Insert Running Workflow Large

GRAYSCALE NAME:

HARDCOPY NAME:

17[_!7_7
NASA Contractor Report 4111

Design Considerations of

Manipulator and Feel System

Characteristics in Roll Tracking

Donald E. Johnston and Bimal L. Aponso

Systems Technology, Inc.

Hawthorne, California

Prepared for

Ames Research Center

Dryden Flight Research Facility

under Contract NAS2-12221

National Aeronautics

and Space Administration

Scientific and Technical

Information Division

1988
FOREWORD

This research was accomplished at Systems Technology, Inc.,

Hawthorne, CA, under Contract NAS2-12221 for the NASA Ames Research

Center, Dryden Flight Research Facility. The NASA Contract Technical

Manager was Mr. D. T. Berry. Mr. D. T. McRuer was the STI Technical

Director and D. E. Johnston was the Principal Investigator. The

experimental setup, operation, and data reduction was accomplished by

Mr. B. L. Aponso. Other staff members who contributed in the

experiments were Messrs. R. H. Hoh, J. R. Hogue, and R. E. Magdaleno.

The control sticks and force loader system were furnished by the

McFadden Systems, Inc.

ii
T&BLE OF CONTENTS

I. INTRODUCTION ............................................... 1

II. BACKGROUND ................................................. 5

A. Some Simple Pilot/Vehicle Considerations ............... 6

B. The Neuromuscular System Model ......................... I0

C. Limb Actuation/Feel System Interface ................... 16

III. EXPERIMENT GOALS AND SETUP ................................. 21

A. Experiment Setup ....................................... 21

B. Disturbance (Target) Input Function .................... 24

C. Protocol ............................................... 25

IV. DISPLACEMENT SIDESTICK AND FEEL SYSTEM ..................... 27

A. Influence on Pilot's Neuromuscular System Mode ......... 27

B. Influence on Closed-Loop Tracking Performance .......... 37

C. Influence on Pilot Rating .............................. 41

D. Addition of 2nd Order Command Filter Lag ............... 47

E. Summary ................................................ 52

V. CENTER STICK WITH DISPLACEMENT VS FORCE SENSING ............ 55

A. Configurations Investigated ............................ 55

B. Force vs Displacement Sensing .......................... 59

C. Influence of Dynamic Lag Location ...................... 64

D. Influence of Cumulative Effective Time Delay ........... 67

E. Summary ................................................ 71

VI. PILOT DYNAMIC MODEL WITH FORCE VS

DISPLACEMENT CENTER STICK .................................. 73

A. High Order Dynamic Model ............................... 74

B. Crossover Model Approach ............................... 84

C. Roll Ratchet Potential ................................. 88

VII. CONCLUSIONS AND RECOMMENDATIONS ............................ 89

REFERENCES ........................................................ 93

iii
TABLE OF CONTENTS (Continued)

APPENDIX A ........................................................

APPENDIX B ........................................................

APPENDIX C ........................................................

LIST OF FIGURES

i. Roll Ratchet During Banking Maneuver ........................

2. General Compensatory Control Situation and

Basic Human Operator Subsystems .............................

3. Ideal Crossover Model .......................................

4. Closed-Loop Neuromuscular System Model Fit to Y_ for

Pressure and Free-Moving Manipulator (Yc = I/s)F

from Ref. 2.................................................

5. Example Describing Function Showing High Frequency Peak

(Taken from Gordon-Smith, Ref. 12) ..........................

6. Neuromuscular Subsystems for Free-Moving and Pressure

Manipulators and Central Equalization for Rate Dynamics .....

7. Bode and Root Locus for Spindle Feedback (Some Effects of

Joint Sensor Feedback Shown) (Ref. 2) .......................

8. Bode Amplitude Ratio Plot for Neuromuscular System

Contribution to Roll Ratchet Potential ......................

9. Pilot Closed-Loop Limb Actuation

With Feel System Dynamics ................................ -..

i0. Closed-Loop Limb/Manipulator System with Force Sensing ......

Ii. Experimental Setup ..........................................

12. Display Format ..............................................

13. Definition of Flying Qualities Levels .......................

14. Example Bode Plot With K/S Fit to Obtain AAR ................

Page

95

102

104

1

II

12

13

15

16

18

19

22

22

26

29

iv
LIST OF FIGURES (Continued)

15.

16.

17.

18.

19.

20.

21.

22.

23.

24.

25.

26.

27.

28.

29.

Amplitude Deviation From Yc Asymptote Due to

Neuromuscular Mode .........................................

Influence of Feel Gradient and Effective Time

Delay on Neuromuscular Mode Peaking With Stick

Displacement Sensing .......................................

Comparison of Neuromuscular Peaking With

Displacement vs Force Sensing ..............................

Typical Describing Function Plots for

Displacement vs Force Sensing ..............................

Roll Ratchet Tendency With Variation of Roll Lag and

Command Gain; Displacement vs Force Sensing ................

Roll Tracking Crossover vs Command Gain

With Displacement Sensing ..................................

Roll Tracking Crossover vs Command Gain

With Force and Displacement Sensing ........................

Relationship Between Tracking Performance and

Loop Closure Parameters ....................................

Variation of CHPR With Position Sensing

Command Gain and Feel Gradient .............................

Control Force-Response Gains,

Up-and-Away Flight (Ref. 16) ...............................

Comparison of CHPR vs Command Gain for Flight and

Lab With Feel Gradient 1.2 - 1.3 ib/deg ....................

Comparison of CHPR vs Command Gain for Flight and

Lab With Feel Gradient 0.65 - 0.7 ib/deg ...................

Effective Controlled Element Dynamics for 2nd Order

Command Filter Experiment with Sidestick ...................

Influence of Effective Time Delay and Command Gain

on CHPR (0.65 lb/deg Sidestick) ............................

Influence of Effective Time Delay and Command Gain on

Roll Tracking Performance (0.65 ib/deg Sidestick) ..........

Page

30

31

34

35

36

38

39

41

42

43

45

46

48

49

51

v
30.

31.

32.

33.

36.

37.

38.

39.

40.

41.

42.

43.

44.

LIST OF FIGD_.ES (Continued)

Command Gain Normalized Tracking Performance Measures

Variation With Effective Time Delay

(0.65 ib/deg Sidestick) ....................................

Command Gain Normalized CHPR Variation With Effective

Time Delay (0.65 ib/deg Sidestick) .........................

Comparison of CHPR With MIL-F-8785C Effective

Time Delay Criteria (0.65 ib/deg Sidestick) ................

Effective Controlled Element Configurations

Employed -- Roll Tracking With Center Stick ................

Roll Acceleration Time Response to Step Stick Force ........

Influence of Feel and Filter Dynamics on Tracking

Bandwidth With Center Stick Displacement

and Force Sensing ..........................................

Influence of Command Gain on Tracking Bandwidth

(Center Stick) .............................................

Influence of Feel and Filter Dynamics on Tracking

Performance With Center Stick Displacement

and Force Sensing ..........................................

Influence of Feel and Filter Dynamics on CHPR With

Center Stick Displacement and Force Sensing ................

Influence of Dynamic Lag Location on Tracking Bandwidth

(Center Stick) .............................................

Influence of Dynamic Lag Location on Tracking Error

(Center Stick) .............................................

Influence of Dynamic Lag Location on CHPR

(Center Stick) .............................................

RMS Tracking Error vs Effective Time Delay,

Subject A (Center Stick) ...................................

RMS Tracking Error vs Effective Time Delay,

Subject B (Center Stick) ...................................

CHPR Variation With Effective Time Delay and

Command Gain (Center Stick) ................................

Page

51

53

53

57

58

60

61

62

63

64

65

66

68

68

69

vi
45.

46.

47.

48.

49.

50.

51.

52.

53.

A-I.

A-2.

A-3.

A-4.

A-5.

A-6.

B-I.

LIST OF FIGURES (Concluded)

Comparison of CHPR with MIL-F-8785C Effective Time Delay

Criteria (4 ib/in Center Stick) ............................

Example Root Locus of the Limb/Manipulator

Positional Loop Closure ....................................

Dynamic Model Match to Describing Function Data, Force

Sensing Center Stick, Subject A, _FS = 26, _CF = 14 ........

Dynamic Model Match to Describing Function Data, Force

Sensing Center Stick, Subject A, mFS = 14, mCF = 26 ........

Dynamic Model Match to Describing Function Data,

Displacement Sensing Center Stick, Subject A,

mFS = 14, mCF = 26 .........................................

Dynamic Model Match to Describing Function Data,

Displacement Sensing Center Stick, Subject A,

mFS = 26, mCF = 14 .........................................

Pilot Generation of Lead With Displacement Sensing

Center Stick, Subject A, _FS = 14, _CF = 26 ................

Pilot Generation of Lead With Displacement Sensing

Center Stick, Subject B, _FS = 14, _CF = 26 ................

Effective Controlled Element Time Delay

Summation for Crossover Model ..............................

Sidestick Tracking Station Set-Up ..........................

Small Displacement (Stiff) Sidestick Calibration ...........

Large Displacement (Medium) Sidestick Calibration ..........

Very Large Displacement (Soft) Sidestick Calibration .......

Center-Stick Calibration ...................................

Force Sensing Electrical Signal Breakout Calibration .......

Roll Analog Simulation .....................................

Page

70

74

76

77

78

79

80

81

85

96

97

98

99

I00

I01

103

vi i
LIST OF TABLES

Io

2.

3.

C-I.

C-2.

Page

Stick/Feel System Configurations ............................ 23

Roll Tracking Forcing Function .............................. 24

Configurations Fitted With High Order Models ................ 75

Closed-Loop Tracking Measures, Center-Stick ................. 107

Closed-Loop Tracking Measures, Side-Stick ................... iii

vi ii
A

ADI

CHPR

CRT

D

DFA

OFRF

dB

e

F

Gm

GM

H

i

K

Kc

KF S

m

ms

n

NM/L

P

PIO

RMS (rms)

LEST OF SYMBOLS

Analog

Attitude Director Indicator

Cooper Harper Pilot Rating

Cathode Ray Tube

Digital; Displacement

Describing Function Analyzer

Dryden Flight Research Facility

Decibel, 20 log10() for Amplitude Ratio; i0 log10( ) for Power

System Error

Roll Stick Force

Combined Muscle-Manipulator Dynamics

Gain Margin

Feedback Loop Dynamic Transfer Functions

Forcing Function Input

System Gain

Controlled Element Gain; Command/Force Gradient

Feel System Force/Displacement Gradient

Closed-Loop System Output

Millisecond

Number of Runs

Neuromuscular and Limb System

Roll Rate; First Order Inverse Time Constant of Transfer

Function Denominator

Pilot Induced Oscillations

Root Mean Square

ix
s

t

T

X

Y

Z

_AR

O

%

T

¢

%

coC 0 e

%

#

LIST OF SYMBOLS (Continued)

Laplace Transform Variable

Time

First-Order Time Constant

Internal Muscle Length

Forward Loop Transfer Function Particularized by Subscript

First Order Inverse Time Constant of Transfer Function

Numerator

Alpha Motor Neuron

Stick or Control Surface Command

Neuromuscular Peak Amplitude Ratio Departure from YpYc Bode

As ymp tote

Damping Ratio Terms

Standard Deviation

Task RMS Error

Time De lay

Roll Response Angle

Phase Margin

Angular Frequency Term

Closed-Loop Bandwidth (Crossover of the 0 dB Gain Line with

the K/s Amplitude Ratio Plot)

Product of Crossover Frequency and RMS Error (Performance

Measure)

Frequency at Which System Becomes Unstable (_ = 180 °)

Angle

Unit of Pound Force

x
Subscripts

a

c

CF

e

FS

i

J

N orNM

P

r orR

sp

LIST OF SYMBOLS (Concluded)

Muscle Actuation Open-Loop Elements

Controlled Element; Command; Crossover

Command Filter

Error;

Feel System

Input or Forcing Function

Joint Receptor

Neuromuscular System Closed-Loop Dynamic Mode

Pilot

Roll Mode

Muscle Spindle

Roll Angle

xi
Almost every new aircraft with fly-by-wire or commandaugmentation

in the roll axis has encountered either Pilot Induced Oscillations (PIO)

or roll ratchet (or both) in early flight phases. PIO has typically

been associated with high gain, neutrally stable closed-loop pilotvehicle control oscillations with a frequency of about I/2 Hz (3 tad/

sec). The "roll ratchet" has been somewhatmore obscure and idiosyncratic, appearing most often in rapid rolling maneuvers. Ratchet frequencies are typically 2-3 Hz (12-19 rad/sec). Figure i illustrates

this seldom recorded phenomenon. The frequency alone indicates that the

PIO and ratchet situations are different phenomena,yet both clearly

involve the closed-loop pilot/vehicle system.

(Ib) 241 _ osAx(deg) 0 '

-90

180

120 _/_

/

Fas O r

_ 60

(deg) -600

-120

-180

P

deg i

I00

-50 ]

oy

(g) -0.2 J

} ! i i I i i I i

0 2 4 6 8

Time (sec)

a) F-/6

P I00 _ _ _//_ (deg / 0

_sE_f -I00

ay

(g)

I I I I I I I I 1 I I

0 2 4 6 8 I0 12 14 16 18 20

Time (sec)

b) A/rcrof! A, CAS I

Figure I. Roll Ratchet During Banking Haneuver
From the earliest studies on the interaction between the human

pilot's neuromuscular system and aircraft control devices (e.g., Refs. i

and 2) the presence of a neuromuscular system limb manipulator dynamic

resonance peak at 14-19 rad/sec has been well known. In Ref. 3 the

neuromuscular system characteristics are cited as "exceptionally important and critically limiting in such matters as

• control precision where limited by the pilot's

neuromuscular system dynamics.

• effects of control system nonlinearities, including their connections with control system sensitivity requirements."

Reference 4 and other summaries place great stress on the importance of

considering these characteristics even though this frequency range of

major activity may be well above the bandwidth associated with the

"usual" control task.

It is becoming more and more apparent that modern, high performance,

high gain, command response flight control system bandwidths may be

encroaching on the neurological system. Advances in flight control

system fly-by-wire technology permit new manipulation devices, for

example force sensing sticks, at the pilot output/effective-vehicle

interface. These have thus far been generally successful in application, but have introduced or enlarged some pilot-vehicle flying qualities problems. Particular problems include:

o high roll control sensitivity and PlO's in precision maneuvering;

• roll ratchet or jerkyness in otherwise steady

rolling maneuvers;

• sensitivity to the way the pilot grips the stick

or to location of his hand/arm support;

• effective time delay associated with stick filters, with attendant increase in pilot remnant;

• biodynamic interactions, e.g., hand/arm stick

bobweight effects.
Attempts to alleviate these effects have involved adjustments in stick

force gradients, filtering, and sensitivity. Thesehave included introduction of various nonlinear elements such as commandgain reduction as

a function of pilot input amplitude or frequency (Ref. 5), filter time

constant changes with sense of input (increase or decrease) (Ref. 6),

and different force gradient for right and left roll commands(Ref. 6).

These adjustments have generally involved complex ad hoc empirical modifications in the course of the aircraft development. Muchof this has

been accomplished in flight test with correspondingly large cost.

Other modifications have included reverting from stick force sensing

to displacement sensing (Ref. 7) to take advantage of the natural

filtering of the associated mechanical systems. This approach may

involve increased complexity in the manipulator and feel package due to

the necessary pivots, detents, springs, dash pots, etc., but the dynamic

lag of the quasi-linear elements can be compensated for by reducing

other intentional forward path filter elements. Furthermore, recent

flight experience (Refs. 8, 9) has suggested that the pilot may not be

as sensitive to this feel system lag contribution as he is to other

downstream lags from electronic filters, computational delays, etc.

