashowninthespacecraftbody-xedbasisare hW=Iw1(h1^eb1+h2^eb2+h3^eb3)(A) andhG=Ig3_(d1^eb1+d2^eb2+d3^eb3)(A) whereIw1andIg3aretherstandthirdcomponentsoftheywheelandgimbalinertias.TheangularmomentumfromthespacecraftisexpressedasthetensorproductofthespacecraftcentroidalinertiadyadicJcwiththeinertialspacecraftangularvelocity!. hS=C=Jc!(A) Thespacecraftcentroidalinertiadyadicis Jc=IG+J0+mGW(rcrc1)]TJ /F8 11.955 Tf 11.96 0 Td[(rcrc)(A) wherercisthepositionofthecmofaCMG'scmfromthecmofthesystemexpressedasrc=rc1^eb1+rc2^eb2+rc3^eb3(A) andthestaticspacecraftinertiadyadicJ0ismadeupofconstantinertias(i.e.,assumingthatthecmoftheCMGsliesalongthegimbalaxis)andtheinertiasdue 148

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tothegimbal-wheelassemblyIGaretimevaryingduetotherotationaboutthegimbalaxis.Theexpressionofthestaticspacecraftinertiadyadicis J0=3Xi=13Xj=1Jij^ebi^ebj(A) where(^ebi^ebj)^ebi=0and(^ebi^ebj)^ebj=^ebi.Itisassumedthatthegimbal-wheelassemblyinertiaisalignedwiththeprincipleaxesandcanbeexpressedas IGW=Ig1^h^h+Ig2^^+Ig3^^(A) where^h^h=3Xi=13Xj=1hihj^ebi^ebj(A) ^^=3Xi=13Xj=1titj^ebi^ebj(A) ^^=3Xi=13Xj=1didj^ebi^ebj(A) Theequationsofmotion(EOM)assumingtorquefreemotion(i.e.,noexternaltorques)arefoundthroughtakingtheinertialtimederivativeofthetotalsystemcentroidalangularmomentuminEq.( A )as dHc dt=XMc=Hc+!Hc=0(A) whereHc=[(a11+a21)_+b1+c1_]+Jc_!(A) ThenalEOMforasingleCMGthathasasinglegimbalis[(a11+a21)_+b1+c1_]+Jc_!+!Hc=0(A) TheJacobianmatricesa11,a21,b1,andc1are 149

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a11=@h @(A) a21=@IG @!(A) b1=@h @_(A) c1=@h @(A) wheretheCMGangularmomentumh=hW+hG. ForasystemofCMGswithasinglegimbal,theEOMconcatenatedintomatrixwhichisaconsequenceofEq.( A ),isexpressedas [(A1+A2)_+B+C_]+Jc_!+!Hc=0(A) wherefornCMGstheJacobianmatricesarerepresentedasA1=[a11,a12,a13,...a1n](A) A2=[a21,a22,a23,...a2n](A) B=[b1,b2,b3,...bn](A) C=[c1,c2,c3,...cn](A) ThisconcludesthedevelopmentoftheEOMforarigidbodyspacecraftsystemofnCMGswhichcontainasinglegimbal. 150

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APPENDIXBMOMENTUMENVELOPECODE 1 function [hx,hy,hz]=Momentum Envelope PM(th,si,h0,int ext) 2 3 % % 4 % 5 % This code is generate the singularity surfaces for a 6 % any general SGCMG cluster with skew angle theta or 7 % inclination angle phi ( i ) and spacing angle si ( i ) 8 % where i = num CMG. 9 % 10 % The angles th ( i ) and si ( i ) are the Euler angles relating the 11 % spin axis of each CMG to the body frames X )]TJ ET 0 G 0 g 0.133 0.545 0.133 RG 0.133 0.545 0.133 rg BT /F35 9.963 Tf 286 -278.35 Td[(axis 12 % % 13 % 14 % Frederick Leve 15 % Last updated : 07/08/08 16 % % 17 % 18 % This function simulated the CMG algorithms 19 % % 20 % INPUTS : 21 % h0 = nominal SGCMG angular momentum ( could be vector if each 22 % CMG does not have the same angular momentum 23 % 24 % th = vector of inclination angles 25 % si = vector of spacing angles 26 % % 27 % OUTPUTS : 28 % hx = angular momentum of envelope in the x )]TJ ET 0 G 0 g 0.133 0.545 0.133 RG 0.133 0.545 0.133 rg BT /F35 9.963 Tf 286 -614.71 Td[(direction 29 % hy = angular momentum of envelope in the x )]TJ ET 0 G 0 g 0.133 0.545 0.133 RG 0.133 0.545 0.133 rg BT /F35 9.963 Tf 286 -634.5 Td[(direction 151

