UNCLASSIFIED AD NUMBER LIMITATION CHANGESnozzle thrust vector control system for tactical missiles. This investigation was conducted during the period 22 April 1976 to 19 October 1976,

UNCLASSIFIED

AD NUMBER

LIMITATION CHANGESTO:

FROM:

AUTHORITY

THIS PAGE IS UNCLASSIFIED

ADB019655

Approved for public release; distribution isunlimited.

Distribution authorized to U.S. Gov't. agenciesonly; Test and Evaluation; MAY 1977. Otherrequests shall be referred to Naval WeaponsCenter, China Lake, CA.

NWC ltr 22 Sep 1977

V

THIS RCPOUT HAS BEEN OEUMITEO

AND CLEARED FOR PtOLIC RELIilE

UNDER DOD DIRECTIVE 5200,20 ARD

NO rtESTRICTIONS ARE IMPOSED UPON

ITS USE AND DISCLOSURE.

DISTRIBUTION STATEMENT A

APPROVED FOR PUBLIC RELEASE;

DISTRIBUTION UNLIMITED.

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/ NWC TP b9ü2

10

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Low Cost Servo Actuator for Tactica

Missile Thrust Vector Control (TVC)

a- o o • D D C

j Uu i

by R. Lwi

P. H. Siahlhuth

Aeronutronic Division

Aeronutronic Ford Corporation for the

Propulsion Development Department

MAY 1977

Distribution limited to U.S. Government agencies only; test and evaluation! 9 February 1977. Other requests for this document must be referred to the Naval Weapons Center.

/

Naval WeaponsCenter CHINA LAKE, CALIFORNIA 93555

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Naval Weapons Center AN ACTIVITY OF THE NAVAL MATERIAL COMMAND R. G. Freeman, III. RAdm., USN Commander

G. L. Hollingsworth Technical Director

FOREWORD

This report describes the design, development, fabrication and testing of cold gas driven, closed-center servo actuator systems for operation of a trapped ball type of movable nozzle thrust vector control system for tactical missiles. This investigation was conducted during the period 22 April 1976 to 19 October 1976, The work was sponsored by the Naval Weapons Center (NWC), China Lake, California, under Navy Contract N00123-76-C-0286 and supported by the Naval Air Systems Command under AirTask A03W3300/008B/ 6F31330300.

Mr. Michael J. Ripley was the Navy Technical Coordinator and has reviewed this report for technical accuracy.

This report is released for information at the working level and does not necessarily reflect the views of NWC.

Released by B. W. HAYS, Head Propulsion Development Department 4 May 1977

Under authority of G. L. HOLLINGSWORTH

Technical Director

NWC Technical Publication 5902

Published by Propulsion Development Department Collation Cover, 45 leaves First printing J 90 unnumbered copies

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4 Chief of Naval Material MAT-030(1) MAT-032 (1) NSP.27 (1) NSP-2731 (1)

5 Naval Sea Systems Command SEA-03 (1) SEA-031(1) SEA-033 (1) SEA-04H(1) SEA-6531 (1)

3 Marine Corps Development and Education Command, Quantico (Marine Corps

Landing Force Development Center) 1 Air Test & Evaluation Squadron S (LT Karl Kail) 1 Naval Air Development Center, Warminster (Code 3014) 1 Naval Ammunition Depot, Hawthorne (Code 05, Robert Dempsey) 1 Naval Explosive Ordnance Disposal Facility, Indian Head 1 Naval Intelligence Support Center (OOXA, CDR Jack Darnell) 1 Naval Ocean Systems Center, San Diego (Code 133) 1 Naval Ordnance Station, Indian Head (Code FS, A. T. Camp) 3 Naval Surface Weapons Center, White Otk

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1 Naval Intelligence Support Center Liaison Office (LNN) 1 Army Armament Command (AMSAR-SF) 4 Army Armament Research & Development Center (SMD, Concepts Branch) 3 Air Force Systems Command, Andrews Air Force Base

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1 Defense Advanced Research Projects Agency, Arlington 2 Defense Documentation Center

42 Chemical Propulsion Mailing List No. 271 dated October 1975, including Categories 1,2,3,4,5

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2. COVT ACCESSION NO

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^T-^CIPIENT'S CATALOG NUMBER

TITLE fand Sublllle)

LOW COST SERVO ACTUATOR FOR TACTICAL MISSILE JHRÜST VECTOR CONTROL (JYC)

7. AuTHORfaJ

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9. PERFORMING ORGANIZATION NAME AND ADDRESS

Aeronutronic Division Aeronutronic Ford Corporation Newport Beach, California 92663

II. CONTROLLING OFFICE NAME AND ADDRESS

Naval Weapons Center China Lake, California 93555

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U. MONITORING AGENCY NAME & ADDRESSf/f dlllermnl from Conltolllnt Otllc»)

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10. PROGRAM ELEMENT, PROJECT, TASK AREA ft WORK UNIT NUMBERS

AirTask A03W3300/008B/

6F31330300

May #77 r m IS. NUMBER OF PACES

110 IS. SECURITY CLASS, fo

UNCLASSIFIED

ISa. DECLASSIFICATION/DOWNGRADING SCHEDULE

16. DISTRIBUTION STATEMENT (at IM» Report)

Distribution limited to U.S. Government agencies only; test and evaluation; 9 February 1977. Other requests for this document must be referred to the Naval Weapons Center.

17. DISTRIBUTION STATEMENT (ol th» mbilrmet antarad In Block 30, II dlllmtml from Rmpotl)

IB. SUPPLEMENTARY NOTES

19. KEY WORDS (Contlnua on ravaraa «Ma II nacaaaary and Identity by block number)

Servo Actuator Movable Nozzle Thrust Vector Control Tactical Missiles Position Transducer

Electronic Controller Helium Performance Actuator Performance

20. ABSTRACT CContlnua on ravaria «Ida II nacaaaary and Identify by block number)

See back of form.

DD 1JANM73 1473 EDITION OF I NOV 65 IS OBSOLETE S/N 0102-014-6601 |

UNCLASSIFIED SECURITY CLASSIFICATION OF THIS PACK flWian Data «i»lar*J

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(U) Low Cost Servo Actuator for Tactical Missile Thrust Vector Control (TVC), by R. Levi and P. H. Stahlhuth. China Lake, Calif., Naval Weapons Center, May ll)77, 110 pp. (NWC TP 5902, publication UNCLASSIBIED.)

(U)^rhis report discusses the design, analysis, and performance evaluation of a helium driven closed center servo actuator system. The actuator was sized to operate a trapped ball type of movable nozzle thrust vector control system for tactical missiles. Predicted and actual performance are compared and typical test data are presented.

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CONTENTS

•

Technical Summary 4

System Description 10 Load Definition 10 Mechanical Description 15 Control Valve 15 Actuator Cylinder Assembly 15 Position Transducer 16 Actuator Operation 17 Valve Operation 17 Electronic Controller 19 Weight i 2A

Design Analysis 25 Actuator Sizing 25 Control System 37 Test Rig Design 38

Test Results 43 Helium Performance A3 Nitrogen Performance 46

Gas Supply Sizing 78

Appendix: A. Acceptance Test Procedure, NWC Nozzle Actuator 81 B. Cost Study Technical Report . . 89

Figures: 1. Servo Actuator Assembly 5 2. NWC Actuator Frequency Response on Helium 7 3. Low Cost Servo Installation - Rear Mounted Helium

Power Supply 8 4. NWC Servo Actuator Schedule 9 5. Closed-Center Pneumatic TVC Actuator 11 6. Low Cost Actuator Assembly 12 7. Servo Actuator - Exploded View 14 8. Block Diagram 20 9. Existing Laboratory Controller 21

10. Closed-Center Actuator Schematic 22 11. Stinger Alternate Flight Control Electronics 23 12. Command and Position Relationship .... 27 13. Actuator Stroke and Area 30 14. Actuator Mounted on Load Fixture 39 15. Gas Supply System 40

