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NASA CR 174788

NASA-CR-174788

19860017810 N/_A

LARGE-SCALEADVANCED

PROP-FAN (LAP)

PITCH CHANGE ACTUATOR AND

CONTROL*

DESIGN REPORT

By: Robert A, Schwartz

Paul Carvalho

Mark J, Cutler

HAMILTON STANDARD DIVISION

UNITED TECHNOLOGIES CORPORATION

Prepared For

National Aeronautics and Space Administration

NASA-LewisResearchCenter

Contract NAS3-23051

FE6 1 8 198'5'

LA! ;'___;; ,d._.:_,'-,J,fi-_'_CENTER

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' 'A,_6PTg__l.VIRGINIA ,

January31, 1986

DISPLAY E!£/6/!

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MAJS',_ i.I_!.i:)ESIGN ANALYSISi._.HYDRAULIC CONTROLi-_':'PROP""'FAN TECHNOLOGYi.':_.STRUCTURAL

i'.)ESIGN!.!4.VARiABLE PITCH PROPELLERS

H]:NS_', / AERODYN¢'JMIC CONF!GURATIONS/ PROPELLER BLADES! Sb..IEEP EFFECT/ THIN

AIRFOILS

ABAI', A,,..vt;h {:.,'r"

NASA CR 174788

LARGE.SCALE ADVANCED

PROP.FAN (LAP)

PITCH CHANGE ACTUATORAND

CONTROL

DESIGN REPORT

By: Robert A. Schwartz

Paul Carvalho

Mark J. Cutler

HAMILTON STANDARD DIVISION

UNITED TECHNOLOGIES CORPORATION

Prepared For

w National Aeronautics and Space Administration

NASA.Lewis Research Center

Contract NAS3-23051

January 31, 1986

CONTENTS

' SUBJECT PAGE

" 1.0 SUMMARY 1

2°0 INTRODUCTION 3

3.0 PITCH CHANGEOPERATION 5

3.1 P_tch Change Actuator and Control Schematic 5

3oioi Pitch Change Actuator and Pitchlock 5

3.1.2 Servo Ballscrew & Delta Press. Reg. Valve 6

3.1.3 Governor & Feather Valve 7

3oio4 Hydraulic System 8

3.2 Design Constants and Accuracies 9

3.2.1 Design Constants 9

3.2.2 Design Accuracies 9

3.2.3 Actuator Mechanical Hysteresis I0

4.0 DESIGNREQUIREMENTS 13

4.1 Detail Requirements 13

4°2 Design Philosophy 13

4.3 Airframe/Engine Interface 14

5.0 PITCH CONTROLDESIGN 17

5.1 General Description 17

5.2 Main, Standby, and Scavenge Pumps 17

5°3 Main and Standby Regulating Valve with

Orifice Pack 18

5.4 Governor 19

5.5 Feather Valve 20

5.6 Governor and Feather Linkage 21

5.7 Transfer Bearing and Seals 21

5.7.1 Transfer Bearing 21

5.7.2 Rear Barrel Support 22

5.7.3 Lip Seal 22

5.8 Gearing 23

- 5.8.1 Description 23

5.8.2 Stress Calculations 23

iii
CONTENTS (C0ntinued)

SUBJECT PAGE

5.9 F1lters 24

5.10 High PressureRelief Valve 24

5.]I AuxiliaryPump and Motor 25

5.12 Pump and Valve Housings 25

6.0 PITCH CHANGE ACTUATORDESIGN 27 

6.1 General Descriptlon 27

6°2 Servo and Ballscrew 28

6°3 P|tchlockScrew 29

6°4 Beta Valve 30

6.5 StructuralAnalysis 31

6.5.1 Typical StructuralAnalysis 32

6.5°2 Axial Deflectionof Actuator 33

6.6 TrunnlonAssembly 33

6°7 Delta P RegulatingValve 34

6.8 Actuator Load Deflnitlon 35

6.8.1 Total TwistingMoment (TTM) 35

6°8.2 ActuatorEff|ciency 36

6.8.3 ActuatorMargin 36

6°9 Sizing Crlteria 36

6o]0 Stress|ngCriteria 37

6o11 TypicalMaterialAllowables 38

7°0 DYNAMICANALYSIS 41

7o1 Introduction 41

7o2 ModeI Simulation 41

7°3 Prop-FanControl Performance 42

7.4 Prop-FanControl Limitations 43 

7°5 SynchrophaserPerformance 44

7°6 Summary 44

8°0 BETA CONTROL 47

9°0 TRUNNIONDESIGN 49

9o1 TrunnlonLoading 49

9.2 TrunnlonStressAnalysls 50

9°3 Blade Shank Threads 52

iv
CONTENTS (Continued)

SUBJECT PAGE

" 10. 0 CONCLUSIONS 55

II.0 REFERENCES 57

APPENDIX PAGE

A DESIGNREQUIREMENTS 155

B DESIGNLAYOUTS 187

v/vi

LISTOFTABLES

TABLE PAGE w

5.1 Speeds and Capacities for the LAP and E2/C2 61

5.2 Main and Standby Regulating Valve Data 61

5.3 Orifice Pack Data 62

5.4 Governor Performance Data 62

5.5 Governor Material List 63

* 5°6 Feathering Valve Data 63

5°7 Front and Rear Lip Seal Data 64

5:8 Idler Gear Support Data 64

5.9 LAP Gear Data 65

6.1 Pitchlock Screw Assembly - Material List 66

6.2 Trunnion Loads 66

6.3 Actuator Ear Near Plate 67

7.1 LAP System Model Variables 68

LISTOFFIGURES

2°I Large Scale Advanced Prop-Fan (LAP) 71

3.1 LAP Control Schematic 73

3.2 Pitch Change Actuator & Pitchlock 75

3.3 Schematic of Half-Area Servo Valve and aP

• Regulating Valve 76

3.4 Servo Governor and Feather Valve 77

3.5 Pitch Control Pumping System 78

3.6 Main & Standby Regulating Valve Low DemandCondition 79

3.7 Main & Standby Regulating Valve High DemandCondition 80

3.8 Pitch Change Actuator Constants 81

3.9 Servo Governor Constants 82

3.10 Governor Loop Fr_ction •Hysteresis 83

4°I Modification to ANDIOI52-60 Propeller Shaft 84

4.2 Input Lever Interface for Pitch Change Control 85

5.1 Main & Standby Regulating Valve 86

5.2 Main & Standby Regulating Valve Supply Pressure

as a Function of Net Flow to Actuator 87

5.3 Orifice Pack 88

5.4 Governor 89

5.5 Governor Pressure Gain Test and System Requirements 90

5.6 Governor Gain Flow Requirements 91

5.7 Feathering Valve 92

5.8 Governor Linkage Schematic 93

5.9 E2/C2 Governor and Feather CamSchedule 94

5.10 LAP Governor and Feather CamSchedule 95

5.11 Transfer Bearing Modifications 96

5.12 Lip Seals 97

5.13 Gear Tram Schematic - View Looking Aft 98

5.14 E2/C2 and LAP Main PumpGears - View Looking Aft 99

vii
LISTOFFIGURES (Continued)

FIGURE PAGE

5o15 Modification to E21C2 Rot° Transfer Sleeve Gear lO0

5oi6 Idler Gear Support I01

5.17 High Pressure Relief Valve Data 102

6oi Pitch Change Actuator 103

6.2 Pitch Change Actuator Materials 104

6.3 Pitch Change Actuator Clearance 105

6°4 Half Area Servo 106

6°5 Servo Double Seal Configuration 107

6.6 Half Area Servo Materials 108

6°7 Pitchlock Screw 109

6°8 Beta Valve II0

6°9 Beta Valve Materials II0

6o10 Pressure Gain Characteristics of Beta Valve III

6o11 Flow Gain Characteristics of Beta Valve 112

6o12 Stationary Piston Loads 113

6o13 StationaryPistonInput Geometry ]14

6.14 StationaryPistonRadialDeflections@ 1375 PSI ll4

6.15 StationaryPistonStresses 115

6o16 StationaryPistonS/P Curve If5

6.17 Axial Deflectionsof Actuator- NormalOperating

Condit|on I16

6o18 Axial Deflectionsof ActuatorP|tchlockCondition If6

6o19 Drag and PropellerSpeed for PitchlockCondition 117

6.20 Blade TrunnionAssembly If7

6o21 TrunnionRollerand Near Plates ll8

6.22 TrunnionRollerand Near Plate Materials I18

6°23 Trunnion Roller and Near Plate Stresses 119

6.24 LAP vs. QCSEEStress @104~ 119

6°25 Actuator Ear Stress 120

6.26 AP Regulating Valve 120

6°27 AP Regulator Characteristics 121

6.28 AP Regulating Valve Materials 122

6°29 Centrlfuga] Twisting Moment (CTM) - I00% Speed 123

6°30 Aerodynamic Twisting Moment (ATM) - I00% Speed 123

6.31 Actuator Efficiency (Toward High Pitch) 124

6°32 Actuator Capacity 125

6.33 Actuator Pressure Spectrum for a Typical Flight 126

6°34 Calculation of A11owable Cycles 127

6.35 Stress Allowables as a Function of Stress

Concentration Factor (KT) 128

7oi LAP System Functional Block Diagram 129

7.2 Dynamic Engine Model of Allison Gas Turbine 501-M78 130

7°3 501-M78 Gas Generator Fuel Control 131

7°4 501-M78 Electronic Control Block Diagram (NI Overspeed

N20verspeed, and ITI Limiting Funct|ons) 132

viii
LISTOFFIGURES (Continued)

FIGURE PAGE

7.5 "LAP" Prop-Fan Control Block Diagram 133

7.6 "LAP" Prop-Fan Model Block Diagram 134

7.7 54H60 (C-130) S/nchrophaser Block Diagram 135

7.8 Simulation Res onse to a 1.5% NGRef Step Increase 136

7.9 Simulation Res)onse to a 1.5% NGRef Step Increase 137

7 I0 Simulation Res)onse to a 2°0 Sec Rampof NGRef 138

7 II Simulation Res)onse to a 2°0 Sec Rampof NGRef 139

7 12 Simulation Res)onse to a 0.5 Sec Rampof NGRef 140

7 13 Simulation Res)onse to a 2.0 Sec Rampof NGRef 141

7 14 Simulation Res)onse to a 2.0 Sec Rampof NGRef 142

7 15 54H60 Synchrophaser Limit Cycle in /8 X Gain 143

7 16 54H60 Synchrophaser Performance 2.0 Hz, + 2.5 V Dither 144

7 17 54H60 Synchrophaser Performance with 2.0 Hz,

+ 3.5 V Dither 145

8.1 Motor Driven Pitch Control Mechanism 146

8.2 Blade Angle Slewing Capacity on Beta Mode vs.

Pitchlock Screw Friction 147

9.1 Actuator/Bearing Relationship 148

9.2 Trunnion Loads at 38.5 DEG_ 3/4 149

9.3 Blade Shank Bottom View 149

9.4 GoodmanDiagrams for Trunnion Sections 150

9.5 Summaryof Bolt Loads and Stresses 151

9.6 Tapped Hole Load Distribution 151

9.7 Tapped Hole Stresses 152

1.0 Summary

The pitch controlsystemprovidedfor the LAP programwas designedto provide

a low risk developmentprogram. This was accomplishedby using existing

hardwarewhere possibleand conservativestateof the art design andmanufacturing practices. The use of an existinghydrauliccontrolresultedin less

than optimum transientresponse. This was not consideredcriticalas the

primary purposeof this program is to evaluatethe structuralintegrityand

aeroacousticperformanceof the blades° This approachalso resulted in a

heavier design than would have been expectedin a productionconfiguration._

The calculatedweightsof the pitch control systemare:

Actuator: 214 Kilograms(472 pounds)

Control: 61.2 Kilograms(135 pounds)

Fluids

Control: 0.017 meter3, 15.3 kilograms(18.0 quarts,33.75 pounds)

Actuator: O.Oll meter3, I0.2 kilograms(12.0 quarts,22.5 pounds)

The pitch control system is designedfor fail-safeoperation. Any malfunction of the controlwill either pitchlockthe bladesat their present blade

angle (less approximatelyone degree)or feather the blades. In addition,

the pitch controlsystem providesthe followingfeatures:

• Constant speed governing

• Mechanical in-place pitchlock

• Feathervia mechanicalinput signal

• Feathervia electricalinput signal

• Ground adjustable low pitch stop

• Fixed reverse angle capability

• Electrical aux feather/unfeather motor

• Single lever input

• Individual replacement of blades

• Self-contained oii management(except for heat exchanger)

1/2

2.0 Introduction

In recent years, considerable attention has been directed toward improving

aircraft fuel consumption. Studies have shown that the inherent efficiency

advantage that turboprop propulsion systems have demonstrated at lower cruise

, speeds may now be extended to the higher speed of today's turbofan and turbojetpowered aircraft. To achieve this goal, new propeller designs which feature more blades with thin airfoils and aerodynamic sweep are required.

Since 1975, Hamilton Standard has been deeply involved with the NASALewis

Research Center in the development of the advanced turboprop or Prop-Fan.

" Many aircraft system studies have been accomplished for a variety of subsonic

air transport applications and all these studies have shown significant fuel

savings with Prop-Fan propulsion. The fuel savings potential of future

Prop-Fan powered aircraft is generally 15-20% for commercial applications and

25-35% for military patrol aircraft compared to equal technology turbofan

systems, depending upon the specific application, cruise speed, stage length

and other requirements.

To date, several propeller models have been designed, manufactured and subjected to a number of tests. A series of small-scale 0.6223 meter (24.5

inch) diameter model tests have been conducted in both UTRCand NASAwind

tunnels and on a modified NASAairplane. These tests have shown that propellers with 8-10 swept blades, high tip speeds and high power loadings can offer increased fuel efficiencies at speeds up to 0.8 Mn.

Under the NASAsponsored Large-Scale Advanced Prop-Fan (LAP) Program,

Hamilton Standard has recently designed a 2.743 meter (9-foot) dia. singlerotation Prop-Fan (Figure 2.1) and is now in the process of manufacturing

this system. This hardware will be tested at Wright Field and in the ONERA

S1 wind tunnel in France° The hardware will then be used in a follow-on program where it will be run with an engine on a static test stand, and on a research aircraft. The major objective of this testing is to establish the

structural integrity of large-scale ProprFans of advanced construction in addition to the evaluation of aero-acoustic performance.

Previous reports have described the design of the blades, the spinner and the

hub & retention (references I-3).

This report is limited to the operation and design description of the LAP

pitch control system. The pitch control system consists of two separate assemblies. The first is the control unit which provides the hydraulic supply,

the speed governing and the feather function for the system. The second unit

is the hydro-mechanical pitch change actuator which directly changes blade

angle (pitch)as scheduledby the control. This report describesthe operat_on, the requirementsand actual hardwarethat are used in the LAP pitch

control system.

3/4

3.0 Pitch Change Operation

3.1 Pitch Change Actuator and Control Schematic

Figure 3.1 is a complete schematic that shows all of the operational features

of the pitch change actuator and control. The following sections break the

control schematic into more easily understood subsystems.

The overall system was based on using an existing 54460 control with a pitchlocking propeller. The system was carefully reviewed to ensure all control

malfunctions will either cause the system to pitchlock or go to feather. The

- actuator itself is considered prime structure and designed to conservative

loads and allowable stresses.

3.1.1 Pitch Change Actuator and Pitchlock

The hub mounted pitch change actuator assembly (See Figure 3.2) consists of

an internal stationary piston, a translating outer cylinder with an integral

yoke, a centrally located pitchlock screw, and a four-way metering actuator

valve. The pitch change actuator is the prime mover for blade angle change.

The translating motion of the cylinder is converted to rotary motion at the

blades through an integral yoke on the cyl'inder and a trunnion attached to

the base of each blade. The blade rollers are mounted on a pin which is attached to the butt of the blade (trunnion) and eccentric to the blade retention axis. Translation of the yoke moves the eccentric roller and converts

the linear motion and force of the actuator to a rotary output at the blades.

To change blade pitch, a pressure signalis generated by the propeller control and converted to a rotary signal by the servo and ballscrew (Operation

of the screw and ballscrew is described in Section 3.1.2). This rotary input

is applied to a pitchlock screw. To increase pitch, the rotary signal drives

the screw rearward a small amount relative to the cylinder assembly. This

rearward motion temporarily decreases the pitchlock gap and moves the fourway

valve rearward relative to the valve housing and sleeve. The valve stroke

directs the supply oil to the increase pitch chamber of the actuator, and

ports the decrease chamber to drain. This causes the actuator cylinder to

move forward to increase the blade pitch toward feather. The motion of the

actuator to increase pitch also carries the pitchlock screw forward returning

the four-way valve to null and re-establishing the pitchlock gap. Thus, the

pitchlock gap, in steady state operation, is independent of actual blade angle. The no load pitchlock gap is the equivalent of .6 ° of blade angle, approximately .762mm(.030in) and is always ready to prevent uncontrolled decrease pitch if oil pressure is lost. To decrease pitch the rotary input

" signal is reversed.

p,
3.1.2 Servo Ballscrew & Delta Press. Req. Valve

The rotary input to the pitchlock screw comes from the .servo and ballscrew

(Figure 3.3). The servo consists of a half area actuator with supply pressure always acting on the smaller area. The control or metered pressure is

normally half supply pressure and acts on the twice area side of the servo.

The servo piston is attached to the screw element of the ballscrew through a

set of ball bearings. The nut elementof the ball screw is groundedwith the 

stationaryelementof the servo. Axial motionof the servo piston is converted to both rotarymotionof the ballscrewand an axial translationof the

screw. This output drives the pitchlock screw through a sliding spline thus

obtaining an equivalent of a pure rotational motion of the pitchlock screw.

The metered hydraulic pressure from the pitch control to the servo is normal-

]y half of the supply pressure. For this condition the servo is in equilibrium and there is no output. Increasing or decreasing the metered pressure

will cause the servo to translate and the ballscrew will convert this to the

rotary signal to the pitchlock screw. The system has been designed such that

an increase in metered pressure will cause the servo to call for a decrease

in blade angle. Conversely, a decrease in metered pressure will call for an

increase in blade angle. Thus, shutting off the metered pressure to the servo will cause the blades to feather.

The servo was designed to have the maximumarea possible to improve dynamic

response and accuracy. Under normal operating conditions, the metered pressure from the control is approximately half of supply pressure and the force

output of the servo is only enough to overcome the friction forces in the

system which are less than 1334 Newtons (300 Ibs). However, when the feather

valve is actuated,metered pressurecan drop to drain pressureresultingin a

potentialforce output of the servoof as much as 30700 Newtons (6900 Ibs).

Another case where high loads could be generatedby the servo is when the actuator is driven against the adjustablelow pitch stop. In this case, metered pressurecould equal supplyand the servo would generatean output of

almost 26700 Newtons (6000 lbs). The delta P regulatingvalve controlsthese

loads to less than 6230 Newtons (1400 IbSo) and results in an appreciably

lighterand more efficientlinkagebetweenthe servo and pitchlockscrew.

The purposeof the delta P regulatingvalve is to limit the output force of

the servo under extremesof metered pressure. The delta P regulatingvalve

(see Figure 3.3) is a double-actingreliefvalve. It has the same area ratio

as the servo and has metered pressureactingon the largerarea and supply

pressureon the small area. The valve is preloadedsuch that for small

changes in controlmetered pressurenothing changes. If the ratio of metered

pressureto supply increasesbeyond 59%, the valve opens (i°e.,moves to the

right). Pressure to the servo is than regulated between the valve and the

fixed orifice to approximately 59%of supply pressure. Conversely, if the

signal pressure changes below 42%of supply, the valve moves to the left
3.1.2 (Continued)

and maintains pressure to the servo at approximately 42%. By controlling the

relationship between metered and supply pressure in this band, the output

force of the servo can also be controlled at the desired level.

3.1.3 Governor & Feather Valve

The only control functions provided by the LAP pitch control system are constant speed governing via a pilot selected schedule and feathering.

The constant speed servo governor is provided which automatically controls

blade angle under all aerodynamic conditions. Changing blade angle changes

the power absorption of the blades and thereby matching power absorption to

input torque at a fixed RPM. (See Figure 3.4)

The governor consists of a flyweight assembly, a pilot valve, and a speeder

spring. The flyweights are geared to the propeller shaft so that their rotation develops centrifugal force in direct relation to the Prop-Fan speed

squared. This centrifugal force is converted by the assembly to an axial

load at one end of the pilot valve. The speeder spring imparts an axial

force on the other end of the pilot valve. The position of the pilot valve

is controlled by maintaining equilibrium between the flyweight force and

speed set spring. In this condition, the pilot valve is at its null position

and the metered pressure is at I/2 supply pressure resulting in a zero force

output at the ballscrew servo.

An error in governing speed alters the force balance between the flyweights

and speeder spring moving the governor valve off null. This valve stroke results in a change in metered pressure which changes the force balance in the

half area servo, thus driving it in the direction of the unbalanced load.

The half area servo moves thus changing blade angle in a direction that corrects the speed error and returns the governor valve back to null. For example, an increase in RPMwill lower metered pressure causing the blade pitch

to increase, slowing the RPMto its scheduled governing speed.

Thus, the blade pitch is continually adjusting slightly about the scheduled

speed condition.

The force on the speeder spring can be changed via the input linkage and can

thus allow the governor to schedule different propeller operating speeds.

The feather valve enables the control to bypass the governor and feather the

Prop-Fan. The feather valve can be actuated mechanically through the input

linkage or hydraulically through the electric feather solenoid valve. The

electric signal to the solenoid is either a result of a pilot commandto

feather or activated by the engine overspeed governor.
3.1.3 (Continued)

The basic function of the feather valve is to dump metered pressure to

drain. Dumping metered pressure drives the servo and the beta valve and

moves the actuator in the increase pitch (feather) direction° In addition,

valve actuation blocks main and standby pumpflow to the governor system,

pressure regulating valve, and orifice pack. With the pressure regulating

valve blocked, both main and standby flows are mixed, and the supply pressure

is controlled by the high pressure relief valve. The feather valve is spring

loaded in the unfeather position, and removal of the input signal will restore the control to normal (unfeathered) operation.

3.1o4 Hydraulic System

The hydraulic system is essentially the same as the 54460 propeller control.

The schematic of the pumping system is shown on figure 3_5. The pump housing

assembly contains three mechanically driven pumps (main, scavange, and standby) and an electrically driven double element auxiliary pump. The purpose of

the two mechanically driven supply pumps (main & standby) is to minimize heat

generation during non-transient conditions. In addition, a measure of redundancy is achieved by separating the main and standby flows with two check

valves which allows the system to operate with one pumpoperational. The

combined flow of the two pumps is only required during transient high demand

conditions° To minimize the power requirements of the pumping system during

low demand conditions the standby pumpoperates at low pressure with its output dumped to drain. Operation of the two stage flow system is controlled by

the main and standby regulating valve shown previously in Figure 3.4. The

regulating valve operation is illustrated on figures 3.6 and 3.7. It provides the system with a relatively constant supply pressure over a wide range

of system flow conditions. The valve balances the force to supply pressure

on one end of the valve against a spring and a fixed reference pressure on

the other end. The fixed reference pressure is set by the orifice pack which

meters a fixed percentage of Ps to the pressure regulating valve° This is

accomplished by connecting two orifices in series between supply pressure and

drain, and the reference pressure is taken between them.

Under conditions of low flow demand, most of the combined flow of the main

and standby pumps is diverted to the pressurized sump through separate metering windows. With an increase in demand supply pressure drops and plunger

motion closes the main pump metering window resulting in an increase in supply flow to the actuator. At a maximumslewing condition the plunger has

displaced sufficiently to close the main metering window and meter across the

standby land. Standby pressure increases and opens its check valve thereby

augmenting main pumpflow.

A flow switch is located downstream of the standby pump; a positive indication occurs when the pumpoutput is below an acceptable level signifying

leakage or a malfunction of the standby pump.
3.1.4 (Continued)

The auxiliary pumps(scavenge & supply) provide ground handling and feather

capability in the event of a main hydraulic system malfunction. The auxiliary pumps are separated from the rest of the hydraulic system, during normal

operation, via two check valves. The auxiliary pumps are driven by an electric motor on the control.

The pump housing also contains two oil sumps, a pressurized sump and an atmospheric sump. The pressurized sump is located in the rear of the pump

housing, upstream of the main pump, standby pump, and auxiliary pump. The

pressurized sump provides a boost to the pump inlets in high altitude conditions. The pressurized sump is fed by the scavenge pump, which draws from

the atmospheric sump. The atmospheric sump is located in the front of the

pump housing and its pressure is maintained with an atmospheric breather.

All output flow is protected from contamination with a filter located in the

control. The hydraulic system is protected from excessive pressures with a

high pressure relief. This valve also regulates supply pressure when the

feather valve is actuated.

The hydraulic oil is cooled by an airframe supplied heat exchanger and circulation pump. For cold temperature starts the heat exchanger is bypassed by a

check valve.

3.2 Design Constants and Accuracies

3.2.1 Design Constants

Key dimensional and operational constants were generated at the early stages

of the design to coordinate consistent criteria between Design and Analysis.

The values include such data as areas, strokes, time constants, and spring

rates and gains. The actuator constants are shown on Figure 3.8 and the governor constants are shown on Figure 3°9.