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30 % hz = angular momentum of envelope in the x )]TJ ET 0 G 0 g 0.133 0.545 0.133 RG 0.133 0.545 0.133 rg BT /F35 9.963 Tf 286 -13.85 Td[(direction 31 % % 32 33 % epsilon parameter vector for surface generation 34 % to show internal singular surface make one epsilon )]TJ /F35 9.963 Tf 7.31 0 Td[(1 instead of 1 35 36 37num CMG=length(si); 38 39 if length(h0)==1 40 if int ext==0 41 % external singular surface 42eps=ones(num CMG,1); 43 44 elseif int ext==1 45 % internal singular surface 46 % eps = [ ones ( num CMG )]TJ /F35 9.963 Tf 7.47 0 Td[(1,1);)]TJ /F35 9.963 Tf 7.48 0 Td[(1]; 47 % eps = [1 1 )]TJ /F35 9.963 Tf 7.31 0 Td[(1 1]; 48eps=[11)]TJ /F35 9.963 Tf 7.3 0 Td[(11]; 49 50 else 51display( int ext must be either 0 or 1 ... 52 for internal or external singular surface ) 53 end 54 else 55min h0=min(h0); 56 if int ext==0 57 % external singular surface 58 for i=1:num CMG 59eps(i)=h0(i)/min(h0); 60 end 61 elseif int ext==1 62 % internal singular surface 152

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63 for i=1:num CMG)]TJ /F35 9.963 Tf 7.16 0 Td[(1 64eps(i)=h0(i)/min(h0); 65 end 66eps(num CMG)=)]TJ /F35 9.963 Tf 7.16 0 Td[(h0(num CMG)/min(h0); 67 else 68display( int ext must be either 0 or 1 ... 69 for internal or external singular surface ) 70 end 71 end 72 73 for l=1:num CMG 74 % The transformation C1 is about the inclination angle phi ( i ) 75C1(:,:,l)=[cos(th(l)+3*pi/2)0)]TJ /F35 9.963 Tf 7.16 0 Td[(sin(th(l)+3*pi/2); 76010; 77sin(th(l)+3*pi/2)0cos(th(l)+3*pi/2)]; 78 79C2(:,:,l)=[cos(si(l))sin(si(l))0; 80)]TJ /F35 9.963 Tf 7.16 0 Td[(sin(si(l))cos(si(l))0; 81001]; 82 83g(:,l)=transpose(C1(:,:,l)*C2(:,:,l))*[1;0;0]; 84 end 85 86 87 % total angular momentum at the singular states corresponding to singular 88 % direction u 89H=zeros(3,1); 90n=100; % number of grid point for unit sphere 91[x,y,z]=sphere(n); % generate the unit sphere ( domain of u ) 92 93red light=5; 94traffic light=zeros(n+1,n+1); 95 153

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96 for i=1:n+1 97 for j=1:n+1 98 99u=[x(i,j);y(i,j);z(i,j)]; 100 101 for k=1:num CMG 102 103 % this is the cosine of angle 104 % between vectors since 105 % both of unit norm 106u dot gk=abs(u*g(:,k)); 107 108 if (u dot gk0.95) 109traffic light(i,j)=red light; 110 end 111 112 end 113 end 114 end 115 116 117 for i=1:n+1 118 for j=1:n+1 119 120u=[x(i,j);y(i,j);z(i,j)]; % compose the 121 % singularity vector u 122 123 for k=1:num CMG 124 125H=H+eps(k)/norm(cross(g(:,k),u)) ... 126*cross(cross(g(:,k),u),g(:,k)); 127 128 end 154