This has led to somecontroversy (Ref. 9) concerning intent and interpretation of the response time delay criteria of the current flying

qualities specification (Ref. I0) and is discussed in the next section.

Thus there is a need to revisit and expand the earlier neuromuscular

system experimentation with a focus on now practical flight control

system configurations in order to quantify possible interactions between

these and the neuromuscular system. The result should provide

manipulator/flight control system design guides and criteria to minimize

roll control problems. A first step was taken in the Ref. ii investigation of side-stick force sensing and the interaction with the neuromuscular system. The experimental program with results documented

herein was taken to further satisfy these goals.

In the section which follows, a fairly broad and simplified view of

somepertinent manual control considerations are presented as background

for the experiment. These include the pilot/controlled-element

3
crossover model and the basic humanoperator subsystems of particular

interest (arm neuromuscular subsystem model, limb/manipulator and feel

system interaction, etc.). This sets the stage for definition of experiment setup and goals which is presented in Section III.

The experiment encompassed some 48 manipulator/filter/aircraft

configurations involving 320 runs by two subjects. Key findings are

summarized in Sections IV through VI. Section IV concentrates on the

displacement side-stick experiment results and compares these with the

Ref. II force sidestick results. Attention is focused on control bandwidth, excitement (peaking) of the neuromuscular mode, feel force/

displacement gradient effects, time delay effects, etc. Section V is

devoted to experiments with a center-stick in which force vs displacement sensing, feel system lag, and commandprefilter lag influences on

tracking performance and pilot preference are investigated. Results in

these two sections are summarized in numerousplots which are intended

to serve as guides in the design of future control systems.

Section VI concentrates on extraction of dynamic models for the

pilot and closed-loop arm/manipulator/feel system from the describing

function data obtained in the Section V experiments. Parameters suitable for detailed models and for the simple crossover model are derived.

Section VII summarizes the findings and conclusions. Details concerning the manipulators, feel gradients, etc. along with run logs and

data summariesare presented in the several Appendixes.
SECTION II

BACKGRODR_D

In the preceding Ref. ii study a fixed-base simulation was performed

to identify and quantify interactions between the pilot's hand/arm

neuromuscular subsystem and various effective controlled element aspects

representative of typical modern fighter aircraft high-response, rollrate-command control system mechanizations. In particular:

• force-sensing side-stick type manipulator

• command augmentation prefiltering

• flight control system effective time delay

The simulation results provided insight to high frequency PIO (roll

ratchet), low frequency PIO, and roll-to-right control and handling

problems previously observed in experimental and production fly-by-wire

control systems. The simulation configurations encompassed and/or

duplicated several actual flight situations, reproduced control problems

observed in flight, and validated the concept that the high frequency

nuisance mode known as "roll ratchet" can derive from the pilot's own

neuromuscular subsystem. The results demonstrated that roll ratchet

tendency, difficult to detect in fixed-base simulation display or time

traces, is readily apparent from application of frequency response

spectral analysis techniques to the experimental data with adjustment

for the pilot's motion sensing neurological apparatus. Results show

that force-sensing sidestick manipulator force/displacement/command

gradients, command prefilters, and flight control system time delays,

need to be carefully tuned to minimize neuromuscular mode amplitude

peaking (roll ratchet tendency) without restricting roll control bandwidth (with resulting sluggish or PIO prone control).

One means of reducing control sensitivity problems and roll ratchet

tendency in the F-18, X-29, and other aircraft, has been to replace

stick force sensing with stick displacement sensing. This results in a

natural smoothing (filtering) of the command input via the manipulator
dynamic elements. This also brings into play a different set of receptors in the pilot's hand/arm neuromuscular subsystem. Results obtained

from the X-29 aircraft flight test (Ref. 8) also cast doubt on the

response time delay criterion of the flying qualities specification,

MIL-F-8785C. This specification states that the effective time delay

between application of force at the manipulator and appropriate vehicle

response shall not exceed 0.i sec for Level I and 0.2 sec for Level II

flying quality ratings. Due to the use of stick displacement sensing

and commandprefiltering in the X-29 plus the presence of computational

time delay, surface actuation and aircraft lags, etc., the total effective time delay is predicted to produce Level II pilot ratings. However

the aircraft is rated a solid Level I in flight. If the feel system

dynamic delay is excluded from the effective time delay calculation,

Level I response is predicted.

The question raised is whether the dynamic lag of the feel mechanism

is transparent to the pilot (and therefore not to be included in the

calculation of forward loop effective time delay, or alternatively, the

specification criteria to apply from the source of the electrical command-- force or displacement) or whether the specification criteria is

too stringent and should be relaxed to allow for the dynamic lag contribution of the displacement mechanismand feel package.

The relative merits of stick displacement commandversus prefiltered

stick force commandhave not been quantified in terms of effective time

delay, maximumcontrol bandwidth, neuromuscular peaking, low and high

frequency PIO tendency, etc. The goal of the experimental program

reported herein is to provide such information, to compare results with

the preceding force sensing investigation, and to provide design

guides/tradeoffs for minimizing roll control problems in future high

performance aircraft.

A. SOME SIMPLE PILOT/VEHICLE CONSIDERATIONS

A block diagram for a general compensatory control situation showing

the basic human operator subsystems is presented in Fig. 2 (from

Ref. 2). The major elements are the controlled element dynamics (Yc) ,
0

l-w

0_0T IIL

09

c --._c

0W

C._ r7

°_

_J

o ._o

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°r'4 00U00

64

{",1

c_O
the manipulator dynamics, and the human operator which encompasses the

central sensory, equalization, and neuromuscular (N.M.) command elements, the muscle actuation dynamic element, and the various N.M. force

and displacement feedbacks.

Simple tracking task pilot model forms and associated pilot vehicle

system properties begin with the ideal crossover model of Fig. 3 (see

e.g., Ref. 4). In this model the pilot adjusts his dynamic characteristics so that the open-loop pilot-vehicle dynamics are approximately K/s

over the frequency band immediately above and below the gain crossover.

The prescription for K/s-like controlled element dynamics in the region

of pilot vehicle system crossover as an often desirable form stems from

the fundamental feature of human dynamics that no pilot lead is then

required to establish good closed-loop dynamics over a wide range of

pilot gains. The basic recipe is almost invariably conditioned by such

statements as "in the frequency region about crossover." Such statements are made to restrict the form of the pilot model to that required

only in the crossover region. In particular, the cases covered are such

that an effective time delay term in the pilot model is an adequate

approximation to the high frequency effects. The model also indicates

that in full attention tracking operations the pilot will adjust his

gain to offset any variation in controlled element gain in order to

maintain a nearly fixed control system bandwidth. Thus the full attention closed-loop bandwidth _c (identified as the crossover of the 0 dB

gain line with the K/s amplitude plot) is independent of the controlled

element gain. Furthermore, the pilot tends to keep the product of the

crossover frequency and the task RMS error, _cae , constant.

In the crossover model the exponential term with time delay, T,

approximates all the lag contributions due to pilot and vehicle high

frequency dynamic modes. The effective time delay is a function of,

among other things, the force/displacement characteristics of the manipulator. As shown in Fig. 3, an isometric (force) stick results in less

lag than does an isotonic (free moving) stick. Past experimentation has

identified the difference to be approximately 0. i sec (e.g., Ref. 4).

8
-'- constant _c

f(K c )

o_c o-e = constant

1.0 I0

I I

_o (rad/sec)

Ratio Kpl OdB _,wcj

(dB) Kp 2

Yp Yc :

-'rjw

Kc Kp e

j(_

Phase

(deg)

-9O

-180 ....

GM

I0.0

I

Isotonic

Stick

( free moving)

Isometric

Stick

( force )

AT : r isotonic - r isometric --"0.1 sec

Figure _. Tdeal Crossover _fo_el
In Fig. 3 if the pilot gain were set at the value represented by Kp2

with an isometric stick, the bandwidth would be indicated by mc2and

would result in a system stability phase margin _m2,and gain margin,

GM. If this samegain were employed with the isotonic stick, the phase

margin would be 0, and a low frequency continuous oscillation (PIO)

would result. This oscillation can then be alleviated by the pilot

reducing his gain to the value represented by Kpt and accepting the

reduced bandwidth. Thus Fig. 3 can be used to demonstrate the common

low frequency PIO problem which generally occurs in the vicinity of

0.5 Hz and which is relieved by reducing pilot gain. Note that in the

crossover model an 0au of 4 rad/sec corresponds to T = _/2_" 0.4 sec

for the total pilot, control system, aircraft, etc., latency.

S. THENEUROMUSCULAR SYSTEM MODEL

As previously noted, early studies on the neuromuscular system

(e.g., Ref. 1 circa 1908) lloted the presence of a ne_r_)musc_11ar system

or limb-manipulator peak at 14-19 tad/see, well past ti_e usual "crossover region." The effects of various restraints on the Limb/

neuromuscular system are described in detail in Ref. 2. Figure 4 (from

Ref. 2) shows closed-loop neuromuscular system ,nodel fits to pilot/ controlled-element describing fdnction measurements for pressure and free

moving manipulators. .-_n important part of the neuromuscular dynamics in

each case is a peaking quadratic mode with damping and natural frequency

of

HANI PULATOR NH/L DYNAMICS

Free Hoving

Isometric or Pressure

[0.07, 17]

[0. 138, t8.61

The experiments which allowed identification of these modes used forcing

fdnctions having a low power shelf extending to higher frequencies than

uormally utilized in tracking tasks. The human pilot describing function data of Fig. 5 from Ref. [2, provides an example of the range of

frequencies needed to completely define the resonant peak. Note in

Fig. 4 that there is a second neuromuscular mode for the pressure

i0
m

"0

0

-40

-80

"0

-120

-160

-200

0.1

20dB

OdB for Yp,

(3

r l

[]

OdB for/o/nt sensor _.

I_-_ --

e

Manipulator Curve --it*

e-j(_(.143)

\

Central Time Delay

Included in Fit

82 ms

66 ms

i

(.07, 17)(.96,26.7)

Free Moving

• Pressure

e-joJ(.064)

(11.8)(.138,18.6)

1.0 _ ( rad/sec) I0

t

Free

Moving

50

"Note: To simplify the notation, (s 2 + 2_s +_) is written (_ , w) and

(s+a_ is written (a].

Figure 4. Closed-Loop Neuromuscular Nystem MoSel Wit to Y_ for

Pressure and Free-Moving Manipulator (Yc = I/s) from ReX. 2

ii
40

8

O_ eo

>-

-6O

Fp_

LO

| I I I i i I

I 1 i

O.

>-

--J _z_

i I i i

-360

m FITTED

0-1.

I I | J I I I | I I I | _ !

1.o (.0 (RAr'I/SEC) lo.o

RUN _0?9. SUBJECT LR. _I = _.0 RAD/SEC. F_T,E-NOV'[NG P_IFULATOI_.

Figure 5. Example Describing Function Showing High Frequency Peak

(Taken from Gordon-Smith, Ref. 12)

12
manipulator which is approximated by a first-order lag break at 11.8

rad/sec. This mode is also somewhat dependent on the nature of the

manipulator restraints (Refs. 13 and 14). The free moving manipulator

results in a second, highly damped second-order mode at about 27

rad/sec.

The reason that the neuromuscular actuation system dynamics differ

when the manipulator restraints are changed is physiological -- the

neuromuscular apparatus involved depends on the restraints and limb

movements. While greatly oversimplified, the neuromuscular actuation

elements of the human may be viewed in Fig. 6 as a two loop system. The

inner loop feedback principally involves muscle spindle receptors (and

Golgi) with short pathways directly to the spinal level and back to the

musculature. Viewed from the output end this loop is primarily sensitive to forces, and because of the short neural pathways the time lags

of information flow are small.

can, therefore, be quite high.

Retinal and

Central

Equalization

The effective bandwidth of this loop

The second or outer loop includes joint

Muscle/Manipulator

r i Spinal Cord Dynamics

LManipulator

I ' I rL--I

I AI l Spindle Feedback

s I

[ ._J •

Effective

Joint Sensor Feedback

i

Hj = Kj e-TJ'

/

Figure 6. Neuromuscular Subsystems for Free-Moving and Pressure

Manipulators and Central Equalization for Rate Dynamics

13
and other (e.g., peripheral vision) receptors as major feedback

elements. Their neural pathways, and associated delays, are longer,

leading to a lower outer loop bandwidth. Note that this simplified

model does not include dynamic elements associated with the manipulator

itself and the form of the retinal and central equalization term in

Fig. 6 is appropriate for controlled element dynamics of the form Yc =

Kc/s. If the controlled element differs appreciably from this ideal the

pilot's central equalization can include an additional lead, lag, or

both.

Figure 7 (from Ref.2) presents root locus and Bode-Siggy plots showing typical closed-loop N.M. actuation subsystem root migrations for

successive closures of the Fig. 6 inner (spindle, force) and outer

(joint sensor, displacement) feedback loops. Increasing force or tension in the inner loop increases the frequency but decreases the damping

of the complex modeand decreases the inverse time constant of the first

order mode. Closure of the joint displacement loop results in a slight

decrease in the complex modefrequency and further decreases modedamping. Peaking of the closed-loop N.M. mode increases with increasing

joint displacement feedback gain.

In isometric (force-stick) manipulator conditions, there is little

or no joint movement, so the inner loop elements should be dominant.

With isotonic (free-moving stick) conditions, on the other hand, the

joint receptors are major elements. As already indicated in connection

with Fig. 3 the net difference, in terms of an effective latency, is

approximated at low frequencies by a difference in effective time delay

of about 0.i sec.

If we now employ the Fig. 4 detailed model of the neuromuscular

system (instead of only approximating its phase lag contribution as in

Fig. 3) and superimpose it on the controlled element K/s as in Fig. 8,

we see an open-loop resonant peak in the 2 to 3 Hz frequency range due to

the neuromuscular system. The correspondence of the neuromuscular/limb

quadratic mode numerical values and roll ratchet frequencies is very

unlikely to be a coincidence. So, the neuromuscular/limb mode clearly

should be taken into account in the analysis of roll ratchet. Since the

14
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Line _ ./

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YpYc=('_x (TN)[_N,WN ]

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(_C = KC Kp)

(JU

Figure 8. Bode Amplitude Ratio Plot for Neuromuscular

System Contribution to Roll Ratchet Potential

primary effect is a resonant peak from which "Gain Margin" might be

measured, it is quite apparent that these properties will be of central

importance for high gain situations. While the "Gain Margin" shown in

Fig. 8 indicates the magnitude difference between the mN peak and the 0

dB line, the phase at or near this frequency may differ appreciably from

that required for instability. Thus when the "Gain Margin" shown is

zero only one of the two conditions for instability may be satisfied.

Consequently this is not necessarily a true gain margin in the conventional sense. It does, however, indicate a resonant tendency contributed by the pilot.

C. LIMB ACTUATION/FEEL SYSTEM INTERFACE

In the world of real aircraft, control manipulators (sticks) are not

pure isotonic or isometric devices. Even the F-16 force stick has

limited movement. The reason for this is to provide the pilot with some

predetermined "feel" relating to aircraft state, mission task, comfort,

etc. These feel systems generally involve stick displacement response

16
to force application through a spring-damper (dash-pot) type arrangement. They also incorporate nonlinear elements such as detents, breakouts, stops, etc. Detents and breakout are incorporated around stick

neutral to provide positive centering and minimize inadvertent stick

inputs due to vehicle motion, turbulence, cross axis coupling, etc.