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NWC TP 5902

(Contd.) Cold Gas Servo Gas Supply Breadboard 41 NWC Actuator Frequency Response on Helium

(Gain 0.58o/100%) - Low Load 44 NWC Actuator Flow Rate on Helium - High Load 45 NWC Actuator Flow Rate on Helium - Low Load 47 High Load Helium Performance -12°, 1 Hz Sine

Wave Commands 48 High Load Helium Performance - ±2°, 5 Hz Sine

Wave Commands 49 High Load Helium Performance - ±2°, 10 Hz Sine

Wave Commands 50 High Load Helium Performance - 12°, 15 Hz Sine



Wave Commands 53 High Load Helium Performance - 12°, 1 Hz Square

Wave Commands 54 High Load Helium Performance - +2°, 5 Hz Square




Wave Commands 58 High Load Helium Performance -'120°, 5 Hz Square

Wave Commands 59 Low Load Helium Performance - 12C, 1 Hz Sine

Wave Commands 60 Low Load Helium Performance - 12°, 5 Hz Sine




Wave Commands 64 Low Load Helium Performance - t20o, 5 Hz Sine

Wave Commands 65 Low Load Helium Performance - 12', 1 Hz Square

Wave Commands 66 Low Load Helium Performance - 12°, 5 Hz Square

Wave Commands 67 Low Load Helium Performance - 12°, 10 Hz Squace

Wave Commands 68

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Figures 41.

42.

43.

44.

45.

46. 47.

Tables; 1. 2. 3. 4.

5.

(Contd.) Low Load Helium Performance - 12°, 15 Hz Square

Wave Commands 69 Low Load Helium Performance - t20o, i Hz Square

Wave Commands 70 Low Load Helium Performance - 120°, 5 Hz Square

Wave Commands 71 NWC Actuator Frequency Response uti Nitrogen

(with Com* ensation) - High Load 73 NWC Actuator Frequency Response on Nitrogen

(with Compensation) - Low Load 74 NWC Actuator Flow Rate on Nitrogen - High Load 75 NWC Actuator Flow Rate on Nitrogen - Low Load 76

Actuator Performance 6 Actuator Design Summary ..... . . . . V 13 Parameter List, Closed Center Actuator 26 Typical Slew Rates for Square Wave Commands,

High Load, 2,000 psi Helium 72 Typical Slew Rates for Square Wave Commands,

High Loads, 2,000 psi Nitrogen 77 Required Helium Supply for Typical Duty Cycles 80

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NWC TP 5902

TECHNICAL SUMMARY

A cold gas driven, closed-center servo actuator has been developed to operate a trapped ball type nozzle assembly on a tactical missile. One engineering test model and six flight-weight units have been delivered. Figure 1 shows the actuator assembly. Helium at a 2000 psi regulated pressure drives the balanced area actuator. The helium flow to each side of the piston is controlled by a pair of three-way electro- magnetic flapper valves which are Pulse Duration Modulated (PDM) at 100 Hz. When fully "on", a valve vents its end of the cylinder, closing the inlet port. When fully "off", the cylinder is pressurized at the valve inlet pressure. A linear position feedback transducer (potentiometer) is mounted on line with the piston. The control valve PDM ratios are commanded by an electronic controller in response to errors between the input command and the output position. Table 1 compares the specified performance requirements and the measured performance of these actuators. Figure 2 shows a Bode plot of the actuator operating at full load on room temperature helium. The servo system performance was char- acterized for operation on both helium and nitrogen gas supplies. The actuator met predictions of slew rate, frequency response, and helium flow rate.

A proposed installation of this type of actuator on a trapped ball nozzle and blast tube is seen in Figure 3. This drawing shows one means of packaging multiple small high pressure (10,000 psi) helium bottles in a cluster around the blast tube, along with the required squib valve, pressure regulator, and other attendant hardware.

This program was conducted under NWC contract N00123-76-C-0286, according to the Milestone Chart shown in Figure 4.

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TABLE 1. Actuator Performance.

| Item Specification Actual |

Nozzle Deflection ±20° ±20°

Stroke ±0.82 in. ±0.82 in. j

Moment Arm 2.40 in. 2.40 in.

Load Inertia 13-15 lb-in2 14 lb-in2 j

Slew Rate 400 deg/sec 500 deg/sec max for hard over step command

Servo Bandwidth 15 Hz 16.5 Hz | j ±1070 Amplitude (-3db or -90° (-3db and -30°

phase lag) phase lag) |

Minimum Damping Ratio \ (200 to 700 in-lb 1 load) 0.3 0.38

(+2.4db peaking | at 200 in-lb j load)

Flow Rate at 15 Hz, ±10% Amplitude Sine Commands

Hysteresis-Resolution (700 in-lb load)

Maximum Electric Power Consumption

Actuator Weight (ea)

Nozzle Friction Load

Actuator Stall Torque

Specified Duration (at -40oF)

Specified Duty Cycle for Two Actuators - Total Activity

0.015 lb/sec

10 max

35 watts

3.00 lb max

200-700 in-lb

17 sec

625c

0.0125 lb/sec

0.4° (typical)

34 watts

1.39 lb

200-900 in-lb

2928 in/lb

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1976

APR MAY JUN JUL AUG SEPT OCX

GO AHEAD

RELEASE LONG LEAD MATERIAL

TECHNICAL COORDINATION WITH NWC

PRELIMINARY DESIGN

DELIVERY PRELIMINARY DESIGN DRAWINGS TO NWC

COST STUDY

DELIVER COST STUDY REPORT TO NWC

FINAL DESIGN, ACTUATOR AND POWER SUPPLY

DELIVER FINAL DESIGN DRAWINGS TO NWC

FABRICATE ENGINEERING TEST ACTUATOR (I)

VERIFICATION TESTING OF ENGINEERING TEST ACTUATOR

DELIVER ENGINEERING TEST ACTUATOR TO NWC

FABRICATE FW ACTUATORS (6)

ACCEPTANCE TEST FW ACTUATORS

DELIVER FW ACTUATORS TO NWC (SIX)

SUBMIT TEST REPORT TO NWC

SUBMIT MONTHLY PROGRESS REPORTS TO NWC

SUBMIT FINAL REPORT DRAFT TO NWC

SUPPORT TO NWC TESTS

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SYSTEM DESCRIPTION

The closed-center pneumatic nozzle actuator is designed to operate a trapped ball nozzle assembly in response to an applied electrical command. An integral position transducer is used as the feedback element in a closed loop control system. Helium was selected as the working fluid because of the higher response characteristics it provides.

The actuator is shown schematically in Figure 5. The drawing of Figure 6 shows the details of the control valves and the actuator assembly. Design and operating characteristics of the actuation systems are listed in Table 2,

The individual actuator assemblies each comprise two three-way valves, a double-acting piston/cylinder assembly, and a potentiometer position transducer. A spherical rod-end bearing on the piston rod and a clevis on the housing provide means for attachment to the nozzle and motor aft closure, respectively. An exploded view is seen in the photo- graph of Figure 7.

To obtain pitch and yaw control of the TVC nozzle, two actuators are used. The power supply comprises one or more hermetically sealed pressure vessels (depending on the installed location), an explosive actuated valve, and a pressure regulator (see Figure 3).

The two actuators are each mounted on the blast tube with clevis fittings that allow for alight angular motions of the actuators resulting from X and Y coordinate motions of the nozzle attachment fittings. Since the nozzle lugs are located in the same plane as the nozzle spherical seal centerline, each actuator's angular motion occurs only in one plane. No significant cross talk occurs between the actuators. The slight angular motion (±0.6 degree) of each actuator requires that the gas inlet tubes have sufficient length to permit a small amount of flexing. The package arrangement requires inlet tube lengths from the gas distribution manifold which are adequate to provide the required flexibility.