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129 130hx(i,j)=H(1); % parse out the components of the 131hy(i,j)=H(2); % momentum vector for later 132hz(i,j)=H(3); % surface or mesh plotting. 133H=zeros(3,1); 134 end 135 end 136 137surfl(hz,hy,hx); 138alpha(0.05); 155

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APPENDIXCCONTROLMOMENTGYROSCOPEACTUATORSPECIFICATIONS TableC-1. Off-the-ShelfCMGSpecications CMGOutputTorque(Nm)Mass(kg))Power(W) Honeywell M5074.633.175.0M95128.838.56129.0M160216.944.0217.0M225305.154.0305.0M325441.061.2441.0M325D441.061.3441.0M600813.581.6814M715969.489.8949.0M13001762.6125.21716M14001898.1131.51899.0EHCMG2304.9146.44822306 SSTL MicroWheel-10S-E0.011.15.0 SSTL CMG15-45S45.018.425.0 156

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REFERENCES [1] JohannBohennberger.Descriptionofamachinefortheexplanationofthelawsofmotionofeartharoundanaxis,andthechangeinorientationofthelatter.TubingerBlatterFurNaturwissenschaftenundArzneikunde,3:72,1817. [2] Simeon-DenisPoisson.Memoironaspecialcaseofrotationalmovementofmassivebodies.Journaldel'EcolePolytechnique,9:247,1813. [3] DonaldMacKenzie.Inventingaccuracy:Ahistoricalsociologyofnuclearmissleguidance.MITPress,pages31,1990. [4] P.C.HughesandP.Carlisle.Spacecraftattitudedynamics.Numberp.156-184.J.Wiley,1986. [5] H.Krishnan,N.H.McClamroch,andM.Reyhanoglu.Attitudestabilizationofarigidspacecraftusingtwomomentumwheelactuators.InInitsEfcientReorientationManeuversforSpacecraftwithMultipleArticulatedPayloads20p(SEEN93-2998811-18),1993. [6] J.D.AdamsandS.W.McKenney.Gyroscopicrollstabilizerforboats,June42003.USPatentApp.10/454,905. [7] IISkeltonandC.Eugene.MixedControlMomentGyroandMomentumWheelAttitudeControlStrategies.PhDthesis,VirginiaPolytechnicInstituteandStateUniversity,2003. [8] APothiawalaandMADahleh.HooOptimalControlfortheAttitudeControlandMomentumManagementoftheSpaceStation.MITPress. [9] F.Wu.Fixed-StructureRobustCMGMomentumManagerDesignfortheInternationalSpaceStation.InAIAAGuidance,Navigation,andControlCon-ferenceandExhibit,Denver,CO,Aug.14-17,2000. [10] S.R.VadaliandH.S.Oh.SpaceStationAttitudeControlandMomentumManagement-ANonlinearLook.JournalofGuidance,Control,andDynamics,15(3):577,1992. [11] L.R.Bishop,R.H.Bishop,andK.L.Lindsay.ProposedCMGMomentumManagementSchemeforSpaceStation.InIN:AIAAGuidance,NavigationandControlConference,Monterey,CA,Aug.17-19,TechnicalPapers,volume2,1987. [12] P.Hattis.PredictiveMomentumManagementfortheSpaceStation.JournalofGuidance,Control,andDynamics,9(4):454,1986. [13] B.Wieandetal.NewApproachtoAttitude/MomentumControlfortheSpaceStation.JournalofGuidance,Control,andDynamics,12(5):714,1989. 157