These nonlinearities are kept as small as possible and once overcome,

the stick/feel system can be modelled by 2nd order dynamics.

The limb-actuation, feel system model including the linear elements

is shown in the block diagram of Fig. 9 allowing for the two types of

manipulator sensing (force or displacement)•

1. Force Sensing

Closing the muscle actuation/spindle feedback loop produces the

effective inner loop transfer function

F •

_c

KNM (I + p_p) e-rNMs

2_s s 2

(i+ T_is)[1+-_ + I_) ]

If the manipulator is fixed (no displacement) this simple transfer function then reflects the pilot's N.M. subsystem• However if stick displacement is incorporated through a feel system, the feel dynamics are

encompassed in the joint receptor feedback path and contribute to the

effective N.M. system as shown in Fig. 10. The pilot/manipulator closed

loop transfer function then is of the form

F •

e

(i+T_lS)[1+-T+ 2_s (_)s 2][I+ 2_ss 4--7+ (_s)2]

The feel system contribution is seen to produce a complex pole/zero

dipole• The residue of this dipole depends upon the separation and

hence effective loop gain. Any lag or delay contribution to the effective pilot dynamics (and hence transparency to the pilot) will depend

upon the location of the feel system dynamics with respect to the N.M.

17
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Figure i0. Closed-Loop Manipulator System With Force Sensing

system (ie., lower or higher frequency) and dipole residue. If the feel

dynamics are well above the N.M. mode and the stick displacement small,

then the effective closed-loop dynamics reduce to the third order noted

earlier.

2. Displacement Sensing

When displacement sensing is employed the feel system dynamics

become a part of the forward loop in Fig. 9 and the pilot/manipulator

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S-s -rpS --)e

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D .

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lags. A question remains however as to whether this lag is partially or

totally compensated for by a change in the pilot's central equalization

(lead generation or decrease in latency) and therefore is transparent to

the pilot. If this should be the case then there is room for argument

19
that the flying qualities time delay criterion should not be referenced

to force application but rather to manipulator displacement. On the

other hand, past experimentation (e.g., Fig. 4) has indicated an

increase in pilot central latency with pure displacement (inertia)

manipulators. This would portend a compounding effect for pilot/feelsystem lags with displacement sensing. Thus the need for experimentation to resolve as many of these issues as possible.

20
SECTION III

EXPERIMENT GOALS AND SETUP

A. EXPERIMENT SETUP

The experimental setup is shown in block diagram form in Fig. Ii. For

the greater part it is identical to that employed in the Ref. II experiments. A roll tracking task is used in which the pilot matches the bank

angle of his controlled element with that of a "target" having pseudorandom rolling motions. The random motions are obtained via a computer

generated sum of sine waves. The error is presented as an Attitude

Director Indicator (ADI) type display with a rolling reference line

(target) implemented on a circular CRT (Fig. 12). A grid was available

in the background for estimating angular deviation of the reference line

and a prominant aircraft symbol was available in the foreground. The

pilot attempts to null the error by applying force to the manipulator,

the output of which becomes the command to the controlled element, Yc"

The controlled element consisted of a classical airframe roll transfer function (Y_) with time delay and an optional command filter (YcF)

of the form shown in Fig. 11. Y_ approximates a high gain roll rate

command system. The lag parameter (T) may be considered to be the

effective roll subsidence time constant or a first order flight control

system prefilter (between the pilot's stick command and the flight control system), whichever is larger. For very small values of T the pure

time delay may be a realistic approximation to digital flight control

system sample and hold dynamics. More generally it is a low frequency

approximation for all the high frequency lags in the system which are

not covered by the first order lag T. The parameter values for T and T

used in the experiment are generally consistent with values that would

be present in systems designed to be Level I on the basis of flying

qualities specifications. Thus, the parameter values used for Y_, in

the main, should produce excellent effective controlled elements providing the gain is appropriately adjusted.

21
+

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Analyzer I

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Yc : Y_YcF

Kc e-rs

Y'/': s(Ts+l)

2

(:JCF

r : 0--'-0.1 sec

T : O-,--0.4sec

YCF =

S 2+ 2 _CF(JJCF S + (SJC2F

_CF: 0.7

(_CF= O, 14, 26 rad/sec

(optional)

Figure ii. Experimental Setup

Figure 12. Display Format

22
An optional response command2nd order prefilter was incorporated

for someof the experiments so that the manipulator feel system and commandprefilter dynamic lag contributions could be interchanged. In this

manner it was possible to investigate pilot sensitivity to location of

lags within the forward loop path.

The manipulators were McFaddenforce loader systems used in many

aircraft research and development simulations. Two stick types were

employed (Table I). One was a right hand mounted side stick. Three

different force gradients which fairly well encompassedthose used in

other experiments (Refs. 15, 16) were evaluated with this manipulator at

one set of feel dynamics. The second stick was a conventional fighter

type center stick. Two different sets of feel system dynamic characteristics were evaluated with this stick at one representative force

gradient. Both displacement and force sensing were available as output

from the stick electronics. Plots and time traces of the feel characteristics together with details of their mounting are given in

Appendix A.

TABLEI. STICK/FEELSYSTEMCONFIGURATIONS:

SIDESTICK YFS 1 YFS 2 YFS 3

KFS (#/in) 14 7.5 3.75

(#/deg) 1.22 0.65 0.33

(deg/#) 0.82 1.53 3.1

_FS (rad/sec) 31.4 22.4 18.0

_FS 0.7 0.6 0.6

Breakout (#) 0.75 >

CENTERSTICK YFS 4 YFS 5

KFS (#/in) 4 4

_FS 26 14

_FS 0.7 0.7

Breakout (#) 1.0 1.0

23
Analog signals from the manipulator sensors were passed through to

the controlled element, resident in an analog computer. The controlled

element roll response was in turn passed through an A to D converter to

a digital computer where YpYc describing functions and various performance measures were cor_uted using STI's Describing Function Analyser

(DFA) program. The computations were essentially on-line and printed

out at the conclusion of each run. Some 380 data runs were accomplished

which provided a suitable data base from which to determine or identify

the various interactions of interest.

B. DISTURBANCE (TARGET) I_PUT FUNCTION

Key aspects of the experimental program were that the describing

function measurements must cover the limb-neuromuscular peaking frequency region, and the forcing functions should emphasize good data in

the neuromuscular subsystem region. The experimental runs were accomplished using the summation of sine waves presented in Table 2. This

input function was scaled to give an rms disturbance of 18.60 deg roll

ove_ a run length of approximately 27 sec. There was an approximately

II sec "warm-up" and 1.5 sec "cool-down" period at the beginning and

end, respectively, of each run where no data is taken by the DFA. This

is necessary to allow for desirable initial conditions to be met before

starting the describing function analysis and to assure the run length

covered a full cycle of the lowest frequency sine wave.

TABLE 2. ROLL TRACKING FORCING FUNCTION

Sine Wave (i)

Frequency (m i)

Amplitude (A i)

Relative

Amplitude

1 2

0.467 0.701

15.2 15.2

1 1

3 4 5

1.17 1.87 3.51

15.2 7.6 3.04

1 0.5 0.2

IA i cos _i t (deg)

7.01

0.76

0.05

7 8 9

11/2

0.38

0.025

14.0 18.7

0.228 0.152

0.015 0.01

24
C. PROTOCOL

The primary subject (Subject A) completed the whole test matrix. A

second subject (Subject B) was employed on approximately half of the

configurations to provide a set of check results. Subject A is a flying

qualities engineer with previou3 experience as a military pilot. Subject B is a flying qualities engineer and licensed pilot with experience

as a test pilot in both simulator and actual flight situations (light

and heavy fixed wing plus rotary wing aircraft).

The subjects were required to perform a minimum of three runs at

each configuration before allocating a pilot rating. The Cooper-Harper

Pilot Pating (CHPR) scale shown in Fig. 13 was used. Desired performance level required extended time periods of less than I to 2 deg roll

error and peak errors during reversals of less than [I deg. Adequate

performance required short periods of 1 to 2 deg tracking error and peak

errors on reversals of less than 22 deg. The background grid on the

display (Fig. 12) was used to judge these angles.

The subjects were instructed to maintain tracking error as small as

possible throughout the entire run,

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26
SECTION IV

DISPLACEMENT SIDESTICK AND FEEL SYSTEM

One of the key findings of the previous investigation (Ref. Ii) was

that the combination of a stiff sidestick manipulator and force sensing

could contribute to roll ratchet tendency. A principal goal of this

continuing investigation is to determine if the use of displacement

sensing alters the tendency to roll ratchet, and if so, to what extent.

Another goal is to identify handling performance metrics for the roll

control task and to determine the influence of sidestick manipulator,

feel system, and response command characteristics on task performance.

Roll ratchet tendency was related in Ref. Ii to the open-loop YpYc

frequency response amplitude ratio and phase lag at the pilot's neuromuscular (refered to hereafter as N.M.) subsystem dynamic mode

frequency. Therefore, we will first look at the influence of position

sensing on N.M., peaking tendency, and the system phase lag at that

frequency. Results will be compared with those obtained with force

sensing in Ref. ii.

Attention will then be focused on roll tracking with various feel

system force/displacement gradient and response command gains in order

to identify the influence of these design parameters on tracking performance and pilot acceptance.

A. INFLUENCE ON PILOT'S NEUROMUSCULAR SYSTEM MODE

A simple block diagram representation of the effective controlled

element is shown below.

Feel Command Ve hicle

System Gain Dynamics

[0.6,22.4] KFS Kc s (T R s ÷ I)

27
For this series the experimental matrix encompassed:

o Feel gradient, KFS = 0.33, 0.65 and 1.22 Ib/deg

o Command gain, Kc = 5 to 20 deg/sec/ib

o Time delay, T = 0, 0.067, and 0. i sec

• Roll subsidence, TR = O, 0.2, and 0.4 sec

A total of 179 data runs were required to complete this series.

1. Amplitude RatLo Peaking

Amplitude ratio peaking of the N.M. mode was determined by drawing

straight line asymptotes of the effective controlled element (Yc) and

moving the asymptote up or down to fit the YpYc describing function data

points and satisfy the crossover model in the region of gain crossover

as shown in Fig. 14. The amplitude deviation from the asymptote in the

region of II to 19 rad/sec is then taken as the pilot's N.M. mode contribution.

Average amplitude ratio deviations obtained with an ideal K/s roll

task across all command gains and two different feel system gradients

(0.65 and 1.22 ib/deg) are shown in Fig. 15. The range of peak readings

obtained are indicated by the bars and the number of runs by the number

above the bar. The neuromuscular subsystem natural frequency is seen to

be in the vicinity of 14 rad/sec for the two test subjects. The peak

magnitude and form seems to be independent of the feel system gradient.

The lone point which disagrees with this observation (1.22 ib/deg

gradient at 19 rad/ sec) was not thought to be significant due to the

relatively small number of runs made with this configuration.

The latter finding is reinforced by the data of Fig. 16 which show

the effects of varying feel system gradient and controlled element time

delay. In essentially, all cases, the neuromuscular mode peak frequency

remains at approximately 14 rad/sec. However, there are some changes in

peak amplitude. Increasing stick stiffness and/or controlled element

time delay may be seen to produce a weak increase in the peaking

tendency.

28
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Figure 14. Example Bode Plot With K/S Fit to Obtain AAR

29
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Asymptote Due to Neuromuscular Mode

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w (rad/sec) w (rod/sec)

r=O r = 0.067 r=O.l

Figure 16. Influence of Feel Gradient and Effective Time

Delay on Neuromuscular Mode Peaking With

Stick Displacement Sensing

31
The influence of force vs displacement sensing on peaking was

checked with the 0.65 Ib/deg feel gradient over the range of time

delays. These results are also shown in Fig. 16. As expected from

theory, an increase in peaking with displacement sensing can be seen,

however the difference is not nearly so pronounced as in the Fig. 4

cases of pure force vs. free moving manipulators.

2. Roll Ratchet Tendency

The Ref. II experiment indicated that neuromuscular peaking could

lead to a roll resonance (ratchet) at the N.M. frequency if the 0 db

gain line of the YpYc Bode plot cuts through the N.M. mode amplitude

peak while the phase angle is very close to 180 deg. To satisfy the

phase criteria, a first order correction to the measured describing

function phase was required in Ref. II to account for the difference

between no motion and the rapid rolling motion in flight where the

phenomenon actually occurs. The correction is derived from Ref. 4 where

it is shown the pilot's angular motion sensing neurological apparatus

acts very much like a rate gyro inner loop in the frequency range near

and slightly above crossover. This inner loop, present when super

threshold rolling velocities are imposed on the pilot, has the effect of

reducing the effective time lags in the pilot's visual input/manipulator

output response. The reduction can be as much as 0.i sec from the

fixed-base data. When a phase lead of 0.io is made in Fig. 14 (which in

effect "bends" the 180 deg phase line with respect to the YpYc

describing function data points as shown by the solid line) it is

apparent that the phase angle at the N.M. mode is still considerably

greater than 180 deg. The rapid phase roll-off in YpYc is due to the

feel system 2nd order mode lag contribution (in this case tat 22.4

rad/sec) and to the low damped N.M. mode itself. Thus, in spite of the

increased N.M. peaking tendency, shown in Figs. 15, 16, we would expect

no roll ratchet tendency with this stick displacement sensing responsecommand system. This was indeed the case with all such configurations

investigated.

32
3. Comparison Between Displacement And Force Sensing

It was noted above that there is slightly greater peaking of the

N.M. mode with displacement vs force stick sensing. Further comparison

between the two is made in Figs. 17, 18, 19. Figure 17a presents the

1.22 and 0.65 ib/deg feel gradient displacement sensing cases of Fig. 16

compared to the almost identical configurations from the Ref. Ii experiment with stick force sensing (17b). While the N.M. mode peaking

generally may be greater with displacement sensing the difference is

negligible.

Figure 18 allows direct comparison of typical describing function

data for the same controlled element dynamics with displacement vs force

sensing being the only variables. The most notable difference is the

more rapid YpYc phase roll-off above roughly 5 rad/sec with displacement

sensing (as predicted by Eq. 3) which reduces the roll ratchet tendency

as noted above. But, it is also worth noting, that this additional

phase lag results in the open-loop YpYc phase data points crossing the

180 deg line at a lower frequency. This could result in the pilot being

forced to reduce his gain in order to maintain adequate gain and phase

margins, thereby resulting in a slightly lower control bandwidth. This

decrease in bandwidth would in turn result in degraded tracking performance. It could also result in the pilot generating slightly more lead

to offset this added lag in the region of intended crossover, thereby

maintaining control bandwidth and hence performance at the expense of

extra pilot co_ensation.

Figure 19 presents survey plots of predicted roll ratchet tendency

across different K c and TR combinations for the sidestick manipulator/

feel configurations investigated in this and the Re f. II force sensing

experiments. The open symbols represent no roll ratchet tendency, the

filled symbols reflect roll ratchet. The dashed line boundary separating ratchet and no ratchet is derived from the Ref. 16 NT-33 flight test

data also shown in Fig. 19. It appears from this data that the boundary

could either be moved to the right by a considerable amount are liminated entirely by use of displacement sensing.

33
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I0 04 02 017 0.I TR (sec)

[ , i , _ [ , J, , , I

I 2 3 4 5 6 7 8 9 I0 lIT R(l/sec)

b) Force (Ref II)

Legend :

C) No Ratchet I Ref.]6 Flight

• Ratchet I

• Ratchet- Ref. 17 Flight

O Peak < Gain Line I Simulation

• Peak ; Gain Line I

S = Small Displacement Slick

L = Large Displacement Stick

F = Fixed Stick

Figure 19. Roll Ratchet Tendency With Variation of Roll Lag

and Command Gain; Displacement vs Force Sensing

36
B. INFLUENCE ON CLOSED-LOOP TRACXING PERFORMANCE

Tracking performance metrics obtained directly from the describing

function measures are rms roll error (o e) and crossover frequency (_c).