LOAD DEFINITION

Two identical actuators are used to operate the trapped ball nozzle in two planes. The nozzle deflection can be up to ±20 degrees in any plane. The moment arm at the nozzle attachment point is 2.40 inches, requiring an actuator stroke of ±0.82 inch. The specified nozzle inertia is 13 to 15 in^-lb or 0.0363 in-lb-sec^. The specified maximum slew

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TABLE ? . Actuator Design Summary.

Item Value

Nozzle Travel ±20 degrees

Moment Arm 2.40 in

Stroke ±0.82 in

Load Inertia 14 lb in.2

(0.0363 in-lb-sec2/rad)

Load Torque 200-700 in-lb

Maximum Slew Rate 500 deg/sec

Servo Bandwidth (±10% Amplitude) 16.5 Hz (-3db)

Position Resolution 0.4 degree

Piston Diameter * 0.937 in.

Rod Diameter 0.312 in.

Effective Area 0.61 in.2

Inlet Gas Pressure 2000 psig

Maximum Stall Force 1220 lb

Control Valve Orifice 0.038 in. diam

Orifice Flow Coefficient 0.65

Design Flow Rate 0.015 lb/sec

PDM Frequency 100 Hz

Valve Coil Resistance 23 ohm

Coil Voltage 28 vdc

Valve Current Per Actuator 1.22 amperes

Actuator Assembly Weight 1.39 lb

Installed Length 8.75 ±0.81 in.

Duration ■

17 sec

Duty Cycle - Total Activity For Two Actuators 625°

Storage Volume - Helium at 10,000 psi, 140oF 30-60 in,^ depending on

type of duty cycle

Regulator Flow Rate 0.030 lb/sec

13

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NWC TP 5902

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rate is 400 degrees/second. The trapped ball nozzle torque load is primarily coulomb friction ranging from 200 to 700 in-lb. The minimum required frequency response of the nozzle actuator system is 15 Hertz at ±10 percent amplitude command (±2 degrees) with a maximum of 90 degrees phase lag or -3db amplitude ratio, whichever occurs first. The operating duration of the servo system is 17 seconds and the duty cycle requirement includes a total activity for the two actuators of 625 degrees of nozzle motion.

MECHANICAL UESCRIPTION

The closed-center nozzle actuator assembly is shown in Figure 6. Electrical connections are required for the position transducer and the control valves. The cold gas power supply will be connected at the point marked "inlet" on the drawing.

CONTROL VALVES

Two control valves are used on each actuator. These are three-way valves with the inlet-to-cylinder passage normally open and the cylinder- to-vent passage normally closed when pressurized.

Each valve has a flexurally mounted armature (flapper) within a cylindrical body which incorporates magnetic pole pieces and restrictors, The flapper is an assembly made of iron/cobalt magnetic material resistance welded to a heat-treated stainless stfal flexure and tapered mounting piece. The valve body is made of type 3A7 stainless steel with magnetic polo pieces brazed in place. Valve seat orifices of 347 stainless steel are threaded into the body, staked, and sealed with Loctite to prevent movement and leakage. The orifices have a throat diameter of 0.038 inch.

The valve coil assembly consists of two individual coils, each wound with 417 turns of #33 magnet wire. The two coils are connected in series to provide the actuating force for the flapper. The total coil resistance is 23 ohms. The coils are encapsulated in epoxy potting com- pound to prevent damage to the coils and to help anchor the lead wires.

ACTUATOR CYLINDER ASSEMBLY

The actuator uses a balanced-area double-acting piston operating in an aluminum cylinder block. The cylinder block has provision for mounting the two control valves and contains internal flow passages to duct gas to valves and cylinders. O-ring seals are used on the piston, around the piston rods, and on the cylinder end caps. Teflon slipper seals are used to minimize breakaway friction on the piston and rod O-rings. The effective piston area is 0.61 in^. The housing and cylinder is 7075 T6 aluminum with a solid film lubricant deposited on the cylinder bore.

15

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The control valve mounting seats are an integral part of the cylinder block. The end of the valve bodies are tapered and fit into matching tapers in the cylinder block. Teflon tape seals are installed between the valves and the cylinder block to prevent leakage and cross-flow between ports.

Lock nuts on the outside of the cylinder hold the valves in place and maintain a preload on the seals. Pressure taps in the ends of the valve bodies are provided to monitor valve body pressure for testing purposes. Gas flow to and from the valves is through passages drilled in the cylinder block.

The actuator piston is made from type 416 stainless steel, heat treated to a hardness of 34 Rockwell C. The nozzle-end piston rod is machined as an integral part of the piston. The outer end of this rod is threaded to accept a rod end bearing. A dummy piston rod is used at the other end to pressure balance the effective piston area so that equal forces are generated by equal pressures acting on both sides of the piston. This second rod, which is threaded onto the piston, is integral with the position transducer, as described below.

All dynamic 0-ring seals are Viton with Shamban Glyd rings.

The NMB piston rod bearings are Teflon impregnated fabric mounted in aluminum housings. The sleeve bearings for the piston rods are mounted in the cylinder end caps with a lifcjht press fit.

POSITION TRANSDUCER

The position transducer is a linear potentiometer mounted concen- trically with the piston in the lower end of the actuator, as shown in Figure 6. A unique feature of this installation is that the potentiometer shaft is the same size as the nozzle-end piston rod; when it is threaded onto the piston the area of this shaft reduces the piston area by the same amount as the piston rod. This provides equal areas on both sides of the piston. The potentiometer was made by Bourns for this application.

The transducer uses a ceramic resistance element and a precious metal slider pickoff, both mounted in a common housing. The housing mounts in a bore in the lower end of the actuator; it is retained by a shoulder at one end of the bore and by the cylinder cap at the other end. Electrical leads are brought out of the actuator next to the clevis at the lower end of the unit.

16

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ACTUATOR OPERATION

The closed-center actuator concept was selected for this application because it minimizes gas consumption by requiring gas flow only when a new position is commanded, providing the most efficient gas usage. In addition, higher servo bandwidth can be obtained than with a similar size open-center actuator.

The actuator schematic, Figure 5, shows the actuator in the pressurized and de-energized condition (no electrical command applied). For this condition the control valve flow passages from inlet to cylinder are normally open, and the vent ports are normally closed. Gas from the power supply enters each actuator inlet and flows through the valve inlet ports into the cylinder cavities on both sides of the piston. Since there is no vent flow the pressure on each side of the piston builds up to regulated supply rressure. Both sides of the piston have equal area, therefore there is no net force generated by the actuator and the piston will remain in position. When the cylinder is fully pressurized there is no further flow demand on the power supply, hence the description "closed-center" actuator.

When a new actuator position is required, one control valve vents gas from its side of the piston in response to an electrical command. The other valve has no command applied. It remains as shown in Figure 5, and the pressure on that side of the piston remains at the original supply pressure level. The resulting pressure differential across the piston generates the force necessary to move the piston to the commanded position. When the new position is reached the electrical command is terminated so that both valves are again de-energized, the previously vented cylinder pressure builds up to supply pressure level, and the piston remains at rest with no gas flow demand.

Since the load characteristic of the trapped ball nozzle is primarily coulomb friction with essentially no spring component, a significant actuator force is always required to move the nozzle or, conversely, the nozzle will remain at its last commanded position unless acted upon by a significant actuator force. This load characteristic permits use of the closed-center actuation system to full advantage.

VALVE OPERATION

The moving element in the control valves is a cantilevered flapper which is pressure loaded against «-he vent orifice in the absence of a command. When an electrical command is applied to the valve the magnetic force generated by the solenoid pulls the flapper away from the vent orifice and seats it on the inlet orifice. This seals the inlet and opens the vent port. When the command is removed, the combined action of the inlet pressure and the cantilever spring force move the

17

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flapper away from the inlet orifice. The pressure times vent nrea force then forces the flapper against the vent orifice.