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[14] R.H.Bishop,S.J.Paynter,andJ.W.Sunkel.AdaptiveControlofSpaceStationwithControlMomentGyros.IEEEControlSystemsMagazine,12(5):23,1992. [15] R.H.Bishop,S.J.Paynter,andJ.W.Sunkel.AdaptiveControlofSpaceStationDuringNominalOperationswithCMGs.InProceedingsofIEEEConferenceonDecisionandControl,30th,Brighton,UnitedKingdom,Dec.11-13,volume3,1991. [16] J.ZhouandD.Zhou.SpacecraftAttitudeControlwithDouble-GimbaledControlMomentGyroscopes.InRoboticsandBiomimetics,2007.ROBIO2007.IEEEInternationalConferenceon,pages1557,2007. [17] T.Yoshikawa.ASteeringLawforDoubleGimbalControlMomentGyroSystem.NASATM-X-64926,1975. [18] B.K.Powell,G.E.Lang,S.I.Lieberman,andS.C.Rybak.SynthesisofDoubleGimbalControlMomentGyroSystemsforSpacecraftAttitudeControl.InAIAAGuidance,NavigationandControlConference,Snowmass,CO,August19-21,1985,pages71s. [19] H.F.Kennel.AControlLawforDouble-GimballedControlMomentGyrosUsedforSpaceVehicleAttitudeControl.Technicalreport,NASA,1970. [20] J.AhmedandD.S.Bernstein.AdaptiveControlofDouble-GimbalControl-MomentGyrowithUnbalancedRotor.JournalofGuidance,Control,andDynamics,25(1):105,2002. [21] D.J.Richie,V.J.Lappas,andGPrassinos.APracticalSmallSatelliteVariable-SpeedControlMomentGyroscopeForCombinedEnergyStorageandAttitudeControl.InAIAA/AASAstrodynamicsSpecialistConferenceandExhibitAug18-21,Honolulu,HI,2008. [22] D.J.Richie,V.J.Lappas,andB.Wie.APracticalVariable-SpeedControlMomentGyroscopeSteeringLawforSmallSatelliteEnergyStorageandAttitudeControl.InAIAA/AASAstrodynamicsSpecialistConferenceandExhibitAug18-21,Honolulu,HI,2008. [23] H.YoonandP.Tsiotras.SpacecraftLine-of-SightControlUsingaSingleVariable-SpeedControlMomentGyro.JournalofGuidanceControlandDynamics,29(6):1295,2006. [24] H.Yoon.SpacecraftAttitudeandPowerControlUsingVariableSpeedControlMomentGyros.PhDthesis,GeorgiaInstituteofTechnology,2004. [25] H.YoonandP.Tsiotras.SpacecraftAngularVelocityandLine-of-SightControlUsingASingle-GimbalVariable-SpeedControlMomentGyro.In2005AIAAGuidance,Navigation,andControlConferenceandExhibit,pages1,2005. 158

PAGE 159

[26] HSchaubandJLJunkins.SingularityAvoidanceUsingNullMotionandVariable-SpeedControlMomentGyros.JournalofGuidance,Control,andDy-namics,23(1):11,2000. [27] H.Lee,I.H.Lee,andH.Bang.OptimalSteeringLawsforVariableSpeedControlMomentGyros.In2005AIAAGuidance,Navigation,andControlConferenceandExhibit,pages1,2005. [28] H.Schaub,S.R.Vadali,andJ.L.Junkins.FeedbackControlLawforVariableSpeedControlMomentGyros.JournaloftheAstronauticalSciences,46(3):307,1998. [29] H.SchaubandJ.L.Junkins.CMGSingularityAvoidanceUsingVSCMGNullMotion(VariableSpeedControlMomentGyroscope).InAIAA/AASAstrodynamicsSpecialistConferenceandExhibit,Boston,MA,pages213,1998. [30] J.Y.Shin,KBLim,andDDMoerder.AttitudeControlforanAero-VehicleUsingVectorThrustingandVariableSpeedControlMomentGyros.InAIAAGuidance,Navigation,andControlConferenceandExhibit,pages1,2005. [31] D.DeVon,R.Fuentes,andJ.Fausz.Closed-LoopPowerTrackingforanIntegratedPowerandAttitudeControlSystemUsingVariable-SpeedControlMomentGyroscopes.InAIAAGuidance,Navigation,andControlConferenceandEx-hibit,AIAA,Aug,2004. [32] H.YoonandP.Tsiotras.SpacecraftAdaptiveAttitudeandPowerTrackingwithVariableSpeedControlMomentGyroscopes.JournalofGuidanceControlandDynamics,25(6):1081,2002. [33] D.A.DeVonandR.J.Fuentes.AdaptiveAttitudeControlandClosed-LoopPowerTrackingforanIntegratedPowerandAttitudeControlSystemusingVariableSpeedControlMomentGyroscopes.InAIAAGuidance,Navigation,andControlConferenceandExhibit,Aug15-18,SanFrancisco,CA.,pages1,2005. [34] J.L.FauszandD.J.Richie.FlywheelSimultaneousAttitudeControlandEnergyStorageUsingaVSCMGConguration.InProceedingsofthe2000IEEEInterna-tionalConferenceonControlApplications,2000.,pages991,2000. [35] D.A.Bearden.Small-satellitecosts.Crosslink,2(1):32,2001. [36] E.B.Tomme.TheParadigmShifttoEffects-BasedSpace:Near-SpaceasaCombatSpaceEffectsEnabler,2005. [37] K.Schilling.DistributedsmallsatellitessystemsinEarthobservationandtelecommunication. [38] W.J.LarsonandJ.R.Wertz.SpaceMissionAnalysisandDesign.Microcrosm,3rdedition,1999. 159