These are inversely related, in that, generally for systems which obey

the crossover law, higher crossovers result in lower rms error. Another

measure obtained by the analysis software is rms manipulator amplitude

(stick displacement or force). This measure is inversely related to

command gain (Kc). Combined, the rms error and manipulator effort constitute a measure of tracking workload. A final measure is pilot rating

(CHPR), which is the pilot's opinion of the overall system (pilot and

controlled element) closed-loop performance and task workload.

As noted previously, there were 3 runs per configuration, with the

configurations being presented to the subjects in a random order. The

subjects were highly trained in the tracking task, having accomplished

several hundred runs each. However, the first run for each configuration often resulted in the subject adapting to the new effective controlled element dynamics by varying his control technique (gain, lead,

lag) during the course of the run. This is evidenced by scatter in the

first run data points. The second and third run data would then be

quite consistant. Thus, only data from the last two runs for each configuration are included in the results reported herein.

1. Bandwidth

The performance measure for bandwidth is crossover frequency (_c).

It was found in the Ref. II stick force sensing simulation that tracking

bandwidth remained relatively constant for command gains above i0 deg/

sec/lb., but decreased linearly for gains below I0 deg/sec/Ib.+ A similar trend was observed with stick displacement sensing. Fig. 20 presents a plot of crossover frequency versus command gain obtained for the

1.22 and 0.65 Ib/deg feel gradient sticks over all values of • and TR.

Boundaries for data from the Ref. ii experiments are also shown. The

data from Ref. ii (force sensing) and the new displacement sensing

results show so little difference that all the data points at each

37
8

7

6

a

" :3

(J

3 2

I

0

0

z" T R Pos. Force

o o C5 O"

0.067 0 C) •

- o.,o 0 g"

0.067 0.2 []

0.067 0.4 A ,-Ref. II

Subject _ /

Force Sensing

[]

[]

I I I I I

5 Z5 I0 15 20

Kc(deg/sec/Ib)

a) L22 Ib/deg Stick

8

7

•-. 6

5

-o 4 O

L_

% 3

3 2

0

0

m

m

m

D

B

B

m

B

o @

Ref. II Force Sensing

t_ rn

I I I I I

5 7.5 10 15 20

Kc(deg/sec/Ib)

b) 0.65 Ib/deg Stick

Figure 20. Roll Tracking Crossover vs Command Gain

With Displacement Sensing

38
command gain have been averaged in Fig. 21. This represents a summation

of all values of T and roll subsidence, the 1.22 and 0.65 ib/deg feel

gradients, and force and displacement command sensing. The data means

are represented by the open symbol and the la range by the lines. The

data show quite remarkable consistancy. It may be noted at the lowest

values of K c that the mean crossover frequency is slightly higher for

force sensing than for displacement sensing. This does not hold true

for higher command gains. However, it was found that the values for the

1.22 ib/deg feel gradient generally lie within the +io bound from the

overall mean and the values for the 0.65 Ib/deg feel gradient (displacement and force) lie within the -io region. Thus, the stiffer feel gradient results in a slight increase in closed-loop bandwidth, but there

is no significant difference between displacement and force command

sensing.

m

7-

_-_ 6 --

5--

O

U

3 2-

I -

O

O

GENERALLY:

Mean value for KFS = 1.22 lies within + la- value

Mean value for KFS =0.65 lies within - Io- value

SUMMATION OF:

KFS = 1.22 and 0.65 Ib/deg

All values of r and T

Force and Displacement sensing

mean

(_ (_ Force and Displacement

Displacement

I I I I

5 7.5 I0 15

Kc(deg/sec/Ib)

Figure 21. Roll Tracking Crossover vs Command Gain

With Force and Displacement Sensing

I

20

39
It has been established in this and the previous experiment that the

crossover "droop" at low command gain is due principally to a physical

inability on the part of the pilot to overcome the high stick forces and

move the manipulator far enough to generate the required correction

signal in a timely manner. There is also evidence from pilot commentary

that at the high end of the gain scale the problem of PIO begins to

arise and any further increase in command gain may result in a decrease

in crossover. On this basis, it would appear that command gains between

10 and 20 deg/sec/ib may be optimum (subject, of course, to flight verification).

2. Tracking Error And Pilot Effort

A measure of pilot effort in achieving either desired or best

obtainable performance was selected to be the product of rms tracking

error (o e) and rms manipulator displacement (a c) or force when using

force sensing. This product of performance and effort might be considered a measure of workload. It will be shown later that this parameter correlates well with Cooper-Harper pilot ratings (CHPR) obtained.

Figure 22 presents a plot of the workload parameter (o e x Oc) versus the

reciprocal of the product of crossover (mc) and command gain (Kc). As

should be expected from the crossover model (Ref. 4), this indicates

good correlation between the parameters. Decreasing crossover and/or

command gain results in increasing workload -- or vice-versa. However,

the data also tend to indicate a relative insensitivity of workload,

over the range of i/(gain x crossover) between about 0.75 and 2 ib/(rad/

sec) 2. The data points within this region generally reflect a command

gain of 20 deg/ sec/ib and this may suggest that this gain is sufficiently high that tracking effort becomes minimal and other factors such

as higher frequency time delays begin to dominate.

40
60

58

40

-o

_- 30,

b_

• 20, bo

I0,

O

Subjects A and B

TR:O, 0.2 13 ..'-'_"'_

r_ = 0, --'-- 0.1

_,,a,r 13

0 | -"_"_" I +

13 2

.__ - = 0.82 P

|,,

I I I I I I I

2 4 6

I/Kc wc ) Ib/(r/s) 2

Figure 22. Relationship Between Tracking Performance

and Loop Closure Parameters

C. INFLUENCE ON PILOT RATING

I. Co_mandGainAnd Feel Gradient Variation

Previous plots indicated the influence of command gain (K c) on

tracking bandwidth and workload. It is of interest to determine if and

how this may correlate with the Cooper-Harper pilot ratings obtained.

Figure 23 presents a crossplot of CHPR and command gain for the two best

roll subsidence and the three feel system configurations. These data

represent only one pilot subject because only the one was given the

entire matrix of configurations. However, it will be shown subsequently

that ratings were quite consistent between the two subjects.

The trend with CHPR reflects that of crossover (Fig. 20) in that

there is a general degradation with decreasing command gain. Command

gains less than I0 deg/sec/ib were unacceptable for anything other than

a K/s equivalent aircraft. K c = I0 deg/sec/ib was marginal depending

upon feel system gradient and roll subsidence time constant. There was

41
-0.067s

Kce

Yc -

S(TRS+I)

Sidestick- Displocement Sensing

Subject-A

CHPR

7-

6-

5-

4-

3-

2-

I0

Feel Gradient:

(Ib/deg)

1.22 (_ _] 0.33 Oscillatory

\ \ PIO Tendency

0.65 Z_\\\ \\\ (_ "K_

0.65 _1_ _,, _ITJ (_ '_

%.

"""X ..... --)( T R = 0

I I I I I

5 7.5 I0 15 20

Kc ( deg/sec/Ib)

Figure 23. Variation of CHPR With Position Sensing

Command Gain and Feel Gradient

a general complaint about the stiffest feel gradient (1.22 Ib/deg) that

forces were too high and tiring, especially at gains below i0 deg/sec/ib

and at the higher gains there was some tendency to low frequency PIO

which degraded ratings significantly. The lowest feel gradient

(0.33 ib/deg) seemed to induce larger stick deflections and required a

high command gain to avoid hitting the stick travel limits in the larger

maneuvers. It therefore, gave the impression of inadequate control

authority, and was downgraded accordingly. Also, this configuration

would not be expected to be acceptable in flight, because it would be

too difficult to avoid inadvertent stick inputs in turbulence. The best

compromise feel gradient was found to be 0.65 lb/deg.

42
There was a significant improvement in CHPR in going from a TR of

0.2 to zero (K/s) with the 0.65 ib/deg stick gradient and higher command

gains. This may reflect a requirement for pilot generated lead and/or a

tendency to oscillation at the higher T R.

2. Comparison With NT-33 Flight Data

It should be recalled at this point that the response command gradient gain (K c) generally is not a fixed parameter in a flight control

system, but is changed as a function of stick force or displacement

input. A low response gain is provided around stick neutral to avoid

inadvertent command inputs and/or over control in situations requiring

small precise command inputs. For larger stick inputs as required for

gross maneuvering, the response command gradient is increased to obtain

faster motion response with lower stick force or displacement. In most

instances two or more discrete gradient are employed to cover the range

of flight tasks from air refueling or air-to-air tracking to gross

maneuvers as in air combat. The Ref. 16 flight test in the NT-33 investigated several response command gradients (Fig. 24) in air-to-air

tracking and gross maneuvering tasks.

f..

12

10

4

2 : : ,

Breakout = l.O Ib

0 ' : i

0 20 40 60 80 lO0 liO 140 160

_, deg/sec (Steady State)

Figure 24. Control Force-Response Gains, Up-and-Away Flight (Ref. 16)

43
In general the first gradient was dominant in the tracking task and

the second gradient was dominant in the gross maneuvers. One objective

of the test was to select the best gradients. The flight test also

investigated three different feel system configurations: fixed stick,

"small" deflection (1.3 ib/deg measured at the center of the stick

grip), and "large" deflection (0.7 Ib/deg).

Figure 25 presents a plot of CHPR vs. K c made up from data for the

1.3 ib/deg stick in Ref. 16. Across the bottom of the figure is a

superimposed guide to the NT-33 first and second command gradients and

associated evaluation tasks. The average pilot rating is identified (X)

along with the range obtained. Also shown in the figure are the ratings

obtained with the 1.22 ib/deg feel configuration (0) in this simulation.

For the region corresponding to the NT-33 ist gradient (tracking) there

is excellent agreement between the two. For command gains appropriate

for the 2rid gradient, the CHPR for our tracking task was downgraded

because of a tendency to PIO. The existance of this PIO tendency in the

tracking task is probably related to the relatively high bandwidth,

closed-loop nature of the task when compared with a relatively open-loop

gross maneuvering task. The "target" which was being tracked did

produce some gross maneuvers (60 deg right to 60 deg left and vice

versa) but the subjects were not required to rate these separately.

Figure 26 presents similar data for the NT-33 0.7 ib/deg ([7) and

our 0.65 ib/deg feel configurations (/_). Here the correlation in the

Ist gradient range (only two data points) is not good because in-flight

tracking produced PIO tendencies (Level If) whereas in the simulation at

similar command gains the task was rated Level I. For the gross maneuvering range the correlation is much better (Level I for both). This

implies that our subject pilot was able to track the gross and small

maneuvers well with no tendency to PIO.

The data points for the NT-33 1.3 ib/deg (X) are also shown in

Fig. 26 and interestingly match our simulation results for the 0.65 ib/

deg feel system quite well over the command gain range from 10 to

20 deg/sec/ib. However, the match below 10 deg/sec/Ib is not as good as

in Fig. 25.

44
CHPR

B

7-

6-

5-

4-

3-

2-

I -

0

Gradients at center of stick grip

(_) 1.22 Ib/deg simulation

X 1.30 Ib/deg T-35 (Ref. 16)

(__-- Simulation ; All cases TR = 2 sac

I I I I I

5 I0 15 20 25 Kc(deg/sec/Ib)

VH H M L H M L

NT-33 I st Gradient --.l---

(tracking)

2 nd Gradient

(gross maneuvering)

(Ref. 16)

Figure 25. Comparison of CHPR vs Command Gain for Flight

and Lab With Feel Gradient 1.2 - 1.3 ib/deg

45
8

7

6

CHPR

5

4

3

2

I

0

Gradients at center of stick grip

A 0.65 Ib/deg simulation

_ X 1.30 Ib/deg T-33 t Ref.16

n 0.70,b/ egT-33)

All case.____sT._RR_--2_se__.c_

J_ (OSC and PIO)

Flight -/ J_

m

I I I I I

5 I0 15 20 25 K c(deg/sec/Ib)

VH H M L H M L

NT-33 Ist Gradient

(tracking)

2 nd Gradient

(gross maneuvering )

(Ref. 16)

Figure 26. Comparison of CHPR vs Command Gain for Flight

and Lab with Feel Gradient 0.65 - 0.7 Ib/deg

46
There are several potential design lessons in these results:

The fixed base simulation indicated the 0.65 ib/

deg feel to be the best overall force/

displacement gradient but flight results show the

stiffer 1.3 ib/deg feel to be the better. The

difference may well be due to motion effects and

minimization of inadvertent inputs in turbulence,

etc., but may also have been related to the particular stick design or installation. On the

other hand, a feel gradient somewhere between the

two may offer a better compromise. Such factors

should be checked in flight.

Flying quality rating degrades very rapidly as

response command initial gradient is decreased

below i0 deg/sec/ib. For Level 1 rating the

initial gradient should be above i0 deg/sec/ib.

Flying quality rating does not change appreciably

for 2nd gradient command gains appropriate for

gross maneuvering although there is flight evidence from Figs. 25, 26 that K = 20 deg/sec/Ib

or higher is to be preferred wheCn the stick force

gradient is high.

D. ADDITION OF 2ND ORDER COMMAND FILTER LAG

Using the "best" feel gradient (0.65 Ib/deg) and command gains (i0

and 20 deg/sec/ib), a set of 57 additional runs was made with a second

order filter inserted in the command path. The basic purpose was to

determine the effect on tracking performance and pilot rating of such a

command filter tuned to the pilot's N.M. subsystem. For this set of

runs the effective controlled element was as shown by the block diagram

of Fig. 27. Both displacement and force stick sensing were employed.

Thus, four configurations were involved; the force to roll response

equivalent time delay of the four varies from 0.033 sec to 0.207 sec.

I. Influence of Effective Time Delay On Pilot Rating

Figure 28 presents a summary of pilot ratings versus effective time

delay across the four configurations and two command gains (I0 and 20

deg/sec/Ib). Ratings by the two subject pilots are within one rating

point of each other for each gain and all but one configuration (DI).

47
Command

Gain

F __L_ c F

Fee System ]°Lt

[o.6,22.4]I

re = 0.074

Filter

I

ro.7, 14]

re =0.1

s(O.15s+l) e-°°33s >

Configuration Sensing Filter _ .r "re

(sec) (sec)

Ai Force OUT 0.055 0.053

Bi Disp. OUT 0.074 + 0.055 0.107

Ci Force IN 0.1 + 0.0:55 0.155

DI Disp. IN 0.074 + 0.1 + 0,055 0.;>07

Figure 27. Effective Controlled Element Dynamics for

_nd Order Command _ilter Exneriment _',TithSidestick

48
I0

9

8

7

CHPR

6

5

4

5

2

I

0

1":'i Subject A

(_) Subject B

Open -K c = 20 deg/sec/Ib

Solid-Kc- I0 deg/sec/Ib

KFS = 0.65 Ib/deg

m

I

----- []3

Ref. 18 (F-8)

0

/HI

/f

o

Stress

•-.- LO Stress

AI Bi Cj DI

I I I

0.1 0.15 0.2

re(sec)

0.05

Configuration

Figure 28. Influence of Effective Time Delay and

Command Cain on CHPR 0.65 Ib/deg Sidestick)

49
The ratings show a general degradation with increasing effective time

delay similar to that reported in Ref. 18 for a low stress task. However, it will be noted that the 20 deg/sec/Ib gain fared better than the

I0 deg/sec/ib gain by about 1.5 to 2 rating points across all configurations. This trend is opposite that shown in Fig. 23 but is supported by

the flight data of Fig. 26. Thus , it is suspected that the K c = 20

(0.65 ib/deg and T R = 0.2) data run in Fig. 23 may have been adversly

influenced by some unknown factor.