The command used on the control valves is a Pulse Duration Modulated (PDM) signal which is proportional to an input analog voltage command. In PDM operation the valve flapper cycles between the vent seat and the inlet seat at a constant cyclic rate (the PDM frequency). The flapper dwell time on either seat can be controlled from 0 to 100 percent of the cycle period. The PDM command is expressed on a scale from 0 to 100 percent where 0 percent represents continuous dwell on the vent seat (no command applied), 100 percent represents continuous dwell on the inlet seat (full dc command), and any command between 0 and 100 results in the flapper dwelling part of the cycle time on the inlet seat and the remain- ing part of the cycle time on the vent seat. Since the flapper motion time is extremely short, the relationship between percent PEM command and dwell time on the inlet seat may be expressed as

%PDM = in. t. in.

+ t or vent

t. in.

1/f

where

t. = dwell time on the inlet seat, sec m.

t = dwell time on the vent seat, sec vent

f = PEM frequency, Hz

A zero percent PDM command applied continuously will result in full cylinder pressure. A 100 percent command applied continuously will result in zero cylinder pressure since the vent is open continuously. A conniand between zero and 100 percent, applied continuously, will result in an intermediate cylinder pressure since the control valve modulates the flow into and out of the cylinder.

Remembering that each control valve establishes pressure on one side of the piston, it can be seen that, with appropriate commands to the two control valves, the actuator piston can be moved in either direction at any desired rate up to the maxinum permitted by gas flow capability. A 100 Hz PDM frequency has been selected to meet the performance requirements and to provide smooth actuator motion.

18

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NWC TP 5902

ELECTRONIC CONTROLLER

Figure 8 shews the control system block diagram used for the closed center actuator. Each valve is controlled by a separate electronic driver (a transistor switch with current control) and PDM modulator with gain control. The separate modulators provide the capability of controlling the loop gain, depending on the direction of motion. Since there will be some variation in the performance of any two control valves, the controllable electrorrc gain permits matching of the overall loop gain regardless of which valve is modulating. This provides symmetrical actuator performance without imposing the requirement of precise control valve matching.

During actuator operation one valve is modulated to cause the piston to extend and the other valve is used for retract motion. A position error signal is formed by taking the difference between the commanded position and the actual position; the magnitude or absolute value of the error signal is applied to each modulator through the appropriate gain. The valve to be modulated is selected by the error sign detector which inhibits the valve driver for the "off" valve. The sign detector operation is designed to be consistent with negative feedback. The sign detector switches command from one valve to the other as the error signal passes through zero and changes sign.

Each modulator also contains a "crossover bias" adjustment that is used to prevent any valve operation until the error signal reaches a preset value. This capability can be used to create a deadband to prevent sporadic switching between valves for small error signals.

Figure 9 shows the laboratory controller used. The mechanization of the controller is shown schematically in Figure 10. This controller used discrete components with hand wired circuit boards and plug-in oper- ational amplifiers. Since space is not at a premium in a laboratory controller, the convenience of having access to the circuitry dictates this type of construction.

For bench testing or static firings it is suggested that a controller similar to this be used by NWC. For flight use, the controller would be redesigned to use hybrid circuits, integrated circuits, and other space- saving, . environmental resistant techniques.

For small quantities of flight hardware, a package similar to that shown in Figure 11 might be developed. This figure shows the Flight Servo Control assembly for the Aeronutronic Stinger Alternate missile. This package is 3.3 inches OD, 2.0 inches ID, and 2 inches long, and contains all the control circuitry for four individual fin actuators. Printed circuit boards are used, with hand wiring between individual components, hybrid circuit packages, and film resistors.

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In production quantities, this package size can be further reduced by the use of integrated circuit chips and circuit boards designed for more automated assembly. In production, these circuits would probably be integrated into other electronic modules on the missile.

WEIGHT

Each actuator assembly includes the inlet manifold from the 1/4-inch inlet fitting, the actuator housing, values, output rod and rod end bearing, and position feed back potentiometer. This assembly weighs 1.39 pounds.

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DESIGN ANALYSIS

The actuator sizing analysis is presented in this section.

The Parameter List of Table 3 is used to identify variables in the calculations presented below.

ACTUATOR SIZING

Gas Flow Rate Determination

To determine the size of the control valve orifice, it is necessary to determine the maximum instantaneous flow rate required by the actuator. Each component of the required gas flow is calculated below.

Friction Flow Rate - Direction Reversal. When operating the servo in response to a sine wave command, it is necessary to reverse direction at the end of each cycle. With a large coulomb friction load, the differential cylinder pressure must be reversed at the end of a cycle from a value to overcome friction in one direction, to a value to overcome friction in the opposite direction. The friction load always acts to oppose motion. If it is desired to follow a sinusoidal command of a certain frequency with an effective phase lag of 90 degrees, then a required pressurizatlon rate for the cylinder is defined. Figure 12 illustrates the command and actual position at the frequency where 90 degree phase lag is desired.

The required pressurizatlon rate can be defined as follows: The applied torque necessary to overcome friction must change from -Kp to -Tp as the position changes direction (velocity changes sign). For the

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TABLE 3. Parameter List, Closed Center Actuator.

Symbol Item Value

9 m Maximum slew rate (specified) 400°/sec

T F

Load friction torque 200 to 700 in-lb

J Nozzle inertia io i u -i—^ *-.— r\ no/. iu 2 -lb-sec

li ib-m or ü.U3^ rad

e m

Maximum vector angle ±20°

i Moment arm (3 0° 2.4 in.

X Linear stroke ±0.82 in.

T Rankine gas temperature 460° + 0F

R Specific gas constant 4635 in./0R (Helium)

f Frequency Hz

tu Angular frequency rad/sec

W Gas weight lbs

ft Gas weight flow rate lbs/sec

A Effective piston area 0.61 in.2

dp Piston diameter 15/16 in.

dr Rod diameter 5/16 in.

dv Valve orifice diameter 0.038 in.

Av Valve orifice area -3 2 1.13 x 10 J in/

C m

Mass flow coefficient (helium) 14.26 Tin/sec

cd Orifice discharge coefficient 0.65

£© Unchoking factor 1.0 for IÔ^S d

Ps Supply pressure 2000 psi

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COMMAND

TIME

FIGURE 12. Command and Position Relationship.

output position to be only lasting the command by 90 degrees, the torque reversal must occur in a quarter of a cycle or in t/4 seconds (t * 1/ frequency).

The required torque rate (T) is

^ or T = 8 f TF

4f

where

f = frequency, Hz

T = coulomb friction torque, in-lb

For the piston actuator, the output torque is defined as

T = APA£

I

where

T = applied torque (in-lb)

A? = differential pressure across piston (psi)

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NWC TP 5902

2 2 A = effective piston area (in. ) ■ 0.61 in.

i = moment arm • 2.40 in.

differentiating

T = APAÄ

for the closed center actuator

Ap ss p . p S

| AP | ~ P

combining equations

PAC = 8fT

8Jh PJL

from the ideal gas equation of state

PV + VP = WRT

V -♦ 0 for small amplitude motion

P = WRT V

28

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combining Che P equations

HSI = 8fTF

V = AÄ

V = XA

X = midstroke head clearance

A = effective area

The flow rate required to overcome friction at a certain frequency is

W„ = 8fTFX

F £ R T

This equation defines the flow rate needed to follow a small amplitude sinusoidal command with 90 degree phase lag at a frequency, f.

For the worst case conditions

T = 700 in-lb F

T = 420oR (-40oF)

f = 15 Hz

i = 2.4 in.

X = 0.82 in.

Substituting:

wF = 8\15 CIec/(700 in-lb) <0-82 ln-)

2.4 in. (4635 ^)(420oR)

WF = 0.0147 lbs/sec (helium)

29

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m/C TP 5902

Note that this factor is independent of actuator area. This is explained by the following example. Consider three piston actuators of different areas as shown in Figure 13 but all with the same stroke, moment arm, and coulomb friction load. For Figure 13a, if the required torque rate, T, to move the output at a certain frequency f = PA, for small motions the flow rate required is

W = P V RT

1 x_ A RT

In Figure 13b, to maintain the same t, the pressure rate must be doubled, But, since the volume is halved, the flow rate remains the same. Like- wise, if the area is doubled (Figure 13c), the pressure rate is halved, but the volume is doubled. Hence, the flow rate again remains constant.