PAGE 160

[39] B.Wie,D.Bailey,andC.Heiberg.SingularityRobustSteeringLogicforRedundantSingle-GimbalControlMomentGyros.JournalofGuidance,Control,andDynam-ics,24(5):865,2001. [40] B.Wie,D.Bailey,andC.Heiberg.RapidMultitargetAcquisitionandPointingControlofAgileSpacecraft.JournalofGuidanceControlandDynamics,25(1):96,2002. [41] H.Kurokawa.AGeometricStudyofSingleGimbalControlMomentGyros.PhDthesis,UniversityofTokyo,1998. [42] G.MarguliesandJNAubrun.GeometricTheoryofSingle-GimbalControlMomentGyroSystems.JournaloftheAstronauticalSciences,26(2):159,1978. [43] N.S.Bedrossian,J.Paradiso,E.V.Bergmann,andD.Rowell.SteeringLawDesignforRedundantSingle-GimbalControlMomentGyroscopes.JournalofGuidance,Control,andDynamics,13(6):1083,1990. [44] BWie.SingularityEscape/AvoidanceSteeringLogicforControlMomentGyroSystems,August2003. [45] B.HamiltonandB.Underhill.ModernMomentumSystemsforSpacecraftAttitudeControl.AdvancesintheAstronauticalSciences,125:57,2006. [46] FBAbbott,B.Hamilton,T.Kreider,P.DiLeonardo,andD.Smith.MCSRevolution.AdvancesintheAstronauticalSciences,125:99,2006. [47] H.Kurokawa.SurveyofTheoryandSteeringLawsofSingle-GimbalControlMomentGyros.JournalofGuidanceControlandDynamics,30(5),2007. [48] T.A.Sands,J.J.Kim,andB.Agrawal.2HSingularity-FreeMomentumGenerationwithNon-RedundantSingleGimbaledControlMomentGyroscopes.In45thIEEEConferenceonDecisionandControl,pages1551,2006. [49] B.UnderhillandB.Hamilton.MomentumControlSystemandLine-of-SightTestbed.AdvancesintheAstronauticalSciences,125:543,2006. [50] D.BrownandM.A.Peck.Scissored-PairControl-MomentGyros:AMechanicalConstraintSavesPower.JournalOFGuidance,Control,ANDDynamics,31(6),2008. [51] D.Brown.ControlMomentGyrosasSpace-RoboticsActuators.AIAA/AASAstrodynamicsSpecialistConferenceandExhibit18-21August2008,Honolulu,HI,2008. [52] H.Kurokawa.ConstrainedSteeringLawofPyramid-TypeControlMomentGyrosandGroundTests.JournalofGuidance,Control,andDynamics,20(3):445,1997. 160