3.2.2 Design Accuracies

The constant speed accuracy of the governor is dependent on the governor

pressure gain, the servo area, and system friction as seen by the servo. (A

" detailed description of the governor pressure gain is covered in Section

5.4) The components of the governor loop friction hysteresis are shown on

Figure 3.10. Assumptions on the coefficients of friction and the ballscrew

efficiency are based on previous experience with similar hardware. A statistical summation yields a 1380 Newtons (310 lb.) maximumfriction load at the

" servo. The following calculation results in a .205_ maximumspeed error due

to the governor loop friction hysteresis. This error manifests itself as a

.205_ speed dead band. A change in propeller speed, exceeding the dead band,

is required prior to the pitch change system reacting to maintain constant

speed.
3°2.2 (Continued)

GOVERNORLOOPFRICTIONHYSTERESIS

%e :aFp(aP/SFp)(axg/aP)(aFg/axg)

(aNg/aFg)(lOO%/Ng)

a Fp : Max. Delta Force Required To Move Servo : 310#

a P/a Fp : I/(Pm Servo Area) = I/I0.22in

a xg/a P : 11(Pressure Gain) = lll8,000psilin @170°F

a Fgl Xg = Governor System Spring Rate = 98.31#/in

aNg/ Fg = I/(Flyweight Sensitiv%ty) = I/4.86 X IO-6Ng#/RPM

Ng : Governor RPM: 4072 RPM@100%Ng

I 1 1 I00

%e=(310) (98.31)

10.22 18,000 (4.86XI0-6)(4072) 4072

%e = .205% Max. Speed Error

The calculated maximumspeed error is less than the original estimate. The

original estimate was used in the dynamic simulation stability and the potential to syncrophase. Since the final estimate is lower than the value used

in the simulation, the simulation is considered valid.

3.2.3 Actuator Mechanical Hysteresis

Mechanical hysteresis of the actuator is the amount of servo motion required

before blade angle change occurs. The majority of this hysteresis is a function of the backlash and spring rate of the linkage that connects the servo

to the beta valve. The spring rate reacts against the friction loads described in Section 3.2.2o The linkage consists of the ballscrew, pitchlock

screw, and beta valve connecting rod. Using a statistical summation, the mechanical hysteresis was determined to be ±.25 degrees of blade angle. This

is higher than the target hysteresis of ±.15 degrees. The increase was primarily due to higher friction loads at the servo and the relatively soft

spring rate of the ballscrew and duplex bearing. Hysteresis values of ±.08

and +°24 degrees were used for the dynamic analysis. The analysis indicates

a decrease in stability margin at the ±.24 degree hysteresis; however, experience with other propeller systems has shown that the actual control system

is more stable than analysis predicts. The potential destabilizing effect of

the additional hysteresis is counteracted by additional servo friction. If

the stabilizing effect of the servo friction is less than anticipated, the

10
3.2.3 (Continued)

" dynamic stabilitycan be improvedby decreasingthe governorgain. Analysis

indicatesthat a 25% decreasein gain shouldbe sufficientto insure stability duringflight conditions. A decreasein governorgain can be obtainedby

changingthe governormeteringwindowconfigurationand has been providedfor

in the LAP program.

11/12

4.0 Design Requirements

4.1 Detail Requirements

The overall aim of the LAP program is to evaluate the structural integrity

and aeroacoustic performance characteristics of the blades. The pitch change

and control system must provide those features and characteristics necessary

to properly control blade angle and propeller speed. A detailed design specification similar to one expected on a production installation was generated

by Hamilton Standard. This specification integrated the entire Prop-Fan Pro-

, pulsion requirements and was used as the preliminary interface document between the propeller, the hub and retention, and the pitch control system.

The actual design specification is included in Appendix Ao The following

summarizes some of the key specific requirements of the pitch control system.

• Governing Speed' 75-105%, Tip Speed" 183-256 meters/sec. (600-840

feet/sec.)

• Ground Adjustable Feather Stop" 87.5 ± 2.5 Deg (6 3/4)

• Ground Adjustable Low Pitch Stop" -I0 to +40 Deg (6 3/4)

• In-Place Pitchlock: 1.5 Deg. Max.

• Slew Rate (Max)" 9 Deg/Sec

• Steady State Accuracy: .22 % Max. RPM

• Hysteresis: .15 Deg.

4.2 Design Philosophy

The basic design philosophy was to provide a fail-safe propulsion system.

This is done by insuring that any malfunction of the propeller will cause the

blades to either pitchlock or feather. Section 6.5.2 describes a typical

p_tchlock case. The basic actuator design is similar to pitchlocking actuator installations on the 14RF and 14SF commuter propellers. All changes in

hardware execution or conceptual approach were reviewed to ensure that the

operation was fail-safe. A detailed Failure Mode and Effect Analysis (FMEA)

was conducted but is not a part of this report.

Since LAP is a demonstration test program, the design effort concentrated on

providing hardware with minimum manufacturing and development risk. This was

accomplished by using conservative design and construction approaches and by

" using service 0roven concepts and existing hardware wherever possible. The

13
4.2 (Continued)

basic concept for the actuator is used on several aircraft like the

deHavilland DHC-7 (24PF). The pitchlock and actuator concept is almost identical to the EMBRAEREMB120 (14RF) and the deHavilland DHC-8 (14SF). In addition, the ballscrew is identical to the 14SF ballscrew. The control is a

modified 54460 which is used on the GrummanE2/C2.

4.3 Airframe/Engine Interface

The Prop-Fan has several interfaces with the engine and airframe. These are:

• Mounting

• Cooling

• Overspeed governor

• Input lever

• Solenoid (electrical signal)

• Aux pump (electrical signal)

Mounting

The propeller is mounted on a modified ANDIOI52-60 propeller shaft as shown

in Figure 4.1. Dimensionally, the configuration is compatible with the T56

gear box. The bulkhead will not have to be modified for the higher operating

pressure of 3000 psi.

Cooling

Due to the higher normal operating speeds, pressures and flows of the LAP as

compared to the 54460, an external heat exchange is required. The mechanical

interface is defined on L/O L14325-20 (Appendix B). The airframe supplied

heat exchanger must provide oi1 to the control at 77°C (]70°F) maximumafter

removing 7°68 kilowatts (10.3 horsepower) from the control oi]. Inlet and

outlet pressure shall vary between 138,000 and 275,000 PA (20 and 40 psi).

Normal flow shall vary between .028 and .056 cubic meters/sec (30 and 60

qts/min). During transients or actuation, the heat exchanger bypass valve

may drop the heat exchanger flow to zero.

Overspeed governor

An overspeed governor is required to protect the system in case of malfunction of the normal control systems. The engine overspeed governor will

14
4.3 (Continued)

Q

activate the feather solenoid causing the propeller to increase pitch. When

the propeller slows sufficiently, the signal will be shut off and the PropFan may modulate at the overspeed setting. At this point, pilot evaluation

is required to determine if the system should be shut down.

The overspeed governor should be wired as part of the electrical feather circuit to the feather solenoid.

Input lever

The control accepts a single mechanical signal from the cockpit to control

the operation of the Prop-Fan. Figure 4.2 shows the input lever interface.

Section 5.6 describes the function of the linkage and Figures 5.8 and 5.10

describe the schedule of operation.

Feather Solenoid

The feather solenoid will cause the Prop-Fan to go to feather (high) pitch

whenever energized. If the solenoid remains energized, the Prop-Fan will go

to feather. The solenoid operates on a 28VDCsignal and has a resistance of

22.5-26.5 ohms at 24°C (76°F). The electrical connector conforms to

MS3102R-12S3P.

Auxiliary Motor and Pum_

The auxiliary motor and pump isdescribed in Section 5.11.

15/16

5.0 Pitch Control Design

5.1 General Description

The control is required to provide the speed governing and feathering functions required for this program. To min|mize cost and development time, it

was decided to use an existing propeller control. Based on the type of engine and gearbox anticipated to be used on the LAP, the control selected prior to the start of the program was a modified 54460 control. The 54460 is a

hydromechanical control in use on the GrummanE2 and C2 aircraft which was

modeled after the Lockheed C-130 and P3 controls° Since the first production

unit in 1956, there have been over II,000 built and they have logged over

73,000,000 hours. The control will provide the constant speed governing

function in the range of 75% to 105%of design speed and the capability to

either manually or electrically feather the propeller.

The use of a modified 54460 control instead of a completely new and optimized

control configuration resulted in certain restrictions on the design. The

54460 control and tailshaft arrangement resulted in a nonmodular design where

the pitch change system must be built up inside the hub. Because of physical

restraints on the installation, no beta (i.e., direct blade angle) control is

provided. However, for ground testing, modification to the system by disabling most control functions permit limited blade angle control via an electrical signal. In addition, the maximumpressures and flows are limited by

the 54460 control. This means a large actuator and a low pitch change rate.

This, in conjunction with the 54460 transfer bearings, will limit transient

response and the ability to synchrophase under turbulent conditions.

An aircraft supplied overspeed electrical signal is required in the event of

a malfunction of the on-speed governor. The control will utilize this signal

through the feather solenoid to cause the blade angle to increase until the

propeller speed does not exceed the overspeed setting and will govern there.

A number of modifications in the control components were necessary to accommodate the increase in RPM, the opposite hand rotation, and different type of

actuator. A description of both the existing components and the design

changes follows.

5.2 Main, Standby, and Scavenge Pumps

The main, standby, and scavenge pumps are modified E2/C2 pumps. The modifications were necessary to accommodate the change in direction of the pump

drive (see section5.8). This was accomplishedby interchangingthe drive

and driven gears on the pumps and by making a new front cover for the pump

housing. Table 5.1 comparesthe LAP and the existingE2/C2 pumps. The main

and scavengepumps are runningabout 12.4% higher RPM than the E2/C2. Both

the LAP and the presentpumps have the same maximumoperatingpressures.

1Y
5.2 (Continued)

The LAP pump will operate continuously at 7.6 x 106 PA (llO0 psi) while the

present pumps normally operate at about 4.14 x I0 _ PA about (600 psi). It

is felt that for the shorter life required for LAP, the pumps should be acceptable. An accelerated fatigue test is planned for the pumps to justify

this assumption.

5.3 Main and Standby Regulating Valve with Orifice Pack

The purpose of the main and standby regulating valve is to provide the system

with a relatively constant supply pressure over a wide range of system flow

conditions. The valve balances the force due to supply pressure on one end

of the valve against a spring and a fixed reference pressure on the other

end. Pressure regulation occurs in response to a change in actuator line

flow requirements. Under conditions of low flow demand, most of the combined

flow of the main and standby pumps is diverted to the pressurized sump

through separate metering windows. The separate windows allow the main pump

to operate at high pressure while the standby pumpoperates at low pressure

I.I x IO6 PA (approximately 160 PSI).

With an increase in demand the supply pressure drops slightly causing a displacement of the plunger. The plunger motion closes the main pump metering

window, restricting flow to drain, and results in an increase in supply flow

to the actuator. When the flow to the system approaches the flow capacity of

the main pump, the plunger has displaced sufficiently to completely close the

main metering window and begins to meter across the standby land. Under this

condition the standby pressure increases to the same value as supply pressure

and opens the check valve located upstream of the regulating valve. In this

manner the standby pump augments main pump flow to meet high system flow demandwhile allowing it to operate at lower pressure and heat generation levels during low demand conditions.

There are only two differences between the LAP and E2/C2 regulating valves.

The LAP regulating valve must be shimmed at assembly to meet system operating

requirements as discussed above. The mating diameters of the plunger and

sleeve of the LAP regulating valve are held closer than the existing valve by

matching the spool and sleeve. This is done to minimize the effect of leakage across.the standby to reference pressure land, which would affect the

reference pressure.

Figure 5.1 and Table 5.2 show the valve and summarize its materials and dimensional characteristics. Figure5.2 shows a typicalpressurevs flow characteristicexpectedfor this valve.

The purposeof the orifice pack is to provide the desired referencepressure

across the regulatingvalve under all conditionsof systemoperation. This

is accomplished by connecting two restrictors in series with the reference

18
5.3 (Continued)

Q

pressure being taken between the two orifices. The orifice pack is designed

to occupy the space normally reserved for the selector valve of the E2/C2

. propeller control which is not required for this design. The orifice pack

consists of new hardware with the exception of the restrictors and seals,

which are off-the-shelf items. Figure 5.3 and Table 5.3 summarize the ori-

" fice pack design details.

5.4 Governor

The purpose of the governor is to maintain a desired propeller speed under

all aerodynamic conditions. This is accomplished by maintaining equilibrium

i between the flyweight force and speed set spring. The flyweight assembly,

driven by the transfer bearing gear, rotates in direct proportion to the propeller speed (Figure 5.4). This rotation of the flyweights develops a force

as a function of the rpm squared. An error in governing speed alters the

force balance between the flyweights and speeder spring moving the governor

valve from null. This valve stroke results in a change in metered pressure

which changes the force balance in the half area servo, thus driving it in

the direction of the unbalanced load. The half area servo indirectly controls blade pitch via the ballscrew, pitchlock screw, beta valve, and actuator (ref. section 3.1). The blade angle change is made in the direction

that corrects the speed error returning the governor valve back to null. For

example, an increase in rpm will lower metered pressure causing the blade

pitch to increase, slowing the rpm to its correct governing speed. The performance data of the governor is listed in Table 5.4.

The governor's ability to accurately govern at a constant speed is dependent

on the effective governor pressure gain at the servo. Before motion occurs

at the servo to correct speed errors the pressure level must be increased to

overcome system friction. The pressure gain of the valve is the increase in

metered pressure as a function of valve stroke and is metered pressure leakage.

The pressure gain of the valve without any external leakage is Io56 x lO_

PA/MM(576,000 psi/in); however, the transfer bearing leakage reduces the effective pressure gain to 4°89 x 106 PA/MM@77°C (18,0OO psi/in @170°F).

This effect is illustrated on Figure 5.5. An attempt was made to minimize

" this reduction in pressure gain by reducing the E2/C2 transfer bearing clearances (ref. section 5.7).

With system fr_ction levels based on experience and the above pressure gains

the calculated governing accuracy was determined to be .205% (ref. Figure

3.10).

19
5°4 (Continued)

The flow gain of the valve is the volume of metered flow supplied by the governor to the servo valve as a function of va]ve stroke (ref figure 5.6).

There are two requirements that have to be met by the flow gain. Based on

experience and the dynamic analysis of section 3.2 the flow gain should result in a blade angle rate of 2.5 deg/sec/%rpm around the null position. In

addition, with the valve wide open the flow of the valve should provide for a

slewing rate of the servo equal to or slightly higher than the actuator slewing rate of 9 deg/sec. To limit the magnitude of the dynamic overspeed the

system should reach slewing rate as soon as possible. Based on experience,

this is a valve stroke corresponding to a 3%speed error. These two requirements result in the stepped window design shown on figure 5.4. Each window

is diametrically opposed to eliminate side loading due to fluid momentum

forces.

The governor is designed to incorporate as much of the E2/C2 hardware as possible and occupies the same cavity of the valve housing° There are a number

of differences between the two governors. The sleeve and servo of the former

governor have been removed and replaced with a single valve housing having

screened ports. The spool was also redesigned, incorporating a lubrication

path to the ball bearing at the speed set end of the governor. Spool lands

were also repositioned as required. The flyweights have been machined from

existing E2/C2 flyweights, and generate approximately the same force and

spring rate as the original flyweights.

Table 5.5 lists the materials used for the governor.

5°5 Feather Valve

The feather valve enables the hydraulic system to bypass normal control signals, from the governor, to feather the propeller (reference Figure 5.7).

The feather valve can be actuated mechanically through the input linkage or

hydraulically through the electric feather solenoid valve. The electric signal to the solenoid is either a result of a pilot commandto feather or activated by the engine overspeed governor.

The basic function of the feather valve is to dump metered pressure to

draino Dumping metered pressure drives the servo and beta valve which moves

the actuatorin the increasepitch (feather)direction.

In addition, valve actuation blocks standby and main pump flow to the governor system, pressure regulating valve, and orifice pack. With the pressure

regulating valve blocked both main and standby flows are mixed and controlled

by the high pressure relief valve° The va]ve is returned to normal operation

via the spring loaded spool.

The feather valve hardware is identical to the E2/C2 except for one seal

groove on the sleeve that was removed to accommodate plumbing changes necesSary for the LAP control design (see Table 5.6).

2O
5.6 Governor and Feather Linkage

Figure 5.8 is a schematic representation of the control linkage, which is

identical to that of the E2/C2 with the exception of the governor speedsetting cam.

The governing range is 1273 to 1783 RPM, (75 to I05%), as defined in section

5.4.3, (Appendix A) of the Design Requirements for the SR-7L Prope!ler. Figures 5.9 and 5.10 give a comparison of the operating schedules for the E2/C2

and the LAP feather and speed-setting cams. The total range, reverse to

feather, remained the same as the 54460. The governing range, however, was

extended into the beta range of the E2/C2 schedule in order to allow for more

degrees of input rotation per percent change in RPM. There are no provisions

for beta control.

The aircraft interface is the control input lever. The input schedule was

designed to be linear over the required 30%speed variation, with 2.5356 ° of

input lever rotation equivalent to a I% change in RPM. The input lever rotation is transmitted to the speed-setting cam by a set of spur gears such that

l° of input rotation results in 2.8824° of cam rotation. The speed-setting

cam rise is translated to the speeder spring by the same series of levers

utilized in the E2/C2 control, where 2.54mm (oi inch) of cam rise yields

3.238mm (.12746 inches) of speeder spring displacement. In order to obtain

the speeder spring loads that would keep the governor at null, for an operating range of 75-105% RPM, the speed-setting cam had to be designed to incorporate a total rise of 3o5738mm(.1407 in). Beyond the governing range the

cam radius remains constant.

Mechanical feathering of the system is accomplished through the same feather

cam and linkage presently utilized in the E2/C2 valve housing. Rotating the

control input lever to the feather position causes the feather cam to activate the feather linkage and mechanically displace the feather valve spool.

Reverse is obtained by setting the speed-setting cam for 105%RPMand limit-

%ngthe fuel flow to the engine. The reduction in fuel flow prevents the engine from supplying enough power to the system to allow it to rotate at 105%

RPM. In this condition, the governor runs underspeed and causes the servo to

call for a lower blade angle. The system will continue to decrease pitch until it reaches the low pitch stop and will remain in reverse until the governor runs overspeed and signals for increase pitch. This overspeed condition

" can be accomplished by increasing engine power until the propeller speed is

above 105%RPMor by lowering the governor setting to a value below the propeller RPMin reverse.

5.7 Transfer Bearing and Seals

5.7 l Transfer Bearing

The transfer bearing supplies hydraulic fluid to the rotating propeller from

the stationary pumphousing. The transfer bearing has three passage across

21
5.7.1 (Continued)

the rotating interface, supply pressure, metered pressure, and drain pressure. The transfer bearing configuration is identical to the E2/C2 design

except for the diametral clearance at the metered pressure land which was reduced, per Figure 5.11, to minimize the leakage effect on pressure gain (ref.

section 5.4). The E2/C2 has a history of seizing at land A. This condition

is created by the combinedeffectof minimal lubrication,from using this

land to transferdrain pressure,and radial barrel deflectionschanging the

diametralclearanceof the bearing. The use of this land for metered pressure increasesthe lubricationflow and the barrel supportring (see section

5o7.2) stabilizesthe diametralbearingfit. These two changes from the

E2/C2 design will minimizethe possibilityof a bearing seizure. A test was

conductedto determinethe effectsof the clearancechangeand rpm increase

(If30 rpm vs. 1698 rpm) on the transferbearing. Preliminaryanalysisof the

test indicatesthat this configurationwill be successful.

The bearing inner sleeveincorporatesa pump drive gear. The gear was modified from the E2/C2 design to accommodatethe change in speed and direction

of the propellerwhile still maintainingapproximatelythe same pump output

(ref. section 5.8).

5.7.2 Rear BarrelSupport

The rear barrel supportwas stiffenedin order to prevent the radialdeflections of the barrelfrom reducingthe transferbearing clearances. Using deflectionsfor shellsand vessels as a means of calculatingstiffness,the

stiffnessof the rear barrel supportwas found to have increasedby a factor

of 3.59 over the present E2/C2 support. Stiffening the rear barrel support

also enabled the use of the same lip seal in the rear as in the front of the

control.

5.7.3 Lip Seal

The lip seals (Figure 5.12 and Table 5.7) provide a dynamic seal between the

pump housing cavities and the atmosphere. The sea]s incorporate helical

flute located on the sealing face of the ]ip which prevent oi] leakage by

providing a pumping action back into the sump. The direction of these flutes

was originally designed to seal with a right hand rotation propeller. With

left hand rotation the front seal will be operating in a direction that will

permit some leakage. To avoid the cost of new lip seal dies, tests were run

to determine if the extent of the leakage problem is significant. This test

also determinedthe effectof operatingat an increasein speed. Initial

tests have shown that the front seal runs hot and requires additional cooling

and lubrication. Features have been added to the final design to correct

this situation.

The-stiffened rear barrel support (ref. Section 5.7.2) enabled the use of the

front lip seal at the rear of the control. In this new location the lip seal

will have the proper orientation for left hand rotation and provide pumping

action back to the sump.

22
5.8 Gearing

5.8.1 Description

The configurationof the LAP geartrainis similarto the E2/C2 geartrain.

The systemwas designedwith the intentionof keeping the speedsof the pumps

and governoras close as possibleto the present speedsof the E2/C2 compon-

- ents. The system was also designed to accommodate left hand rotation, which

is opposite of the E2/C2, with minimum control modifications.

Plumbing changes to the control were minimized by maintaining the same flow

direction in the LAP pumps as in the E2/C2 pumps. This was accomplished by

: reversing the input gear on the pumps. Figures 5.13 and 5.14 illustrate the

i changesthat were made to the LAP pumps in comparisonto the E2/C2 pumps.

The transferbearinggear, used to drive the pumps and idler,was enlargedin

order to achieve the desired gear ratios. This was accomplishedby machining

off the old gear and rivetingon a new gear, as indicatedin Figure 5.15.

Twelve I/8 in. rivetswere used and the resultingstresson each rivet, based

on a transferbearinggear torqueof 89NM (789 in.-Ibs),is I07PA (1450

psi).

The differentialgeartrain,used to run the governorand for beta feedbackin

the E2/C2 control,was replacedby a compoundidler gear. To minimizemodificationsto the governorassembly,the governorgear utilizedin the E2/C2

control is incorporatedinto the systemand the resultingincreasein RPM is

only 15.78_.

The idler gear is held in place by an eccentric support shaft, located in the

same position on the control as was the differential gear shaft. Anti-torque

for the idler is achieved by a dowel pin pressed into the gear support. The

bearings used on the support are identical to the ones found on the 14RF/14SF

control and are preloaded by a wave spring with a working load of 337-365N

(75-82 IbS)o Lubrication for the bearings is provided by plumbing in the

support shaft, (Figure 5.16) and the expected life for these bearings is

19500 hours.

5°8.2 Stress Calculations

" The control gears have been sized per allowables established by Hamilton

Standard for gears in this type of application based on our experience in

both large aircraft gearbox design and for control hardware, such as the

- 54460 control. The stress levels established represent a X - 3.0 sigma deviation from data, or 1 failure in I000.

The gear material for the transfer bearing, pumps, and idler is AMS6260, all

carburized with the exception of the idler gear that meshes with the governor

" gear (Table 5.8). This is because the governor gear is not presently carburized. The gears are splash lubricated with MIL-H-5606.

28
5.8.2 (Continued)

The dynamic hertz and dynamic being stresses in the gear teeth were determined using the following formulas;

Scs=(.70(Nt)(I/PD_+]/PD2))/(L(sin PA)(cos PA)(I/E,+I/E_))

where Sos = statichertz stress

Wt = tangentialtooth load

PD, and PD2 = pitch diametersof mating teeth

PA = pressureangle

L = min. axial engagement

E, and E2 = modulus of elasticity

SoD = NdlNt (Sos)

where SoD = dynamic hertz stress

Nd = dynamic load

S_o = 1.5(Wt)(Kt)IFW(x)

where SBo : dynamic bending stress

Kt = Heywood stress concentration factor

FN = min. face width

x : Lewis geometryfactor

Table 5.9 lists the gear data for the gears used in the control and the calculated stress levels.

5.9 Filters

The main and standby flows are protected by a double element filter located

in the valve housing. The filter is the same as the one used on the E2/C2

control. The filter consists of a phosphor bronze screen mesh (120 X 120)

with a .08]mm (.0032 inch) nominal screen opening.

5o10 High PressureRelief Valve

A high pressurerelief valve is providedas a backup to the pressureregulating valve. It only operatesduring start-upand shut=downor if the pressure'

regulatingvalve malfunctions. The high pressurereliefvalve is set to operate between7.93 x lO6 and 9.48 x lO6 PA (ll50 and 1375 PSI). The

valve is identicalto the present E2/C2 hardware. The valve design details

are summarizedon Figure5.17.

24
5.11 AuxiliaryPump and Motor

" The auxiliarypump and motor are the same hardwareused on the E2/C2

control. The pump output is used for groundhandlingand to feather the

propellerin case of main and standbypump failure. The auxiliarypump

supplieshydraulicsto the actuatorduring staticoperation. The auxiliary

pump is a two element pump. The main element supplies .0132M3/MM @ 9.48 x

lO6 PA (14 qpm at 1375 psi) to the supply systemand the secondelement

acts as a scavengeto return drainoil from the atmosphericsump to the

pressurizedsump. The auxiliarypump is driven,through a geartrain,by a

200 volt 400 Hz 3 phase electricmotor. The motor limitsthe duty cycle to

" ,20 secondson at 3350watts (4.5 horsepower),and 15 minutesoff.

5.12 Pump and Valve Housings

The pump housing is the same as the E2/C2 housingexcept for the hydraulic

plumbingin the transferbearing area. The housing is made of cast magnesium.

The valve housing is cast aluminumand is the same as the E2/C2 housing

exceptfor clearancecuts made to accommodatethe modifiedgear support.