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2P, 1/2A

I ^ ! z^

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I (a)

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Inertia Flow Rate. Similar logic can be used to determine the flow required for controlling the load inertia.

32 n2 j 9 f3 x W 2 1 WJ £ R T

where 9o is the peak sinusoidal position amplitude for a ±10 percent excursion

9o = — = 0.0349 rad

57.3"/rad

Substituting:

V 32 ^ I0'034 rad1"—) <0-Q349 rad) (15 Hz/sec)J (0.82 in.)

(2.4 in.) (4635 in./0R) (420OR)

W = 2.22 x lO"4* lbs/sec which is additive with the direction reversal flowraf.e, Wp.

Sinusoidal Velocity Flow Rate. The flow rate required while following a small amplitude sine wave is given by:

2TT p0 A f e £ A. = § o_ W9 R T

The basis for selecting the actuator area will be discussed later,

31

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NWCTP 5902

For A- 0.61 in.', substituting

W^ =

/ 2Wn fti in 2U15 ^-)(0.0349 rad)(2.4 in.) 2TT (2000 lbs/in. )(0.61 in. ) U3 sec^ |

(4635 in./OR)(420OR)

W. =4.95xlO-J lbs/sec =0.005 lb/sec 9

Since direction reversal and maxiraum sinusoidal velocity occur 90 degrees apart when following a sine wave, the larger of the two terms govern the requirements.

The direction reversal terms are

i + « - (2.22 X IC"4) + (1." X IQ'2) lb./.« F J

= 0.0149 lbs/sec

The velocity term is

WA = 0.005 lbs/sec

It ls obvlous ^ .e „c^ ^„..X ^s Sover„ (« - 0.0.3 ./see,

■ T?I„„ Rai-P The required maximum F.uU^roUeJlewRat^^ uo Let a given slewTSTe is 400 deg/sec or 7 rad/sec. slew rate is:

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Substituting:

2000 i~- (0.61 in.2)(7 rad/8ec)(2.4 in.) in.

(4635 in./0R) (420OR)

Wi " 0.0105 lbs/sec which is less than the 0.015 lb/sec M required for direction reversal.

The available pressurization rate in a hardover position is given by

WR T

«here V is double the midstroke volume (total displacement) and W is maximum available flow rate.

(0.015 lbs/sec)(4635 in./0R) 420OR

(0.61 in. ) (1.64 in.)

P ■ 2.92 x 10 psi/sec

T B P A £ » (2.92 x 104 ^ (0.61 in.2) (2.4 in.) sec

T = 4.3 x 10 in.-lbs/sec

The time to generate 700 in-lbs is

..? 700 in-lb 4

4.3 x 10 in-lb/sec

At = 16 milliseconds.

33

NWC TP 5902

Since the maximum available flow rate is greater than the minimum required (due to frequency response considerations) just to meet specified slew rate of 400 deg/sec, the slew rate for a hardover valve command is given by

(0.015 lbs/sec)

e = W RT . ? k i'

h t) (420OR) 2000

lbs

in/ •\(o- 61 in. )(2.4 in.)

fi =9.97 rad/sec or 571 deg/sec.

The no load time from stop to stop is 40 deg/571 deg/sec, or 70 milliseconds. If the effect of friction on slew time is considered, then the total slew time for 40 degrees is 0.016 + 0.070 ■ 0.086 sec, for a slew rate of 40/0.086 = 465 deg/sec.

Notice that the assumed pressure rate is dependent only on the venting cylinder chamber. This is not true in reality, since gas must enter the pressurized cylinder chamber to make up for volumetric flow used in slewing. Since this flow rate (and the cylinder pressure) depend on the slewing velocity, the problem becomes closed loop and must be solved by computer simulation to include this term.

Valve Orifice Sizing

The calculations above have computed the maximum instantaneous flow rate required co achieve a 15 Hz bandwidth at -40oF. This flow rate, 0.015 lb/sec, defines the valve orifice size. It is assumed the servo control electronics have been adjusted so that the valve is achieving maximum flow as demanded by the servo.

The valve orifice diameter may be calculated from the flow equation

W = CJ Cm A P f(Pu/Pd) d v s

34

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NWCTP 5902

The orifice area is given by

TTd Av

and since the valve orifice is venting a pressurized cylinder chamber to atmosphere, choked flow nearly always occurs so that f(Pu/Pd) "1.0.

,2 4 WVR T C,, C PU rr d m

4 (0.015 lbs/sec) \(4635 in./OR)(420OR)

(0.65)(14.26)Vin^/sec (2000 lb/in.2) (n)

1.44 x 10"3 in.2

0.038 in.

The maximum size of the valve orifices are limited by the magnet and spring force available to overcome the P x A force created by the orifice area and the supply pressure. The flapper valve in this configuration has been found to perform best when the vent and inlet orifice areas are the same.

<

Piston Area Determination

Piston area is determined largely by the need to accommodate a widely varying friction load. Of course, the minimum stall force of the actuator must be sufficient to overcome the largest value of coulomb friction. Excess force capability over and above this value is dictated by servo stability.

From results obtained in a feasibility study simulation, it was found that scall torque to coulomb friction torque ratio should be 3.7:1 to provide sufficient force margin to be insensitive to fricLion load variations. This provides high load iesponse without incurring low load instability.

35

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NWC TP 5902

Since

T stall = 3.7

If TF = 700 in-lbs, then T3tall = 2590 lbs. In order to keep the piston

size relatively small, it is desirable to use a high system supply pressure. It is also necessary to minimize the piston diameter to permit packaging the actuator in the annular envelope available. With Pg = 2000 psi and a moment arm of 2.4 in., the required area is

A = 2590 in-lbs

P I s 2000 ^^2 (2-4 in'>

in.

= 0.54 in.'

From mechanical considerations it is desirable to have a 5/i6-incn diameter piston rod. Therefore the piston size is calculated as

A =

2 2 TT (d^ - d^)

d2 ^ A ^ ,2 P = ' T" U

TT

d2=4_i0^4I+ (0.3125)2

P "

d = 0.886 in. P

The closest convenient diameter that will yield at least as much area as a 0.886 diameter piston tive area of the piston is as a 0.886 diameter piston is dp = 15/16 in. diameter so that the effec-

IT (0.93752 - 0.31232) n ,, . 2 A = —* ; *• ■ 0.61 in. . P *

This diameter permits packaging in the envelope,

36

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NWC TP 5902

CONTROL SYSTEM

The basic operation and block diagram for the Control System has been shown in Figure 8. The circuit schematic that mechanizes the block diagram is shown in Figure 10.

A single integrated circuit generates the 100 Hz carrier triangle wave used by both the PDM modulators. One hundred Hz has been found to be the highest PDM frequency that is practical for the 5/16-inch diameter flapper valves. The valve coils are designed so that 28 Vdc power can be used as the input voltage for the driver stage; ±12 Vdc is required to power the integrated circuits and feedback pot.

Each valve coil has a dc resistance of 23 0. With a supply voltage of 28 Vdc, the total power dissipated per actuator is 34 watts when a valve is hardover. No external power resistors are required. For missile battery sizing, the duty cycle required should be taken into account since, when the servo error signal goes to zero, no dc power is used except that required to maintain ±12 Vdc across the integrated circuits.

There are three loop gain adjustments. Each modulator has its own gain for equalizing differences between valves, the third gain control operates directly on the error signal and allows setting the overall gain with one control.

The "cross-over" adjustments are essentially dc bias terms that are added to the modulator input. Since the valves normally require 20 to 25 percent PDM command to begin pressure modulation, these bias controls can be used for varying the system pressure deadband. To maintain closed center operation, there must be some finite deadband. In practice, it has been found that a 5 to 10 percent overlap produces the best results; the deadband is at a virtual minimum.