PAGE 161

[53] FLeve,GBoyarko,andNFitz-Coy.Optimizationinchoosinggimbalaxisorientationsofoptimizationinchoosinggimbalaxisorientationsofacmgattitudecontrolsystem.Infotech@AerospaceConferenceApril6-10,SeattleWA,2009. [54] B.WieandJ.Lu.FeedbackControlLogicforSpacecraftEigenaxisRotationsUnderSlewRateandControlConstraints.JournalofGuidance,Control,andDynamics,18(6):1372,1995. [55] Honeywell.Dynamiccmgarrayandmethod.Patent20070029447,2009. [56] J.Baillieul,J.Hollerbach,andR.Brockett.ProgrammingandControlofKinematicallyRedundantManipulators.InDecisionandControl,1984.The23rdIEEEConferenceon,volume23,1984. [57] M.A.Peck,B.J.Hamilton,andB.Underhill.MethodandSystemforOptimizingTorqueinaCMGArray,July302004.USPatentApp.10/903,774. [58] D.A.Bailey,C.J.Heiberg,andB.Wie.ContinuousAttitudeControlthatAvoidsCMGArraySingularities,October102000.USPatent6,131,056. [59] M.R.Elgersma,D.P.Johnson,M.A.Peck,B.K.Underhill,G.Stein,B.G.Morton,andB.J.Hamilton.MethodandSystemforControllingSetsofCollinearControlMomentGyroscopes,November302005.USPatentApp.11/291,706. [60] L.JonesandM.Peck.Ageneralizedframeworkforlinearly-constrainedsingularity-freecontrolmomentgyrosteeringlaws.InAIAAGuidance,Naviga-tionandControlConference,209. [61] D.E.Cornick.SingularityAvoidanceControlLawsforSingleGimbalControlMomentGyros.AIAA,pages20,1979. [62] R.D.HefnerandC.H.McKenzie.ATechniqueforMaximizingTorqueCapabilityofControlMomentGyroSystems.Astrodynamics,54(83-387):905,1983. [63] B.Wie.SpaceVehicleDynamicsandControl.AIAA,1998. [64] D.JungandP.Tsiotras.AnExperimentalComparisonofCMGSteeringControlLaws.InProceedingsoftheAIAAAstrodynamicsSpecialistConference,2004. [65] M.D.KuhnsandA.A.Rodriguez.APreferredTrajectoryTrackingSteeringLawforSpacecraftwithRedundantCMGs.InProceedingsoftheAmericanControlConference,1995.,volume5,1995. [66] J.A.Paradiso.Globalsteeringofsinglegimballedcontrolmomentgyroscopesusingadirectedsearch.JournalofGuidance,Control,andDynamics,15(5):1236,1992. [67] J.A.Paradiso.ASearch-BasedApproachtoSteeringSingleGimballedCMGs.Technicalreport,NASAJohnsonSpaceCenter(JSC),Houston,TX,CSDL,1991. 161