As a further check on the CHPR differential for the two command

gains, a plot of effective workload (Oe x Oc) versus effective T is

shown in Fig. 29 for the four configurations. This also shows a distinct separation for the two gains. This might be expected since a

doubling in gain should reduce the rms stick input by roughly a factor

of two and this is the case here. If the workload parameter is normalized with command gain as shown in Fig. 30 the resulting error measure

is shown to be independent of command gain. Thus, the CHPR rating differential between the command gains can be related to amplitude of stick

activity (effort) required. (It also becomes apparent that the pilots

were employing different error criteria in performing their tracking

task. Subject A consistently tracked tighter than Subject B).

The influence of the effective time delay now becomes more clear in

Fig. 30. Displacement sensing added a small increment in tracking error

to that obtained with force sensing, and the 2nd order filter added a

significant increase. If we now eliminate the effect of command gain on

CHPR in Fig. 28 and concentrate on change in CHPR with increasing

(referenced to the initial ratings for each subject in each case) we

obtain the plot of Fig. 31. The end result is that addition of the

second order filter at the pilot's N.M. mode frequency to the force

sensing stick produced essentially the same decrement in rating as did

changing to displacement sensing but without the filter (increase in

is almost the same in both cases). Adding the filter to the displacement sensed input degraded the rating by about 1 to 1.5 rating points.

Obviously, since the use of displacement sensing alone essentially

eliminates any tendency to roll ratchet (Fig. 19), there is little need

for this added lag.

50
50

20

me o"c

(deg-lb)

IO

O

[] Subject A

G Subject B

KFS = 0.65 Ib/deg

/

/

/

/

•J

/

/

/

K c = IO deg/sec/Ib

K c = 20 deg/sec/Ib

[] E3

At Bi Ci

Force Displ. Force

I I I

0.05 0.1 0.15

Di

Displ.

I

0.2

re(sec)

Figure 29. Influence of Effective Time Delay and Command Gain

on Roll Tracking Performance (0.65 ib/deg Sidestick)

15

Kclo

Ge GClo x --KC2o

ond IO

Cre O"Cz O

( I b 2 sec )

5

[] Subject A

Q Subject B

Open- Kc = 20 deg/sec

Solid- Kc = IO deg/sec

No Filter

[]

With Filter

• .._.---C]

[]

[]

At Bi Cr De

Force Displ. Force Displ.

O I

0.05

I I I

0.1 0.15 0.2

re(sec )

Figure 30. Command Gain Normalized Tracking Performance Measures

Variation With Effective Time Delay (0.65 ib/deg Sidestick)

51
2. Comparison With MIL-Spec Time Delay Criteria

There is some controversy over whether the allowable time delay criterion of MIL-F-8785C should be referenced from the force input to the

manipulator or the resulting manipulator deflection (Ref. 9). Figure 32

presents the average values of rating degradation from Fig. 31 for each

configuration but with an initial CHPR rating selected to be the average

of the two ratings given for • = 0.033 and K c = 20 in Fig. 28. This

configuration was chosen as it is the baseline airframe which received

Level I ratings i.e., Kc was optimal and not a factor in the ratings.

The data presented in Fig. 32a, where the feel system lag is included in

the calculation of the effective time delay, seems to support an

increase in the allowable Levels I and II time delays from the present

MIL-F-8785C limits. In Fig. 32b the feel system lag has been excluded

from the effective controlled element time delay. In this case the data

points lie within the appropriate boundaries.

This limited set of data supports the contention that the allowable

time delay criterion should not require inclusion of the feel system

dynamics. If the feel system dynamics are included, it supports an

increase in the allowable Levels I and II time delays from their present

(MIL-F-8785C) limits. It is also worth noting that this data was

obtained with, and is therefore valid for, a sidestick manipulator.

The goals of this experiment were to determine whether stick displacement sensing decreases the tendency to neuromuscular (NM) system

related roll ratchet and to assess the influence of displacement sensing

and related feel system and command filter dynamics on tracking performance and pilot acceptance. It has been shown that stick displacement

sensing decreases the tendency to NM mode induced roll ratchet due to

the additional forward loop phase lag introduced by the feel system

dynamics. This additional lag, in effect, serves the same smoothing

purpose as the command prefilter lag required with stick force sensing

configurations.

52
5

A 2

CHPR

I

0

I]] Subject A

Q Subject B

Open - K c = 20 deg/sec

Solid-K c = I0 deg/sec

AI

I

[3

9_

Bi Ci Di

I I I

0.05 0.1 0.15 0.2

"re (sec)

Figure 31. Command Gain Normalized CHPR Variation With

Effective Time Delay (0.65 ib/deg Sidestick)

I0

9

8

7

CHPR

6

5

4

3

2

I

0

MIL-F-8785C Boundaries

LEVEL m

LEVEL TT

0

AI

LEVEL I

lo

]BI

I

I

0.1

0

Cl

Te(Sec)

I

I

I I

I I

I

I I

i I

I I

0.2 0.25

Feel System Included

m_mm_m UmB,

umnmm_mnm

- D I

- 0

- B, I o

- I

I

0.1

---'-1

I

I

I I

I I

I I

I I

I I

I

I i

0.2 0.25

Te(Sec)

Feel System Excluded

Figure 32. Comparison of CHPR With MIL-F-8785C Effective

Time Delay Criteria (0.65 ib/deg Sidestick)

53
It was found that displacement sensing does not result in appreciable change in the NMmodefrequency (about 14 rad/sec) and amplitude

peaking from that noted previously with force sensing. As with force

sensing, the peak amplitudes tended to increase with stiffer feel gradients (0.65 to 1.22 ib/deg) and increasing effective time delay (up to

0.I sec). However, this is not as significant with displacement sensing

comparedto force sensing because of the concomitant phase lag contribution of the feel system in the frequency region near the NMmode.

As with force sensing, tracking bandwidth increased as controlled

element commandgain (Kc) increased. The sensitivity to commandgain is

quite high for Kc less than I0 deg/sec/ib, with bandwidth, tracking performance, and pilot acceptance all degrading rapidly as commandgain is

decreased. Performance and pilot acceptance improved slightly with

increasing commandgain in the range Kc = 10 to 20 deg/sec/ib.

Tracking bandwidth, performance, and pilot acceptance were influenced somewhatby feel system force/displacement gradient. A gradient

of 0.65 ib/deg appeared best for this fixed base tracking task. The

0.33 Ib/deg gradient was too weak and allowed excessive stick deflection

when trying to follow large changes in target motion. The 1.22 ib/deg

gradient was considered too stiff and tiring. However, under flight

motion conditions, a gradient between 0.65 and 1.22 Ib/deg might be more

appropriate.

It was found that use of displacement sensing without command prefiltering was rated about the same as force sensing with a second order

command prefilter tuned to attenuate the NM mode. Addition of the

second order prefilter to displacement sensing produced excessive high

frequency lag and degraded handling quality ratings by I to 1.5 rating

points. Again, there is little or no need for such filtering with stick

displacement sensing.

Concerning the -8785C flying qualities specification for allowable

effective time delay, data from this limited experiment with a sidestick manipulator supports the Ref. 9 contention that the feel system

dynamics should not be included in calculation of the effective forward

loop time delay or, if the feel system is included, the -8785C criteria

boundaries should be relaxed.

54
SECTION V

CENTER STICK WITH DISPLACEMENT VS FORCE SENSING

A NASA DFRF flight test program is planned to be accomplished in the

USAF NT-33 aircraft to help resolve several issues concerning the influence of force vs. displacement stick sensing, feel system dynamic lags,

and command prefiltering on the roll control characteristics of highly

augmented aircraft. Among the issues is whether the calculated time

delay for the effective aircraft dynamics as viewed by the pilot should

be referenced to force applied to the manipulator (and hence the feel

lag included) as called for in MIL-F-8785C, or whether the feel system

dynamics are transparent to the pilot (and hence not to be included in

the calculated time delay). Other issues pertain to the influence of

roll mode time constant and stick command gain (or sensitivity) on roll

ratchet. Since this simulation preceeds the flight test, one set of

objectives was to:

o duplicate a few typical configurations to be

tested in flight

preview possible flight results

obtain describing function and other closed-loop

task performance measures which might provide

insight to results

suggest possible improvements to the flight

experiments.

This section will be devoted to a description of the configurations

investigated, the influence of force versus displacement sensing on

closed-loop tracking performance and pilot rating, the influence of

interchanging the feel and command filter lag contributions, and finally

an assessment of results based on overall effective time delay.

A. CONFIGURATIONS INVESTIGATED

A block diagram showing the simulation of the feel system, the force

vs displacement option, the command gain and filter option, the NT-33

55
aileron servo effective time delay and simplified augmentedroll dynamics (roll subsidence of 0.15 sec.) is presented in Fig. 33. The matrix

of configurations investigated is shown in the accompanyingtable. The

configurations are coded in the left column with the subscript 2 to

indicate use of the fighter type center-stick. The type of stick sensing is indicated in the second column. Twosets of feel characteristics

were employed: one with the regular NT-33 system second order dynamics

(26 rad/sec frequency and 0.7 damping ratio) and the second with

14 rad/sec frequency and 0.7 damping (to approximate the X-29 aircraft

lateral feel). The latter actually is 13 rad/sec but this could not be

duplicated in the lab. In both of our cases the force/displacement gradient was 4 ib/in measuredat the top of the stick.

The commandfilter second order lag and feel dynamics were interchanged in four cases (E2 - H2) but with the total lag between force

application to the stick and roll response held constant. Howeverthe

effective time delay between force input and roll response varied with

the sensed stick input as reflected by the effective time delay shownin

the right hand column.

Roll acceleration time responses to step stick force inputs are presented in Fig. 34 for the 26 and 14 rad/sec feel system with no command

filter lag (representing configurations C2 and F2 respectively) and for

the combined feel and filter lags (representing configurations G2 and

H2). These traces demonstrate the large differences in response characteristics produced by the various lags and the quite sluggish response

which results when both the feel system and commandfilter lags are in

the commandpath.

To provide additional comparison between force and displacement

sensing and the influence of low and high frequency feel dynamics, the

second order filter was switched out in four cases (A2 - D2). Finally,

two commandgains were employed (Kc of i0 and 20 deg/sec/ib) for each of

the 8 configurations. Accomplishment of this overall test matrix with

the two test subjects required a total of 141 runs.

56
F

Feel System

_'t KFsKc

KFS = 4 Ib/in.

F

l°

1

 CommonO Filler

Kc = I0 and 20

deg/sec/Ib

Aileron

Servo

-0.0335

e

O_FS _CF r(F-"Sa) "re

Configuration Sensor (rad/sec) (rad/sec) (sec) (sec)

A 2 F 26 - 0.033 0.033

B2 F 14 -- 0.033 0.033

C z D 26 -- 0.033 + 0.055 0.088

D 2 F 14 26 0.033 + 0.055 0.088

E z F 26 14 0.033+ 0.I02 0.135

F 2 D 14 -- 0.033 + O. 102 O. 135

Gz D :>6 14 0.035 + 0.055 + 0.102 0.190

Hz D 14 26 0.033 ÷ 0.102 + 0.055 0.190

Figure 33. Effective Controlled Element Configurations

Employed -- Roll Tracking With Center Stick

57
2

(deg/sec 2 )

@

-Ig

I

I

iII / ,

_,.:...

./",,\_ _FS = 26 rad/sec

"\

7"-'_'",":, '''''''',_--I OJFS = 14- rad/sec

,,.\.,,\ , \..., ,.. _OFs -- 26 rad/sec ' ' " "'" "' S coCF = 14 rad/sec

",. ,

". °.

-.

• ",% ',

i.° ,,.° ',

"°,,.. "°% ',

I J I J I I [.

,2 ,4 ,6 ,8

Time (sec)

I

q Feel p

System

OJFS

q Command

Filter

WCF

Airframeyc

Yc =

-0.033s

Kc(O)2e

(0)(6.67)

Figure 34. Roll Acceleration Time Response to Step Stick Force

58
B. FORCE VERSUS DISPLACEMENT SENSING

For this comparison, attention is focused on closed-loop crossover

frequency (_c), tracking error (Oe), and pilot ratings as metrics to

identify the possible differences in performance resulting from force

and displacement stick usage in closed-loop roll tracking tasks.

1. Tracking Performance

Figure 35 presents a plot of average tracking bandwidth (crossover

frequency, _c) achieved by each subject pilot for the eight configurations and two command gains. Each data point reflects an average taken

over three runs. The configurations are paired to emphasize any difference due to the different stick input sensing methods. It is apparent

from this plot that there is no consistent difference due to the type of

sensing. The only clear trend shown is that Subject A consistently produced a higher crossover which in turn indicates tighter tracking.

This trend is further demonstrated by Fig. 36 which presents averages across all data points for each subject at each command gain. The

symbols reflect the average and the bars reflect the maximum and minimum

values, respectively. This plot also reveals that both subjects tended

to achieve higher crossover with the higher command gain, however the

difference is more pronounced for Subject A.

The influence of the two different types of stick input sensing on

rms tracking error is shown in Fig. 37 where each data point represents

an average of the rms error measures from three tracking runs. The difference obtained is quite striking. Tracking error is consistently

greater with the displacement stick. The degradation varies from

approximately 5 to 25 percent. The influence of command gain on tracking error is small but it is again apparent that subject A is consistently using a tighter tracking technique.

59
U

4

3

3 2

Center Stick

[] Subject A

(_) Subject B

Q,

K c = I0 deg/sec/Ib

Q, (5.

,, "O "o

I I I I I I I I

3

0

3 2 m

Kc = 20 deg/sec/Ib P

/

13- -El //

\ d \

\

I I I I I I I I

26 26 14. 14, 26 26 14 14.

0 0 0 0 14 14 26 26

F D F D F D F D

A2 C2 B2 Fs E2 G2 D2 H2

(A)FS ( tad/sec)

OJCF (rad/sec)

Sensing

Configuration

Figure 35. Influence of Feel and Filter Dynamics on Tracking

Bandwidth With Center Stick Displacement and Force Sensing

6O
L>

G>

IO

O

t=.

t_

3

4

3

Center Stick

Iq Subject A

E) Subject B

4_No. of

runs averaged

0 I I

I0 20

K c ( deg/sec/Ib)

Figure 36. Influence of Command Gain on Tracking Bandwidth

(Center Stick)

2. Pilot Rating

Pilot ratings elicited by the different configurations are summarized in Fig. 38. There are three specific points to be noted from this

plot. First, the ratings given by the two subjects are very consistent,

the average difference being two-thirds of a rating point. Second, the

CHPR ratings tend to reflect tracking error performance and show a preference for force sensing. Third, with the i0 deg/sec/ib command gain

almost all configurations lie in the Level II handling rating range.

When the gain is increased to 20 deg/sec/ib all configurations with

force sensing move into the Level I range.

In summary, the data show a decided preference on the part of both

subjects for force sensing and the 20 deg/sec/ib command gain with this

center stick. This preference would appear to be due to better tracking

performance.

61
_e

(deg)

13

12

II

I0

9

8

7

6

5

Kc : I0 deg/sec/Ib

f)

/

!