25/26

6.0 Pitch Change Actuator Design

6.1 General Description

The pitch change actuator is the prime mover for blade angle change. The

- pitch changeactuatorconsistsof a stationarypistonand a translatingcylinder. The translatingmotionof the cylinderis convertedto rotarymotion

. at the bladeswith an integralyoke on the cylinderand a trunnionattached

to the base of each blade. The _runniondesign includesa pin mounted roller

offset from the centerof the blade retention. The stationarypiston is

groundedaxially to the hub with a retainingnut.

The force exerted by the trunnionrollercombinedwith its rollingmotion

during pitch changecreatesa tangentialfrictionforce on the wear plates°

This force at each actuatorear results in a twistingmomenton the cylinder. This moment is reactedby an anti-torquequill which grounds the cylinder rotationallyto the dome. The quill also acts as an indexingdevice to

ensurealignmentof the actuatorears with the trunnionrollers.

The main componentsof the actuatorare shownon Figure6.1 and the materials

selectedfor the main structuralelementsare shown on Figure6.2.

This type of pitch changeactuatorand materialcomponentsare typicalof

several HamiltonStandardproductionactuatordesigns.

The cylinderand bulkheadassemblypilotson Torlonwear stripsat both ends

of the actuatorto preventmetal to metal contactat the sealing surfaces.

The actuatorwas designedso that under the worst tolerancecondition,includingdeflections due to pressure loads, there will be a minimum of .O127mm

(°0005 in.) radial clearance between the main piston seal lands and the cylinder ID. (ref Figure 6.3)

The dynamic seals for this design were chosen for their successful performance in similar applications. The ID seals have a "D" cross section to prevent rolling in the seal grooves. Teflon backup rings are located on the atmospheric side of the seals to control the extrusion gap. This combination

is used successfully on the 14RF and 14SF commuter propeller actuators.

The main piston seal consists of a double-delta seal loaded by a standard

• O-ring. This seal was designed for the maximumoperating radial extrusion

gap of .762mm(.030 in.) and at an operating pressure of 9.48 x I06 PA

(1375 psi). The seal material, Turcon/Moly, was chosen for its resistance to

extrusion, low friction, good wear characteristics, and its ability to seal

under the operating range of the actuator. This type of configuration has

- been tested on the 54H60 dome l.Omm (.040 in.) rad. extrusion gap with good

results after 30,000 cycles.

27
6.2 Servo and Ballscrew

The servo and ballscrew convert a hydraulic signal from the control to a rotary input to the pitchlock screw (see Figure 6°4). For example, for an underspeed condition the governor would increase metered pressure which in turn

changes the force balance in the servo driving it forward° The forward motion of the servo, which is axially attached to the ballscrew through a duplex ball bearing, rotates the ballscrew. The ballscrew is directly Connected to the pitchlock screw through a splined quil] and the rotary motion of

the ba]]screw is imparted to the pitchlock screw. The pitchlock screw translates forward displacing the beta valve which routes hydraulics to the decrease side of the actuator. Speed is increased to the set speed of the governor nulling the governor va]ve and returning the force balance in the serVOo For an overspeed condition this process is reversed. The governor at

null regulates metered pressure at one-half of the supply pressure, therefore

to maintain a force balance at the servo in an on-speed condition the area of

the metered pressure side of the servo was designed to be twice that of the

supply pressure side.

The servo areas were sized to minimize the loss in governing accuracy due to

the low effective pressure gain of the governor caused by transfer bearing

leakage. As a result, the servo areas are oversized for load capacity requirements. Under normal operating conditions, the metered pressure from the

control is approximately half of supply pressure and the force output of the

servo is only enough to overcome the friction forces in the system which are

about 1335 Newtons (300 Ibs). However, when the feather valve is actuated,

metered pressure can drop to drain pressure resulting in a potential force

output of the servo of as much as 30,700 Newtons (6900 pounds). Another case

where high loads could be generated by the servo is when the actuator is

driven against the adjustable low pitch stop. In this case, metered pressure

could equal supply and the servo would generate an output of almost 26,700

Newtons (6000 pounds). In order to incorporate an existing ballscrew that is

currently used on the 14RF and 14SF propeller controls, a delta-P regulating

valve was added to the system. This valve prevents overloading of the ballscrew by limiting the difference in supply pressure from metered pressure.

The delta P regulating valve controls the ballscrew loads to less than 6230

Newtons (1400 pounds).

The servo is similar to the 14SF and 14RF designs except for the double seal

configuration separating supp]y pressure from metered pressure. (Figure

6°5). This additional seal will prevent supply pressure leakage from increasing metered pressure which would drive the actuator toward low pitch

creating a possible overspeed condition. The sleeve and piston are heat

treated steel (AlSl 4340)(See Figure 6°6). Running surfaces are shot peened

and chrome plated° Rulon glide rings are used to pilot the piston _n the

sleeve. Glass filled teflon s]ipper seals are used in all dynamic sealing

locations to min|mize friction. The servo is retained in the valve housing

w|th six bolts. Backlash in the system is controlled by preload!ng the duplex bearingand ballscrew. The ballscrewis mountedon a sphericalwasher

to provide self-centeringand to prevent bindingdue to misalignmentbetween

the ballscrewsupport and duplex bearingcenters.

28
6.3 Pitchlock Screw

" The pitchlockscrew assemblyconsistsof the rotatingpitchlockscrew,a

pitchlocksupport (nut),an adjustablelow pitch stop, a drive quill from the

ballscrew,and an adjustableconnectingrod to drive the beta valve. Figure

6.7 shows the pitchlockscrew assembly. The pitchlockassemblyconvertsthe

rotationalsignalfrom the ballscrewto axial motion of the beta valve. The

screw maintainsa constantgap with referenceto actuatorground independent

of blade angle. This gap is set to .660- o762mm(.026 - .030 inch)with the

beta valve at hydraulicnull (i.e.,metered pressureequal to half supply).

Loss of hydraulicswill cause the actuatorto shift thru this gap lockingthe

actuator and transferring the blade twisting moment load directly to ground.

For normal operation the pitchlock screw is driven thru a quill shaft from

the ballscrew. The quill shaft accommodates any angular or radial misalignment due to manufacturing tolerances between the ballscrew and the pitchlock

screw. The quill shaft has a sliding spline at the pitchlock screw to accommodate the relative axial stroke of the ballscrew. The sliding spline on the

pitchlock screw quill is nitrided and coated with a dry film lubricant. The

mating internal spline is chrome-plated. The fixed spline on the quill is

also nitrided. This combination for the quill shaft was selected for its

compatibility with the existing ballscrew spline and is based on experience

with similar applications.

The pitchlock screw and nut threads are 30° modified stub acme threads with a

55o626mm(2.19 in.) P.D. and a 12.7mm(.5 inch) pitch (Table 6.1). This lead

was selected to match the axial stroke of the actuator to the rotational output of the ballscrew. The screw is designed to prevent the actuator from

backdriving under positive load with a coefficient of friction of .03. A

similar screw on the 14RF and 14SF propellers was designed at a coefficient

of .05. These screws were tested under various conditions including a full

vibration spectrum. For light loads, under 200 Ibs., the 14RF and 14SF screw

showed some slippage when shaken at their natural frequency. Once the loads

were increased above this value, all slippage stopped. The minimum load expected in flight is 164,600 Newtons (37,000 Ibs.) at 25° _ 3/4, Sea Level

Standard Day, 75% RPM. In addition, the screw on the LAP has been designed

to be less susceptible to vibrational loads than those used on the 14RF and

14SF propellers.

. The screw and nut are carburized steel to prevent wear and to maintain good

surface quality under all operating conditions. The screw and nut are a

matched set, lapped together so as to limit the maximumbacklash to .0508mm

(.002 inches) and the maximumdriving torque to .565NM(5 inch lbs.) under a

66.7N (15 lb.) axial load. The pitchlock screw and support were analyzed using a finite element body of revolution computer program (H727). (This program is similar to H561, see Section 6.2) The sizing criteria was a proof

load of 140%overspeed including a 1.2 contingency factor. The stress levels

for this condition were compared to the allowable yield strength of the material.

29
6.3 (Continued)

The pitchlockscrew is attachedto the beta valve through the connecting

rod. The effectivelength betweenthe screw and beta valve is adjustedby a

screw mechanismon the rod assemblyat the screw end. This adjustmentallows

the pitchlockgap calibrationpreviouslydescribed. The rod is supportedat

either end by sphericalbearingswhich accommodateany radialor angular tolerance between the screw and valve. The first critical speed of the rod is

slightly less than twice the operating speed (1700 RPM)o

The pitchlock screw assembly also incorporates the adjustable low pitch operating stop. This stop limitsthe minimum allowableblade angle by limiting

the rotationof the screw with respect to the nut. This preventsfurthermotion of the servo, screw,and beta valve in the decreasepitch direction.

The stop is ground adjustablefrom -lO degree blade angle to +40 degree blade

angle. For flight conditions,the stop is a safetydevice which limitsthe

minimum flightangle to a specificlevel independentof any other signal including a control failure. For reverse thrust testing, the stop will be the

reverse blade angle.

The stop is matched to the screw and nut to ensure proper timing and engagement at the stop surface. The stop is designed to react both the inertial

load of the system when hitting the stop plus the stall load of the servo.

The control the inertia loads, one element of the stop is machined as a cantilever off the pitchlock support (nut) and its spring rate will soften the

impact loads.

6.4 Beta Valve

The beta valve is a four way, force compensated, spool valve with direct mechanical feedback with the actuator (Figure 6.8). The steel sleeve and spool

are manufactured as a flow matched set. Their running surfaces are

carburized to prevent erosion wear at the metering lands (Figure 6.9).

Motion of the valve aft of the null position meters oii to the increase pitch

side of the actuator cylinder, and allows flow out of the decrease pitch side

of the cylinder. This causes blade motion to increase pitch through the yoke

and blade trunnion rollers. Since the valve's inner spool is attached to the

actuator cylinder via the connecting rod, motion of the actuator brings the

beta valve back to the null position. Conversely,forwardmotion of the

valve moves the blades towarddecreasepitch. When the valve lands are near

the null position,the properblade angleis attained,and the bulk of the

supplyflow is bypassedby the pressureregulatingvalve as describedin section 5.3.

Pressuregain and flow gain are the two importanthydrauliccharacteristics

of the valve (Figures6.10 and 6.ll). The pressuregain of the beta valve

determinesthe dead band hysteresisof the actuator. To minimize this hysteresisthe pressure gain was made as large as possible within the constraints of the desired flow gain and envelope restrictions. Experience with

3O
6.4 (Continued)

" this type of valve combined with a detailed analysis of its structure restricts the maximumallowable window to 60%of the sleeve internal circumference. The resulting peripheral length of the window, combined with tight diametra] clearances and an underlapped spool, yields apressure gain of 1.25 x

108 PA/MM(460 psi/.O01 in). Flow ga!n is the change in supply flow to the

. actuator from the beta valve as a function of valve stroke (Figure 6.11). A

flow gain of 3.5xlO6mm3/sec/mm (5°433 in3/sec/oOOlino) was required to

obtain the desired time constant of .020 SeCo Experience with propeller

control systems have shown that this time constant is required for stable

dynamic response. A detailed analysis of the dynamic response can be found

in section 3.2.

Force compensation has been provided by contouring the spool. The contours

will reduce the closing force, inherent in four-way valves, to 8.9 Newtons (2

Ibs.). The contour geometry required is based on both theoretical analysis

and empirical data derived from similar hardware. The force thus obtained is

appreciably less than the 53 Newtons (12 lb.) expected from the valve seal

frlction and will have a negligible effect on hysteresis. These seals are

required at both ends of the spool to prevent valve leakage from filling the

dome.

The beta valve, ballscrew servo, and delta-P regulating valve are packaged in

an aluminum valve housing which is mounted in the stationary piston I.Do. The

valve housing acts as a manifold that routes hydraulics from the hub to the

beta valve and then to the actuator. In addition, there are plumbing lines

that connect the servo and delta-P regulating valve hydraulics to the hub.

The actuator dome incorporates high and low pitch pressure sensors that will

be used to analyze the control and determine actual blade loads during

flight. Low pitch and high pitch pressure are routed to the pressure transducers at the front of the dome through tubes that are connected to the valve

housing. Locating the pressure transducers in the front of. the dome provides

better accessibility for maintenance.

6.5 Structural Analysis

The actuator was structurally analyzed with a finite element shell of revolu-

- tion computer program (H561). This program performs linear elastic analysis

of axisymmetric structures which can be modeled as collections of thin shells

of revolution. Basically, the finite element displacement method is used to

convert the mathematical mode] of the physical problem from one of solving

the simultaneous partial differential equations of thin shell theory to one

of forming and solving a large set of simultaneous linear algebraic equations. This algebraic set represents equilibrium and compatibility equations

at a number of discrete points in the structure, called nodes, written in

. terms of unknown displacement and rotation values at these same points.

81
6.5 (Continued)

Output consists of printed listings of displacement, load and stress values

at a large number of points throughout the structure. Plots of input geometry, deflected shape, and stress distributions are also available to assist

in interpretationof the results.

There were four major considerationsin determiningthe wall thicknessesof

the actuator.

I. allowable stress levels for the materials selected.

2. radial extrusion gaps at dynamic seal surfaces

3. axial deflection in the pitchlock condition

4. radial threadengagementof the retainingnuts

6.5.1 TypicalStructuralAnalysis

I. pressure spectrum on the increase side of the actuator with the cylinder at mid-stroke.

2. pressure spectrum on the decrease side of the actuator with the cylinder at mid-stroke.

3. pressureon increaseside of actuatorwith cylinderagainst the

feather stop.

4. pressure on decrease side of actuator with cylinder against the reverse stop.

The results of a sample calculation for the actuator piston is shown in Figures 6.12 thru 6.16. This is typical of the analys_s performed on each

structural item of the actuator assembly. The loading condition of case 3

(above) for the stationary piston is illustrated on Figure 6.12o Figure 6.13

is a plot of the input geometry for the H561 model (ref. section 6.2).

Figure 6.14 is an exaggerated plot of the deflected piston shape showing the

areas with the largest radial deflections. These deflections were used to

determine the actual operating piston clearances. The high stress areas in

the piston and their respective stress concentration factors are shown on

Figure 6.15. The high stress areas are below the allowable S/P values for

the piston (refo Figure 6.16). The S/P allowables are based on the load

spectrum and the ma- terial allowables. A typical derivation of S/P

allowablescan be found in Section 6.11.

32
6.5.2 Axial Deflection of Actuator

" In designing the actuator, axial deflections were of prime concern. The axial deflections are a result of pressure induced deflections in the cylinder

and load dependent deflections of the blade, trunnion, and actuator ears.

The axial deflections affect operation in three ways.

. I. There is a shift in actual blade angle with respect to null of the

beta valve. During governing this has no effect on operation as the actual

blade angle is not required to maintain constant rpm.

2. The adjustable low pitch stop blade angle setting changes with actuator deflection. A typical flight load case (refo fig.6.17) of 318,000

Newtons (71,500 Pounds) would yield .762mm(.03 in.) axial deflection at the

pitchlock screw and results in a .56 degree increase in the low pitch stop

blade angle.

3. Actuator deflections add to the loss of blade angle in a pitchlock

condition. Figure 6.18 tabulates the total blade angle loss, including the

pitch]ock gap, for various loading conditions. Figure 6.19 shows the effective drag or loss in thrust vs. loss in blade angle for a pitchlocked

propeller operating in a typical cruise condition. This drag on the aircraft

is not considered to be significant°

6.6 Trunnion Assembly

The trunnion assembly consists of an eccentric arm (trunnion) bolted and doweled to the bottom of the blade shank. On the end of the trunnion is a

crowned roller bearing that engages the yoke ears of the actuator. The linear motion of the actuator working against the roller on the trunnion arm

causes the blade to change pitch in the barrel retention bearing. (figure

6°20)

The blade trunnion rollers and wear plates are similar to the design found on

the 14SF and 14RF propellers° The angular misalignment between the trunnion

shaft and yoke, under load, requires a crowned roller to control the location

of the hertz contact pattern on the wear plates. This misalignment was minimized by matching the stiffness of the yoke with the combined stiffness of

the trunnion and blade retention. The crowned roller is thru hardened 52100

bearing steel. A teflon filled polyimide journal bearing, which interfaces

with the chrome plated trunnion shaft, is pressed into the roller. The roller assembly is retained with a pin and sleeve combination (ref. figure 6.21

& figure 6.22)

33
6.6 (Continued)

The loads used for sizing the trunnion are summarized in Table 6.2. Figure

6.23 lists the critical stress areas of the trunnion roller assembly. The

wear plate subsurface shear stress is comparable to similar pitch change roller designs tabulatedin Table 6.3. The comparisonrequestedby NASA between

the LAP and Quiet Clean Short-HaulExperimentalEngine (QCSEE)roller designs

is illustratedon figure6.24. The stressingon the LAP is shown to be

greater than the QCSEE but the higherhardnessof the wear plateson the LAP

comparedto the hardnessof the cam trackson the QCSEE yield higher subsurface shear stressallowables. In additionto the higherallowables,the LAP

roller is also immersedin oil to reducewear while the QCSEE roller ran dry.

The actuatorears were stressedusing the pressurespectrumdevelopedto analyze the cylinderand piston. This is conservativesince the actuatorears

are not subjectedto the start-stoppressurereliefvalve load. The resulting S/P values are summarizedon figure 6°25. The S/P allowablesare based

on the load spectrumand materialallowables. A typicalderivationof the

S/P allowablesis illustratedin Section 6ollo

6.7 Delta P RegulatingValve

The purposeof the delta P regulatingvalve is to limit the output force of

the servo under extremesof metered pressure. The servo was designedto have

the maximum area possibleto improvedynamic response(see Section 6°4). Under normaloperatingconditions,the metered pressurefrom the control is approximatelyhalf of supplypressureand the force outputof the servo is only

enoughto overcome the frictionforces in the systemwhich are about 133 PA

(300 ]bs). However,when the feather valve is open, metered pressurecan

drop to drain pressureresultingin a potentialforce outputof the servo of

as much as 30,700Newtons (6900 Ibs)o Another case where high loads could be

generatedby the servo is when the actuatoris driven against the adjustable

low pitch stop. In this case, metered pressurecould equal supplyand the

servo would generatean output of almost26,700 Newtons (6000 Ibs)o The

valve controlsthese loads to less than 6230 Newtons (1400 Ibs) and results

in an appreciably lighter and more efficient linkage between the servo and

pitchlock screw. Lower loads also allow the use of existing ballscrew hardware presently used on the 14SF and 14RF propeller controls.

The delta P regulatingvalve (see Figure 6°26) is a double-actingrelief

valve. It has the same area ratio as the servo and has controlmetered pressure actingon the largerarea and supply pressureon the small area. The

valve is preloadedsuch that for small changes in controlmetered pressure

nothing changes. If the ratio of metered pressureto supply increasesbeyond

59%, the valve opens (i.e.,moves to the right). Pressureto the servo is

then regulatedbetweenthe valve and the fixed orifice to approximately59%

of supplypressure. Conversely,if the signal pressurechanges below 42% of

supply,the valve moves to the left and maintainspressureto the servo at

approximately42%. By controllingthe relationshipbetweenmetered and supply pressurein this band, the output force of the servo can also be controlledat the desiredlevel.

34
6.7 (Continued)

Figure 6.27 shows the characteristics of this valve and the force output of

the servo for various conditions. The minimum output is approximately three

times the maximumexpected servo friction forces. The maximumoutput is

within the structural capability of the linkage.

The valve is installed in the actuator on the center line of rotation to eliminate centrifugal loading on the valve. The preload and areas of the valve

are sized to permit the valve from stroking at the normal pressure variations

and to provide enough force to overcome friction or break small chips. The

" valve is carburized to prevent wear and is further protected by screen meshes

on both metered and supply flow. Both screens have .1676mm(°0066 in) diameter holes and 28%open area. The supply pressure screen also acts to prevent chips from getting to the beta valve. The valve is internally damped by

an orifice between the metered pressure signal and the metered pressure load

cavity. The materials used for the delta P valve are shown in Figure 6.28.

6.8 Actuator Load Definition

6.8.1 Total Twisting Moment (TTM)

The forces that are reacted by the pitch change system consist of a Centrifugal Twisting Moment (CTM) and an Aerodynamic Twisting Moment (ATM). In additlon, when changing blade angle, the actuators must overcome blade retention friction (FTM).

The CTMis the moment produced by the centrifugal forces acting on the blade

which tends to align the blade parallel to the plane of rotation. The CTMis

the dominant loading component. Figure 6.29 is a plot of CTMat 100%RPM.

CTMis proportional to RPMsquared. A positive CTMis a load tending to

twist the blade toward a lower blade angle.

The ATM is the result of the summation of unbalanced aerodynamic forces about

the pitch change axis. The ATM is a function of the effective air speed,

shaft horsepower, air density, blade angle and RPMo A positive ATM is a load

toward increase blade angle. Figure 6.30 shows typical ATMcharacteristics

for the type of blade used on LAP (SR-7)o The ATMfor an SR-7 type blade is

usually less than 15%of the CTMand can either add to or subtract from it.

Increasing power or decreasing speed decreases the summation of the CTMand

ATM.

The FTM Is a constant friction torque vs. RPMdue to blade bearing frlction

and blade seal friction. FTMis a function of the centrifugal load, blade

pitch radius and coefficient of friction (assumed to be 0.005 based on previous experience). This relationship has been determined from past experience

with similar blade retentions. The FTMis the same for all blade angles and

always adds to the load required to move the blades, The FTMat 100% speed

is 1256 in-lbs/blade (1.419 x IOs MM-Newton/blade) and is proportional to

RPM squared.

35
6.8.2 Actuator Efficiency

The piston area output force must be combined with the actuator efficiency to

determine the effective actuator capacity. The efficiency is dependent on

•the actuator seals friction, antitorque spline friction, trunnion journal

bearing friction, and the effective trunnion arm radius which changes with

blade angle. Figure 6.31 il]ustrates their combined effects in a plot of efficiency vs.. blade angle.

6.8.3 Actuator Margin

Figure 6.32 demonstrates that the effective actuator capacity of the LAP

meets or exceeds the three design case requirements. This capacity was obtained with a 6.96 x 10-2 M2 (108 in.2)-area on the increase pitch side

of the actuator combined with 76.2mm (3.00 inch) moment arm at the trunnion.

The requirements of the decrease pitch side of the actuator are minimal, but

to simplify the design its area is approximately the same as the increase

side of the actuator 7:35 x lO-_ M2 (114 in2)o

The stall output is calculated at 7.03 x I06 PA (1020 psi), or 7.24 x IOb

PA (1050 psi) at the increase pitch side of the actuator and .20 x 106 PA

(30 psi) drain pressure on the decrease pitch side of the actuator. The

actuator output for a pitch rate of 9 deg./sec, is shown at 4.83 x 106 PA

(700 psi) to reflect the following full flow pressure drops through the

system.

I. B-valve windows (supply & drain) - 1.21 x 106 PA (175 psi)

2. Check valves and filter - .517 x lO6 PA ( 75 psi)

3. Plumbing to B-valve - .345 x 106 PA ( 50 psi)

4. Drain line from B-valve to sump - .345 x 104 PA ( 50 psi)

Total Pressure Drop: 2.413 x 106 PA (350 psi)

6.9 Sizing.Criteria

The first step in sizing the actuator is to determine which operating conditions require blade angle control. The LAP pitch control consists of a single hydraulic actuator with in-place pitch lock. The in-place pitch lock allows loading conditions to exceed actuator capacity without loss of blade angle and the corresponding propeller speed increase. Based on Hamilton

Standard's experience with propeller systems, the following conditions were

chosen for actuator sizing criteria.

The first criteria is a sea level dive, zero horsepower, maximumvelocity,

with a 10%overspeed. At this condition the actuator shall be capable of

changing pitch. This is an extreme condition and could only happen in an

emergency or test condition. In addition, since maximumvelocity at zero

power could not be maintained, slowing the aircraft would reduce actuator

loads. A second overspeed condition looked at is somewhat more realistic but

is also conservative. It assumes a more typical flight condition of Vmax. at

1520 meters (5000 ft.) and 2237 kilowatts (3000 horsepower). An extreme

36
6.9 (Continued)

transient or failure causes the propeller, to then overspeed to 125%RPM. The

actuator should still be able to hold blade angle using only hydraulics at

this condition. Slowing the aircraft would allow the actuator to increase

blade angle and bring RPMdown to 100%without pitch locking. The third sizing criteria was actuator slewing capability at a maximumvelocity and 100%

RPMcondition. Since this is an experimental aircraft and a new blade configuration, a 20% margin was applied to all calculated blade loads.

The following table summarizes the actuator design sizing criteria.

Pitch Change Actuator Sizing Criteria

V-Kts % Kw

Case EAS RPM Aft. (SHP) Criteria

I. Overspeed 33i I10% SLSD 0 Actuator margin

2. Vma× 333 100% SLSD 3580 KW Pitch rate :

(4800) 9°/sec

3. Overspeed 333 125% 1520 meters 2237 Kw Actuator margin

(5,000 Ft) (3,000)

6.10 Stressing Criteria

To structurally analyze the actuator, a life/load spectrum must be determined. Since this is only a test program, Hamilton Standard made an approximation of a typical flight and the resulting loads generated on the actuator. It was conservatively assumed that there would be I0,000 flights or

start/stop cycles.