PAGE 162

[68] S.Asghar,PLPalmer,andM.Roberts.Exactsteeringlawforpyramid-typefourcontrolmomentgyrosystems.InAIAA/AASAstrodynamicsSpecialistConferenceandExhibit,Keystone,CO,2006. [69] Y.NakamuraandH.Hanafusa.InverseKinematicSolutionswithSingularityRobustnessforRobotManipulatorControl.ASME,Transactions,JournalofDynamicSystems,Measurement,andControl,108:163,1986. [70] K.A.FordandC.D.Hall.SingularDirectionAvoidanceSteeringforControl-MomentGyros.JournalofGuidanceControlandDynamics,23(4):648,2000. [71] A.N.Pechev.Feedback-basedsteeringlawforcontrolmomentgyros.JournalofGuidanceControlandDynamics,30(3):848,2007. [72] FLeve,GBoyarko,JMunoz,andNFitz-Coy.Comparisonofstate-of-the-artsteeringlogicsforsingle-gimbalcontrolmomentgyroscopes.InAASAstrodynam-icsConference,Pittsburgh,PA,2009. [73] T.A.Sands.FinePointingofMilitarySpacecraft.PhDthesis,NavalPostgraduateSchool,March2007. [74] S.R.Vadali,S.R.Walker,andH.S.Oh.Preferredgimbalanglesforsinglegimbalcontrolmomentgyros.JournalofGuidance,Control,andDynamics,13(6):1090,1990. [75] H.Leeghim,H.Bang,andJ.O.Park.Singularityavoidanceofcontrolmomentgyrosbyone-stepaheadsingularityindex.ActaAstronautica,2008. [76] D.N.NenchevandY.Tsumaki.Singularity-ConsistentAttitudeMotionAnalysisandControlBasedonEulerAngleParameterization.InSICEAnnualConference,volume1,2003. [77] M.D.KuhnsandA.A.Rodriquez.SingularityAvoidanceControlLawsForaMultipleCMGSpacecraftAttitudeControlSystem.InAmericanControlConference,volume3,1994. [78] J.S.Lee,H.Bang,andH.Lee.SingularityAvoidancebyGameTheoryforControlMomentGyros.InAIAAGuidance,Navigation,andControlConferenceandExhibit,pages1,2005. [79] S.R.Vadali.Variable-StructureControlofSpacecraftLarge-AngleManeuvers.JournalofGuidance,Control,andDynamics,9(2):235,1986. [80] WMacKunis,FLeve,AWaldrum,NFitz-Coy,andWDixon.AdaptiveNeuralNetworkSatelliteAttitudeControlwithExperimentalValidation.IEEETransactionsonControlSystemsTechnology,2009. 162

PAGE 163

[81] W.MacKunis,K.Dupree,N.Fitz-Coy,andW.Dixon.AdaptiveSatelliteAttitudeControlinthePresenceofInertiaandCMGGimbalFrictionUncertainties.InAmericanControlConference,2007. [82] F.Leve.Developmentofthespacecraftorientationbuoyancyexperimentalkiosk.Master'sthesis,UniversityofFlorida,2008. [83] A.FlemingandI.M.Ross.Singularity-FreeOptimalSteeringOfControlMomentGyros.AdvancesintheAstronauticalSciences,123,2006. [84] H.S.OhandSRVadali.Feedbackcontrolandsteeringlawsforspacecraftusingsinglegimbalcontrolmomentgyros.Master'sthesis,TexasA&MUniversity,1988. 163

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BIOGRAPHICALSKETCH FrederickAaronLevewasborninHollywood,Florida,in1981.InAugust2000hewasacceptedintotheUniversityofFloridasDepartmentofAerospaceEngineeringintheCollegeofEngineeringwherehepursuedhisbachelorsdegreesinMechanicalandAerospaceEngineering.AftercompletinghisbachelorsdegreesinMay2005,hewasacceptedintothemastersprograminaerospaceengineeringattheUniversityofFlorida.Whileinthemastersprogram,hereceivedtwoawardsinacademia.InJanuary2007,hereceivedtheAmericanInstituteofAeronauticsandAstronautic'sAbeZaremAwardforDistinguishedAchievementinAstronautics.ForthisawardhewasinvitedtoValencia,Spain,wherehecompetedintheInternationalAstronauticalFederationsInternationalAstronauticalCongressStudentCompetition.HerehereceivedthesilverHermanOberthmedalinthegraduatecategory.HecompletedthemastersprograminMay2008andcontinuedontohisPhD.InMay2006,hewasacceptedtotheAirForceResearchLab(AFRL)SpaceScholarsProgram,wherespenthissummerconductingspaceresearch.Afterspacescholars,hewasemployedasastudenttemporaryemployeeatAFRLwherehereceivedtheCivilianQuarterlyAwardforallofAFRLinhiscategory.CurrentlyheworksintheGuidance,Navigation,andControlgroupatAFRLSpaceVehiclesDirectorate.Hisinterestsinclude,appliedmath,satelliteattitudecontrol,satellitepursuitevasion,astrodynamics,andorbitrelativemotion. 164