One of the primary control considerations is to maintain a high bandwidth, high gain system while keeping a good damping ratio, regardless of a load that can vary between 200 and 700 in-lbs. This problem was addressed by computer simulation. The simplest and most direct way found was to size the actuator so that tae load appeared to be a relatively insignificant percentage of the total stall force capability. Simulations indicated a ratio of stall force to load force of about 3.7:1 produced good results. Electronic compensation aids in stabilizing the servo.

An electronic lead-lag network of the form

10 S + 200 \ S + 2000/

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NWCTP 5902

is used. The equalization is mechanized in state-variable form (analog computer style) to simplify experimental location of the poles and zeros,

TEST RIG DESIGN

The Thiokol trapped ball nozzle assembly was not available for this program, and it was necessary to provide a test rig with similar friction, inertia, and kinematics. Figure 14 shows the servo actuator mounted in the Aeronutronic load fixture. A rigid bracket containing a self-aligning bearing (NMB ABYT4) supports the fixed end of the actuator. The output rod end is attached to a 2.40 inch radius crank arm, simula- ting the nozzle attach point with an 8.75 inch center distance. The crank arm is welded to a 1.00 inch diameter shaft, and is straddle mounted between a pair of pillow block bearings. Cantilevered from one end of the shaft is a 6.00 inch diameter Tol-O-Matic disc brake. Two disc brake caliper assemblies are mounted on the test stand structure and grip the brake disc at two diametrically opposite points. The caliper assemblies are pressurized with nitrogen to establish the desired friction torque levels. A smaller disc is cantilevered from the opposite end of the crank arm shaft. Its purpose is to provide additional moment of inertia to the load assembly. This disc is sized to bring the total moment of inertia of the crank, shaft, and disc assembly up to the specified value of 14 in.2-lb or 0.0363 in-lb-sec^/rad. The moment of inertia was confirmed experimentally by the torsional pendulum technique and the final dimensions of the add-on inertia disc were selected to achieve the required value. The test stand friction torque was calibrated experimentally by hanging known weights on the crank arm and adjusting the brake pressure to just allow slippage. Since this technique basically determined the static friction of the brake assembly, the dynamic friction was obtained during actuator testing by measuring the required F x A forces to maintain constant slew rates. The high load condition of 700 in-lb torque required a brake pressure of 275 psig, while 60 psig provided the 200 in-lb low load torque.

The test setup also included a high pressure gas supply system made up of typical missile system components. It is seen in the photo of Figure 15. The gas supply system schematic is shown in Figure 16. This system utilizes many missile-type flight weight components, combined with necessary instrumentation and lab control elements. Two 71 in.^ gas storage bottles are manifolded together to form the primary storage volume. These are bottles manufactured by Apco for use on the Maverick Missile. Their hermetically sealed fill and run components have been removed to facilitate making repeated tests in the lab. They are filled through a solenoid valve from the facility high pressure (6600 psi) helium supply. The bottles are connected to a solenoid run valve and then to an Altair 215 R 7000 flightweight pressure regulator. The regulator is rated up to 10,000 psi inlet and regulates the downstream pressure to 2000 psi. A 2500 psi relief valve is installed in a tee on the

38

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FIGURE 15. Gas Supply System.

40

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SERVO SYSTEM

HASKELL RELIEF VALVE 2400 PSI

ALTAIR PRESSURE REGULATOR MODEL 2-15R700

MAROTTA SOLENOID VALVE 24 VDC MODEL MV100H 6000 PSI"

PRESSURE DUMP i/^GROVE VALVE

-^^< MODEL 108985

HASKEL CHECK VALVE

/

HELIUM 6000 PSI SUPPLY

.PRESSURE TRANSDUCER VIATRAN MODEL 103 15000 PSI

'THERMOCOUPLE GAS PROBE

KJ APCO RESERVOIR MODEL 808900 7450 PSI

FIGURE 16. Cold Gas Servo Gas Supply Breadboard.

41

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NWC TP 5902

line between the regulator and the actuator inlet fitting. Its purpose is to prevent excessive downstream pressures due to regulator leakage when the closed center actuator valves are shut off. Actuator flow measurements are taken from this system by monitoring initial and final bottle pressure after allowing the system to temperature stabilize.

42

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NWC TP 5902

TEST RESULTS

The servo actuator performance was obtained using the friction- inertia test rig described earlier. Most tests were duplicated using both helium and nitrogen as the working fluids. All tests were run at a nominal gas supply temperature of 60°F, and a regulated servo inlet pressure of 2000 psig. The valve PDM frequency was maintained at 100 Hertz. Two levels of friction torque were established, high load (700 in-lb) and low load (200 in-lb). The test rig moment of inertia was 14 Ib-in^ or 0.0363 in-lb-sec^/rad. The crank arm radius of 2.40 inches resulted in a ±0.82 inch stroke at ±20 degree output angle.

The electronic controller gain settings were optimized for helium operation and the same control settings were used for both the helium and nitrogen tests. Tests conducted included:

1. High and low load frequency response (Bode plots and phase lag) at ±2 degree amplitude commands and frequencies ranging from 0.5 Hz to 50 Hz.

2. Differential pressure versus stroke at low slew rates.

3. Slew rate and gas consumption tests for a series of 10 second runs on the 145 in^ gas supply bottle system. These tests included high and low loads, sine and square wave commands, amplitudes of ±2, ±5, ±10, ±15, and ±20 degrees, and frequencies of 1, 5, 10, and 15 Hertz. Command and output position were recorded on oscillograph traces along with the piston differential pressure. The bottle supply pressures before and after each run were recorded after temperatures were stabilized to permit calculating the flow rate.

4. Valve component response tests using a fixed cylinder volume. Valve output (cylinder pressure) versus % PDM command was presented on a X-Y plotter.

5. Acceptance tests according to the acceptance test plan.

HELIUM PERFORMANCE

Figure 2 shows a typical Bode plot under high load conditions, and Figure 17 shows the low load frequency response and phase lag. Figure 18 shows plots of calculated flow rates under high loads for both sinusoidal

43

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NWC TP 5902

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FIGURE 18. NWC Actuator Flow Rate on Helium - High Load.

45

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NWC TP 5902

and square wave commands. It is seen that flow rate is primarily a func- tion of frequency and is relatively insensitive to command amplitude. The flow rate at ±2°, 15 Hz sine or square wave commands is about 0.012 lb/sec at high load and 0.014 lb/sec at low load. Figure 19 shows a similar plot of flow rates for low load conditions.

Figures 20 through 43 shows a series of typical oscillograph traces for a range of commands. The actuator differential pressure traces are shown below the command and position traces. The pressure trends shown are typical, but the accuracy of the absolute magnitude of pressure is questionable due to a possible error in recording technique. Average slew rates can be obtained from the square wave command data. Table 4 summarizes the slew rates determined from the high and low load helium

tests.

Helium performance tests were also run using large amplitude triangle wave commands, starting with low frequencies, to determine the actuator pressure differential required to slew the load fixture at low slew rates. Higher frequencies commanded higher slew rates, permitting an investigation of the dynamic load characteristics to see if there was any viscous damping content in the applied load.

NITROGEN PERFORMANCE

Figures 44 and 45 show typical Bode plots of frequency response and phase lag for the actuator under high and low load conditions. The gain settings for the controller are those which gave optimum helium performance. The required average flow rates operating with nitrogen for a range of commands were obtained by the same techniques used with helium. Figures 46 and 47 present the nitrogen flow rate data for high and low loads, sine and square wave commands, as functions of frequency and command amplitude. Table 5 summarizes the slew rates obtained from the nitrogen testing.

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FIGURE 19. NWC Actuator Flow Rate on Helium - Low Load.

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0.1 SEC

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EXTEND

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FIGURE 32. Low Load Helium Performance - ±2°, 1 Hz Sine Wave Commands.

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FIGURE 38. Low Load HeLium Performance - ±2°, 1 Hz Square Wave Commands.