Figure 6.33 represents a typical flight. At startup, there is no load or

pressure acting upon the actuator. Whenthe Prop-Fan is started, pressure

reaches the high pressure relief valve (HPRV) setting° this pressure load is

seen by all internal parts even though there is no external load. The actuator is then unfeathered and a typical flight begins. It was assumed that

the operating loads varied from a maximumRPMand effective airspeed condition to a minimum TTM flight condition. This cycle is repeated50 times each

flight varying actuatormetered pressurebetween 60 and I00%of the regulated

supplypressure (PRY). The system also operatesat the PRV settingwith a

_- plus or minus 15% variationin pressure. The conditionsimulateseither pump

pulsationson the pressuresideor cyclidbendingmomentson the blade. In

addition,the actuatormust be capableof meeting proof and burst test requirements. The life/loadspectrumis summarizedbelow.

37
6.10 (Continued)

Pitch'Change Actuator

Life/LoadSpectrum

Start-StopCycles 10,000 0 - 9.48xI0_ PA (0 - 1375 PSI)

FlightCycle 500,000 6.55xi06+ 1.65xlO6 PA (950 + 240 PSI)

PressurePulsations Life 7.93x106. 1.21xlO6 PA (I150 + 175 PSI)

Proof II.72xI0_ PA (1700 PSI)

Burst 17.58xi06PA (2550 PSI.)

Use of this type of pressurespectrumon a11 elementsof the actuatorhas

been found to be a satisfactoryprocedureto analyze pitch control systems

for the type of conditionsactuallyencounteredin operation.

6.11 Typical MaterialAllowables

HamiltonStandardanalysisof low cycle fatigue,where there are a numberof

loadingconditions,is based on Miner's Rule of cumulativefatigue damage.

Each conditionuses part of the tota! life of the part in proportionto the

applied cycles divided by the allowablenumberof cycles. For example (see

Figure6.34), both the steadyand the cyclic stress (includingstressconcentration)is.calculatedfor one condition. The effectivefully reversed

stress is then calculatedand the allowablenumberof cycles is determined

from the S/N diagram. The ratio of actualcycles (example= 10,000)over the

allowablecycles (example70,000) is then calculated(exampleI/7 = .1428).

This is the proportionof fatigue life (14o3%)used by the operatingcondition. Nhen the summationof all the conditionsequalsone, a11 the life is

used up. Expressedmathematically"

CUMMULATIVEFATIGUEDAMAGE

n/N = n,/N, + n2/N_ + n3/N3

n = NUMBEROF APPLIED CYCLES

N = NUMBEROF ALLOWABLECYCLES

Nhen using this rule, HamiltonStandardonly works to a summationof .5, or

50% of the life. The design materialallowablesare based on x - 3a fatigue allowableswhich representa statisticalstressenvelopebased on test

data _hat results in less thanone failure in a thousandtest cases for similar type hardwarein a similarenvironment.

38
6.11 (Continued)

Calculating each individual point for each condition, including any stress

concentration, can be very time consuming and lead to possible errors. To

simplify the process, Hamilton Standard transforms all of the sizing conditions to a single curve called an S/P curve as shown in Figure 6.35. Since

all our loading is pressure related, all stresses are normalized for a single

pressure 6.9 x 106 PA (I000 psi). The stress is then calculated for 6.9 x

106 PA (I000 psi) applied load and the S/P (stress/lO00 psi) determined.

The curve is only good for a particular material and a particular loading

. condition. For example,Figure 6.35 shows that the actuatorcylinderis

burst limitedfor low stressconcentrationfactors (Allowable= 496 x lO6

PA (72,000psi). As the stressconcentrationgets large,such as in a thread

root (Kt = 4.0), the actuatoris fatigue limited(allowable= 345 x lO6 PA

(50,000psi)). Use of curves like this for each componentallows almost

immediatereview of the geometryof a compo'nento determineits structural

integrity.

39/40

7.0 DynamicAnalysis

" 7.1 Introduction

A dynamicanalysisof the large-scaleadvancedProp-Fan(LAP) systemwas per-

- formed to determine the pitch change governor gain needed for acceptable stability and also to investigate the overall dynamic performance of the control

, system. In additionto small signal stabilityand largerpower excursion

transientperformance,the expectedsynchrophaserperformancewas also investigated. Results show stabletransientperformanceat variousflight conditions with the pitch changegovernorgain chosen and also that acceptable

synchrophasingaccuracycan be achieved. This analysiswas specifically

gearedtoward the investigationof this particulartest hardwareand was not

intendedas indicativeof future Prop-Fansystem performance.

The LAP systemdynamic analysiswas accomplishedwith the developmentand use

of a non-linearcomputersimulationmodel. This simulationmodel was needed

to help understandthe dynamic inter-relationshipsamong sub-systemswithin

the overall.LAPsystemand to relate these dynamic inter-relationshipsto

systemstabilityand overall transientperformance. The model was used to

determinethe appropriatevalueof pitch change governorgain and to analyze

dynamicperformancecharacteristicsof this Prop-Fansystem. Previouspropfan and propellercontrolstudieshave shown the need to thoroughlyanalyze

and understandsystem stability,transientovershoot,rate limitationsand

generaltransientperformancein order to assist in the designof the control

hardwareand to definethe capabilitiesand limitationsof each specificcontrol application.

7.2 Model Simulation

A simplifiedLAP systemblock diagram consistingof the Allison Gas Turbine

501-M78engine and controls,a modified54460 prop control,a 54460 type synchrophaser,and the Prop-Fanmodel is shown in Figure 7°I. This functional

signalflow diagram shows the input/outputrelationshipsof each turboprop

sub-system.

The dynamicengine and fuel controlmodelswere created by HamiltonStandard

with the use of a customersuppliedsteadystate engine cycle deck and fuel

control systemblock diagram. The dynamic enginemodel block diagram is

- shown in Figure 7_2. This dynamic engine model format is successfully used

by Hamilton Standard to simulate free turbine turboprop engine. Block diagrams of the gas generator fuel control and the electronic engine control are

. shown in figures 7.3 and 7.4 respectively. The gas generator fuel control

provides transient fuel scheduling and isochronous gas generator speed gov-

- erning and the electronicengine controlprovidespower turbineoverspeed

limiting,power turbineinlet temperaturelimiting,and gas generatoroverspeed limiting.

41
7.2 (Continued)

The LAP Prop-Fan control is a hybrid control which combines 54460 hardware

and commuter prop technology to achieve satisfactory Prop-Fan governing. The

resulting control block diagram is shown in Figure 7.5. The Prop-Fan speed

governor strokes a pilot valve and flow is sent through a transfer bearing to

the servo. The servo stroke rotates a ball screw which commandsthe blade

actuator. The 54460 transfer bearing concept is combined with the servo,

ballscrew, and blade actuator technology of HSDcommuter propeller controls°

The Prop-Fan model, shown in Figure 7.6 uses power and thrust coefficient

maps which are functions of blade angle and advance ratio. The power turbine

torque output, accessory torque, and Prop-Fan torque absorbed are summedto

determine the torque available to accelerate the Prop-Fan, power turbine, and

gearbox inertias to determine Prop-Fan speed.

The synchrophaser model used for this study is a production C-130 electromechanical synchrophaser and is shown in block diagram form in Figure 7.7.

The model consists of a synchrophaser portion with proportional plus integral

control on phase error, an electronic trim portion which provides prop speed

derivative compensation and power lever anticipation, and a motor loop which

is the electro-mechanical interface to the Prop-Fan control.

Table 7.1 lists and defines the major variables used in the LAP system model.

7.3 Prop-Fan Control Performance

To determine an acceptable pitch change governor gain, the system response to

small and large changes in engine speed reference, which is indicative of engine power level, was investigated at various sea level and altitude flight

conditions. Figures 7.8 and 7.9 show acceptably stab]e response to 1.5% step

changes in engine speed reference at sea level, static and I0700M (35,000

Ft), 108 M/S (390 KTS) respectively. These conditions are believed to be the

least stable of those conditions at which the Prop-Fan is to be tested° Figure 7.10 shows the system response to a 2.0 sec engine speed reference increase from 88% to 100% at sea level, static conditions. Figure 7.11 shows a

2.0 sec engine speed reference decrease from 100%to 88% at the same condition. These large power excursion transients exhibit stable performance at

both power levels; however, the power advance of Figure 7.10 shows a Prop-Fan

peak transient speed of 108.8% which is caused by the combination of a lO.O

DEG/SECmaximumblade angle rate and a large blade angle excursion needed to

absorb the power increase. The maximumProp-Fan transient speed is 123.5%

and occurs when the speed reference change from 88% to 100% is made in 0.5

sec or less. The peak transient Prop-Fan speed is limited by the electronic

engine control power turbine overspeed limiter set speed of 120%. If this

power turbine overspeed limiter setting is lowered to 109%, the maximum

Prop-Fan speed is reduced to 112%. Figure 7.12 shows the simulation responses to 0.5 sec engine speed reference advances from 88% to 100%NGat a

sea level, static condition with the power turbine overspeed limit set at

42
7.3 (Continued)

" 120%and 109%. These same types of transients were run at IO,700M (35,000

Ft), 201M/S (390 KTS). A 2.0 sec speed reference increase from 78% to 100%

NGis shown in Figure 7.13 and the corresponding 2.0 sec chop from 100%to

, 78% NGis shown in Figure 7.14. No excessive transient Prop-Fan speed error

occurs at these conditions. The adjustable low pitch stop setting was 20.0

DEGBeta 3/4 for all the transients studied for the determination of the

" pitch change governor gain. Note that 88%NGwas chosen for the above sea

level, static.transients because this is approximately the minimum power required to assure Prop-Fan governing above 20.0 DEGBeta 3/4. The 78%NG setting was chosen for the above altitude transients because this dictates minimumfuel flow.

7.4 Prop-Fan Control Limitations

All engine power setting changes should be input slowly as to avoid large

Prop-Fan speed overshoot. Figure 7.12 shows the Prop-Fan response to a 0.5

sec power setting advance from 88% to 100%NGdone at a sea level, static

condition. The 123.5% Prop-Fan speed peak shown can be reduced to 112.% if

the fuel control power turbine overspeed setting is reduced from 120.% to

109.%. Therefore, Prop-Fan overspeed can be limited either by avoiding quick

power setting advances or by reducing the fuel control power turbine overspeed limiter setting.

It should be noted at this point that the results of this study are valid only for this Prop-Fan control test hardware and must not be viewed as indicative of Prop-Fan control performance potential The low blade angle rate,

large control deadband, and constant minimum blade angle schedule cause limitations in this specific test hardware application that cannot be generalized for all Prop-Fan controls.

Operation of the control in the reverse regime is severely limited in that

the transition into reverse must be done near minimum fuel flow, static conditions to avoid Prop-Fan overspeed and the maximumattainable power at full

reverse (-6°0 DEGBeta 3/4) without exceeding 100%NPROPis approximately

130500 watts (175.0 HP) at sea level, static conditions.

Simulation of reversing transients including simple reversing from flight

, idle and aborted takeoffs were not run because the low pitch stop will not

allow the control to attempt these transients. The low pitch stop acts as

protection against excessive transient overspeed which may occur if the stop

were not there to limit blade angle travel.

43

°
7.5 Synchrophaser Performance

Use of a production 54460 synchrophaser in conjunction with the LAP pitch

change system is impractical because the Prop-Fan control deadband of this

test hardware causes phase error limit cycle of approximately + 50. DEG.

Several possible options to achieve adequate synchrophasing were studied.

One option was to increase the synchrophaser gain to decrease the limit cycle

amplitude. The limit cycle amplitudecan be decreasedto + 26. DEG, as shown

in Figure 7.15, if the synchrophaserproportionaland integralgains are increased by a factor of eight. This does not yield acceptablesynchrophaser

performance. An option studied that does yield acceptable synchrophaser performance is the addition of an electronic dither signal to the synchrophaser

voltage output, as depicted in Figure 7.7. This dither signal moves the control back-and-forth through the deadband so that, on the average, the control

acts as if no deadband exists. If a 2.0 HZ square wave dither voltage signal

of + 2.5V is added which generates an NTRIMsignal of + 8.5 RPM, the limit

cycTe amplitude is decreased to ± 3°4 DEG. This result is seen in Figure

7.16. The 2.0 HZ NTRIMsignal, which is the input to the hydromechanical

Prop-Fan control, is the waveform which results from a square wave voltage of

+ 2.5V being added to the synchrophaser voltage output, as shown in Figure

7.7. If the dither amplitude is increased to + 3.5V, the resulting + 10.75

RPMNTRIMsignal reduces phase error to less than + 0.5 DEG. Figure-7.17

shows that the increased dither amplitude moves both the servo and the blade

actuator back=and-forth through their respective deadbands, thus minimizing

the effects of both non-linearities. Dither has been successfully used in

commuter propeller applications to counteract the negative effects on synchrophasing of prop control non-linearities. The synchrophaser accuracies

defined in this report must be viewed as analytical estimates since random

atmospheric condition changes and engine power variations which add to synchrophasing inaccuracy were not included in this study. All synchrophaser

accuracy investigation was done at I0700M, 201M/S (35,000. FT, 390 KTS)

flight conditions.

7.6 Summarv

The results of this study are valid only for this LAP test hardware and cannot be generalized for future Prop-Fan control applications.

The pitch change governor gain recommendedas a result of this analysis is

2.5 DEG/SEC/%RPM. Simulation studies show acceptable stability at those operating conditions at which the control system was designed to govern.

Since no minimum blade angle schedule exists in this control, an adjustable

low pitch stop is needed to assure that the system will not be forced to govern at low blade angles. Stability margin decreases considerably at low

blade angles.

44
7.6 (Continued)

. Changes in power setting must be input slowly to avoid excessive transient

Prop-Fan speed variation. Since larger power excursions can occur at sea

level than at altitude, the system is much more susceptible to Prop-Fan overspeed at sea level operation.

This control must not be used for gross reversing transients because severe

" Prop-Fan overspeed could result. All transitions to reverse must be done under static conditions at or near minimum power. Operation in the reverse regime with a minimum Beta 3/4 of -6.0 DEGwill only permit approximately

- 130,500 watts (175 HP) before exceeding 100%NPROPat sea level, static conditions.

The modified 54460 hybrid Prop-Fan control used in the LAP system has a ±

.205% RPMgovernor deadband and a + .25 DEGBeta 3/4 actuator hysteresis

band. These control non-linearities have been found to help Prop-Fan control

stability and adversely affect synchrophaser performance. A production 54460

synchrophaser will not accomplish satisfactory synchrophasing accuracy due to

these Prop-Fan control non-linearities. The combination of these Prop-Fan

control non-linearities and synchrophaser integral control causes a phase error limit cycle of ± 50 DEG.

An effective method of reducing the effect of Prop-Fan control hysteresis and

deadband on synchrophaser accuracy is the addition of an electronic square

wave dither signal to the synchrophaser voltage output. The phase error variation is reduced to + 3°5 DEGif a 2.0 HZ square wave dither of ± 2.5V (±

0.5% RPM) is added to'the synchrophaser output. This variation in phase error can be further reduced to less than + 0.5 DEGif the dither amplitude is

increased to + 3.5V (+ 0.63% RPM). Randomatmospheric disturbances and engine power variations-have not been included in this synchrophaser accuracy

analysis. The C-130 synchrophaser was arbitrarily chosen for use in this analysis as a typical 54460 type synchrophasero

45/46

8.0 Beta Control

• The SR-7L will be tested on the static test rig at Wright Field to explore

the stall flutter boundary and in the Modanewind tunnel to explore the high

speed flutter boundary. For these tests it is desirable to operate the

- Prop-Fan with direct blade angle control and not with speed control. These

boundaries are established by varying speed over a range while holding blade

. angle constant. This procedure is repeated for many different blade angles.

To incorporate blade angle control on the LAP, the rotary input to the pitchlock screw from the servo is replaced by a D.C. motor and gearhead. This

arrangement replaces the governing mode configuration and is shown in Figure

8.1.

: The motor is powered by a remote open loop electronic controller. The controller is designed to pulse power in discrete time increments to obtain an

incremental change in blade angle. The actual change in blade angle is dependent on the length of the pulse and the level of friction at the pitchlock

screw. Power is transmitted across the rotating interface through two slip

rings located on the spinner bulkhead. The motor, which is mounted on the

front of the dome, is a D.C. permanent magnet planetary gear motor. The motor operates on 24 volts + 2,5 volts at 6500 RPM(no load speed). The planetary gearhead provides a 780 to 1 speed reduction ratio. A flexible shaft is

mounted on the output shaft of the motor and drives the pitchlock screw. The

output speed of the motor is a function of pitchlock screw friction. This

relationship, expressed in deg/sec of blade angle change vs. screw friction

torque, is shown on figure 8.2. The maximumexpected torque requirement due

to the pitchlock screw friction is 565 mm-N(5 in-lbs). This results in an

actuator slewing capacity of 1.36 deg/sec in the beta control mode. This

rate is sufficient since there are no dynamic response requirements in the

beta mode. The main pumpflow could provide 3 deg/sec, therefore the standby

pump flow is not required and could be removed from the system during beta

control. However, to minimize the modifications necessary to convert from

the governing mode to the beta mode, the standby pumpwill not be removed.

Blade angle control in the governing mode is accomplished by motion of the

servo which changes blade angle via the ballscrew, pitchlock screw, and beta

valve. This sequence of events is detailed in sections 3.1.1 and 3.1.2. In

order to give the motor full control of blade angle the servo and ballscrew

must be disconnected from the pitchlock screw by removing the quill shaft.

This disables the control signal and allows the motor to work directly on the

pitchlock screw without having to fight the servo output.

The servo is loaded against the low pitch stop by setting the governor at

I05% speed via the condition lever. This prevents uncontrolled cycling of

the servo during beta control operation. Metered pressure is regulated by

the delta-P regulating valve at slightly higher than one half of supply pressure. This condition provides the proper amount of lubrication and cooling

flow to the transfer bearing.

47/48

9.0 TRUNNIONDESIGN

" The trunnion assembly provides the link between the blade and the blade pitch

actuator. It transmits the torque generated by the blade about the retention

radial axis to the actuator and translates actuator axial motion to blade to-

, tation about the retentionradial axis. Figure 6.20 is a sketchof the trunnion assembly. The trunnionis boltedto the blade shank at four places.

The trunnionroller bearingmates with "ears"on the actuatoryoke. Figure

9.1 shows the actuator/bearingrelationshipat the extremesof travel and at

mid travel.

The major componentof the assemblyis the trunnionwhich is bolted directly

to the blade shank. This part is sized to withstandthe blade/actuatorloads

imposedon ito Since the blade shank materialis aluminum,threadedinserts

were incorporatedto take advantageof the increasedshear and tensile stress

areas of the oversizedthreadsof the inserts. The threadsin the blade

shank will be rolled to increasethe fatigue life. The bolts are designed to

react the moment caused by the offset of the trunnionbearing reactionplane

and the blade shank face Lockingcups provide a means of lockingthe bolt

heads and also functionas a means of securingthe dowel pins.

The dowel pins are used for torque restraint. They are also used as a pilot

during assemblyto obtain the proper angularrelationshipbetween the blade

airfoil and the trunnionbearingcenterline. This results in the blade and

trunnionbeing a matched set.

Fatigue life is a concernwhenevercomponentsare designedto withstandcyclic loading. One means of increasingthe life of the blade shank and trunnion, is to shot peen highly loadedareas of each component. To increasefatigue strength,both the blade shank and trunnionmountingface are shot

peened,also, the trunnionface has been undercutto increasethe interface

pressureand thus reducethe possibilitiesfor fretting. The remainingshot

peenedarea includesthe trunnionpin radius,and was incorporatedprimarily

for the higher allowablestrengthaccompanyingshot peening.

The trunnionrollerbearing is throughhardenedsteel and mates with surface

hardenedsteel wear plateson the actuator. A detaileddescriptionof the

rollerbearing and wear plate can be found in Paragraph6°6.

, 9oi TRUNNIONLOADING

As stated above, the trunnion reacts the blade forces and transmits them to

. the actuatorsystem. The trunnionwas positionedat a maximumpin radius of

7°62 cm (3.00 in.) from the propellerpitch changeaxis with the actuatorat

- midstrokeand the blade deliveringits maximum forces. Therefore,the summation of the blade loads,includingCTM (CentrifugalTwistingMoment),ATM

(AerodynamicTwistingMoment) and FTM (FrictionalTwistingMoment),and re-

. acted at the maximummoment arm of the trunnion. This summation of blade

49
9.1 (Continued)

loads, TTM (Total Twisting Moment), is applied to the trunnion as a steady

force. In addition, the ATMcomponent of the TTMis applied as a cyclic

load. This combined load was used to determine the physical size of the

trunnion for high cycle fatigue (HCF) (10 _ cycles).

Another loading case, the number of start - stop cycles was considered for

low cycle fatigue (LCF) (104 cycles). These two loading cases wi11 size

the trunnion. Figure 9.2 shows the breakdown of trunnion loading.

The calculation of the bolt loads from the blade TTMwas determined by

straight beam theory. Considering the trunnion as a straight beam fixed at

both ends with an intermediate moment (the momentcaused by the offset from

the trunnion bearing midplane to the blade shank face), the applied bolt

loads were determined to be 21305.9 + 2473.1 N (4790 + 556 lb.). However,

this applied load must be modified by the effects of preload and temperature

difference° The minimum required preload force was calculated using the maximum applied load plus the additional bolt load due to a temperature change

of 43°3 degree C. (110 degree F.)o Bolt friction factors and torque tolerance requirements were used in determining the maximumresulti.ng preloado

The final bolt load used in the stress computations is 34249,6 - 52264 N

(7700 - 11750 lb.) preload force and 21305.9 ± 2473.1 N (4790 ± 556 lb.) applied load caused by external trunnion loadings. The bolts were also analyzed for resistance to shear loading caused by dowel failure of 20967.9 +

5239.7 N (4714 + 1178 lb.) for the outer bolts and 17084.8 + 4283.4 N (3841 +

963 Ibo) for the bolts straddling the trunnion bearing. The tensile bolt

loads are transmitted to the blade shank through threaded inserts, and computerized analyses H466 and H088 were used to calculate thread stress. The

results of these analyses will be presented in a later sectiono

Dowel pin shear 1oadings were resolved in the same manner as the bolt shear

loads. In the analysis, both dowels were considered to carry the same load.

This was determined by applying the direct shear force, then calculating the

dowel pattern c.g. and applying the moment load. The resulting shear load is

then the square root of the sum of the squares of the direct load and moment

load and was calculated to be 27795.5 + 3229°2 N (6249 + 726 lb.). From this

load, dowel pin shear and bearing stresses can be determined.

9.2 TRUNNIONSTRESSANALYSIS

The externally applied design loads, for reference, are as follows"

HCF= 46704 • 5426°6 N (10500 + 1220 Ib.)

LCF = 23352 + 23352 N (5250 + 5250 lb.)

Figure 9.3 is an illustration of the trunnion assembly showing the components

and the sections for which the stresses were calculated. The trunnion sections were analyzed using beam theory, where both ends are fixed and the beam

5O
9.2 (Continued)

" has a length equal to the arc length between the outboard bolts with a moment

applied midway. This loading analysis conservatively neglects the restraint

of the two inner bolts, and also the support provided by the blade shank.

" Trunnion bending stresses were then determined using the resultant bending

moment at each of the four cross-sections of Figure 9.3. Bending stress was

, calculated for the trunnion pin base and was modified by a stress concentration factor to account for the fillet radius between the trunnion pin and the

main load carrying section of the trunnion. The trunnion section bending

stresses are as follows"

SECTA-A HCF = 2 ]lIE8 + 2.312E7 N/M2 (30582 + 3353 psi)

LCF = 1 054E8 + 1.054E8 N/M2 (15289 + 15289 psi)

SECTB-B HCF : 2 742E8 + 3.185E7 N/M2 (39765 + 4620 psi)

LCF = 1 371E8 + 1.371E8 N/M2 (19883 + 19883 psi)

SECTC-C HCF = 3 983E8 + 7.637E7 N/M2 (57774 + 11077 psi)

LCF = 1 992E8 + 3.286E8 N/M_ (28887 + 47663 psi)

SECTD-D HCF = 3 226E8 + 5.435E7 N/M_ (46790 + 7883 psi)

LCF = I.613E8 + 2.339E8 N/M2 (23395 + 33923 psi)

TRUNNIONPIN RADIUSHCF = 4.628E8 + 8.066E7 N/M2

(67124 + 11698 psi)

LCF : 2°314E8 + 8.066E7 N/M2

(33562 + 50343 psi)

Sections C-C and D-D include a bolt hole and dowel hole, therefore a stress

concentration factor has been applied to the cyclic stresses calculated at

these sections. Figures 9-4a and 9-4b are GoodmanDiagrams for the trunnion

material, and are used to illustrate the stress margins available at HCFand

LCF respectively for each of the trunnion sections.

Figure 9.5 presents the bolt loads and equations used in calculating the bolt

shank and bolt head fillet stress. The shank stress is a P/A type stress,

. however, the externally applied loads travel through the entire bolted joint

and are modified by the "joint efficiency". The preload is a direct load

applied to the bolt and therefore is not effected by "joint efficiency."

This "joint efficiency" is a measure of the relative stiffness of the bolt

compared to the stiffness of the joint and is used to modify the external

loads applied to the bolt.

51
9.2 (Continued)

The dowel pin load was calculated at maximumTTMo_ the blade and referenced

as follows:

DOWELLOAD : 27795.5 + 3229.2 N (6249 + 726 Ibo)

Dowel shear was calculatedusing the equationT/A and determinedto be

2o194E8± 2.549E7 N/M2 (31826+ 3697 psi) which has a marginof safetyof

360%. The dowel pins, under sTde loading,bear againstboth the trunnionand

the blade shank° The bearing stress is obtained by dividingthe loadon the

pin bearingagainst the edge of the hole by the bearingarea, where the area

is the productof the pin diameterand shank thickness. Substitutingthe pin

side load in the equation,results in a bearing stressof 1.149E8+ 1.335E7

N/M2 (16664± 1936 psi). This stresshas a margin of safetyof 365% in the

blade shank, > I00% in the trunnionand > fOOD% in the dowel pin.