I

!

d

l

Ef

P

/

/

I I I I

P

/

/

d

I I

Center Stick

l'q Subject A

Q Subject B

13

1 I

o- e

(deg)

13-

12-

III0-

9-

8-

7-

6-

5

K c = 20 deg/sec/Ib P

$

1

/

d

P (D / p / / /

/ /(_ 1

d d

I I I I I I

26 26 14 14 26 26

0 0 0 0 14 14

(9

/

/

I

d

!

I

I

I

d

I I

14 14 Feel

26 26 Filter

F D F D F D F D

A2 C2 B2 F2 E2 G2 D2 H2

Sensing

Configuration

Figure 37. Influence of Feel and Filter Dynamics on

Tracking Performance With Center Stick Displacement

and Force Sensing

62
CHPR

_

7

6-

5-

4-

5-

2-

Center Stick

[] Subject A

C) Subject B

P

/

0-/-0

I

d

I I I I

K c = I0 deg/sec/Ib

o---_

/

EF

P

I

/

/

I I

LEVEL

LEVEL I

CHPR

_

6-

5-

4-

5"

2-

I

P

I

I

,¢

Kc = 20 deg/sec/Ib

I / I

I /

m Er _

P

/

II

/ /

I I IIQ

&,

I I I I I I I I

26 26 14 14 26 26 14 14

0 0 0 0 14 14 26 26

F D F D F D F D

A2 C2 B2 F2 E2 G2 D2 H2

LEVEL E

LEVEL I

QJFS (rad/sec)

(,oCF (rad/sec)

Sensor

Configurotion

Figure 38. Influence of Feel and Filter Dynamics on CHPR

With Center Stick Displacement and Force Sensing

63
C. INFLUENCE OF DYNAMIC LAG LOCATION

The influence of dynamic lag location on tracking performance and

pilot rating may be observed by comparing configurations D2 and E 2

(force sensing) and configurations G2 and H 2 (displacement sensing) in

which the low and high frequency dynamic lags are interchanged between

the feel and command filter locations while keeping the overall forward

loop path lag constant.

l. Tracking Performance

Figures 39 and 40 show the influence on tracking bandwidth and rms

error, respectively. Separate comparisons are made for force and

displacement sensing at each command gain. Results in both figures

4

u_

_D

(D

t)

3 2

Center Stick

0 Subject A

(_) Subject B

Kc = I0 deg/sec/Ib

..,L3

",,

Kc = 20 deg/sec/Ib

r -- -- _.

EL, _

b

q

"',

""

I I I I I I I I

14 26 14 26 14 26 14 26 Feel

26 14 26 14 26 14 26 14 Filter

De E2 H2 Gz De Ez H2 Gz Configuration

Force Displ. Force Displ. Sensing

Figure 39. Influence of Dynamic Lag Location on Tracking Bandwidth

(Center Stick)

64
cre

(deg)

13

12

II

I0

9

8

7

6

5

Kc = I0 deg/sec/Ib

r ii ___

P

I

I

i I

1 1

' 6

d

/

I

I

I

I

6

Kc = 20 deg/sec/Ib

P

I

1

1

I

®'

I

0

.El

s

I I I I I I

14 26 14 26 14 26

26 14 26 14 26 14

Center Stick

E] Subject A

(_) Subject B

I I

14 26 Feel

26 14 Filter

D2 E2 H2 G2 D2 E2 H2 Ga

Force Displ. Force Displ.

Configuration

Sensing

Figure 40. Influence of Dynamic Lag Location on Tracking Error

(Center Stick)

indicate placing the 14 rad/sec lag in the command filter location has a

decidedly adverse effect on these tracking performance measures. The

difference is more pronounced for Subject B than for Subject A. Referring back to Figs. 35 and 37, it may be observed on the other hand that

there is no adverse effect when the 14 rad/sec feel dynamics are substituted for the 26 rad/sec feel system with the command filter lag eliminated. These results support the findings of Refs. 8 and 9.

65
2. Pilot Rating

Figure 41 shows that pilot ratings are not changed significantly by

interchanging the 14 rad/sec lag between the feel system and the command

filter locations. This is probably due to the accumulative system lag

being so great as to result in Level II flying qualities in three sets

of comparisons and borderline Level I-II in the fourth. But, referring

back to Fig. 38 it may be observed again that there is no adverse effect

on pilot rating when the 14 rad/sec feel dynamics are substituted for

the 26 rad/sec feel system. This too supports the flight results of

Refs. 8 and 9 concerning the lack of adverse effect on flying qualities

when a feel system having relatively low frequency dynamics is employed

in the roll axis.

CHPR

u

8-

7-

6-

5-

4-

3-

2-

I

Center Stick

ITI Subject A

0 Subject B

K c = I0 deg/sec/Ib K c = 20 deg/sec/Ib

0.. J3 ¢T

0--O

I I I I

14 26 14 26

26 14 26 14

D2 E2 H2 G2

Force D isp I.

(_ ::;_ LEVEL 11"

E)

LEVEL I

ID

I I I I

14 26 14 26 Feel

26 14 26 14 Filter

D2 E2 H2 G2 Configuretion

Force Displ. Sensing

Figure 41. Influence of Dynamic Lag Location on CHPR (Center Stick)

66
D. INFLUENCE OF CUMULATIVE EFFECTIVE TIME DELAY

Having looked at the various direct comparisons between mechanizational elements, it is of interest to view the results in terms of the

computed cumulative buildup in effective time delay with the different

configurations. The table in Fig. 33 shows this effective delay, from

force application to roll response, extends from 0.033 sec to 0.19 sec

in four nearly equal time increments.

1. Tracking Performance

Figures 42 and 43 present rms tracking error (_e) plotted against

cumulative effective Te for subjects A and B,respectively. The data of

Fig. 42 show for Subject A a linear increase in error with increasing

time delay. There is a distinct difference in performance between the

low and high command gains which begins at the lowest effective time

delay configurations and diminishes as Te increases. The lower rms

error for Kc = 20 was noted in the previous plots but here it becomes

most noticeable that the gain influence decreases at the higher effective time delay values.

Figure 43 for Subject B also indicates a linear increase in tracking

error as controlled element effective time delay increases. However

this plot tends to bring out the wider spread in error due to a less

consistent tracking gain loop closure of Subject B as compared to

Subject A.

In summary, the rms tracking error measures show a linear degradation in performance proportional to an increase in overall controlled

element effective time delay.

2. Pilot Ratings

A summary plot of CHPR versus controlled element effective time

delay is presented in Fig. 44. This shows the mean and I _ values

obtained by combining data from both subjects at each command gain. The

numbers beside the symbols indicate the number of runs involved. Not

only is the data very consistent (one half rating point i o ranges), but

67
Ge

(deg)

I0

9

8

7

6

5

I

III

I I

0.05 _ 0.1

"B'2

B2

Figure 42.

I

G]E 2

[] F2

Solid - Kc : I0

Open - K c = 20

Subject A

I I

0.15 _ 0.2 re(Sec)

C2 Ez Gz

D2 F2 H2

RMS Tracking Error vs Effective Time Delay,

Subject A (Center Stick)

Ge

(deg)

13-

12-

II -

I0-

9-

8-

7-

6-

5

• A2

A2

B2

• C2

E2

c2C D2

F2

C) D2

C) F2

G2

C) H2

H2

Solid - Kc = I0

Open - Kc = 20

Subject B

I I I I

0.05 _ 0.1 _ 0.,5 _ 0.2 re(Sec)

A2 C2 E2 Gz

B2 D2 F2 Hz

Figure 43. RM,% Tracking Error vs Effective Time Delay,

Subject B (Center Stick)

68
CHPR

m

6-

5-

4-

5-

2-

I -

0

(Z)e 1

18

Center Stick

Solid-K c =10

Open - Kc = 20

Subject A and B

I I I I

t 0,05 _ O.I _ 0,15 _ 0,2 re(sec)

Az Cz E_, G2

Bz D2 Fz Hz

Figure 44. CHPR Variation With Effective Time Delay and

Command Gain (Center Stick)

this plot again shows the definite preference for the K

c

gain at T less than 0.19 sec.

= 20 command

3. Comparison With MIL-Spec Time Delay Criteria

It was noted in Section II-D-2 that the question has been raised as

to whether the time delay criterion of MIL-F-8785C should be referenced

to manipulator force or displacement (i.e., whether the feel system lag

should be included when the command input is manipulator displacement).

Figure 45 presents summary plots of average CHPR vs Te data including

the feel system lag (per 8785C) and excluding the feel lag (per Ref. 9).

Since results of this simulation show a sensitivity to command gain,

data are plotted for each gain separately and for the average of the

two. The upper plots reflect K c = i0, the middle plots reflect K c = 20,

and the lower plots the average for both gains. All plots reflect the

averaged CHPRs from Fig. 44.

69
8

7

6

5

CHPR

4

3

2

I

0

8

7

6

5

CHPR 4

3

2

I

0

8

7

6

5

CHPR

4

3

2

I

0

Center Stick

Open -Force Stick

Filled- Displacement

Stick

With Feel

- A2

- B2

- I

I

0.1

(_ H2

E2 G 2

F2

0.2

A2

B2 I

I

I

OI

G21

H21

I

F2 I

E2 I

I

5

4

A2

B2 I

i n I

0.1

3

2

I

0

- Without Feel

- H 2 G 2

F2 •

_) ®D2

_ | E2

cz I

- A2 I

_ B2 I

I

0.1

_

7-

6- H2

5- •

4 - F2

3- oq

2 - A2 D21

I - B2 I

o I

0.1 0.2 0.1

G2

(3

E2

0.2

]

I

I

I

I

I

I

I

0.2

7

-'71 6 H2 G2 ]

I

O i 5 c2 • •

I

E2 I 3 oE2 l

F2 I1 2fl BA2z ,1 II

, _1 0 ,I , I

O.2 0.1 0.2

r e ( sec ) re(Sec )

K c = I0

KC=20

Kc : I0 and 20

Averaged

Figure 45. Comparison of CHPR With MIL-F-8785C Effective

T_me Delay Criteria (4 ib/in Center Stick)

7O
When the effective time delay is computed from force application

(therefore feel dynamics included), there is a steady progression of

increasing CHPR with increasing _ which supports the -8785C Level I

boundary but may indicate the boundary between Level II and III is too

conservative. On the other hand, when the feel dynamics are excluded in

the calculation of the effective time delay, two different CHPRs result

for each value of T and the ratings for the displacement sensing center

stick (shaded symbols) definitely do not fit the criteria boundaries.

The data are not consistent with the contention that the feel system

dynamic lag should be excluded when calculating the overall vehicle

effective time delay. It should also be noted that these results for

the center stick are not consistent with those of Fig. 32 for the side

stick. There was not so much difference in pilot rating between force

and displacement stick sensing for the side stick.

It seems obvious from these comparisons that the more appropriate

criterion is that of MIL-F-8785C which includes the feel lag contribution in the overall vehicle computed effective time delay when stick

displacement sensing is employed.

E. SUMMARY

This set of experiments was performed to obtain lab data for later

comparison with flight. Results of this limited simulation forcast that

for a center stick configuration in a roll tracking task:

® Force sensing results in lower tracking error and

is rated better than displacement sensing.

In the absence of appreciable command filter lag,

there is no difference in tracking performance or

pilot preference between feel system dynamics of

26 and 14 rad/sec.

A 14 rad/sec command filter in series with a

26 rad/sec feel system has a decidedly adverse

effect on tracking error in comparison with a

26 rad/sec command filter and a 14 rad/sec feel

system, however it appears to make no difference

in pilot rating. (The latter probably is due to

the excessive total effective time delay for both

configurations).

71
o Pilot ratings obtained for effective time delays

up to 0.19 sec support the MIL-F-8785C criterion

which requires the time delay be referenced to

force applied to the manipulator regardless of

whether force or displacement sensing is used.

Results also fell within the 8785C Level I and II

boundaries.

72
SECTION VI

PILOT DYNAMIC MODEL WITH FORCE VS DISPLACEMENT CENTER STICK

It was shown in the previous section that stick displacement sensing

(feel dynamics in the forward path of Fig. 9) degraded CHPR an average

of 1 rating point compared to force sensing (feel dynamics in the feedback path). However, based on Refs. 8, 9 and other findings of the

previous section, there also is evidence that use of a comparatively

sluggish feel system does not degrade flying qualities ratings appreciably, providing there is little or no command filter lag in the outer

(tracking) loop path.

An explanation for the latter in the case of displacement sensing is

advanced in Ref. 9 by noting that closure of the limb/manipulator loop

via the pilot's arm joint receptors moves the open-loop feel dynamic

mode to higher frequency and therefore reduces the closed-loop effective

lag contribution of this dynamic element. Figure 46 presents an example

root locus plot for closure of the positional limb/manipulator loop of

Fig. 9 using typical 3rd order NM system parameter values and the

14 rad/sec, 0.07 damped, feel system dynamics. (The second order mode

above 40 rad/sec is the Pade' approximation for the time delays in the

forward and feedback paths of Fig. 9.) This locus shows the feel system

mode moves to higher frequency and lower damping as loop gain is

increased. Conversely, the NM 2nd order mode moves to a lower frequency

but at lower damping. For the maximum stable gain shown, the closedloop feel mode moves to about 17 rad/sec.

Another possible explanation is that the feel system lag contribution is partially or totally compensated for by a change in the pilot's

central compensation (decrease in latency and/or generation of lead).

But, if the feel lag is offset by pilot appreciable lead generation,

then one might expect an accompanying degradation in flying quality

rating. This has not necessarily been noted.

In this subsection an attempt is made to shed light on the subject

by applying two different model fitting approaches to the data obtained.

73
l

M

a

g

30

25

20

t5

t0

P

/

/

/

/

I

I'

I/

It

[0.48 ;17]

/I

w FS ){"/

[0.7;14]

I

coN

X

[0.082 ;9.8]

L

I

Figure 46. Example Root Locus of the Limb/Manipulator

Positional Loop Closure

In the first, YpYc describing function data from several typical centerstick tracking runs were fitted with transfer function models appropriate for the various elements in the loop. Parameter values for the

"known" control system and aircraft were considered fixed while values

for the "unknown" pilot were adjusted until an adequate fit was

obtained. This approach appears to provide some of the needed insight

to the above uncertainties concerning high frequency mode contributions.

Additional insight is obtained by taking a much simplified crossover

model fit to identify dominant factors in the region of tracking loop

closure frequency.

A. HIGH ORDER DYNAMIC MODEL

The STI Describing Function Analyzer (DFA) program computationally

extracts opened-loop describing function amplitude and phase data from

74
the closed-loop tracking task. Therefore it is appropriate to start

with known open-loop dynamic models and parameter values for the feel,

command filter, and aircraft elements and assume models and parameter

values for the pilot. The latter are then iterated until acceptable

fits are obtained to the amplitude and phase data points. Adjustments

to the feel dynamics are considered if adequate matches cannot be

obtained through this method.

fitted are listed in Table 3.

filter.

The example configurations which were

All included the 2nd order command

The matches to the describing function data points are presented in

Figs. 47 through 52. The transfer functions employed in the various

fits are included in each figure. It should be noted that the pilot NM

mode in each instance is modeled by second order rather than third order

dynamics. The first order lag is either beyond the frequencies employed

or is effectively canceled by the NM mode zero (see Eqs. I, 2, 3). In

some cases it is also necessary to include lead/lag terms in the pilot

model to achieve an adequate fit. In all cases excellent amplitude and

phase fits are achieved out through 14 rad/sec. At the highest frequency (19 rad/sec) the amplitude data points generally are somewhat

higher than the models predict and the phase matches vary. This is not

TABLE 3. CONFIGURATIONS FITTED WITH HIGH ORDER MODELS

FIGURE CONFIG.