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FIGURE 39. Low Load Helium Performance - ±2°, 5 Hz Square Wave Commands.

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NWC TP 5902

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FIGURE 40. Low Load Helium Performance - ±2°, LO Hz Square Wave Commands.

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FIGURE 42. Low Load Helium Performance - ±20°, 1 Hz Square Wave Commands,

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TABLE 4. Typical Slew Rates for Square Wave Commands, High Load, 20C0 psi Helium.

Coraraand Output Slew Rate (deg) (deg/sec)

±2°, 1 Hz ±1.83 91.7

±2°, 5 Hz ±1.89 94.4

±2°, 10 Hz ±1,89 94.4

±2°, 15 Hz ±1.46 87.4

±10°, l.Hz ±9.7 391.0

±10°, 5 Hz ±9.7 391.0

±15°, 1 Hz ±15.0 394.0

±15°, 5 Hz ±12.6 420.0

±20°, 1 Hz ±19.5 432.0

±20°, 5 Hz ±11.3 452.0

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NV NWC ACTUATOR FLOW RATE ON NITROGEN

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TABLE 5. Typical Slew Rates for Square Wave Commands, High Loads, 2000 psi Nitrogen.

Command 1 Output | (deg)

I Slew Rate | j (deg/sec) j

| ±2°, 1 Hz ! ±1-9 1 114 i ±2°, 5 Hz ±1.9 125 1

| ±2°, 10 Hz ±0.8 1 25

±2°, 15 Hz ±0.5 27 !

| ±5°, 1 Hz ±4.9 140 j

±5°, 5 Hz ±5.1 152

±5°, 10 Hz ±0.9 26 j

1 ±10°, 1 Hz ±10.0 168 !

| ±10°, 5 Hz ±6.4 188 |

±15°, 1 Hz ±15.2 186 |

| ±15°, 5 Hz ±7.2 212 |

I ±20°, 1 Hz ±19 198 j

j ±20°, 5 Hz ±6.9 222

1

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GAS SUPPLY SIZING

Flow measurements in this report were made by monitoring the initial and final pressures in the 145 in^ gas supply system for 10-second runs at the various operating conditions. To avoid large temperature effect errors, the stable initial and final pressures were used to calculate weight of gas used. For an assumed gas temperature of 80°F, and 10- second runs, the nitrogen flow rate was found as follows:

W GN„

. 3 (P1-P2) psig x 145 in

55.16 ~^ x 12 ^ x 540oR x 10 sec

W GN„

P -P 1 2

24651 lb/ sec

For helium the flow rate is;

W HE

(P^) psig x 145 in"

386.30 TJT^ X 12 7S x 540OR X 10 sec io R rt

W, HE

P -P I 2

172636 lb/ sec

In conducting the flow rate tests, it became apparent that as the gas supply bottle pressure drops rapidly during periods of high actuator activity, the gas storage temperature drops accordingly. Thus, the actual gas flow rate increases because of the higher gas' density flowing into the actuators. For short duration, high activity duty cycles, this sit- uation can have a significant effect on gas bottle sizing because little time is available for the gas supply system components to transfer their heat into the cold helium. Consideration must be given to actual system mass, surface areas, heat capacities, flow rates, and duty cycles to optimize t le storage volume requirements. The flow rates for steady ten- second dut ' cycles, shown in Figures 18 and 19 for helium, and Figures 46

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and 47 for nitrogen, were based on initial and final bottle pressure measurements after temperatures had stabilized for 10 to 15 minutes.

In order to provide some preliminary bottle sizing data for typical duty cycles, the actuator was also run at high load on helium following commands which would result in a total activity of 625 degrees in approximately 16 seconds of operation. Typical commands were ±5 deg at 2 Hz, ±10 deg at 1 Hz, and ±20 deg at 0.5 Hz. Both sine and square wave commands were used. Table 6 summarizes the results of these tests utilizing the 145 in test gas supply system. The actual weight of gas used was calculated from the stabilized pressure data and converted to a required bottle volume, based on 10,000 psi initial storage pressure at 140oF. A more conservative bottle volume was also calculated based on typical bottle pressures immediately after shutdown, before the temperatures restabilized. This approach simulates typical missile installations and requirements.

Note that square wave commands require less gas consumption than sinusoidal commands.

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NWC TP 5902

APPENDIX A

ACCEPTANCE TEST PROCEDURE

NWC NOZZLE ACTUATOR

D120-76-151 Revised 9/30/76

t. QMM/7 UcodO R. Levl 0 Ck H» Lewis, Supervisor

Systems Test Section

MOTE: 1. Tests to be performed using P/N 51040 test stand.

2, Record all data on brown-llna copy of this procedure.

Aeronutronic Aeronutronic Ford Corporation Aeronutronic Division Newport Beach, Ca. 92663

AHDSin V7»

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NWC NOZZLE ACTUATOR

ACCEPTANCE TEST

WN 51030 S/N oo4 A. Proof Test ™A static Leakage

1. Mount assembled actuator on test stand and connect inlet to

N source. Apply 3000 psi for 3 minutes. No visible permanent

deformation allowed.

Technician

9/2L9/ 7e> Date

2. With 2000 psi N2 applied attach a flow meter (rotameter) to

each exhaust port and record the leakage flow rate.

Valve E OjJ^S SCFH e'" ^ (Piston Rod End)

S/N

Valve A O ^ [jt^ SCFM (Clevis End)

S/N 15

Leakage for each valve must be lees than 0.4 SCFM.

Technician

3/2.9 /76 Date

3. Inspect piston rod end seals and flanges for leakage. Also

Inspect valve mountings including nut and seal screw. All

Joints to be bubble tight.

Technician

g /W^ Date

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B. Valve Characteristics NWC TP 5902 P/N 51030

S/N QQ^f

Coll Resistance

1. using a resistance bridge, measire each coll resistance and record,

Valve A "jj-j-* gj-; Q Valve B 2,2-. 6 7 Q

Resistance must be 23 Q + 1 Q.

Technician

9/zv/H Date

2. Static Pressure Characteristics

Install pressure transducers. Mount actuator on test fixture

and lock In mldstroke by use of the centering bar. Apply

brake pressure of 500 psi. Apply 2000 psi N. to inlet. Record

differential cylinder pressure vs. PDM command (+ 1007.) on the

X-T plotter. Comnand frequency 0.02 Hz. Controller must be set for 57. to

10% crossover bias. The trace must be similar to that shown in

Figure A-l. Attach trace to this data sheet.

CJA L^rc^ 5/2.9/n TechnlcIan

C. Performance Characteristics

Date

Mount actuator on test fixture. Connect feedback pot to proper

terminals. Set controller gain for 0.2V error ■ 1007. (.6 degrees

■ 1007.). Crossover bias set for 57, to 107. at modulator output. PDM

carrier frequency 100 Ms, D.C. power 28V and compensation switch se:

to "IM". Gas oupply to be 2000 psi helium.

Frequency Response

1. S< t brake pressure to 275 psi. Connect Weston frequency

response analyzer to controller. Adjust analyzer output

comnand for + 0.7V (+ 2 ), Apply pressure to servo and

record amplitude and phase lag at 15 Hz.

83

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NWC TP 5902

P/N 51030 S/N OüS-

AmplitudeQ.O^ V. Must be less than .98V and greater than ,49V.

Phase -"2.1/.^- deg. Must be less than -90 .

Technician edni^T^ slw/n

Date

2. Set brake pressure to 60 psi and repeat (1).

Amplitude OtbOb y. Must be less than .98V and greater than .49V.

Phase XJLLJ- (êg* Must ê less than -90 .

Technician Date

Slew Rate

3. Set controller to 1 Hz square wave. Adjust analyzer command to

maximum voltage. Set brake pressure to 275 psi. Record on

Sanborn recorder position and command for hardover inputs to

controller. Determine average slew rate by slope of position

trace. Attach recorder trace to original data sheet.

Average slew rate 513 /sec, extend.