In order to properlytransmitthe shear loads, the dowel pins were pressed

into both the trunnionand blade shank with a diametralinterferencefit of

o0381- .0635cm (o0015- °0025 in.). Using the standard,thickwalled pressure vessel equations,the values for radial, shear and hoop stressfor both

the trunnionand blade shank were computedand listedas follows along with

the correspondingmargin of safety:

SHANK STRESS DUE TO PRESS FIT

RADIAL= 1.31E8N/M2 (18937psi) MS = 164%

SHEAR = 1.74E8N/M2 (25250psi) MS = 18.8%

HOOP = 2.18E8 N/M2 (31563psi) MS = 58.4%

TRUNNIONSTRESS DUE TO PRESS FIT

RADIAL= 3.88E8N/M2 (56250psi) MS = 202%

SHEAR = 5.17E8N/M2 (75000psi) MS = 36%

HOOP = 6o46E8 N/M2 (93750psi) MS = 81%

9.3 BLADE SHANK THREADS

The blade shank threads carry the trunnionloads as supplied to them through

the bolts. The threads are manufacturedby a rolling process and they have

the advantageover cut threadsas the threadstrengthand fatigue life are

increased. One factor contributingto higher fatigue strengthis that rolled

threads have a minimum foot radiusof .0762mm (.003 in°)while the minimum

radiusof cut threads is °0254mm (.OOl in.)o The root radius is incorporated in the determinationof the tooth stressKt, therefore,the smaller the

radius, the largerthe Kt.

52
9.3 (Continued)

" From published reports and tests conducted at Hamilton Standard, it was determined that thread loads from the preload and applied load effect the bolted joint through different load paths. As was previously explained, the pre-

, load force is constant and not effected by joint efficiency, whereas the applied load is modified by joint efficiency. The following equation was used

for thread stress calculations and can be summarized into the basic form:

THREADSTRESS= TOTALSTEADYSTRESS+ TOTALCYCLIC STRESS

where the total steady stress is the summation of the preload stress and the

applied load steady stress, and the total cyclic stress is the applied load

cyclic stress. Independent of preload and applied load, the thread stress

has two components, one is the member stress and the other is referred to as

tooth stress. Both the member stress and tooth stress are determined by computerized analyses H088 and H466. H088 calculates the member stress of each

thread considering it as a shell of revolution, and the tooth stress is determined through H466 which basically uses the equation P/A + MC/I for each

thread. For both preload stress and applied load stress, the member and

tooth stress are combined through the use of a proximity factor, A, which accounts for the fact that the member stress and tooth stress occur at different locations on a full radius thread root. The proximity factor, A, is calculated from the equation"

A : I/If + .47(MEMBERa/TOOTH_)]

Nhere .47 is a function of the thread form. Therefore, the preload stress is

determined using the following equation"

TOTALPRELOAD

STRESS= PRELOADMEMBERSTRESS(H088) + A * PRELOADTOOTHSTRESS(H466)

and the applied load mean stress from the equation"

APPLIEDLOAD = APPLIEDLOADMEMBERSTRESS(H088) * KfM * Nf +

STEADYSTRESS APPLIEDLOADTOOTHSTRESS(H466) * KfT * Nf

where Nf is the joint efficiency, KfM is the stress concentration factor in

• fatigue for the member stress and KfT is the stress concentration factor in

fatigue for the tooth stress. It should be noted that no stress concentration factors are applied to the steady stresses resulting from the preload

and applied load. KfM is determined from the case of an internal hole with

an internal notch subjected to tension loading, and is modified three ways,

'- once for the angle of the thread form, once by a multiple notch factor, finally because the material displays a different behavior in the presence of a

notch in fatigue than under static loading conditions. Kt for the tooth

stress is determined from an empirical equation developed by Heywood(Reference 4)"

KtT= I + .26 (e/R) "7

53
9.3 (Continued)

where e is 7/16 times the thread pitch and R is the thread root fillet radius. KfT is determined in the same manner as KfM, as related to notch sensitivity of the material.

The total cyclic stressonly containsthe cyclic componentof the applied

load and is determinedusing the same equationfor applied load steady stress

except the force is the cyclic load"

TOTAL

CYCLIC = CYCLICAPPLIED LOAD MEMBER STRESS (H088)*KfM * Nf + A *

STRESS CYCLIC APPLIED LOAD TOOTH STRESS (H466)* KFT * Nf

Figures9.6 and 9.7 illustratethe thread stressdistributionin the threaded

hole and also the resultingstressmargins.

The trunnionassemblyweight and the resultingcentrifugalload is"

TRUNNION = 7.47 N (Io68 LB)

TRUNNIONBEARING&.WASHERS = 3°75 N (0°85 LB)

SHANK ATTACHINGHARDWARE = 2°85 N (0°64 LB)

TOTAL = 14.10 N (3.18 LB)

RESULTINGC.Fo = 6863 N (1543 LB)

54
10.0 CONCLUSIONS

" Since the primary purpose of the LAP program is to evaluate the structural

integrity and aeroacoustic performance of the blades, the requirements for

the pitch control system were neither critical nor stringent. As shown in

* this report the designapproachused both existinghardware,conservativeanalysis and stateof the art manufacturingpracticesthe origina!assumptions

, for controlgains and accuracywere confirmedby analyticalmodeling. The

calculatedvaluesof these systemparameterswere eitherthe same or better

than originallyassumed. Therefore,the Prop-Fanpitch control system as designed is suitablefor use in the LAP program.

55/56
°
II.0 References

" I. Large Scale Advanced Prop-Fan (LAP) Blade Design Report, NASACR174790

2. Large Scale Advanced Prop-Fan (LAP) Hub and Blade Retention Design

" Report, NASACR174786

. 3. Large Scale Advanced Prop-Fan (LAP) Spinner Design Report, NASACR174785

4. R.J. Roark, "Formulas for Stress and Strain", 4th Edition, 1965

57/58

TABLES

59/60

TABLE! 5.1. SPEEDS AND CAPACITIES FOR THE LAP AND E2/C2

L.A.P. E2/C2

1698 PROPRPM 1106 PROPRPM

LEFT HAND ROTATION RIGHT HAND ROTATION

R.P.M. CAPACITY R.P.M CAPACITY

?2.47 Q.P.M. 20 Q.P.M. MAIN 5294 4712 PUMP AT 1375 PSi AT 1375 PSi

SCAVENGE 5294 32.92 Q,P.M. 4712 29.3 Q.P.M, PUMP

39.86 Q.P.M. 40 (:;).P.M.

STANDBY 4500 4516

._ PUIVP AT 1375 PSI AT 1375 PSI

GOVERNOR 4072 NA 3517 NA

IN ORDERTO ACHIEVE THE SAMEFLOW DIRECTION INTHE

LAHADP.TOPUMPS BE REVERSED, AS IN THE E2, AS/C2INDICATED PUMPS,TIlEIN FIGU ..IPUGu_5.GUREC 1IVE4 GEARS _S

TABLE 5.2. MAIN & STANDBY REGULATING VALVE DATA

THIS VALVE IS SIMILAR TO THE MAIN

AND STANDBY VALVE USED ON E2/C2 PROPELLER

CONTROLS. DEVIATIONS ARE AS FOLLOWS:

-SLEEVE AND PLUNGER ARE MATCHED AT

THEIR MATING DIAMETER.

=USE,OF SHIMS BETWEEN THE SPRING

AND PLUNGER.

MATERIALS

SLEEVE: AMS 6415, 40-44 HRC

PLUNGER: AMS 6270OR AMS 6272

26-44 HRC

SPRING: AMS 5112

,_ DIMENSIONS

SLEEVE & PLUNGER DIACLR: 0.0002-.0005

SLEEVE &HOUSING DIACLR: O.003-.007

PERFORMANCE

" PISTON AREA: 0.3705 IN Z

SPRING RATE: 83 LB/IN

LOAD AT WORKING HEIGHT 56 ± 2.8 LB

OPERATING STROKE: 0.067 IN

61
TABLE 5.3. ORIFICE PACK DATA

ORIFICE PRESSURE RATIO:

(PS-PB)/(P_PD)= 0.15

NET FLOW AT :

PS=1140 PSi

& 70 DEG F. .48 QT/MIN

NET ORIFICE RATING: 4662 I OHMS

(PS_ PD) (D =0.0127 IN)

RATING OF PS_PB ORIFICE: 1800 1 OHMS

(D = 0.0205 IN)

RATING OF PB-D,,,.-PDORIFICE: 4300 ! OHMS

(D = 0.0133 IN)

TABLE 5.4 GOVERNOR PERFORMANCE DATA

GOVERNOR RPM (NG) @ 100%SPEED 4072 RPM

GOVERNOR-PROP SPEED RATIO (NG/Np): 2.3981

FLYWEIGHT FORCE EQ: FFW = 2.43 X I0 (NG) 2

@ 100% RPM: FFW = 40.29 LB

FLYWEIGHT RATE EQUATION: KFW = 1.3 X 10 (NG) 2

@ 100% RPM: KFW = 21.56 LB/IN

SPEED SPRING RATE (Ks): 80 LB/IN

SPEED SET LEVER RATE (Kt): 1042 LB/IN

SPEED BIAS RATE (KBG): 6.99 LB/IN

SYSTEM SPRING RATE: KSy S = !/(!/(Ks+KFw=KBG)+!/K I )

@ 100% RPM: Ksy S = 98.31 LB/IN

SENSITIVITY @ t00% RPM: 0.0082 IN/% RPM

AXIAL ACCELERATION SPEED SHIFT, 0.3565 % RPM/G

i

62
TABLE 5.5. GOVERNOR MATERIAL LIST

MATERIALS

FLYWEIGHT: MIL-T-21014 TYPE I! CL 1

SPOOL: AMS 6260, 30-43 HRC

SLEEVE: AMS 6260, 30-40 HRC

SCREEN: AMS 5510

SPRING: HS 261

DIMENSIONS

DIA. CLR. SLEEVE & SPOOL: 0.0003-0.0005 IN,

SURFACE FINISH, SLEEVE/SPOOL MATING DIA: 8

MAX SCREEN HOLE DIA.: 0.007 IN.

SCREEN OPEN AREA: 28%

TABLE 5.6. FEATHERING VALVE DATA

FEATHERING VALVE

THE FEATHER VALVE IS SIMILAR TO THE ONE

USED ON THE E2/C2 PROPELLER CONTROL. THE

DIFFERENCE BEING THAT ONE SEAL GROOVE HAS

BEEN MACHINED OFF THE SLEEVE.

DIMENSIONS

DIACLR, SLEEVE & SPOOL: .0004=.0008 IN

DIACLR, SLEEVE & HOUSING: .004-.0061N

SPRING FREE LENGTH 2.091 + .04 IN

MATERIALS

SPOOL: AMS 4150, 46HRBMIN

SLEEVE: AMS4122, 88 HRB MIN

SPRING: AMS 5673

SPRING SEAT: AMS 5504

PERFORMANCE

_ PISTON AREA: =30498 IN 2

SPRING RATE: 14.6 LB/IN

SPRING LOAD:

, . VALVE OPEN 24LB

• VALVE CLOSED 10LB

SENSITIVITY : 47.87 LB/IN

63
TABLE 5.7 •FRONT AND REAR LIP SEAL DATA

FRONT

LIP SEAL

MATERIAL: VlTON

DIA CL/I NTF'.

SEAL/RETAINING RING: 0.0015-0.0185 INTF.

SEAL/OiL SHIELD: 0.019-0.097 INTF.

SEALANT: MIL-S-880Z CLB-2

CONTROL OIL SHIELD

MATERIAL: AMS 6415

HARDNESS: 26-33RC 

BOND: HS 812 GRADE 3

PLATING: NICKEL PER AMS 2404

PLATING THICKNESS: 0.0008-0.0012

REAR

LiP SEAL

MATERIAL: VITON (SAME AS FRONT)

DIA CL/INTF.

SEAL/RETAINER: 0.001 5-0.01 65 INTF.

SEAL/BARREL SUPPORT: 0.019-0.0 97 INTF.

SEALANT: MI L-S-8802 CLB-2

LIP SEAL RETAINER

MATERIAL: AMS 5616

HARDNESS: 32-38 RC

D|A CL/INTF.

RETAINER/PUMP HSG 0.001-0.009 CL

RETAINER/HEAT EXCHANGER 0.001 MIN CL

ALL DIMENTIONS ARE GIVEN IN INCHES

TABLE 5.8. IDLER GEAR SUPPORT DATA

SUPPORTSHAFT

MAT'L: AM56415

HARDNESS: 26-33 RE

DIA. C:L/INTF:

SHAFT/HSG.BORE: 0.0002-0.0015 CL

SHAFT/GEAR SUPPORT: 0.0002-0.0015 CL

SHAFT/SPAC_ER. 0.0(X)5-0.008 CL

SHAFT/GEAR Si..I=R)RTPLATE: 0,035 MIN. RAD. CL.

SHAFT/SHIM : 0.072-0.078CL

WAVEWASHER

MAT'L:. AMSSI21

HARDNESS: 45-50RC

WORKINGHGT: 090".094

LOADATWORKINGHGT: 82-75LBS.

FREEHGT: .137_+;005

THICKMESS: .068+_.003 

BALL BEARINGS

MAT'L/HARDNESS:

RING: SAE 52100/60 RC MIN

BALLS: SAlE52100/60 RC MIN -

SPEED: 3529 RPM

NO. OF BALLS: 36

BALL DIA: .1875

RADIAL PLAY: .00050.0015

LUBRICATION: MIL-H-5606 ,_

THRUST LOAD: 100 LBS

EXPECTEDLIFE: 19500HRS

DIA. CL./INTF.

BRG./GEARSUPPORT: .0001-0.0015 CL.

BRG./IDLER GEAR: ,0005-0.0013 CI,..

ALL DIMENTIONS ARE GIVEN IN INCHES.

64
TABLE. 5.9. LAP GEAR DATA

STANDBY TRANS. LARGE SMALL GOV.

MAIN PUMP PUMP SCAV.PUMP BRG. IDLER IDLER

_" NO. TEI_TH 34 40 34 106 51 60 52

DiA. PITCH 12 12 12 12 12 16 16

PRESS. ANG. 22,50 ° 22.50 " 22.50 o 22.50 o 22.50" 20 o 20 •

FACE WIDTH 0.20 0.20 0.20 0.22 0.20 0.15 0.20

(IN)

RPM 5294 4500 5294 ! 698 3529 3529 4072

TORQUE 76 161 2.1 427 • 57 11.54 10

(IN-LBS)

HS 256

MAT'L AMS 6260 AMS 6260 AMS 6260 AMS 6260 AMS 6260 AMS 6260 PER HS 178

H CASE 685VHN 685VHN 685VHN 685VHN 685VHN _

D

N S CORE 30-42HRC 30-42 HRC 30-42 HRC "30-42 HRC 30-42 HRC 30-42 HRC 32-35 HRC

TOOTH WT 53.7# 96.6# 1.5# 96.6 # 26.84# 6.2 # 6.2#

LOAD WD 333# 373# 277# 373# 290# 180# 180#

|

DYNAMIC

HERTZ 214,705 PSI 213,950 PSI 195,840 PSI 213,950 PSI 207,085 PSI 183,780 PSI 183,780 PSI

DYNAMIC "68,1"15 PSI 64,555 PSI 56,660 PSI 63,580 PSI 50,190 PSI 51,925 PSI 40,500 PSI

BENDING

• SIZED BY STANDBY PUMP MESH

65
TABLE 6.1 . PITCHLOCK SCREW ASSEMBLY - MATERIAL LIST

PITCHLOCK SCREW

MATERIAL: AMS 6265

HARDNESS: 35-43 (CORE), 90HR15N (CASE) 

THREAD: 2.19 P.D.

BACKLASH: 0.002 IN.

FRICTION TORQUE 5 IN.-LB

PITCHLOCK SUPPORT

MATERIAL: AMS 6265

HARDNESS: 35-43 (CORE), 90HRISN (CASE) 

BALLSCREW QUILL

MATERIAL: AMS 5616

HARDNESS: 32-38 RC 

BETA VALVE CONNECTING ROD

MATERIAL: AMS 6415

HARDNESS: 36-40 RC

LOW PITCH STOP

MATERIAL: AMS 6415

HARDNESS: 36-40 RC

ADJUSTMENT 40° _3/4 TO -10 ° _3/4

TABLE 6.2. TRUNNION LOADS

MAX DESIGN LOAD -

CONDITION: (100% SPEED; SLSD; 0 POWER; 333 KEAS; 440 _3/4)

CTM: 22,130 IN# X l.Z* = 26,556 iN@/BLADE

ATM: 3,052 IN# X 1.2" = 3,662 IN#/BLADE

FTM: CL X PD/2 X 0.005 = 1,256 iN#/BLADE

TTM: 31,474 IN#/BLADE

*(CONTINGENCY FACTOR)

TRUNNION LOADS (@ 3.00 MA):

(LCF) START-STOP I 04''" 0 TO ! 0,500#

(HCF) VIBRATORY 107'_' 10,500# + 2,625# "

NO ) YIELD CASE (125% RPM) 16,000#

FTM _ ULT. CASE (140% RPM) 20,000#

66
TABLE 6.3. ACTUATOR EAR WEAR PLATE SUBSURFACE SHEAR STRESS SUMMARY

EXPERIMENTAL DUCTED EXPERIMENTAL EX,'ER, V PROPELLER FOR / TOL O ER LAP

V/STOL PROPELLER VARIABLE VEREIGTE FLUGTEDINISCHE pROPELLER

FOR BELL AIRCRAFT CAMBER pROPELLER WERKE

MATERIAL AMS 6260 AMS 6270 AMS 6285 AMS 8260 AMS 6268

CASE: 90RI 5N 90R ! 5N 90RI 5N 90R! 5N 59-64 R C

• CORE: 30-40 R C 30-40 R C 35-43 R C 35-43 RC 38-43 R C

CASE ,,_EPTH: 0.058-0.076 0.060-0.075 0.130-0.170 0.035-0,048 0,080-0.090

GEOMETRY

PLATE THICKNESS: 0.195 0.217 0.378 0. 156 0.200

ROLLER O.D.: 1.400 1.800 Z.800 1.378 2.000

ROLLER CROWN RADIUS: 6.0 14.0 24.0 30.0 20.0

MAX DESIGN LOADS 5,000 9,180 28,000 3,900 10,800

SHEAR STRESS

(@ MIN CASE DEPTH) 93,300 86,400 80,100 94,300 75,000

ALL DIMENTIONS ARE GIVEN IN INCHES

ALL LOADS ARE GIVEN IN POUNDS

ALL STRESSES ARE GIVEN IN POUNDSPER SQUARE INCH

67
TABLE 7.1. LAP SYSTEM MODEL VARIABLES

DESCRIPTION OF VARIABLES

NI, NG - GAS PRODUCER SPEED (100%- 14300. RPM)

N2, NPT - POWER TURBINE SPEED (100%- !!500o RPM)

NPROP " PROP-FAN SPEED (100%- 1698. RPM))

WF o METERED FUEL FLOW

P3, CDP - COMPRESSOR DISCHARGE PRESSURE

QPT - POWER TURBINE TORQUE

T45, TIT - TURBINE INLET TEMPERATURE

NREF - PROP-FAN SPEED REFERENCE

NMAS'rER o PROP-PAN SPEED REFERENCE FOR SYNCHROPHASER

NTRIM ° PROP;FAN SPEED TRIM FROM ELECTRONICS

XACT - BLADE ACTUATOR STROKE

XE - SERVO STROKE

THETE - SYNCHROPHASER PHASE ERROR

ACCHP - ACCESSORY HORSEPOWER

PLA - POWER LEVER ANGLE (REPRESENTATIVE)

i

68
FIGURES

69/70

FIGURE 2,1

LARGE SCALE ADVANCED PROP-FAN (LAP)

71/72

L-- h

• ,., _._

' tr

"111_ ____J

/

5")/

rrl _ _ ..__ 0 0 ::

1>

° Z_

,'-I,

o _:_.

r" =__,=,--

it

"

Actuator valve

I

LAP

CONTROL

SCHEMATIC

Atmospheric

sump

Sump relief

valve

,-----------------,

I I

I

II

I

II

I

I

I

I

I

I

II

I

I

I

r---Hxbypa-;S-------

I valve

I aIIII

I Feather

---..

Feather

@) 0 0 0

cam /

/ Feather

_

Main

filter

Standby

filter

ow pitch stop

7283

FIGURE 3.1 LAP CONTROL SCHEMATIC

73/74

FIXED LOW PITCH STOP

TRUNNION

FEATHER INC. REVERSE DEC.

TRANSLATING

CYLINDER

\

FEATHER STOP :ENTRIC

PIN AND

INC. DEC. ROLLER

PITCH PITCH

YOKE

.,..3

ROTARY INPUT

STATIONARY PISTON

--'" _ ACTUATOR

VALVE

_A PITCHLOCK

SCREW

PITCHLOCK GAP = 0.028+0.002 IN.

DJUSTABLE LOW PITCH STOP

FIGURE 3.2. PITCH CHANGEACTUATOR & PITCHLOCK
P METERED FROM

P DRAIN P SUPPLY CONTROL

\ \ \ _ BEARING

\\\\'

\

BALL BEARING

BALLSCREW

SERVO

/_ P REGULATING VALVE

FIGURE 3.3. SCHEMATIC OF HALF-AREA SERVO VALVE AND Ap REGULATING VALVE
METEREDPRESSURE

(T0-C(_NTROL)

L

ATMOSPHERIC _

SUMP. I|I

• MAIN PUMP I/I

SUPPLY _

FEATHERING II

SOLENOID II

SERVO

FEATHER VALVE GOVERNOR

i sTMp • [)RIVE GEAR

A MOS

MAIN PUMI

suPPLY ii PI.OVTALVE

PRESSURE

MAIN &

STAND BY STANDBY _ /SPEEDER SPRING PUMP REGULATING

VALV E __::::::=:_'"_ I_'_,_ _11

'SUPPLY _:::::::::__

PRESS. SPEED

SET

INKAGE

TRIM

SPRING

MAX

MECHANICAL.f RPM

FEATHER

INPUT

/_P

VALVE

DRAIN I_

_, PRESSURE

o

FIGURE 3.4. SERVOGOVERNORAND FEATHERVALVE

77
STANDBY FLOW

MAIN FLOW

TO ACTUATOR TO ACTUATOR

STANDBY MAiN

FILTER FILTER

FROM ACTUATOR

HIGH PRESS. AUX. lira-J[}

RELIEF VALVE STANDBY HX BYPASS_VALVE

PUMP

CHECK VALVE ,AIRFRAME

SUPPLIED

MAIN PUMP HEAT FROM GOVERNOR

.CO_ "CHECK EXCHANGER &,_EGULATING

Z8V r %, _[!AU..X SCAVENG..E PUMP VALVE

._ J FLOW VALVE

SWITCH

AUX.

AUX SCAVEN_SE ATMOSPHERIC STANDBY MAIN :R

PUMP PUMP PUMP

PUMP PUMP

MOTOR

)SPHERIC SUMP

| PRESSURIZED SUMP SUMP RELIEF

VALVE

FIGURE 3.5 PITCH CONTROL PUMPING SYSTEM
i  ̧ !

i

TO GOVERNOR & ACTUATOR T/

1-20 QT/MIN 120 PSI@60 QT/MIN

PS-1150 PSI MAX 51 PSI@40 QT/MIN

r

(PS-lss)

(PsTBY-PD - 65PSi)

HP RELIEF VALVE .48 QT/MIN

V

CHECK VALVE ORIFICE PACK j

O

TANDBY PUMP

O O 39°86 QT/MIN

100% RPM

PRESSURIZED

SUMP

I

MAIN PUMP I

ZZ.47 QT/MIN I

!00% RPM

FIGURE 3.6. MAIN & STANDBY REGULATING VALVE LOW DEMAND CONDITION

79
TO GOVERNOR & ACTUATOR

Z0-60 QT/MIN J__ _F

PS°I040PSI MAX

(PS-150)

(PSTBY-PS)

HP RELIEF VALVE o48 QT/M|N

3 _ -.--. ",& •

CHECK VALVE ORIFICE PACK _

L 0 •

STANDBY PUMP

0 o

39.86 QT/MiN

100% RPM

PRESSURIZED

SUMP

MAIN PUMP I

22.47 QT/MIN

@100% RPM

FIGURE 3,7.. MAIN & STANDBY REGULATING VALVE HIGH DEMAND CONDITION

8O
1 ! 3.7 IN 2

107.7 IN Z

1_ PITCHLOCK GAP BETA VALVE

= 0.028+0.002 IN. WINDOW WIDTH: !:88 IN. (60% OF CIRCUMFERENCE REF)

PITCHLOCKSCREW FLOW GAIN (UNDER NO LOAD): 5430 IN3/SEC/IN

oo LEAD = 0.5 IN./REV.

STROKE = 4.656 IN. (9.46 EFFECTIVE REV'N) BALL SCREW

LEAD = 0.2 IN/REV ....