COMMAND

SENSING

FEEL

SYSTEM

COMMAND

FILTER

SUBJECT

47 E2 Force 26 14 A

48 D2 Force 14 26 A

49 H2 Displacement 14 26 A

50 G2 Displacement 26 14 A

51 H2 Displacement 14 26 A

52 H2 Displacement 14 26 B

75
_u(radlsec)

÷

-_n13

X

o,I U

6667x I

¥c: co-6_7.__ _ : ,o

,, ^z -O.07s °aFS = 26 +1"

r, px I/ e ..... I

YP : _[0.05; wCF = 14

12"]'-- wCF = 14

e-O.O33s [-0.886; 105] _,

[0 886 ;105] _,.

It

Figure 47. Dynamic Model Match to Describing Function Data,

Force Sensing Center Stick, Subject A, mFS = 26, mCF = 14

76
_D

[]

0.'i.0

_J cu(radlsec )

X

Y¢ =

6,667 x 262e -0"033s

(0)(6,667) [0.7 ;26]

Kp x 122e O'07s

[0.05 ; 12]

e-O.O33s : [-0.886; 105]

[0.886 ;105]

Force Sensing:

K c : 20

wFS: 14

_CF = 26

X

Figure 48. Dynamic Model Match to Descri6ing Function Data,

Force Sensing Center Stick, Subject A, _FS = 14, mCF = 26

77
D.10 1.00 w (rad/sec) 10.00

,,-o _J ! I

O3

._ ×

2O _... 4"

D

--I0(3

o \

O_

--20O

Yc (0)(6.667)[0.7 ; 14] [0.7; 26][0.866; E05

Kp x 13Ze -O'06s

YP = [0.07_ 13]

Displocement Sensing:

Kc = IO

wFS = 14

eJCF = 26

÷

1

Figure 49. Dynamic Model Match to Describing Function Data

Displacement Sensing Center Stick, Subject A

_FS = 14, _CF = 26

78
1,00 w ( rod/sec ) lO.O0

I I

rm

= 2D --

.o-

.--

"E

E

D

"0

--IOD --

u_

0

tQ_

--200 --

6.667x t4Zx 26 z [-0.866; 105]

Yc = (0)(6.667)[0.7; 14][0.7; 26][0.866; 105]

Kp × 15ze -O'06s

YP = [0.07; 1:5]

Displacement Sensing:

K c =20

WFS = 26

_CF = 14

+

0

Figure 50. Dynamic Model Match to Describing Function Data,

Displacement Sensing Center Stick, Subject A,

_FS = 26, mCF = 14

79
0.1.n 1.aCl _ ( rod/sec ) ln.mi3

_J I I

m

x

E

c0C

r'l _

II

a __

0

--200 --

6.667 x 14

Yc : {o}{s.TT[oT;_T,_] _;iB_] X

Kp×ls2e°lfs { _{s_} ]

YP -- [0.03 ;13] x 3{-_}

Displacement Sensing:

K c = 20

_FS = 14

WCF = 26

Figure 51. Pilot Generation of Lead With Displacement

Sensing Center Stick, Subject A, _FS = 14, mCF = 26

80
4-D

1.013 w(radlsec) _13.013

I I

¥

x*

6.667 x 14z x 262 [-0.866 ; 105]

Yc =

(0)(6.667)[0.7 ; 14] [0.7 ; 26] [0.866 ; 105]

YP = Kpxll2e -O'lls 17 (3)

[0.03;11 ] x .5 (17)

Displacement Sensing:

K c =20

WFS = 14

o_CF = 26 J

Figure 52. Pilot Generation of Lead With Displacement

Sensing Center Stick, Subject B, _FS = 14, _CF = 26

81
too disconcerting as the forcing function amplitude is quite small at

this frequency and the closed-loop response power is also very low

allowing system noise and other sources of remnant to dominate. Therefore there is considerably more scatter in the data at this frequency.

Comparison between the figures reveals the following:

I. Force Vs Displacement Sensing (Figs. 47, 48 vs 49, 50)

The fits to the force sensing configurations are shown in Figs. 47,

48. These configurations reflect command filters of 14 and 26 rad/sec

and gains of I0 and 20 deg/sec/ib, respectively. Both are characterized

by the absence of the feel dynamics lag. The pilot model consists of a

time delay of 0.07 sec and the 2nd order NM mode at 12 rad/sec and 0.05

damping. In both cases excellent matches are obtained. The gain line

crossover with the amplitude fit (_c) is above 3 rad/sec, the phase

crossover (_u) is close to 5 rad/sec, and the phase margin at _0c is

30 deg in both cases. These were given CHPR ratings of 3 and 4.

Two matches to displacement sensing configurations in which the subject pilots did not employ lead/lag compensation are shown in Figs. 49,

50. These are characterized by the presence of the feel system lag

modeled with the open-loop parameter values. Figure 49 is for the 14

rad/sec feel, 26 rad/sec filter, and I0 deg/sec/ib gain. Figure 50 is

for the 26 rad/sec feel, 14 rad/sec filter, and 20 deg/sec/ib gain. In

both of these cases an excellent match is obtained with a pilot time

delay of 0.06 second 2nd order NM mode of 13 rad/sec and 0.07 damping.

Thus there is very little difference in the pilot model parameters

between this displacement and the previous force sensing fits. But it

will be noted that the amplitude crossover is barely 3 rad/sec in both

of the displacement stick cases, the phase crossover is less than

4 rad/sec, and the phase margin is 20 deg or less. These elicited CHPR

ratings of 5 and 6.

It should be pointed out that the two force and the two displacement

sensing cases shown represent an interchange of feel and command filter

dynamic lag values. In these examples there obviously is no influence

on the pilot model or parameter values.

82
2. Pilot Generation Of Lead

Although the interchange of the 14 and 26 rad/sec dynamic elements

had little influence, there was considerable evidence in the describing

function amplitude and phase data that the additional phase lag of the

two elements caused the subjects to adopt lead/lag compensation in many

runs with the displacement sensing stick. Two of the more extreme examples of this behavior are presented in Figs. 51, 52. Figure 51 shows

Subject A adopted a lead at 3 rad/sec and lag at 10 rad/sec. His central latency increased to 0.Ii sec but there was no change in his

NM mode. This particular set of runs indicates gain crossover very near

to the phase crossover frequency (zero phase margin), a PIO condition.

It resulted in a CHPR of 5.

Another set of example runs for the same controlled element but with

Subject B is presented in Fig. 52. This was again fit with a lead at

3 rad/sec but a lag at 17 rad/sec. Note that this resulted in a little

too much phase lead in the region around crossover. Thus a slightly

higher lead break would probably produce a better match. This Subject

also increased his central latency to 0.11 sec and reduced his NM mode

to 11 rad/sec, 0.03 damping. The equalization increased the phase

crossover (_u) to almost 5.5 rad/sec but the accompanying increase in

amplitude ratio at frequencies above 3 rad/sec caused a gain regression

which resulted in an amplitude crossover of only 2.5 rad/sec. Also the

amplitude peaking of the NM mode is greater than in any of the previous

plots. Despite these adverse influences, the Subject gave the configuration a CHPR rating of 4. Quite possibly this is due to the phase margin at gain crossover being increased to about 60 deg.

In both Figs. 51, 52 it is apparent the adoption of lead causes a

flattening of the amplitude ratio plot in the region of desired crossover (e.g., 2.5 to 4 rad/sec) and therefore non-K/s like characteristic.

This results in increased sensitivity of crossover frequency and/ or

phase margin to pilot gain and may help to explain the Level II ratings

given to these configurations in Fig. 41.

83
3. Summary

Results of these higher order mode data fits have shown no appreciable difference in pilot model form or parameter values for configurations in which the dynamic lag contributions of the feel and command

filter elements were interchanged. Presence of the 14 rad/sec 2nd order

lag (either feel or filter) with a displacement sensing stick either

reduces closed-loop bandwidth and phase margin (if Yp is unchanged from

the force sensing cases) or induces the pilot to adopt lead/lag compensation in an attempt to improve bandwidth and phase margin. The latter

results in an offsetting increase in pilot central latency and increased

sensitivity of closed-loop performance to pilot gain.

Excellent high order fits have been obtained over the entire

frequency range through the use of feel system parameter values representative of open-loop limb/manipulator characteristics. This tends to

indicate that the limb/manipulator position loop is not closed sufficiently tight to appreciably modify either set of dynamic characteristics.

B. CROSSOVER MODEL APPROACR

The foregoing high order models provide excellent fits to six

example cases but one might question whether these results are representative of all 105 describing function data runs. Detailed fits to

accommodate the nuances of all runs could not be accomplished within

time and budget constraints. Therefore the much simpler crossover model

approach was employed. This approach focuses on the region around

crossover (mc) since this is the region which dominates the pilot's perception of system performance (e.g., bandwidth, phase margin, gain

margin, need for lead compensation).

In this simplified model the lag contribution of all higher

frequency dynamic modes are approximated by the summation of their

equivalent time delays (see Fig. 53). The describing function data

points reflect the total open-loop pilot/feel/filter/aircraft system,

84
O3

Q.

E

<

_D

_D

O

D._m

..J

"J..n 0 u_ (rodlsec) lO.oO

I I

-20dBldec

¢,

_ X

+ _ @ O

t-

(3_

-180 deg

@

: Yp YFSYCFYc

-'- Tp + rps+ TCF+rC

Pilot Feel Filter Airframe

It

I3

,4-

÷

rl

÷

X

Figure 53. Effective Controlled Element Time Delay

Summation for Crossover Model

85
YpYFSYCFYc . Since the dynamic parameters of YCFYc are known, the equivalent time delay contribution of these elements can be subtracted and the

residual for the unknown YpYFS is left. The average values for the

"residuals" from the 105 data runs are indicated below for the force and

displacement sensing sticks.

_p(Sec)

0.076

0.08

_FS (rad/sec)

14

26

(_p + _FS)

0.125

0.105

_FS

14

26

Force sensing sticks

Force _._

Displacement

Displacement sensing sticks

For the force stick the residual consists of the pilot effective

time delay Cp. The values obtained with the 14 and 26 rad/sec feel

dynamics are essentially the salne. If we calculate the total effective

pilot time delay based on the high order model of Figs. 47, 48 we

obtain:

2 ENt4

re = _p +-----_q_M

2(0_7_

= O. 07 + >-:-'-:-:---"

12

= 0.082 sec

This is within about 5 percent of the average of the two values noted

above. Thus the higher order model fit derived for the force sensing

stick is consistent with all force sensor runs.

86
For the displacement sensing stick the residual reflects the pilot

and feel dynamics (Tp + TFS). If we calculate the total effective time

delay using the dynamic models of Figs. 49, 50 with the 14 rad/sec feel

system we obtain:

2_NM 2_FS

Te = Tp + -- + --

mNM _FS

2(0.07) 2(0.7)

= 0.06 + +

13 14

= 0.06 + 0.0108 + 0.I

= 0.1708 sec

This is about 37 percent higher than the 0.125 sec extracted from the

data runs. If the difference (0.0458 sec) were to be attributed to a

shift in the feel system frequency and damping due to closure of the

pilot limb/manipulator loop, it would require nearly doubling the feel

system frequency. This is not likely based on the root locus of

Fig. 46. On the other hand, the differential could be accounted for by

a relatively modest first order pilot lead/lag. This also is consistent

with the phase plots of Figs. 49, 50 where it can be noted that the

describing function data points are about i0 deg above (lead) the fitted

curve over a small range at, and slighly above, the crossover frequency.

A similar calculation of the effective time delay with the 26 rad/

sec feel system shows:

Te = 0.06 + 2(0.07) + 2(0.7)

13 26

= 0.06 + 0.0108 + 0.055

= 0.1258 sec

This is about 20 percent higher than the 0.105 sec from the data runs.

Again this can be accounted for by a small pilot lead.

Thus it appears the subject pilots were consistently adopting lead

to partially offset the lag introduced by the displacement sensing stick

87
configuration. But the lead time constant is so small that it may not

adversely effect the handling quality rating.

C. ROLL RATCHET POTENTIAL

Consistent with the side-stick portion of the experiment, it is of

interest to look at the roll ratchet potential for the center-stick configurations. A comparison of the typical phase lag data points in the

vicinity of the NM frequency in Figs. 47 through 52 with the similar

phase data points of Fig. 18 (sidestick) shows the controller/aircraft

configurations used with the center-stick have much greater phase lag

and therefore definitely are not candidates for roll ratchet. This

excess lag is due to the cumulative effect of the command filter, the

aircraft roll subsidence (T R = 0.15 sec), and the feel system lag (with

displacement sensing).

88
SECTION VII

CONCLUSIONS AND RECOMMENDATIONS

Results of the experiments discussed herein together with those of

Re f. II provide insight to key interactions between the pilot s neuromuscular system and the manipulator, feel system, control response command gain, and command prefilter elements in roll tracking type control

tasks. The relative merits of stick displacement command versus stick

force command have been quantified in terms of effective time delay,

maximum control bandwidth, tracking performance, low and high frequency

PIO tendency, pilot's neuromuscular mode peaking, etc. Potential tradeoffs between response command gradient (gain), feel system force/

displacement gradient, and effective time delay have been noted. The

overall results provide insight and design guides for minimizing roll

control problems in future high performance aircraft. Specific conclusions are:

o Manipulator Force Vs Displacement Sensing

o Force Sensing

Minimimes forward loop dynamic lag

because the manipulator/feel system lag

is relegated to the feedback path of the

pilot's neuromuscular limb position

system

-- Results in lower tracking error in the

simulation task

-- Minimum phase lag plus the associated

decrease in signal attenuation at frequencies near the neuromuscular mode

amplitude peak increases susceptibility

to roll ratchet

Command prefiltering is generally

required to prevent roll ratchet and

other high frequency extraneous inputs

Caution must be exercised in selecting

the prefilter time constant because

excessive lag leads to low frequency PIO

89
-- As a rule of thumb, the prefilter inverse

time constant should be double the

closed-loop control bandwidth required.

Displacement Sensing

-- Moves the manipulator/feel system dynamic

characteristics to the forward loop commandpath

-- Increases (slightly) the neuromuscular

mode peaking over that obtained with

force sensing

-- Increases high frequency phase lag and

commandattenuation

-- Reduces or eliminates any tendency to

neuromuscular system induced roll ratchet

if the manipulator/feel system dynamics

lie in the region of the neuromuscular

system 2rid order mode

-- Reduces or eliminates the need for commandprefiltering

-- As a rule of thumb, the manipulator/feel

system principal dynamic mode should be

greater than the pilot's neuromuscular

mode (approximately 12-13 rad/sec) to

avoid adverse effect on tracking performance and flying qualities rating

• Side- vs Center -Stick

-- The above apply equally to both types

ResponseCommandGradient (Gain)

-- All results (side- and center-stick)

showed a preference for commandgain of

about 20 deg/sec/Ib in both tracking and

gross maneuvering in the simulation

-- Based upon comparison with previous NT-33

flight experiments, it appears I0 deg/

sec/ib is more appropriate for actual

flight with small precise control input

tasks (formation, tracking) but the

20 deg/sec/Ib remains appropriate for

gross maneuvering

90
o

-- Command gains of less that I0 deg/sec/Ib

result in rapid degradation in control

performance (bandwidth, tracking error)

and flying qualities rating due to the

high forces involved

-- Degradation of CHPR with increasing

effective time delay was less with command gain of 20 deg/sec/ib than with

I0 deg/sec/ib

Feel System Force/Displacement Gradient

side-stick)

(for

-- 0.65 Ib/deg was rated best for precise

tracking and following gross attitude

changes

-- 1.22 ib/deg was found to be stiff and

tiring

-- 0.33 ib/deg was too soft, producing a

tendency to excessive stick deflection

-- Based upon flight test in the NT-33, it

appears that gradients as high as

1.33 ib/deg may be acceptable in the

presence of motion and atmospheric

disturbances

Manipulator/Feel System Dynamics

-- In the absence of appreciable command

filter lag, there appears to be no difference in tracking performance or pilot

preference (CHPR) between feel system

dynamics as low as 14 rad/sec and

26 rad/sec.