Average slew rate 5 /3 /sec, retract.

Must be 400 /sec, minimum.

eMia^ ^l^ln Technician Date

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NWCTP5902 ,/M K,rt,rt P/N 51030

HvsteresIs-ResolutIon S/N ^ ^ ^

4. using a Wavetek or HP low frequency command generator run the

following plots on the X-Y plotter recording position and

command. Brake pressure 275 psi, command 0.05 Hz. Attach

plotter charts to this data sheet.

a) Command amplitude + 7.5V. Trace should be similar

to that shown In Figure A-2.

b) Command amplitude + .TV. Trace must be similar to that

shown in Figure A-3 (1 maximum deviation from command).

Technician ^k^t^ y/tt/y'

Date

D. Potentiometer Characteristics

Serial Number 32.7^ ' 2.3 b Orientation Black A/WA -White

Red

Shaft Extended

Measure resistance between black and red leads and record value.

0. 179 Kß (nom. 3KQ)

Position actuator at mid-stroke and apply excitation voltage. Read

excitation voltage and wiper voltage.

Excitation: Positive 11. 2.7- ypc

Negative \{i1h VDC

Wiper: " I 7-^ mVDC (Voltage should be within + 1000 mVDC)

Mi Ua TL^ Technician

Actuator Performance Acceptable

Date 2±i±l2L

1°/(/?(> Bnglneet Date

NOTE: All data must contain acule factors and identification of all conditions including actuator serial numver.

85

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\'V.T(: 1 I' 5902

APPENDIX B

COST STUDY TECHNICAL REPORT

INTRODUCTION

Aeronutronic Division of Aeronutronlc Ford Corporation is currently iôrking under contract N123-76-C-0286 from the Naval Weapons Center to complete the design and development of a Servo Actuator initiated under prior contracts, and to fabricate, assemble, check-out: and deliver six flight iêms in addition to one development model. Prior design and development work was performed for the Naval Weapons Center under contracts N6n53n-75-M-82c)(; and N6053O-76-M-525V.

The Servo Actuator is a cold gas pneumatic actuator which provides thrust vector control ol a tactical missile by positioning the rocket motor nozzle. A Servo Actuator missile system consists of a cold gas reservoir, squib valve and pressure regulator (these items comprise the helium power supply) and four solenoid valves, two piston cylinder assemblies, and two potentiometers (the actuator) with appropriate plumbing and wiring.

There are two configurations - a rear--mounted helium power supply and a forward-mounted helium power supply - as illustrated in Figure 1 of the main text and Figure B-l. Figure B-2 shows the Servo Actuator, and Figure B-3 shows one actuator mounted on an NWC-furnished trapped ball nozzle. Figure B-4 shows a cross-section view of the Servo Actuator.

The forward-mounted helium power supply requires one large, spherical reservoir, and the rear-mounted helium power supply requires three smaller, cylindrical, manifolded reservoirs, as constrained by the allowable envelope space.

/ This cost study presents cost estimates for production of Servo

Actuator systems at the rate of 2,000 systems per year. This report is a studv only and shall not be construed as a proposal on the part of Aeronutronlc. Costs are presented in 1976 dollars for recurring effort only and do not include fee.

Costs are summarized as follows:

2,000 missile systems/year

Rear-mounted Forward-mounted

Total cost SI,852,602

Missile-system cost $941

$1,7 55,436

$878

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SQUIB ACTUATED CUTTER

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\ V-HEL1UM PRESSURE BOTTLE

(4S.46 1W3 TOTAL)

FIGURE B-l. Servo-Actuator, ForwB Helium Power Supply.

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TRAPPED BA^LL M07.Z.LE

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to-Actuator, Forward-Mounted

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NWC TP 5902

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NWC TP WOl

HISTORICAL COST DATA

Aeronutronir Ford has gathered related historical cost data on a high-volume production program, the Shillelagh Missile System, which includes a control component designated the JRCA (Jet Reaction Control Assembly) , a warm gas control unit with some similarities to the Servo Actuator as described below. The JRCA cost data are presented to show actual cost reduction history on a large-quantity program and Aeronutronic's high-volume-production experience.

The Shillelagh JRCA is a three-element (pitch, roll and yaw) warm gas control assembly, which includes three solenoid-operated flapper valves, a single housing, electrical coil assembly and complex plumbing. The JRCA is shown in Figure B-5. The JRCA is powered by a hot gas generator, which is not Included in the JRCA costs shown below.

The Servo Actuator system includes four similar solenoid-operated flapper valves, two piston assemblies, two potentiometers, and one helium power supply assembly. The potentiometers, pistons and helium power supply are not common to the JRCA, but the Servo Actuator plumbing is less complex. A direct cost comparison is not meaningful.

The Shillelagh JRCA experienced five production buys, totalling approximately 75,000 units produced, with a yearly rate reaching 24,000 (100/day). The cost of the Shillelagh JRCA was approximately $l,800/unit in the second production buy (unit number 1,500, 1963 dollars). This cost was reduced to approximately $290/unit (unit numbers 40,000 to 58,000, 1968 dollars). These costs data are depicted as follows:

Shillelagh JRCA (actual)

1.500

$1,800

Unit number

4,001-6,000a 40.000-58t000b

$ 290

Servo Actuator (estimate)

Forward-mounted

Rear-mounted

$ 878

$ 941

2,000 systems/year, 1976 dollars.

'24,000 systems/year, 1968 dollars.

94

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NWC TP 5902

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FIGURE B-5. Shillelagh Jet Reaction Control Valve CJRCA)

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NWC TP 5902

COSTING PREMISES

The following costing premises apply:

1. A build-up production period of two years is required prior to manufacturing systems at the quantities and rates specified in this report.

2. Non-recurring costs (special tooling and automatic tester) are not included in unit costs.

Special tooling consists of special fixtures and tools for finish machining, gaging and assembly operation.

A semi-automatic tester will be used for acceptance testing of actuators prior to final system assembly. Tests will include proof pressure, leakage, frequency response and slew rate.

3. Costs are based on preliminary drawings of a system that has not been production-engineered. Some production engineering assumptions, e.g., cylinder castings, have been made. Additional producibility and cost reduction studies are assumed.

4. It is assumed that Servo Actuator syrtems will be delivered without helium, in accordance with common practice to fill or pressurize systems at depots or missile assembly locations.

5. Fabrication and assembly estimates are for the 5,000th unit (average of units 4001 through 6000). Support labor costs were estimated for a typical, high volume production program.

6. Material estimates assume a one-lot buy sufficient for 2000 systems. These estimates were derived from prior quotations for most items using a 90% C.R.C. slope to correct for quantity differentials. Estimates, in lieu of quotations, were used for the cylinder housing, valve body and potentiometer.

7. All costs are in 1976 dollars and do not include fee.

DETAILED COST STUDY

Table B-l summarizes the program costs for the 2,000 systems. Table B-2 is a more detailed summary of the manufacturing material costs and labor categories for the 2,000 forward-mounted systems. Tables B-3 through B-7 show the bill of material make and buy unit costs for the 2,000 units. Table B-8 is a detailed summary of the manufacturing material costs and labor categories for the 2,000 rear-mounted systems. Tables B-9 through B-14 show the bill of material make and buy unit costs.

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NWC TP 5902

Engineering hour^

Manufacturing hours

Total hours

Engineering cost

Manufacturing cost

Total labor cost

Material cost

Total cost

Unit cost

TABLE B-l. Cost Summary.

2000 Actuator Systems

Forward-mounted Rear-mounted helium supply helium supply

1018 1159

20356 23172

21374 24331

$36,221 $41,053

$334,167 $379,546

$370,388 $420,599

$1,385,048 $1,462,003

$1,755,436 $1,882,602

$878 $941

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Documents

UNCLASSIFIED AD NUMBER LIMITATION CHANGESnozzle thrust vector control system for tactical missiles. This investigation was conducted during the period 22 April 1976 to 19 October 1976,