MAX FEATHER BLADE & ACTUATOR RELATIONS STROKE:

9u_ P3/4 X = RSIN (_3/4-38"5)

TOTAL: 1 .9 IN. (9.5 REV'S)

.._r--"NOMINAL FEATHER C)X/_B = RCOS (_3/4 - 38.5)/57.825 EFFECTIVE: !.892 IN (9.46 REV'S) 87-s°_3/4

()X/()B = 0.05524 IN/DEG @ MID STROKE MID STROKE _ .4 SERVO PISTON

OX/OB = 0.0326 IN/DEG @FEA & REV

5°_3/4 ()X/()B = 0.0456 IN/DEG (AVERAGE) AREAS: AMETERE D = !0.2225 IN 2

_"'MAXREVERSE " = S.I_42,N2

I " -1o_3/4" ASUPPLV

x 0.01824 IN/DEG

_ _ _CTUATORSTROKE

BLADE ._._..:.-----'_" "

ROTATION

FIGURE 3.8. PITCH CHANGE ACTUATOR CONSTANTS
ELECTRONICS OFF (K E = O)

YWEIGHTS ELECTRONICS ON (K E = oo1

GOVERNOR RPM {NG) @ 100%SPEED: 4072RPM

DEC. PITCH

GOVERNOR-PROP SPEED RATIO (NG/Np): 2.3981

FLYWEIGHT FORCE EQ: FF w = 0.004029 (%NG) 2

@ 100% RPM: FFW = 40.29 LB

FLYWEIGHT RATE EQUATION: KFW = 0.002156 (% NG) 2

XG @ 100% RPM: KFW.......... = 21.56 LB/IN

SPEED SPRING RATE (Ks): 80 LB/IN

SPEED SET LEVER RATE (Kt): 1042 LB/IN

SPEED BIAS RATE (KBG):

PD

INC. PITCH KBG = (L1/L2) 2 (t/(I/Kl+I/K2))

LEVER TO GOVERNOR RATE (KI) 65.8 LS/IN

PMETERED" LEVER TO GROUND RATE (K2) 28.5 LB/IN

= 2.952 BRAKE LEVER WHEN K E = 0: KBG = 6.99 LB/IN

K E = : KBG = 23.12 LB/IN

PSUPPL = 1.75 IN. I -

SYSTEM SPRING RATE:

G0 KSYS = !/(1/(Ks+KFw+KBG)+ I/K! )

@ 100% RPM:

K E = 0 KSy S = 98.31 LB/IN

PD KE = oo KSYS = 1 _ 1.36 LB/IN

SENSITIVITY = _XG/_N% = (_FG/_N%) (_XG/_F G) = S

@ 100% RPM:

KE (AE)__. K E = 0 S = 0.0082 IN/% RPM KE = oo S = 0.0072 IN/% RPM

AXIAL ACCELERATION SPEED SHIFT, 0.3565 % RPM/G

BIAS SPRINGS

I

.SPEED SET SPRING

SPEED SET LEVE

FIGURE 3.9° SERVO GOVERNOR CONSTANTS
j

I _/ TOGOVERNOR

=

HYDRAULIC FoRcEs -I- 2.0#

PITCHLOCK SCREW DRAG BETA VALVE SEALS

5.0 IN# /,t = 0.3 FRICTION 1Z.0_

SERVO PISTON

QUILL SPLINE FRICTION \ SEALS FRICTION 70#

5.0=# /] = 0.3 \

_BALLSCREW THRUST BEARING DRAG

1.0 IN.#

BALLSCREW DRAG

2.0 IN# 0.85% EFF.

FIGURE 3.10. GOVERNOR LOOP FRICTION HYSTERESIS
THIS AREA MUST MEET

4 03 THE PRESSURE SPECTRUiM/_

F

REF

_THIS DIM SCALED FROM ENGINE_,,',,r.//

MOUNTED ON TEST RIG- 3_7/84J

PRESSURE SPECTRUM ."

0-1375 PS! 10000 CYCLES

!_40 +- 160 c_

2550 BURST

700 PROOF

FIGURE 4.1. MODIFICATION TO AND10152o60 PROPELLER SHAFT

84
9.864

CONTROL

FACE OF REAR - _ I- PLAN CONE & ENGINE . (P/N 789454) THRUST NUT

MICRO ADJUSTING RING {P/N 554880 I

INPUT SHAFT (P/N 75061Z1

SECTION A-A A

FIGURE 4.2. iNPUT LEVER INTERFACE FOR PITCH CHANGE CONTROL
SLEEVE

PLUNGE 

S EAT

SHIMS

COVER

SUPPLY DRAIN

FIGURE. 5.1 MAIN & STANDBY REGULATING VALVE

86
1,500

b

!,000

SUPPLY

PRESSURE:

(-WIN)

500

1o Isi zo s 30 3sI 40 4Is 50 sIs eo

NE_T FLOW TO_ACTUATOR QT/MIN

FIGURE 5.2. MAIN & STANDBY REGULATING VALVE SUPPLY PRESSURE

AS A FUNCTION OF NET FLOW TO ACTUATOR

8'7
B7

/PLUG

SUPP

R ESTR ICTO RS: /

/

SLEEVE'S

4

DRAIN

MATERIALS

RESTRICTORS: "

-SCREENS AISI 304

-REMAINDER AISI 308

DIMENSIONS

R ESTRICTOR SCREENS:

-HOLE SIZE: 0.007- 0.010 IN. DIA.

-OPEN AREA 38.7 %

FIGURE 5,3 ORIFICE PACK

88
DRIVE GEAR 7

SLEEVE INNER SPRING

FLYWEIGHT SEAT

b

SPOOL

. SCREEN ;PEEDER

SPRING

I I

DRAIN SUPPLY

METERED

DRAIN PRING

SEAT

B. GOVERNOR ASSEMBLY

p,To_-.-.-,_oP,TcM,_ __.__

DEC. 0.125 MIN

_-'-_0.0125 + 0.0015

. 0.0105/0.010

A. GOVERNOR METERING WINDOWS

ALL DIMENSIONS ARE GIVEN IN INCHES

FIGURE 5.4.GOVERNOR

89
GOVERNOR PRESSURE GAIN

11.56 X 108 PA/MMI FOR Qt = Q2 = 0

,/

I I I I I I I I "" / _'_ST |QI=QZ=0)'

NOTES: F

- _ _-TRANSFERBEARn"6LEAKAGE _--O.g I I

_ _,-(PsuPPLY-PMET .........ERED) / ___

-- 2) QZ- TRANSFER BEARING LEAKAGE_m0.8 _-_

(PMETERED-PDRAIN)

I I I 0.7 F _'_

_SYSTEM PRESSURE GAIN

J - 14.B9 X tO 6 PA/MM)

o.6 f

PSUPPLY-PMETERED / iNCLUDING EFFECTS

PSUPPLY-PDRAIN. f . OF TRANSFER BEARING m

LEAKAGE.

.i. O.4

J 0.3

11 j

//_ / o.z

0.1 /

-o.ozo -0.016 -0.012 -o.oos -0.004 0 0.004 0.008 0.012 0.016 o.ozo

SPOOL DISPLACEMENT (IN.)

PRESSURE GAIN CHARACTERISTICS

FIGURE 5.5 GOVERNOR PRESSURE GAIN TEST AND SYSTEM REQUIREMENTS
3.0

_ 2.0

" METEREVD __ 1,0 LOW _0

(IN3/SEc) -0.03 00Z o01 _ o.0_ o.1o2 0.03

i. TRANSi_'-RBEARiLEA NG KAGE" " 1O'__

_l-aa =0.46_IN3/SEC.

2. _ = BLADE VELOCITY@ 3/4 STATION

= 5.3632Q (DEG/SEC) - 2

N% = 12t.95X (% RPM)

(/_)/("_% N) = 0.04398 (AQ)/(AX)

(D EG/SIECi°/o RPM) I

X SPOOL

DISPLACEMENT

(IN.)

B. FLOW GAIN CHARACTERISTICS

HALF AREA SERVO PISTON

Q2

, Q

SUPPLY

DRAIN<

-X +X

MODEL

A. METERED FLOW PATH

FIGURE 5.6 GOVERNOR FLOW GAIN REQUIREMENTS

91
COYER

/ FEATHER METERED STANDBY 1 SPRING SEAT 

_ ....

I ..... B. vALvE CLOSED SPRING

,=....-

A. VALVE OPEN

FIGURESoT. FEATHERING VALVE

92
A

;GING GEAR

" INPUT GEAR

49 TEETH_ A

99% RPM "7

" ATHER //

I 'UT // LINKAGE /

/ /

/ /

LEVER FEATHER /

/

CAM /

FEA /

INPUT GEAR /

! 7 TEETH /

/

/

SPLINED ALPHA SHAFT /

/

/

/

/

_SPEED SETTING /

CONTROL INPUT JCAM ....- _/

INDEXINPLATE G _. GOVERNOR

F_ I IHOUSING

I INNER |0 f

ADJUSTING SCREW I BIAS YOKE ! CLICKPER L3 _OSETTING

1/4°/o RPM LEVER 84 _SPEEDER II

'L4 _--" ] _OUTER

I SPRING SEAT

GOVERNOR ROCKER

ARM VIEW A-A

TABLE OF LENGTHS AND RATIOS

L 1 = 2.2037 IN.

L 2 = 1.00 IN.

L 3 = 1.38 IN.

L 4 = 4.18 IN.

84 = 1.3746 81

GEAR RATIO = 2.8824 (o CAM/O INPUT)

FIGURE 5.8 GOVERNOR LINKAGE SCHEMATIC

93
RANGE

FLIGHT

IDLE

BETA RANGE

REVERSE FEATHER

66 ° 45'

INPUT

INDEXING PLATE

32 ° 50'

I.d

n,

,€

U

>

O

TOT,

CAM RISE =

0.0997 IN.

0 CAM ROTATION (DEG) 320.33

Id 97.76 269°79

n,, B. GOVERNOR CAM

<

U

ld _TOT°CAM RISE = 0o3317 IN.

'1- /

/ 

0 CAM ROTATION (DEG) 320.33

97.76 269.79

A. FEATHER CAM

FIGURE 5.9. E2/C2 GOVERNOR AND FEATHER CAM SCHEDULE

94
GOV.

f 75oi0

10/

17 ° 32' _ 17 ° 32'

FEATHER

REVERSE | 26 ° 26'

50 ° Z2'

320 50'

DBETWEEN EG OF CAM75%ROTATION AND 105%

RPM = 219.25 o

/

DEG OF CAM ROTATION

PER % RPM = 7.3083

IdU_ m

=E.

<

U

o

C.

TOT. CAM RISE =

0.1407 IN.

0 CAM ROTATION (DEG) 32:0.3:}

I,I U 50.54 B. 269.79 _

n,

:E

<

u

w

'1" F TOT. CAM RISE = 0.3317 IN.

<

w

0 CAM ROTATION (DE:G) 3;Z0.33

2:69.79

50.54 A.

o ACTUAL CAM PROFILE EXTENDS 10 °BEYOND REVERSE AND 10 ° BEYOND FEATHER (105% AND

75% RPM REF) AT A CONSTANT CENTER DISTANCE BETWEEN THE CAM AND ROLLER

FIGURE 5,10.LAP GOVERNOR AND FEATHER CAM SCHEDULE

95

o
STATIONARY

PUMP HSG

AMS 6415

ROTATING SLEEVE

HS44 36-40 RC REL

SUPPORT

RING

BARR]

6.4996 DIA. REF

BABBIT LANDC LANDB LANDA _///T 10_/

AMS 4800 OR SAE I 2

O ON

ROT. SLEEVE

DIAMETER CLEARANCE (IN)

LAND LAP E2/C2

A 0°00155-0.00175 0.0035-0.0040

B 0.00155-0.00175 0000155-0.00175

C 0.00155-0.00175 0.00165-0.00185

FIGURE 5o11. TRANSFER BEARING MODIFICATIONS

96
ANT

SEAL RETAINING RING

LIP SEAL

CONTROL OIL

RETAINING PLATE

ROTATING FLUID

TRANSFER SLEEVE

BONDED

FRONT LIP SEAL

PROPELLER PUMP LiP SEAL RETAINER

HOUSING _

:OTTER PIN 161

SE ANT

SEAL (SAME AS FRONT SEAL)

STATIONARY FLUID

TRANSFER

3ARREL SUPPORT RING

STIFFNESS INCREASED

3.59 TIMES THAT OF E2/C2 RING

HEAT EXCHANGE REAR LiP SEAL

HOUSING

F'|GURE 5.12. LIP SEALS

97
GOVERNOR

4072 RPM

,COMPOUND IDLER

I 3529 RPM

I

I

TRANSF|

BEARING

1698 RPM

/

I

I

MAiN

5294 RPM STANDBY PUMP

4500 RPM

'SCAVENGE PUMP

529,; RPM

FIGURE 5,13, GEAR TRAIN SCHEMATIC - VIEW LOOKING AFT

98
. ,EZ/CZ

TRANSFER BRG GEAR

PD = ! 2, 98T, 1106 RPM

MAIN PUMP GEAR

PD = 12, 23T

FLOW DIRECTION

LTRANSFER .AIP. BRG. GEAR

I PD -- 12, 106T, 1698 RPM

MAIN PUMF

'RAD. TO CTR°

OF PUMP = 5.4416

DIRECTION oF FLOW

MAIN PUMP GEAR

PD = 1Z,34T

RAD. TO CTR. OF PUMP = 5.4416

FIGURE 5.14. E2/C2 AND LAP MAIN PUMP GEARS, VIEW LOOKING AFT

99
TRANSFER BRG. GEAR

MATERIAL: AMS 6260

HARDNESS: 685 VH MIN (CASE), 30-42 RC (CORE)

RIVETS

MATERIAL: CORROSION RESISTANT STEEL

NO. REQUIRED: !2

CIRCUMFERENTIAL IN./RIVET: 2.01 ,TRANSFER BEARING GEAR*

0o0002- 0.0015 IN

INTF°

ROTATING FLUID

TRANSFER

STATIONARY FLUID

TRANSFER SLEEVE

*MODIFICATIONS TO THE PRESENT TRANSFER BRG° GEAR (P/N 750621) WERE

REQUIRED IN ORDER TO OBTAIN A LARGER DIAMETER DRIVE GEAR

FIGURE 5.15. MODIFICATIONS TO E2/C2 ROT° TRANSFER SLEEVE GEAR

I0O
BALL BEARINGS

EXISTING HARDWARE

USED ON 14RF/! 4SF WASHER

PROPELLERS SHIMMED TO OBTAIN

75 TO 82 LB LOAD

GEAR

MAT'L REMOVED FOR SHIM

WEIGHT REI b03-0.007 GAP

PUMP

ANTI-ROTATION

DOWEL PIN ....

I

i

I

THREADS FOR ,_ u,2,1 I.UJ I.U.I"_

SHAFT REMOVAL | ARING

_ _ _ LUBRICATION HOLES

.... d

Ir31 rrn,

suPPORT-= l _ "1-"r "11

SPACER

IDLER GEAR SUPPORT PLATE

GOV. DRIVE GEAR

bTATING SLEEVE

\

FIGURE 5.16. IDLER GEAR SUPPORT
SLEEVE DRAIN

GUIDE SPRING

SHIMS

MATERIALS

SLEEVE: .AMS. 5612, 30-38 HRC

GUIDE: 'AMS 5612, 30,.38 HRC

PLUNGER: AMS 5610, 30-38 HRC

SHIMS: QQ-A-Z50/4 COND T3

DIMENSIONS

SLEEVE & GUIDE DIA. CLR. 0.00095 +_0.00055 IN

GUIDE 8: PLUNGER DIA. CLR. 0.002:+0o001 IN

PERFORMANCE

SPRING RATE: 130.6 LB/IN

NORMAL OPERATING RANGE ! 150- 1875 PSi

FIGURE 5,17. HIGH PRESSURE RELIEF VALVE DATA

102
DOUBLE

REVERSE DELTA ""

BUL STOP SEAL _.NTI-TORQUE SPLINE

WEAR

I

PLATES

iNC

-- ..... II

DJ USTABLE FEATHER STOP WEAR STRi #

" "D '° SEAL RING

STATIONARY PISTON

FIGURE 6.1 PITCH CHANGE ACTUATOR
j-BULKHEAD

/MATERIAL: AMS6415, 36-40HRC :YLINDER

/

STROKE: 4.656 iN,

== --"--- LJ --

DECREASE

_i__ \,oREASE o ARE .\ \:.o13.7 .IN 2

i \,o,.,,o)' \

STATIONARY PISTON

MATERIAL: AMS6415,40-44 HRC

FIGURE 6.2. PITCH CHANGE ACTUATOR MATERIALS
DOUBLE DELTA SEAL

MATERIAL: TURCON/MOLY

EXTRUSION GAP: 0.030 IN. RADIAL MAX

\ sLY"

On "'A , I["J]_ ,_

BACKUP RING _" SEAL

MATERIAL: TEFLON MATERIAL: VITON WEAR STRIP

THICKNESS: 0.060 CROSS SECTION: 0.230 X 0.275 MATERIAL: TOR_ON

4301

DIAMETRALCL.: 0.010-0.025 SQUEEZE: 7% TO 18% THICKNESS: 0.100 WIDTH: 0.55

DIAMETRAL CL.: 0.002-0.010

ALL DIMENSIONS ARE GIVEN IN INCHES

I

FIGURE 6.3. PITCH CHANGE ACTUATOR CLEARANCE
BEARING

EEVE NUT

PITCHLOCK STC PISTON PISTON RING

PITCHLOCK-SCREW _

/I

RETAINING NUT

BALLSCREW

SPHERICAL WASHE

_,

PITCHLOCK RETAINER_

,: _HOOSING \RETAIN,NNOTG

FIGURE 6.4. HALF AREA SERVO
• PISTON-SLEEVE ,0005-.010 DC

PRE-LOADED

•PRE-LOADED BALLSCREW DUPLEX BEARING

PRE-LOAD 312: LBS 5.15 IN;

10.22 IN 2 1140 PSI

'570 PSI

PISTON RING-SLEEVE .006 MIN DC

RULON WEAR RING-SLEEVE .002-.003 DC

PISTON-SLEEVE .006 MIN DC

FIGURE 6.5. SERVO DOUBLE SEAL CONFIGURATION

w

107
VITON

AM$ 64_S

86 ° 40 HRC

_AI$1

INDUCTION HARDENED BALL GROOVES (56 HRC

CHROME

PLATE

RULON

PLATE

\

17-4 PH AM$6415 AAT07S-l"6

NITRIDE 36-40 HRC ANODIZE

FIGURE 6.6 . HALF AREA SERVO MATERIALS

108
,,SHIM /

ACTUATOR BULKHEAD _ /_

,uoo\ _ _o_, o _ooo,o

SUPPORT

LOCKING

BALLSCREW

_O

PITCH LOCK SCREW

LOW PITCH

BALLSCREW QUILL

ADJUSTING SLEEVE

FIGURE 6.7. PITCHLOCK SCREW
mSLEEVE

RETAINING NUT SEAL RING HOUSING

FIGURE 6,8. BETA VALVE

RETAINING NUT GLIDE SEAL HOUSING

MATERIALS

SPOOL: AMS 6415, 36"40 HRC CARBURIZE 

SLEEVE: AMS 6415, 36"40 HRC CARBURIZE

RETAINING NUT: AA6061-T6 ANODIZE

GLIDE SEAL: TEFLON (GRAPHITE IMPREGNATED)

HOUSING: AA7075-T6 ANODIZE

DIMENSIONS

SPOOL & SLEEVE, DIA. CLR ,0003".0006 IN.

SLEEVE & HOUSING, DIA CLR .002".008 INo

FIGURE 6.9. BETA VALVE MATERIALS

110
.99

.91 -

°05 PS TOLERANt i=' BAND

.50

0

n.,

b.I

n,'

U)

U)

id

n."

D.,

.09

.01

I I I I I I

- .004 -.003 - .002 - .001 0 o001 .002 .003 .004

SPOOL TRAVEL (IN.)

FIGURE 6.10 PRESSURE GAIN CHARACTERISTICS OF BETA VALVE

111
30

FLOW RATES

5.868 (IN/SEC)/.001 (IN) STROKE MAX.

20 4.998 |IN/SEC)/.001 {iN) STROKE MIN.

.0003 MAX UN_ERLAP

10

bj_ 0 _b

_ "

°0003 MAX UNDERLAP

-Z0

-30

I I ! I I !

- .004 -.003 -.-002 -.00 ! 0 ,00 _ .00Z .003 .004

SPOOL TRAVEL fiN.)

FIGURE 6.11 FLOW GAIN CHARACTERISTICS OF BETA VALVE
CASE: nNCREASE PRESSURE WITH ACTUATOR AGAINST FEATHER STOP

I

•i ,,UL

,, i_=...__r__j

FEATHER STOP

FIGURE 6.12 STATIONARY PISTON LOADS
IIIIIIIIIIIIIIIIIIIIII

FIGURE 6.13 STATIONARY PISTON INPUT'GEOMETRY

,_+O.OOZ8

CASE: INCREASE PRESSURE WITH ACTUATOR

AGAINST FEATHER STOP

ALL DIMENTIONS GIVEN IN iNCHES

+0.0008 +0,0008

_-0.0008

-0o0019

-0.0038

FIGURE 6.14 STATIONARY PISTON RADIAL DEFLECTIONS @ 1375 PS!

114
CASE: INCREASE PRESSURE WITH ACTUATOR

AGAINST FEATHER STOP

. S/P = 41.e s/P = 2.9.2

(MERIDIONAL) (MERIDIONAL)

K T = 4.0 KT = 2'5

®

S/P = -24 (HOOP)

KT = 2.0

FIGURE 6.15 STATIONARY PISTON STRESSES

PROOF

10Q_ BURNT

,o x,,,,,,

Io U 50 I

old

O n

<""1" 40 _®

_-2 ffJ

_, in _ _o

m :€®

_" '< 20 AMS 6415 40-44 RC .--

,_ _ BAR-MACHINED -170°F

b. (ADJUSTED FOR K F VS KT)

10

0

1.0 2.,0 3.0 4.0 5.0 6.0

KT

THEORETICAL STRESS CONCENTRATION FACTOR

FIGURE 6.16 STATIONARY PISTON S/P CURVE

115
DESIGN CASE #2

F = 71,500.#"

'L-J

0.03 IN, OR 0.56 °

FIGURE 6.17AXIAL DEFLECTIONS OF ACTUATOR - NORMAL OPERATING CONDITION

DEFLECTION

\

CASE LOAD DEFLo + 0.030 BLADE ANGLE CHANGE

110% O.S, 97,500# 0.082 1.58 °

100% VMA X 71,500# 0.068 to30 °

CRUISE 8{ CLIMB 66,000# 0°065 I,Z4 °

FIGURE 6.18 AXIAL DEFLECTIONS OF ACTUATOR PITCHLOCK CONDITION

116
300

672 (FT/SEC) TIP SPEED

250 / -

_ 200 D1 638 (FT/SEC) TIP SPEE

0 150

r_

r_

100

MACH CASE: 0.80 NUMBER; 57.5° P3/4

35,000 FT

50 SDSL

800 FT/SEC TiP SPEED

o I

0o 1.0 ° 2 .0° 3"0°

BLADE ANGLE LOSS

FIGURE '6.19 DRAG AND PROPELLER SPEED FOR PITCHLOCK CONDITION

BLADE SHANK

BLADE PLUG

_UNNION

TRUNNION BEARING

(ROLLER)

FIGURE 6-20. BLADE TRUNNION ASSEMBLY

117
ROLLER

BEARING

TRUNNION SHAFT

FLAT WASHER AND JOURNAL BEARING 

SPRING WASHER

PLATE

SHIMS

COTTER PINS

RETAINING PIN RETAINING SLEEVE

FIGURE 6.21 TRUNNION ROLLER AND WEAR PLATES

ROLLER

MATERIAL: AMS 0440 59-62 RC

OoD.: 2.000IN. _ THRUST BEARING

CROWN RADIUS: 20.00IN ,: VESPEL (TYPE SPZI)

SURFACE FINISH : 32/

WEAR PLATE JOURNAL BEARING

MATERIAL: AMS 6265 MATERIAL: VESPEL (TYPE SP21)

CASE: 59 RC MIN(.0801NTHICK} FIT W/TRUNNION SHAFT: .0012- DIA. CL.

CORE: 35-43 HRC FIT W/ROLLER: °0035°°0060 DIA. INT

THICKNESS: .200 IN THICKNESS: .090 IN

SURFACE FINISH: 16/ MATING SURFACE: AMS 6415 40-44 HRC ,,,,j

CHROME PLATE

FIGURE 6°22 TRUNNION ROLLER AND WEAR PLATE MATERIALS

118
ROLLER

HERTZ STRESS -

HCF: 370,000 + 31,500 PSI

LCF: I 85,000 + 185,000 PSI

YIELD CASE: 426,400 PSI

ULT. CASE: 459,500 PSI

BENDING STRESS -

HCF: 84,000 + 21,000 PSI SEARING

.LCF: 42,000 + 42,000 PSI BEARING STRESS: 11,000 PSi

YIELD CASE:I25,000 PSI

WEAR PLATE _ '_'_"

JOURNAL BEARING

SUBSURFACE SHEAR -

(AT MIN CASE DEPTH) BEARING STRESS -

HCF: 78,000 + 11,000 PSI HCF: 8,500 -+ 2,100 PSI

LCF: 37,500 + 37,500 PSI LCF: 4,250 + 4,250 PSI

YIELD CASE: 94,000 PSi YIELD CASE: 13,000 PSI

ULT. CASE: 105,000 PSi ULT. CASE: 16,000 PSI

FIGURE 6.23 TRUNNION ROLLER AND WEAR PLATE STRESSES

140 I I I I

,ALLOWAI3 LE: / LAP

AMS 6265 i 1 QCSEE

_. 120 59 RC MIN @ 0.040 MIN I

50 R C MIN @ 0.080 MIN /OCTEHEDRAL /SHEAR

_ _ STR ESS 100 _ .1.,-_.

. 80 ..... _

Id _ ALLOWABLE:

l__ _ AMS 6512

j 60 • • _._ 48 RC MIN

< t •

Ld

Z 40 ,_

Id _., _OC"r EHEDRAL

"_ _ SHEAR STR ESS

O 20 I" _ ,,,

0

0 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.09 0.1

DEPTH (IN.)