Feel/Command Filter Effective Time Delay

-- Flying quality rating degradation with

increasing Te was found to be consistant

with the criterion of MIL-F-8785C with

both the side- and center-stick configurations although there is some indication the Level I and II boundaries could

be relaxed somewhat (especially for the

side-stick)

-- Sensitivity to re decreased as command

gain (Kc) increased

91
-- The distribution of principal lag between

feel and commandprefilter has essentially no influence on flying quality

rating but tracking performance measures

show preference for the lessor lag being

located in the commandfilter

Manipulator Configuration (side- vs center-stick)

-- Use o_ the side-stick resulted in significantly higher closed-loop control bandwidth and lower tracking error in the

roll tracking task simulated, however,

there was no significant difference in

pilot rating

92
REFERENCES

_. Magdaleno, RaymondE., DuaneT. McRuer, and George P. Moore, Small

Perturbation Dynamics of the Neuromuscular System in Tracking

Tasks, NASA CR-1212, Dec. 1968.

2. Magdaleno, R. E., and D. T. McRuer, Experimental Validation and

Analytical Elaboration for Models of the Pilot's Neuromuscular

Subsystem in Tracking Tasks, NASA CR-1757, Apr. 1971.

3. McRuer, D. T., L. G. Hofmann, H. R. Jex, et al., New Approaches to

Human-Pilot/Vehicle Dynamic Analysis, AFFDL-TR-67-150,

Feb. 1968.

4. McRuer, D. T., and E. S. Krendel, Mathematical Models of Human

Pilot Behavior, AGARDograph No. 188, Jan. 1974.

5. Powers, B. G., "An Adaptive Stick-Gain to Reduce Pilot-lnduced

Oscillation Tendencies," AIAA Journal of Guidance, Control, and

Dynamics, Vol. 5, No. 2, Mar.,Apr. 1982.

6. Garland, Michael P., Michael K. Nelson, and Richard C. Patterson,

F-16 Flying Qualities with External Stores, AFFTC-TR-80-29,

Feb. 1981.

7. Burton, R. A. and B. T. Kneeland Jr., "F/A-18 High Authority/High

Gain Digital Flight Control System Development and Flight Testing," AIAA 81-2465, AIAA/SETP/SFTE/SAE/ITEA/IEEE First Flight

Testing ConEerence, Las Vegas, NV, Nov. 11-13, 1981.

8. Sarrafian, S. K., "Flying Qualities Result," in X-29A Symposium,

Initial Flight Results, Sept. 17-19, 1985.

9. Smith, R. E., and S. K. Sarrafian, "Effect of Time Delay on Flying

Qualities: An Update," AIAA 86-2202, Aug. 1986.

I0. Hoh, R. H., D. G. Mitchell, I. L. Ashkenas, et al., Proposed MIL

Standard and Handbook - Flying Qualities of Air Vehicles,

AFWAL-TR-82-3081, Vol. II, Nov. 1982.

11. Johnston, D. E., and D. T. McRuer, Investigation of Interactions

Between Limb-Manipulator Dynamics and Effective Vehicle Roll

Control Characteristics, NASA-CR-3983, May 1986.

12. Gordon-Smith, M., An Investigation into Certain Aspects of the

Describing Function of a Human Operator Controlling a System of

One Degree of Freedom, Univ. of Toronto, UTIAS Rept. No. 149,

Feb. 1970.

93
13. Magdaleno, R. E., and D. T. McRuer, Effects of Manipulator

Restraints on Human Operator Performance, AFFDL-TR-66-72, Dec.

1966.

14. McRuer, D. T., and R. E. Magdaleno, Human Pilot Dynamics with Various Manipulators, AFFDL-TR-66-138, Dec. 1966.

15. Mooij, H. A., W. P. de Boer, and M. F. C. van Gool Determination of

Low-Speed Longitudinal Maneuvering Criteria for Transport Aircraft with Advanced Flight Control Systems, NLR TR-79127 U,

Dec. 1979.

16. Hall, G. Warren, and Rogers E. Smith, Flight Investigation of

Fighter Side-Stick Force-Deflection Characteristics, AFFDL-TR75-39, May 1975.

17. Monagan, Stephen J., Rogers E. Smith, and Randall E. Bailey,

Lateral Flying Qualities of Hi$_hly Augmented Fighter Aircraft,

AFWAL-TR-81-3171, Mar. 1982.

18. Berry, D., "In-Flight Evaluation of Pure Time Delays in Pitch and

Roll," AIAA 85-1852, Aug. 1985.

94
APPENDIX A

TRACKING STATION AND FEEL SYSTEH SET-UP

A sketch of the tracking station set-up with the right hand sidestick is presented in Fig. A-I. Key dimensions, angles, pivots, etc.,

are identified. It should be noted especially that the stick grip

orientation was set to maximize comfort and minimize cross-axis coupling

(although the experiment task was single axis). The center-stick set-up

was essentially the same except for location of the stick and its pivot

near the floor.

Manipulator/feel system force-displacement calibration plots are

contained in Figs. A-2 through A-5. Also shown in each figure is a time

trace of stick displacement to a step force release, i.e., the stick was

held to one side and then the force released.

The force-displacement gradients shown here reflect measurements

made at the top of the stick (for convenience). In the main text the

side-stick gradients are referenced to the center of the grip for comparison with similar data from referenced experiments. The conversion

is:

top of stick (ib/in)

6.39

L

center of grip (ib/deg)

1.22

3.44 0.65

1.71 0.33

When using the force sensing stick configurations it is necessary to

incorporate a force signal deadboard (breakout) which matches the forcedisplacement breakout of the feel system. A calibration plot for this

feature is presented in Fig. A-6.

95
23" -'

DISPLAY

Pivots j

SIDE VIEW

\\ \ \ \ \ \ \ \ \ \ \ \ \ \ \ \ \ \ \ \ \

- I

Arm

Seat Rest

_._..__._ ,

TOP VIEW

l

18"

_L

LOOKING

FORWARD

I Arm Rest

Figure A-I. Sidestick Tracking Station Set-Up

96
.30

.2O

t-

.--

"-" .IO

Q;

E

c.b

o 0

GI.

r_

-.IO

-.20

-2

6.4 Iblin.

Iblin.

-I 0 I 2 3

Force (Ib)

Force Gradient (top of stick)

_p

E

cp

o

I

Q.

¢;)

._

r_

: - !:::: .!:i::I:!!!!! iii[iii _:!i_: :i:_ : iii:!i i : !

0 0.I 0.2 0.3 0.4 0.5

Time (sec)

Step Response [_;_] -" [0.7;31.4]

Figure A-2. Small Displacement (Stiff) Sidestick Calibration

97
.50

25

_p

E 0

c)

o

el.

.25 .--

.50

0.75

0.218

5.44 Ib/in.

2.0 1.5 1.0 0.5 0 0.5 1.0

Force (Ib)

1.5 2.0

Force Gradient (top of stick)

I I i T ..... 1 I

0 0.1 0.2 0.5 0.4 0.5

Time (sec)

Step Response [_;w] -" [0.6;22.4]

Figure A-3. Large Displacement (Medium) Sidestick Calibration

98
2.0

1,0 ....,...

rE 0

_P

t,;'}

.D

r_ -I ,0

-2,0

-3 -2 -I 0

Force (Ib)

I 2 3

Force Gredient (top of stick)

Q;

E

0

_5

::::::::::::::::::::: ::: ; ::I:::: : : !:::: ;:::t: ;: i : :::_:

i illliJil i_.::::.ii_::!i!_i:ii i:.:/il ii _ _iii i

• '-:-: :T:f : :i i i _ _, ; I .,. : .....

........ : : _. • I .} .... _ 1 ..

0 0.1 0.2 0.3 0.4 0.5

Time (see)

Step Response [/_;_] " [0,71;18.0]

Figure A-4. Very Large Displacement (Soft) Sidestick Calibration

99
1.25

1.00

.75

"-'. .50

t-

,Z5

Q;

E 0

0

=25

CI.

(/}

-- -.50

Q

-.75

- 1.00

-I.25

J

Iblin.

-7 -6 -5 -4 -3 -2 -I 0 I 2 3 4 5 6 7

Force (Ib)

Force Gradient (top of stick)

_E

_P

o

GI.

_n

_5

i I

o o, o1203 0.4 0'5

Time (sec)

Step Response [C_w] " [0.7; 14.0],

"E

oo

t_

Cb

i l i i I I

0 0,1 0.2 0.3 0.4 0.5

Time ( sec )

Step Response [_;_] -" [0.7; 26.0]

Figure A-5. Center-Stick Calibration

I00
Force Signal Input (Ib)

Figure A-6. Force Sensing Electrical Signal Breakout Calibration

101
APPENDIX B

ANALOG COMPUTER CODING

The roll tracking controlled element dynamics were implemented on an

EAI-231-R analog computer. The computer coding diagram from the

McFadden force loader analog signal inputs through to the display and

strip chart recorder outputs is presented here. Also shown are the

forcing function inputs from the digitally generated sum of sine waves

and the various signals fed to the analog to digital (A/D) conversion

for calculation of performance metrics and describing functions.

102
J

!

(

0

.t_

.I.J

_--_

0

I

1,4

ORIGINAL PAGE 13

OF POOR QUALITY

103
APPENDIX C

EXFERIMENT DATA

I. Closed-Loop Performance Measures

Open and closed-loop describing function measures and parameters are

calculated at the conclusion of each experiment run. Summary tables of

closed-loop performance measures, interpolated open-loop describing

function parameters, and experiment configuration parameter values

across all experiment runs are contained in Table C-I (center-stick

experiment) and Table C-2 (side-stick experiment). The open and closedloop parameters extracted are based on the extended crossover model

(Ref. C-I) where the plant is assumed to be of the form

Ke -j ( _ m - _/_)

y =

S

Amplitude

(dB)

Phase

(deg)

-180

.,._._._/WC 0 dB

I J :

I04
in the region of the crossover. A best "fit" to the describing function

amplitude and phase data points for each run is madeand the resulting

plant and loop closure parameter extracted. These are identified in the

table (and sketch) as follows:

ALPHA -- plant open-loop low frequency phase droop parameter from

the exponential _/_

CBAR -- average manipulator deflection during the data

(identifies any trim bias)

run

CHPR -- Cooper-Harper Pilot Rating given for run

CSIG -- one sigma rms value for manipulator deflection during the

run

EBAR -- average roll error in tracking run

ESIG -- one sigma rms roll tracking error

F -- force sensing stick (experiment configuration)

GM Bode open-loop gain margin at frequency of 180° phase

crossover, mu; computed from straight line interpolation

between describing function amplitude data points immediately above and below computed _u

K c -- Response command gain (experiment configuration)

P

-- position (displacement) sensing sticks

configuration)

(experiment

-- Bode open-loop phase margin at frequency of closed-loop

gain crossover, _c; computed from the complete extended

crossover model

PML

-- Bode open-loop phase margin at frequency of closed-loop

gain crossover, _c; computed from a straight line interpolation between the two describing function data points

immediately above and below _c

SLOPE slope of Bode open-loop amplitude asymptote between two

data points immediately above and below gain crossover

frequency

Tau -- equivalent time delay set into effective controlled element

TE

-- plant open-loop high frequency time delay parameter from

the exponential T_

I05
Tr

WC

WF

WS

WU

first order lag time constant set into effective controlled

element (represents aircraft roll subsidence mode or first

order command prefilter)

crossover frequency - frequency of crossover between openloop 0 db line and Bode amplitude asymptote calculated from

straight line fit between two describing function data

points immdiately above and below crossover

command filter frequency (second order) (experiment

configuration)

feel system (stick) frequency (second order) (experiment

configuration)

unstable frequency - frequency at which system open-loop

phase plot crosses the 180 deg line (calculated from

complete extended crossover model

2. YpYc Describing Function Plots

The remainder of Appendix C contains the computer generated openedloop YpYc plots for all experimental runs. Runs 01AI through 8A141 were

with the center-stick. Runs 15A142 through 22A379 reflect the sidestick manipulator with the various feel system gradient, filter, and

time delay configurations.

3. Reference

C-1 Jex, Henry R., R. Wade Allen, and Raymond E. Magdaleno, Display

Format Effects on Precision Tracking Performance, Describing

Functions, and Remnant, AMRL-TR-71-6, Aug. 1971.

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227
1. Report No. 2. Government Accession No. 3. Recipient's Catalog No.

NASA CR- 4111

5. Report Date

February 1988

4. Title and Subtitle

DESIGN CONSIDERATIONS OF MANIPULATOR AND FEEL SYSTEM

CHARACTERISTICS IN ROLL TRACKING

7. Author(s)

Donald E. Johnston and Bimal L. Aponso

9. _rformingOrganization Nameand Address

Systems Technology, Inc.

13766 South Hawthorne Blvd.

Hawthorne, California

12. Sponsoring Agency Name and Address

National Aeronautics and Space Administration

Washington, DC 20546

6. Performing Organization Code

8. Performing Organization Report No.

H-1438

10. Work Unit No.

RTOP 505-67-01

11. Contract or Grant No.

NAS 2-12221

13. Type of Report and Period Covered

Contractor Report - Final

14. Sponsoring Agency Code

15. Supplementary Notes

NASA Technical Monitor: Donald T. Berry, Ames Research Center, Dryden Flight Research Facility,

Edwards, California 93523-5000

16. Abstract

A fixed-base simulation was performed to identify and quantify interactions between

the pilot's hand/arm neuromuscular subsystem and such control system features of typical

modern fighter aircraft roll rate command mechanizations as: (I) force versus displacement sensing side-stick type manipulator, (2) feel force/displacement gradient, (3) feel

system versus command prefilter dynamic lag, and (4) flight control system effective

time delay.

The experiment encompassed some 48 manipulator/filter/aircraft configurations summarized in Sections IV and VI. Section IV concentrates on the displacement side-stick

experiment results and compares these with the previous force sidestick experiment

results. Attention is focused on control bandwidth, excitement (peaking) of the neuromuscular mode, feel force/displacement gradient effects, time delay effects, etc.

Section V is devoted to experiments with a center-stick in which force versus displacement sensing, feel system lag, and command prefilter lag influences on tracking performance and pilot preference are investigated. Results in these two sections are

summarized in numerous plots which are intended to serve as guides in the design of

future control systems. Section VI concentrates on extraction of dynamic models for

the pilot and closed-loop arm/manipulator/feel systems from the describing function

data obtained in the Section V experiments. Parameters suitable for detailed models

and for the simple crossover model are derived. Details concerning the manipulators,

feel gradients, etc. along with run logs and data summaries are presented in the

several appendixes.

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

Flying qualities

Limb-manipulator dynamics

Neuromuscular system

Roll control

18 Distribution Statement

Unclassified -- Unlimited

Subject category 08

19 S_urity Clasif.(ofthisreport) 20. Security Clasif.(ofthis _ga) 21. No. of Pages 22. Price*

Unclassified Unclassified 238 A11

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

NASA-Langley, 1988