FIGURE 6.24 LAP VS° QCSEE STRESS @104,-,

119
SECTION A-A S/P = 67

SECTION B-B S/P = 57

SECTION C-C S/P = 39

SECTION D-D S/P = 30 ,-

(K T = 1.00)

ALLOWABLE S/P = 72 

D S/P = ALLOWABLE STRESS/IOOO PSI

FOR ASSUMED FLIGHT SPECTRUM

A

B

FIGURE 6,25 ACTUATOR EAR STRESS

-- HOUSI NG

• FILTER SCREEN

CAP

SPOOL

SEAT

SPRING

FIGURE 6-26° Ap REGULATING VALVE

120
MAX SERVO LOAD FRICTION LOAD AND SPRING

] 1400.# 15._ FRICTION MAX SPRING

< i_ 1 ! 00-# 5# FRICTION MAX SPRING

i_ 800# 0_ FRICTION MIN SPRING

713 __ ._

s9o -- __/f///////////////////j -._

48_ 439----L ------/-<...1-_-,-._----1--_:'...... _ .-""_

o /! i I\ \\ _"_ I / I i i \ _ '--'

I / " ' I I "PM = 574,5 _,,_ '_ / I I I P×='""

_ [ / I I ATNOLLR_F

_ I/ I I

v I I

I

,/I! , I

100 487 687

PMETERED FROM CONTROL (PSI)

FIGURE 6.27. I'_,APREGULATOR CHARACTERISTICS (REF: FIG. 3.3)

/-

121
SEAL .I._ SCREEN

GLIDE RING_ _ SPiNG.

SPOOL

MATERIALS

SLEEVE: AMS 6260, 35"43 HRC, CARBURIZE

SPOOL: AMS 6260, 35"43 HRC CARBURIZE

SCREEN: AMS 5513

SEAL: VITON

GLIDE RING: TEFLON {GLASS IMPREGNATED)

SPRING: AMS 5112

DIMENSIONS

SPOOL & SLEEVE, DIAo CLR (SMALL DIA.} .0004°.0006 IN.

SPOOL & SLEEVE, DIA. CLR (LARGE DIA.) .0005 ° 0007 IN.

SLEEVE & HOUSING .001".005 IN,

SPRING PRELOAD 5 LBS.

FIGURE 6.28 Ap REGULATING VALVE MATERIALS

122
40

30

¢1

,J

20

,J

e'l

i,°j ,

CTM: 22,Z31X COS21_3/4-41.27°1

o

-I0

o 10 20 30 40 50 60 70

BLADE ANG LEi_3/4

FIGURE 6-29. CENTRIFUGALTWISTING MOMENT (CTM)-IOO% SPEED

-4O

-30

ill

r_

<

.I

In -20

Zmi

m

.i

¢q

o -10

333 KEAS

P 256 KEAS _,,, *_ % .---.0 SHP

'_ 0 __100 KEAS _1, -- "%_ "_.......,."3000.6000SHP

_ --.. L , i_,,_ ..- ...

SHP

v

+1o

0 ! 0 20 30 40 50 60 70 80

BLADE ANGLE _3/4

FIGURE 6"30. AERODYNAMIC TWISTING MOMENT IATM) " 1OO% SPEED 15,OOO FT)

(+ TOWARD HIGH PITCH)

123
!00

U 8O

w

_ 7O

,, \

,,, /

o

I- 6O

<

IU

<

5o

4O

-20 -10 0 10 20 30 40 50 60 70 80 90

BLADEANGL__S/4

FIGURE 6-31. ACTUATOR EFFICIENI._Y (TOWARD HIGH PITCH)

124
5O

ACTUATOSRTALLOUTPUT @

" _ 40

•_z, j _:0 \

. 333 KEAS

/ I _.o s-P _'z / 2_6KEAS,_f\ I

o,o / I I_\,

_€ ooo SHP "

o OUTPUT @ 700 PSI _'o,-oA----_®-.J

_-_i 20 ./_I I- '?'I \\*®--•

a _ Q - DESIGCNASE4*-IOVERSPEED

_ / ® _ DESIGN_2CA"VSEM,_

(_ - TYPICAL CLIMB

(_ TYPICAL CRUISE

(_) - DESIGN CASE #3 -OVERSPEED

t0 I I I I

-20 _10 0 10 20 30 40 50 60 70 80 90

BLADE ANGLE _3/4

FIGURE 6-32. ACTUATOR CAPACITY

125
t 375 PSi HPRV 

1 150 PSi

!00°1 PRV

LOAD

PUMP PULSATIONS

6€

u FL!GHT_CYCLE ....

50 X PER FLIGHT

i

PS

STARTUP SHUTDOWN

TIME

FIGURE 6-33. ACTUATOR PRESSURE SPECTRUM FOR A TYPICAL FLIGHT

126
S-N DIAGRAM GOODMAN DIAGRAM

Id

I

" I

N ALLOWABLE I

CYCLES I

! I I I

10 7 1 0 6 ! 0 5 10 4 10 3 ULT.

NUMBER OF CYCLES STEADY STRESS

SS = PS X S/P

SC = PcXS/PXK F

PS = STEADY PRESSURE FOR LOADING CONDITION

PC = CYCLIC PRESSURE FOR LOADING CONDITION

FIGURE 6-34. CALCULATION OF ALLOWABLE CYCLES

127
PROOF,

100_ _ BURST ,

70

0 .5

0 _

0

" 50

tJ)

[" 40 

L"

"30

AMS 6415 40-44 RC

20 BARoMACHINED - 170°F ' °

(ADJUSTED FOR KF, VS KT)

10

0

! o0 2.0 3°0 4,0 5.0 6°0

KT

FIGURE 6.35 STRESS ALLOWABLES AS A FUNCTION OF STRESS CONCENTRATION

FACTOR (KT)

128
ClP

0

CIT 0

P/L

0

m "_D ELECTRONIC _"_"_"NGECF FUEL WF --v 501-M78 ENGINE DYNAMIC

CONTROL CONTROL NGOSS ENGINE

MODEL _, MODEL MODEL

t_ L i

_D

CDP

N1

TIT

N 2 PROP-FAN QPT

PROP-FAN /_3/4

SYNCHRO- CONTROL

N PHASER NTRIM MODEL

MODEL w

FIGURE 7.1. LAP SYSTEM FUNCTIONAL BLOCK DIAGRAM
_- ( J/DELTAE • THETAE ee .551) _ (DELTAE)

_ _ _ NG

_ _I •

/ ' NGC _ [NGC}

/_l F (P3CSS} _"'] _--(|/" VTHETAE)

(DELTAE}_

° ."I_P_ _ + °,,-,' °"°_00 Lr IQPTSSI ;"1 (THETAE ee .851

° _ --_ "'"°'"°' ! JNI/I oj_ ___ TTC S +_1 LT45TC

'

EXP {-.4648 • ALOG (AIRFLOW) +I.1571Z|

(DELTAE / VTHETAE'e .8811 _ I

5.0

FIGURE 7.2. DYNAMIC ENGINE MODEL OF ALLISON GAS TURBINE 501-M78
IF (WFCfoWFGV,KIFC=O.O J

NOTE: WFGOV GAIN CHANGE

=F (NGP, -INFO); EQUALS

UNITY PER NGP='13369.3 RPM

& WFO=2S00 PPH

WFGV

WFPROP+

KPFC

NGECFA

SS

NGECFA=(-SSe MA)jNGP+NGSET)

NGIOLE

'-----_-1 SS=0.7S6Xl 0 6-1,097.2 MAI----l....

}--.-ISS = 0.83X 106 +769.2 MA

I

I

I

I

I

I

I

I MA

&_-- - -----

,-------

I

I r------

I

I

I

I

I

NGSET I

PjL

PjL

NG (Nl)

NGOSS

NGECF NGECF }----~MA = ---

-27.0

i~ WFSCH WFD

CIT ( TS+l) (72 +1) (T3 S+I)

Nl

COP

~b PC

COP

FIGURE 7.3. 501-M78 GAS GENERATOR FUEL CONTROL
_GPO //_

NGECMX NGECMN SWEC

- K,o_ _ " IGp0S0=0. !

I o.o -44oo. j Hi I_

• OSTC _

TPTSET KIA I _ (IF NG ECC <-4400 J_.._,.

0 o.o3s_+ 1 "" --. . L__T._._I |_ ,

'_. |

AND

p._ _\INTC=0.0 / •

Co

bO _ 0.0

KP

NPTOST

,o.,_+,,,o.o,,_.,,+ 1 I

i YES _-1 NGOSS=NGRSET

FIGURE 7.4. 501-M78 ELECTRONIC CONTROL BLOCK DIAGRAM (N10VERSPEED

N20VERSPEED, AND TIT LIMITING FUNCTIONS}
I 0.028 0.028

0NPE RPMER R o o!, o ,

QMAX=.l.96 IN3/SEC ! )

¢.o GKQDE

QMIN='Q MAX

B= 50000 PSi

As=lo.8>lN2 XKVA /_3/_51-5°+SlN'_!--.

_Q IN3/SEC

0--_ "=0"002 PSI _HL

XKL=2.5 IN/IN

HL=0.003 IN

XKVA=40

XAMAX=l.87 IN

XAM1N _-2.7 IN IN/SEC

GKQINC=- GKQDEC----0.000192 RPM

FIGURE 7.5. "LAP" PROP-FAN CONTROL BLOCK DIAGRAM
VKlr QPT

I , I--7 Y+

(cP) (xp_op\Z ! _ I _ ' ---

(PROPHP) --_NPROP ; _ (PIE)('X-J TOT) "

XKJ = ! !.Z6

GR = 6.77

XJTOT = 89.64

XKHP= 118t.6

XKTH = 4334.2

CQSI = 0.5984

CQSZ = 0.4016

CRHO = (PAMB/14.7)+ * (TAMBi518.7)

FIGURE 7.6, "LAP" PROP-FAN MODEL BLOCK DIAGRAM
Id THETRF

_-_ NPROF _ _-_

D. s ! 0.00091 (18.S + 11

0 ESIGC (V }

n,

•r 0.071 S + I 0.0238 S + 1 S

U

Z NMASTER

P

DITHER(V)

o_.oo._ oo.._ H t-4 I

Z (0.019S S+ |l (0.05S+ I,

0 o.o01 S + I 1 s

I_ :E > ESIGB (V)

1-:_, o.ozzs s + I o.ozzs s + t o.ozzs s + 1

IL!P 0.617

J PLA Id 0.1S+

OUTPUT

OUTPU1 LIMITS ----4- 35.67 LIMITS=+!2.8

_n _ __

0 ESIGB e__ _ , S

:E

{I.00764 _.;6_

I

I .

NTRIM

(RPM)

FIGURE 7.7. 54H60 (C-130) SYNCHROPHASER BLOCK DIAGRAM
0.3760E + 02

0.3660E + 02.

,q.

m

0,3560E +02 ,

O.1040E . 03

11.n O._000E . 03

0.9600E + 02

0_7700E + 04

a.

w

_z

_ 0.6900E+04

w _

0.6100E "_

O.1020E + 03

,,,_e

z_=[ 0J00OE+03j X./

i,i

0.9800E + 02

0.0 0.2000E +01 0.4000E + 01 0.600( )E + 01

TIME _ (SEC)

FIGURE 7.8. SIMULATION RESPONSE TOA 1.5% NG REF STEP INCREASE

( S.L. STATIC, 100 % NP)

136
0.5550E + 02

w

g

_ 0.5500E +02

j 0,5450E + O_

O.I020E + 03

j

o_ o.looo+oE : _

a.

0o9800E + 02

0.2800E + 04

i,i

Z 0.2400E+04 --

0.2000E + 04

0.1800E + 0N

0.1600E + 04

Z

0.1400E +

0.0 0.4000E + 01 0.8( )00E + 01 0,12000E . 01

TIME -_ (SEC)

FIGURE 7.9. SIMULATION RESPONSE TO A 1.5% NG REF STEP INCREASE

(35,000 FT, 390 KTS, 100% NP)

137
0.4000E + 02

W

0.3000E + 02

0.2000E + 02

0° 1120E + 03

--I

.J_

o1O4OE.O3 IL

0.9600E . 02

0.8000E + 04

ul

0.4000E . 04:

I.d

0o0

0.1200E + 05

-- 0.8000E + 04 I _ '

° J

I0.4000E + 04

0.0 0,4000E + 01 0.8000E . 01 0.120

TIME ,_ (SEC)

t

FIGURE 7,10.SIMULATION RESPONSE TO A 2.0 SEC RAMP OF NG REF

(S.L., STATIC, 100% NP,88% -100% NG)

138
0.4000E + 02

_ Q

0.3000E + 02

STOP

0.2000E + 02

0.1040E + 03

w_a:_ ___. ,___._ 100%

o.9eoo02E. IL

IL

0_88_E + 02

0.8000E + 04

Z

0;4000E +04

Z

bJ

0.0

O.!200E + 05

0 8000E + 0,_ "

0.4000E + 04

0.0 0.4000E + 01 0.8000E + Ol 0.120

T_ME,_ (SEC)

FIGURE 7.11. SIMULATION RESPONSE TO A 2.0 SEC RAMP OF NG REF

(S.LoSTATIC, 100% NP, 100%-88% NG)

139
0.4000E + 02

0.3000E + 02

_NPT LIMIT= 120°/6

I

i .... NPT LIMIT= 109 % 

0.2000E + 02

0. I300E + 03 

_jlZ3.5 0/°

_ 0.1100E + 03

o \,,,_

n " _,,"%,_.

-- 100%

0.9000E + 02

0. IO00E + 05

.1-

i,d

Z 0.6000E + 04

Ill

I

0.2000E + 04

0.1200E + 0_

Z •

I- 0.8000E -_04 _ _ "'---

'J"

0.4000E + 04

0.0 0.4000E + 0! 0.8000E + 01 0.120

"TIME _ (SEC)

FIGURE 7.12. SIMULATION RESPONSE TO A 0.5 SEC RAMP OF NG REF

(S.L. STATIC, 100%NP, 88-100% NG)

140
0.5900E + 02

_ 0.5400E "_O_

0.4900E + O;

0.1040E + 03

0.1000+E03 'f'_. _

hlrIP't'Y .

.J

..i

w

0 0.9600E + 02

IL

O.4000E + 04

'I"

2:

0.2000E + 04 f

ILl

0.0

0.2200E + 04

0.1200E . 04

,, P'

0.2000E + 03

- 0.0 0.4000E + 01 0.8000E + 01 0.12U

TIME _(SEC)

FIGURE 7.13. SlMULATON RESPONSE TO A 2.0 SEC RAMP OF NG REF

(35,000 FT, 390 KTS, 100% NP, 78%-100% NG)

141.
0,5600E + 02

0.5200E + 02

0.4800E . 0_

0,1040E + 03!

g

o.

Wn"0.1000E + 03 ' _F -''_._

..I

..I

Id

n

IL

0,9600E + 02

0.4000E + 04

Z

0.2000E + 04

Ill

0.0

0.1800E + 04

0. !000E + 04

n,.

I,,-

0.2000E . 03

0.0 0,4000E + 01 0,8000E + 01 0.120

TIME _ (SEC)

FIGURE 7.14. SIMULATION RESPONSE TO A 2.0 SEC RAMP OF NG REF

(35.000 FT, 390 KTS, 100% NP, 100%-78% NG)

142
0.7500E - 01

(n 0.6500E - 01

X

0.5500E - 01

0o1800E + 00

_ 0.1600E + 00

×

0.1400E + 00

0.4000E + 02

g

"' 0.0

-1-

-0.4000E + 02

0,5000E + 01

_-_ o.o 1

-0.5000E + 01

0.0 0.2000E + 01 0.4000E + 01 0.600

- TIME ,-_.(,SEC)

FIGURE 7.15. 54460 SYNCHROPHASER LIMIT CYCLE IN/8 × GAIN

(35,000FT, 390KTS, 100% NP, 100% NG)

143
0.6800E -- 01

_" 0.6400E- 01 _ _

0.6000E - 01

0.1640E + 00

) i

;"o._6OO+ooE \ \--

u

" l

X

)

0.1560E + 00

0.5000E + 01

"

i.-' b.o

,,e

i-,

0.5000E + 0t

0o1000E + 02

_ 0.0

= I

I-

"

0. ! 000E + 02

\LL .\\ LLL

0.0 0.I 600E + 01 0.3200E + 01 0.480

• TIME ,-., (SEC)

FIGURE 7.16 54460 SYNCROPHASER PERFORMANCE 2.0 HZ, + 2.5 V DITHER

(35,000 FT, 390 KTS, 100%NP, 100% NG)

'144
0.6600E -- 01

• 0.6200E --01

O.! 596 E + O0

o.o

0.2000E + 02

0.0 0.1600E + 01 0.3200E + 01 0.480

TIME _ (SEC)

FIGURE 7.17. 54460 SYNCROPHASER PERFORMANCE WITH 2.0 HZ,+3.5 V DITHER

(35,000 FT, 390 KTS, IO0%NP, 100% NG)

/

145
DC MOTOR

AND GEARHEAD

r---

i

1-

SQUARE DRIVE FITTING

CONNECTING RO'

FIGURE 8.1 MOTOR DRIVEN PITCH CONTROL MECHANISM

".
3.0

Q

°oo/oooi

0 I | ! I I I I I I

0 4 8 12 16 20 24 28 32 36 40

FRICTION TORQUE (IN LBS)

FIGURE, 8.2 BLADE ANGLE SLEWING CAPACITY IN BETA MODE

VS PITCHLOCK SCREW FRICTION

147
REVERSE (-6 DEG @_) "- .

I

%

ACTUATOR MIDSTROKE

\ / FEATHER (90 DEG @ ,_s/4)

/ \

FIGURE 9.1 ACTUATOR/BEARING RELATIONSHIP

148
CTM: 22,130 IN-LB X !.2 (ALLOWANCE)

ATM: 3,052 IN-LB X 1.2 (ALLOWANCE)

FTM: 1,256 IN-LB

TTM: 31,474 IN-LB

TRUNNION LOAD @3.00 IN PIN RADIUS:

*CYCLIC STRESS OF ATM ONLY

31,474

= 10,500 LB + 1,220 LB

3.00

HCF = 10,500 + 1,220 LB VIBRATORY

LCF = 5,250 +5,250 LBSTART-STOP

FIGURE 9.2 TRUNNION LOADS@ 38.S DEG/33/41100%, SEA LEVEL)

REF. "SYSTEM DESIGN AND INTEGRATION OF THE LARGE-SCALE

ADVANCED PROP'FAN" SECTION 3o3ol. PG. 12.

LOCKING CUP

C

B

D

A

, 'TRUNNION

OWEL PIN

INSERT

e

FIGURE 9.3 BLADE SHANK BOTTOM VIEW

149
lOO

AMS 6415 - 40°44 RC, 250°F -- BAR-MACHINED

' 108 CYCLES

80

. 6o

,J

U

a . 4o

w

N 

SECTION C-C TRUNNION PIN RADIUS

< SECTION D-D I 0 0 20

0 SECTION B-B

z O 0

SECTION AoA

0

0 0 20 30 40 50

NORMALIZED MEAN STRESS - %

AiHi-Gi_ CYC LK FATIGUE:

AMS 6415

40-44RC, 250 F

BAR-MACHINED

5XI04 CYCLES 50 ....

4O

U ,

"; _o

>" SECTION C-C U

Q

Id TRUNNION PIN RADIUS

20

.I

<

SECTION B-B 0

SECTION D-D

2: 10

F' :

SECTION A-A

0

0 10 20 30 40

NORMALIZED MEAN STRESS- % o

B) LOW CYCLE FATIGUE

FIGURE 9.4 GOODMAN DIAGRAMS FOR TRUNNION SECTIONS

150
BOLT LOADS

BOLT TORQUE = 600 - 660 IN°LB

" BOLT PRELOAD = 7700 -I 1750 LB

APPLIED BOLT LOAD = 479'0 + 556 LB

BOLT STRESS

<.

E _ APPLIED CYCLIC LOAD * N F

PRELOAD MEAN APPLIED LOAD * NF + = +

SSHANK AREA AREA AREA

KBOLT

WHERE N F = JOINT EFFICIENCY =

KBOLT + K JOIN T

SSHAN K = 72668 + 412 PSI MS= 2084%

APPLIED CYCLIC LOAD *NF*K F

SFILLET = MEAN SHANK STRESS +--

AREA

SFILLET = 72668 + 824 PSI MS = 992%

FIGURE 9.5 SUMMARY OF BOLT LOADS AND STRESSES

60

4o

bl

MEAN PRELOAD + THRUST

.-.--.---

o

CYCLIC APPLIED LOAD _-

o I I I I

1 2 3 4 s 6 7 8 9 10 11 12 13 14 is

* THREAD NUMBER

TOTAL BOLT LOAD = 11750 + 4790 + 556 LB

JOINT EFFICIENCY = 0.126, UNEQUAL HOLE SPACING = 1.00

TOTAL MEAN STRESS = MEMBER STRESS + A * THREAD STRESS

, TOTAL CYCLIC STRESS = KFM * MEMBER STRESS + A * KFT * THREAD STRESS

KFM = 2.45, KFT = 2.05

FIGURE 9.6 TAPPED HOLE LOAD DISTRIBUTION

151
100 _ 

80 _'oe

I1:

_ 60 "

.1

U YTS

U 40

°

w

N

<

:E 20 15

0 10 20 30 40 50 60 o 70 '80 90 ! 00

NORMALIZED MEAN STRESS-,%

FIGURE 9.7 TAPPED HOLE STRESSES

152
Z

APPENDIXA

DESIGNREQUIREMENTS

153/154

DESIGN REQUIREMENTS

FOR

SR-7LPROPELLER

267X-I

May 6, 1983

Revised 12-1-83

Revised 4-I 1-84

Preparedby:..... ..

. _ro_by:_ _.,"_

' FZ_ZZ_ _

/j ' ,

155/156
II
LIST OF FIGURES

I. Velocity vs. Altitude

2. Power vs. Altitude

3. Temperature vs. Altitude

q

4. Blade Design Cases

5. Blade Analys_s Cases

6. Spinner Contour

7. CTMvs. Blade Angle

8. ATMVs. Blade Angle

9. Input Lever Schedule

I0. T-56 Installation Sketch

157/158
,O
1.0 INTRODUCTION

In recent years, considerable attention has been directed toward reducing

aircraft fuel consumption. Studies have shown that the inherent efficiency "

o advantage that turboprop propulsion systems have demonstrated at 0.65 Mn may

now be extended to the higher cruise speeds of today's turbofan powered

aircraft. To achieve this goal, propeller designs require advancements such

• as thin high-speed airfoils and aerodynamic sweep.

A series of sma11-scale (24.5 inch diameter) model tests have been conducted

- in both UTRCand NASAwind tunnels and on a modified NASAairplane. These

tests have shown that propellers with 8-10 swept blades, high tip speeds and.

high power 1oadings can offer increased fuel efficiencies at speeds up to 0.8

Mn." The next logical step is to test this advanced propeller concept

(Prop-Fan) in a larger scale.

NASAContract NAS3-22394 (Propeller Blade Structure Design Study) covers the

detail design (thru layout) of a large-scale blade (SR-7L) and a preliminary

design of a Prop-Fan suitable for flight test in a future program. NASA

Contract NAS3-23051 (Large-Scale Advanced Prop-Fan Program) covers the

detailing of the large-scale blade, completion of the design of the Prop-Fan,

fabrication of hardware and testing of the isolated assembly at Wright Field

and in the Modane wind tunnel in France. The hardware will then be used in a

follow-on program where it wiii be run with an engine on a static test stand,

in a low speed wind tunnel and on a research aircraft.

This document sets forth the requirements for the design of the SR-7L

Prop-Fan.

I.I Scope

This specification defines the requirements for the testbed Prop-Fan system

(SR-7L) for the Large-Scale Advanced Prop-Fan program (LAP)°

1.2 Classification

The testbed Prop-Fan shall consist of blades, disc/retention, actuator,

trunnion, constant speed control, and spinner. The assembly shall mount on a

• 60A spline shaft. It is intended for testing at Nright Field and in a high

speed wind tunnel and on a testbed airplane with an engine.

1.3 Features

• Constant speed governing

• Mechanical in-place pitchlock (14SF type)

159
• Feather via mechanical input signal

• Feathervia electricalinput signal

• Ground adjustablelow pitch stop

• Fixed angle reversecapability

• Electricalaux feather/unfeathermotor

° Instrumentationslipring

• Single lever input

• Individual replacement of blades

° Aerodynamicspinner

2.0 APPLICABLEDOCUMENTS

2.1 HS Proposal8lAl2

2°2 LAP ProgramPlan

2.3 LAP Work Plan

2°4 MIL-STD-810C

2.5 MIL-H-5606

2.6 L-14325-I and -2 Prop-Fan concept

2.7 Design Requirements for Advanced Turboprop Blades (SR-7) dated February

1983

2°8 MIL-P-26366(vibrationenvironment)

2°9 FAR-25 (deicing)

2.10 FARAdvisory Circular 33-1B dated 4/22/70 (FOD)

160
3.0 PROPELLERDESCRIPTION

3.1 General - This section defines the general requirements for the SR-7L

Prop-Fan.

3.2 Description

Type" Tractor

No. of blades: 8

Blade: SR-7-21 (aero version I00)

Diameter: 9.00 ft (108 inches)

Direction of rotation Conterclockwise (left-hand blade)

Note: Unless otherwise specified, all blade angles are specified at

the 3/4 station (40.5 inch radius), per aerodynamic conic.

4.0 Velocity vs. altitude: See Figure I.

4.2 Power" See Figure 2.

4.3 Tip Speed & RPM:

The normal max tip speed is 800 fps (1698 RPM) and is defined as 100% RPM.

Provision shall be made to test up to 105%or 840 fps (1783 RPM). The min

RPMto be tested is 600 fps (1273 RPM)o As a design goal this range shall be

provided without hardware changes.

4.4 Operating Condition

The Prop-Fan shall be designed to operate satisfactorily within the

temperature versus altitude limits defined in Figure 3. The system shall be

capable of rotating after soaking for a period of three hours in ambient air

temperatures of -65°F to +I30°F. Operation is permitted as soon as engine

oii temperature reaches normal engine operating lim|tSo Maximum

161
4°4 (cont'd)

temperatureof oil providedshallbe 170°Fo

4°5 EnvironmentConditions-- The Prop-Fanshallbe designedto satisfythe

requirementsof MIL-STD-810Cforhumidity,fungus,salt,spray,sandanddust°

4.6 OperatingFluid: MIL-H-5606

4.7 Oil Management-- The pitchchangeactuatorand controlshallhavea selfcontainedoil system.

The hub shallhavea separateoil supplyto lubricatethe retentionbearing

races_the actuatoranti-torquesplineand thetrunnionbearings.

The capacityof theoil tankin the controlis 18 quarts.

The actuatorrequires 12 quartswhileoperating.When thepropellershuts

down (atfeather), 6 quartswilldrainbackto thecontrol.

4.8 Deicing: Deicingheatersshal!be incorporatedin theblades.Thereshallbe no

provisionsfordeicingringor brushblockand connectionsat thistime.

4o9 Max Loadin9:

a) OverspeedLimit-- All elementsof the rotatingpropellerwillbe designed

to withstand125%overspeedor 150%centrifugalloadwithno inelastic

deformation°

All elementsof the rotatingpropellerwillbe designedto withstand140%

overspeedor 200%centrifugalload. Thisincludesthe blade,retention,

disc,and bladeanglecontrolmechanism.Localinelasticdeformationwill

be permittedin all of theseelementsat thisoverspeedbut thepropeller

shallbe capableof featheringafterexposureto 140%overspeed,butmay not

be operational.

162
4.9 b) ProofPressure:1.5 timesnormalpressure (1.5x 1130= 1700psi)

c) BurstPressure:2.25timesnormalpressure(2°25x 1130= 2550psi)

4.10Weight-- The bladesandbladeretentionshallbe designedto haveweight

characteristicsthatare representativeof anticipatedProp-Fansystemsfor future

aircraftapplications.

, For the restof the systemthereis no weighttarget°Thisdesignshallbe

start-of-the-artandaimedat lowdevelopmentrisk.

5.0 SPECIFICREQUIR_

5.1 Blade

5.1.i Configuration-- Basedon SR-7Lversion21 (aeroversion100)°

(ReferencePDRconducted2/25/83.)

5.1.2 Deicin9 -- Heaterincorporatedin inboardleadingedge (ref.FAR-25).

The heaterwillnot be wiredforuse in theProp-Fanbutmay be subjected

to icingtestsbyNASA.

5.1.3 FOD Protection-- Metalsheathon outboardleadingedge°

5.1.4 FlutterMargin

5.1.4.1 Stallm Freeof flutterat 100%of take-offpowerat 100%

designspeed (800fps)at Mn = 0 to 0.2fcr fcrwardthrustand

at 20%of takeoffpowerat 100%designspeedatMn = 0 to 0.2

forreversethrust.

5.I.4.2 Hi-s_eed-- Freeof flutterovertheflightenvelope(reference

Figure 1) and range of power loadings up to 105% of deaign rotational

speed (840 fps) and at 14,000 ft. altitude, the calculated flutter

Machnumbermust be greaterthan0.8at a testrighcrsepow__r

and RPMequivalentto thedesignpowercoefficientandadvance

ratioto allowtestingin theModanewindtunnel.

163
5.1.5 CriticalSpeedMargins-- No IP criticalspeedsshallbe permittedin the

operatingspeedrangeand theminimummarginshallbe 40%of maximum

operatingspeed. For 2P excitation,thegroundoperationcriticalspeedmargin at 800 fpsshallbe a minimumof 20%of Prop-Fanspeedand 2P integercrossover. The flightmarginshallbe a minimumof 10%Prop-Fanspeedand 2P integer

crossover.Thismarginshallbe reducedinverselyas theexcitingorderis

increasedfrom3P up to 5P. In determiningthesemargins,the effectof

bladeangleon frequenciesshallbe determined.

5.1.6 StressMargins-- Thecombinedsteadyand cyclicstressesshallbe plotted

on _DdifiedGoodmanDiagramsfor thematerialsof construction.The strength

boundariesshallrepresenta highprobabilityof survivalderivedfrom

experimentaldataon sp_;imensand full-scalestructures.As a minimum,

theboundariesshallrepresentX- 3.5 o lines. The start-stopstress

rangeshallbe reflectedagainsta boundaryfora lifeof 50XI0_ cycles.

The highcycie'combinedstressesshallbe reflectedagainsta boundary

for 100Xl06cyclesor infinitelife.

The m_ximumelastic(nominalx kT) stressingdue to 125%overspeedand the

nominalstressingdue to a 140%overspeedshallbe keptbelowthe0.2%offset

yieldstrengthforhomogenousmetalmaterials. The changein elasticmodule

shallbe keptbelow5% for fiberreinforcedresinmaterialregardingthese

sameoverspeedrequirements.

5.1.7 AerodlTnamicExcitations-- The equivalentdesignExcitationFactor(EF)shall

be 4.5. The basicEF due to IP only is 3.3.

5.1.8 FODCriteria-- The foreignobjectsare classifiedintothreecategoriesas

follows:minor,moderate,andmajorimpacts.Majorandmoderateimpacts

correlatewithGroupI and IIdefinitionsin FAR AdvisoryCircular33-IB

164
5.1.8 (cont'd)

. datedApril22,1970. Minorimpactsincludesand,smallstones,and birds

up to about4 ounces. Moderateimpactsincludet_-inch"hailstonesand

birdsup to two_mounds.Majorimpactsincludea singlebirdup to fourpounds.

The damagecriteriaare as follows:

MinorImpacts - No damageallowedto basicbladestructure°

Operationwillcontinuewithoutimpediment.

ModerateImpact - Damagecan includelossof materialor airfoil

distortion.Operationshallcontinueat 75%power

minimumfor 5 minutes.No metalfragmentshallbe lost

whichcan penetratetheaircraftfuselagepressureshell.

Roughnessshallbe tolel-ableand as a guide,rotorunbalanc

forceshallbe keptbelow5,000pounds.

MajorImpacts - Damagecan includelossof materialor airfoildistortion.

Abilityto featherthepropellermustbe maintained.A

Shutdownmustbe accomplishedwithoutcatastrophiceffects

on theairframestructure.As a guide,the rotorunbal_%ce

forceshallbe keptbelow25,000Ibs. No metalfragments

shallbe lostwhichcan penetratetheaircraftfuselage

pressureshell.

5.1.9 LifeandReliabilitGoals y -- The bladeshallbe designedforthe following

goals:

MaximumContinuousStressLevel infinitelife

ReplacementLife 35,000hours

MeanTimeBetweenUnscheduledBladeRemovals 50,000hours

(8bladeset) (Inherent)

5.1.10 LightningProtection-- Lightningprotectionshallbe incorporatedin the blade

5.1.11 DesiqnCases-- See Figure4

5.1.12 AnalysisCases-- SeeFigure5

165
5.2 Disc/Retention

5.2.1 Mounting-- Shafttype,sameas 54460propeller.

5°2.2 SlipRing-- Provisionformountingan instrumentationslipring (providedby

the instrumentationsGroup)on _ia_backof U_e-_"-_

5.2°3 Spinner-- Provisionforattachinga spinner(referenceFigure6).

5.2.4 Controlm Provisionformountinga controlon the tailshaft(reference

54460).

5.2.5 StressMar_in-- The_xim_t elastic(nominalX stressconcentratiofact n or)

stressingdue to a 125%overspeedand thenominalstressingdue to a 140%

overspeedshal!be keptbelowthe0.2%yieldstrengthfor homogeneousmetal

material.

5°2°6 Life-- The discshallhavea lifeof 50,000start/stopcycles. The retention

bearingshallhavea lifeof 50,000cyclesof lowcyclefatigueandunlimited

lifeundermaximumcentrifugal,maximumsteadybendingand maximumvibratory

bendinganywherein the flightenvelope.

5.2.7 RetentionBearing-- The retentionbearingshallbe capableof replacz_.ent

or the discshallincludeprovisionsforup to three(3)regrindswitha

maximumtotalstockremovalof 0.04.

5.2.8 SpeedPickup-- Provisionsfor the rotatingportionof a speedpickupshal!be

made. Thissignalshallalsobe conpatiblewithsynchrophasing.Thedevice

shallbe the sameor similarto 54460hardware.

5°2.9 IP ShaftMoment-- ThemaximumIP shaftmomentispredictedto be 76,500in.lbs.

at the followingcondition:

1698rpm

6000SHP

Md = 0°2 (132KEASSoL.S.D°)

EF = 3.3

5.3 PitchChangeActuator/Trunnions

5o3ol GeneralRequirements-- Anymalfunctionof thepitchchangeactuatorwillcause

the systemto pitchlockor feather.The actuatorshallbe modularto theextent

thatis practicalfor the testprogram_

166
5.3.2 In-placePitchlock-- An in-placepitchlockshallbe providedto limitthe

max lossof bladeangleto approximately1 degreebelowthegovernor

scheduledbladeangle. The minimumpitchlockgap shallallowthepitch

changesyst_l_toslew at a rateof approximately9 degreesisec.

5.3.3 GroundAdjustableLow PitchStop -- A groundadjustablelowpitchstop

is providedforreversebladeangle groundtesting.The stopshallbe

set to limitin-flightbladeanglefromgoingbelow+4008. The stop

shallbe adjustablebetween-I0°8and+40°8forreversetesting.The low

pitchstopshallbe readilyadjustablewithoutdismantlingthepitchchange

systemo

5.3.4 Feather-- Thepitchchangesystemshallbe capableof featheringthe

propellerto an angleof 87.5degrees+ 2.5° withan accuracyof 0.1°.

5.3.5 Reverse-- The actuatorshallprovidethetravelnecessaryto givea reverse

bladeangleof -10° setby the adjustablelowpitchstop. The actualminim_n,

bladeanglewillbe establishedby blade-to-bladeinterference.Thereis no

betacontrolto thereverseangle°

5.3.6 ActuatorSizing

5.3.6.1AerodynamicSizingCriteria

CTM (Centrifugaltwistingmoment): SeeFig. 7

ATM (Aerodynamictwistingmoment): SeeFig. 8

FTM (Frictionaltwistingmoment): CL x __PDx .005 2

• CL (CentrifugalLoad): 82,900ibs.at 1698rpm

TIM (totaltwistingmoment): C_4 + ATM + FI%_

A contingencyallowanceof-20%shail-_--add_-_--_eATM andC]!M

. of Figs.7 and 8.

5.3.6.2 The pitchchangesystemshallbe capableof increasingbladeangle

witha i0psi marginover.theoperatingrangeof Figure1 at 110%rl_r

(1868rpm)and 0 SHP.

167
5.3.6°3The pitchchangesystemshallbe capableof changingbladeangle

at 9 deg/secovertheoperatingrangeof Figure1 at 100%rpm

(1698rpm)at 80% (4800)SHP.

5.3.7 Lif__.._e -- Theminimumdesignfatiguelifeshallbe i0,000start/stopcycles

witha flightcycletimeof 1.5hours.

5.3.8 Instrumentation-- Provisionto measurebothhighandlow pitchpressure

duringoperationshallbe provided. (Pressuresensorsto be providedby

the InstrumentationGroup.) Provisiontomeasurebladeangleshallbe

provided.

5.3.9 TimeConstant-- The actuatorlooptimeconstantshallbe 0°25 sec at no load.

5.4 Control

5.4.1 Confi@uration-- Basedon 739000-1(54460/E2).

5.4.2 ZontrolSchematic

a) Present54460Control(SK90621_

b) Modified schematic for LAP (L-14325-3_

5.4.3 Governing Range -- 1273 to 1783 rpm (600 to 840 fps) (75 to 105%

5.4.4 Operating Pressure

Supply - 1050 to 1130 psig

HPRV setting - 1150-1375 psi

Pump Pulsation - 1150 + 175 psi

5.4.5 PumpFlow

62.2qts/minat 100%rpm,1125psi,170°F

46.6 qts/minat 75% rpm,1125psi,170°F

5.4.6 Leaka@e

5.4.6oi Control- 1o5 qts/minat 100%rpm_170°F

5o4.6.2TransferBearing- 3.7 qts/minat 100%rpm, 170°F

5.4.6.3pitchChange- Null 13 qts/minat 100%rpm, 170°F

- - slewing I.1 qts/minat 100%rpm, 170°F

5.4.7 Net Flow- 55.9qts/minslewingat 100%rpm, 170°F

168
5.4.8 CoolingRequirements- Coolingforthecontrolshallbe providedby an

airframesuppliedheatexchanger.Theheatexchangermustremove550btu/_

fromthe controlhydraulicoil. Circulationof thecontroloil throughthe

heatexchangershal!be providedby an airframesuppliedpump. Fittingsfor

hydrauliclinesto andfromtheheatexchangershallbe locatedat therear

of the controlper L-14325-20,

_ Oil outlett_p of heatexchanger: 170°Fmax

Pumpoutletpressure: 30 psi max (@ 1700F)

5.4°9 InputLever- Thecontrolshallhavea singleleverinputfor featherand

speedset. The shaftcan be drivenby eithera mechanicalinputor electric

motor. The driveis aircraftsupplied.The inputlevertorqueis 65 in.lbs.

maXo

5°4.9°I Schedule--Figure9 showstheconditionleverschedule.

............................ 574_.I0FeatherElectrical-- A solenoidactuat_ by theengineoverspeedgovernor

and/ora featherbuttonshallbe provided.

Mechanical-- The inputlevershallbe themechanicalinputfor

feather.

5.4.11Overspe_ protection-- An aircraft-suppliedoverspeedsystemshallactivate

the feathersolenoidwheneverthepropellerspeedexceeds110%of 800FPS.

5.4o12Betacontrol-- Thereareno provisionsforBetacontrol.

5.4.13Unfeather-- The auxiliarymotorshallbe usedto unfeathertheProp-Fan.

_ 5.4.14Reverse-- Provisionto run in reversewiththeadjustablelowpitchstop

set to thedesiredreverseangleshallbe provided.

5.4.15Synchrophasing--Provisionsto incorporatesynchrophasingshallbe provided

withintheconstraintsof easeof assemblyand lowdevelopmentriskthe

hysteresisand backlashbetweenthegovernorand theactuatorshallbe

minimized.

5.4.16DynamicResponse--Governorgain. 2o4O/sec/%RPM at null.

5.4.17Aux Motor& Pump-- An electricalmotor-drivenpump is providedto assistin

feathering,airstart,and forgroundhandling.

169
5.4.17 (cont'd)

o Pressure: 1375PSI

o Flow: 14qts/min

o Voltage: 208VAC,3 phase,400cycles

o Amperage: 15 normal,30 at reliefvalve

o" Dutycycle: 5°5HP fcr 10 sec

4°5 HP for 10sec

15min off

.........5,4.18 Lif_____e -- %_he-minimumdeSi_ fati_e ii-fe-shail_--lO, 000 start _sb0p cycies .....

witha flightcycletimeof 1.5hours.

5.4.19Instrumentation-- Provisionshal!be made to measuresupply,governor

meteredpressure,andstationarytransferbearingand sumptemperature.

(Thesensorswillbe providedby theInstrumentationGroup.)

5.4.20BrushBlock-- Provisionshallbe made to mountan instr_-m_ntation brush

block (providedby theInstrumentationGroup)on theexistingbrushblock

mountingpad.

5o4o21SpeedPickup-- Provisionforthe stationaryportionof a speedpickup

shallbe made. Thissignalshallalsobe compatiblewithsynchrophasing.

The deviceshallbe the sameor similarto 54460hardware.

5o5 Spinner

5o5.i Contour- PerFigure6.

5o5.2 Material5o5.2.1 Forward- Fiberglass

5.5.2o2 Aft- Fiberglass

5.5.2.3Bulkheads- Fiberglass

5°5o3 Deicing- Not required

5.5.4 Length- 44°5in.

5.5.5 BladePlug- A spinnerto bladering is required°

5.5°6 Removal/Installation -- The spinnershallbe designedforeachof installationandremovalQf thef'orwarspinne d rto facilitatesystemtestinqo

170"
6.0 MAINTAINABILITY/RELIABILITY

6.1 M&R

Thereare no specificdesignrequirementsformaintainabilityand reliability.

..... _ "'_ state- The hardwareshouldbe designedfollowingood "-' - "

of-the-arttechnique.

7.0 OTHERCONSIDERATIONS

7.1 Thereis thepotentialfora contractmodificationto addoppositerotation

(clockwise)hardwareto theprogram.Duringthedesigneffort,waystominimize

theeffectof incorporatingoppositerotationshouldbe considered.However,

oppositerotationshallnot be incorporatedat thistime.

7.2 The SR-7Lwillbe testedon the staticthrustrigatWrightFieldtoverifyits

staticperfonmance.For thistestit is desirabletooperatetheProp-Fanas a

fixedpitch/groundadjustableunit. (Thecontrolwillnotbe used,) TherefoLe,

duringthe designof theSR-7L,Designshouldconsiderwaysto incorporatea

fixedpitch/groundadjustableactuatorin theProp-Fan,but thisactuatorshall

not be designedat thistime.

7.3 The SR-7Lwill be testedon thestaticthrustrigat WrightFieldto explorethe

stallflutterboundaryand in theModanewindtunnelto explorethehighspeed

flutterboundary.For thesetestsit is desirabletooperatetheProp-Fanwith

_ directbladeanglecontroland notwithspeedcontro!.To accomplishthisit is

envisionedthatthebladeanglecouldbe variedusingan externa!hydraulicsupply

directedto thehigh or lowpitchchambersof theactuatorthrougha commercial

transferbearingmountedon the frontof theProp-Fan.Designshouldconsiderways

to incorporatethiscapabilityin thehardware,but thishardwareshallnotbe

designedat thistime.

171
7.4 The hardwaredesignedin thisprogramwillbe used in a futureNASAPropeller

TestAssembly(PTA)contractoThe PTAcontractwill includestaticenginetesting.

At thistimeit is notknownwhetherthistestwillbe conductedon an indoorcr

outdocrteststand. The hardwarewillalsobe subjectedto testingwithan engine

andwing in a low speedwindtunnelandwithan engineon a testbedairplane°

8.0 PROPELLERINTERFACES

8.1 Propellercontrolinstallationdrawing(preliminary)(Ref.L14325-I& L14325-2).

Thisdrawingdefinestheconceptand interfacebetweenthepropeller,thepitch

changesystem,thecontrol,and the gearbox.

8.2 Hub/Flange/Control FDuntin_-- SeeFigureI0 (T56)o

8.3 SlipRingu A slipringwillbe providedfor instrumentationly. The brush

blockwillmounton thecontrolin placeof thedeicingbrushblock° The slip

ringand brushblockshallbe suppliedby theInstrumentationGroup.

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FIGURE 1 173
ENGINEPOWERVSo ALTITUDE

S_ __Mn Altitude _ T_ Speed

6000 0-0°2 Sea Level 74oi 800 ft/sec

5400 0.5 10,000ft. 66.7 800

4500 0.6 20,000 55°6 800

3820 008 30,000 47.2 800

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FIGURE 2 174
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FIGURE 3 175-------
DESIGNCASESFORSR-TL BLADE

CA_E CONDITION Mn ALTITUDE HP SHP/D2 TIP SPEED RPM EF

1 Crufse 0.8 35,000 ft. 2592 32 800 fps 1698 4.5

2 M!B Climb 0.2 Sea Level 6000 74.1 800 fps 1698 4.5

.. 3 25% q_speed 0.8 35LgOO,._ftL .............. 0 .................. 0.................. !O00__f_PS 21_ .... 0 ......

4 40% O'speed 0.8 35,000 ft. 0 0 1120 fps 2377 0

:. ANALYSESTO BE CONDUCTED

FLUTTER CRITICAL STRESS

C____EE STALL H! SPEED SPEEDMARGINS MARGINS FOD

1 No Yes Yes Yes No

Yes No Yes Yes Yes

......... 3.................... No ........................... No ........ No ......Yes (steady).... No _ ..............

4 No No No Yes (steady) No

FIGURE4

Revtsed 9115/83
SR-7L BLADE- CONDITIONSTO BE ANALYZED

1. Wfnd Tunnel* 0.8 Fin 14,000 ft. 1050 HP 13 SHP/D2 800 fps 1698 rpm 4.5 EF

2. Statfc 0 Sea level 6000 74.1 800 1698 0

3. Reverse 0 Sea level 1200 14,8 800 1698 0

4. Ceutse-!o epm 0.8 35,000 2025 25 600 1273 4.5

S. C_Jfse-hf rpm & Nn 0.85 35,000 3280 40.5 840 1783 4.5

6. "C1imb O.5 10,000 5400 66.7 800 1698 4.5

7. Dtve 0.6 20,000 0 0 800 1698 0

8. Dive 0.8 35,000 0 0 800 1698 0

These conditionsw411 be analyzed--however,the blade deslgnwill not be modifiedbased upon the results.

, Conditionsshown are for 8 bladeso Will be done for 2 and 4 blades also.

Figure5

Revised 9/15/83
ANALYSESTO BE CONDUCTED

Flutter Cr_ttcal .Stress

_t___ee Stall Ht-Speed Speed Margins Margins FOD

1 No Yes Yes • Yes No

2 Yes No Yes Yes (Steady) No

3 Yes No Yes • Yes (Steady) No

4 No Yes Yes Yes No

No Yes Yes Yes No

6 No Yes Yes Yes No

7 No Yes No Yes (Steady) No .,

8 No Yes No No No

Figure 5 (continued)

-Rev|sed 9/15/83
UNITED

TECHNOLOGIES

J

-X +X f_-

PITCH AXIS _._

,I. X Y

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-34.0000 2.8300

-33.0000 3.9400

-32.0000 4.6000

-31.0000 5.5500

-30.0000 6.2400

-25.0000 0.7000

-20.0000 10.0900

-15.0000 10.7500

_! -10.0000 11.0000

-7.0000 11.1000

-5.0000 !1.3174

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I !1.2000 16,6000

15.0000 17.7098

20.0000 16.6000

25.0000 19.0077

30.3310 19.1750

O O O O O O O O O O O |

0 0 0 0 0 0 0 0 0 0

N • O O O O O O O O O N O O O

_ _ o _ _ _ _ _ _ _ •

FIGURE 6. PRELIMINARY SPINNER/

NACELLE CONTOUR FOR SR-7,PROP FAN

179,/180

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o
APPENDIXB

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1. RepoR No.CR174788 2. GovernmentAccession No. 3. Reclpient's Catalog No.

4. TitleandSubtitle 5. Repo_Date •

Large-ScaleAdvancedProp-Fan(LAP)Pitch Change A,a,_t1. ]q_R

ActuatOrand ControlDesignReport e Pe_orming Oranization Code

73030

7.Authors) 8.Pe_ormlngOrganizationReportNo. ._

RobertA. Schwartz

Paul Carvalho HSER - 9249

Mark Cutler 1owo. rkuniNto.

9. Pe_orming Organization Name and Address

11. Contract or Grant No.

HamiltonStandardDivision,UnitedTechnologiesCorp.

t P.O. Box 1000 NAS3-23051

Windsor Locks, CT 06096 13TypeofRepo . _andPertodCovered

12. Sponsoring Agency Name and Address

ContractorReport

NationalAeronauticsand SpaceAdministration 14. Sponsoring Agency.Code

Lewis ResearchCenter

21000 BrookparkRoad

Cleveland_OH 44135

15. Supplementaw Notes

ProjectManager,DavidA. Sagerser

AdvancedTurbopropProjectOffice

NASA Lewis ResearchCenter

Cleveland,OH 44135

16. Abstraci

In recentyears, considerableattentionhas been directedtowardimprovingaircraft fuel consumption. Studieshave shown that the high inherentefficiency

previouslydemonstratedby low speed turboproppropulsionSystemsmay now be

extendedto today'shigherspeed aircraftif advancedhigh-speedpropellerblades

havingthin airfoilsand aerodynamicsweepare utilized.

HamiltonStandardhas designeda 9-footdiametersingle-rotationLarge-Scale

AdvancedProp-Fan(LAP)which will be testedon a statictest stand,in a high speed

wind tunneland on a researchaircraft. The major objectiveof this testingis to

establishthe structuralintegrityof large-scaleProp-Fansof advancedconstruction in additionto the evaluationof aerodynamicperformanceand aeroacoustic

design.

This reportdescribesthe operation,designfeaturesand actualhardwareof the

(LAP)Prop-Fanpitch controlsystem. The pitchcontrolsystemwhich controls

blade angle and propellerspeed consistsof two separateassemblies. The first is

the controlunitwhich providesthe hydraulicsupply,speed governingand feather

functionfor the system. The secondunit is the hydro-mechanicalpitch change

! actuatorwhich directlychangesbladeangle(pitch)as scheduledby the control.

17. Key Words (Suggestedby Author(s)) 18. DistributionStatement

Prop-Fan

J PitchChangeMechanism GeneralRelease

Structural/HydraulicDesign

19. Security Classif. (of this repot) 20. Security Classif. (of this page) 21. No. of pages 22. Price"

Unclassified Unclassified 200

*For sale by the NationalTechnical InformationService,Springfield,Virginia 22161

199/200
¥
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