Cartridge Pneumatic Starter

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Copyright © 1967 by ASME

Jet Engine Starters, Cartridge-Pneumatic J. A. ANDERSON Group Engineer, Cartridge-Pneumatic Starters, Sundstrand Aviation, A Division of Sundstrand Corporation, Rockford, Ill.

C. R. GALASINSKI Project Engineer, Cartridge-Pneumatic Starters, Sundstrand Aviation, A Division of Sundstrand Corporation, Rockford, Ill.

A typical dual-mode (cartridge-pneumatic) jet engine starter is discussed with respect to its basic design and functional purpose. The establishment of starter output requirements for satisfactory engine starting is presented as an aid to understanding the design problem encountered for a jet engine starter. Discussion of energy conversion and particular design parameters concentrates on the cartridge mode of operation and the handling of the high-temperature, high-pressure gas produced by the starter cartridge.

Contributed by the Gas Turbine Division for presentation at the Gas Turbine Conference and Products Show, Houston, Tex., March 5-9, 1967, of The American Society of Mechanical Engineers. Manuscript received at ASME Headquarters, January 27, 1967. Copies will be available until January 1, 1968.

THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS, UNITED ENGINEERING CENTER, 345 EAST 47th STREET, NEW YORK, N.Y. 10017

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Jet Engine Starters, Cartridge-Pneumatic J. A. ANDERSON

C. R. GALASINSKI

1 INTRODUCTION

There are two particular utilization advantages to the approach of the cartridge-pneumatic starters to jet engine starting. The first utilization advantage is the self-sufficient starting capability. This means that when operating in the cartridge mode, it is possible for an aircraft equipped with cartridge-pneumatic starters to complete its start cycle without additional ground support equipment. The advantages of this are obvious, in that aircraft can be dispersed or operated out of remote or inconvenient areas where ground support equipment either cannot be made available or, at any rate, does not have to be made available. It should be pointed out here that the combination starter device does not have to be operated in the cartridge mode for all starts. This means that when ground support equipment is available, particularly for maintenance starting, the pneumatic mode can be utilized. This extends the service life of the starter and reduces the actual starting costs. The second utilization advantage of cartridge pneumatic starters is the alert or quick-start capability that can be provided. Since the aircraft can be started without any attached ground support equipment, the aircraft can be available for flight just as soon as the engines are brought up to speed. There is no time lost in the disconnecting of ground support equipment and coordination of this starting operation with a ground crew. Fig.l shows a simple schematic of a combination cartridge-pneumatic starter. It can be seen that this starter arrangement is made up of a cartridge gas generating system, a turbine or power-conversion arrangement, and a reduction gear box and disconnector arrangement. The power from either the cartridge gas generator during a cartridge-mode start or from a ground cart or aircraft interbleed system is transferred to appropriately designed nozzles, so that the same turbine is used for power conversion in either mode. A description of a typical cartridge-pneumatic starter is shown in Figs.2 and 3 by a schematic of the functional parts and a cutaway view. Fig.4 shows a list of the leading particulars of this one starter example.

It is the purpose of this paper to present a discussion of the significant parameters which must be considered when sizing or designing a cartridge-pneumatic type of starter. Special attention is given to the design areas which are unique to this particular type of device. To assist and orient the reader who may not be familiar with jet engine starters, an initial section, 2, which describes a typical cartridgepneumatic starter and its functional purpose, is included. 2 FUNCTIONAL PURPOSE AND DESCRIPTION OF A TYPICAL UNIT The primary purpose of all jet engine starters is to provide a torque level sufficient to accelerate the main engine rotor from zero rpm to engine light-off, and to assist the engine to a speed at which it alone can sustain the acceleration to engine idle without excessive internal temperatures. The unit is normally mounted directly on the engine accessory gearbox, which is mechanically connected to the engine rotor. Airframe facilities are provided for pneumatic inlet ducting, electrical power for cartridge ignition, accessory exhaust ducting, and access for cartridge loading.

CARTRIDGE GROUND CART

ENGINE INTERBLEED I 4

11111111P

rPOWER GENERATION

I REDUCTION GEAR - 1POWER CONVERSION BOX AND DISCONNECT

Fig. 1 Cartridge-pneumatic starter schematic

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13

Fig.2 1 2 3 4 5 6 7

8 9

Cartridge Breech cap Breech handle Connector Ignition contact Ground clip Hot-gas nozzles Turbine rotor Turbine exhaust ring

Schematic of functional starter components 10 Overboard exhaust connector 11 Exhaust from turbine and fan 12 Relief valve 13 Compressed-air inlet 14 Aerodynamic-braking fan 15 Air inlet for braking fan 16 Fan exhaust ring 17 Gearshaft 18 Overrunning sprag clutch

3 ESTABLISHMENT AND DISCUSSION OF SIGNIFICANT DESIGN PARAMETERS The sequence of this section presents first a discussion of the parameters which determine the starter output requirements; then in the subsequent sections the energy conversion is followed through the starter from the cartridge to the nozzle, to the turbine rotor, and to the output shaft. In the last section, special items are covered which are unique to cartridge starters, such as hot-gas handling, speed control, and safety devices. A Establishment of Starter Output Requirements The starter output requirements are normally established by the airframe manufacturer and are based upon the engine requirements, the accessory

19 Flyweight 20 Switch actuating rod 21 Switch 22 Adjusting screw 23 Gearbox vent 24 Spline shaft 25 Oil slinger 26 Oil sump 27 Magnetic plug

drag-torque loading, and the desired aircraft performance. The starter output requirements may be presented in specific terms, or in terms of engine or aircraft operational requirements. In specific terms, the requirements are generally presented as an actual "starter torque versus speed envelope," as shown in Fig.5. To present the requirements in this form, the airframe manufacturer must perform the necessary calculations to assure that the minimum torque curve will result in satisfactory engine starts. In many cases, however, the starter output is presented in terms of the engine starting requirements and aircraft operational requirements. In this form the airframe manufacturer takes the engine drag-torque curve from the engine manufacturer and adds the drag-torque loading for the specific hydraulic and electrical accessories.


1. 2. 3. 4. 5. 6. 7. 8. 9.

Electrical Connector 10. Containment Clamp Switch 11. Turbine Rotor Switch Actuating Rod 12. Aerodynamic Braking Fan Lubrication Tube 13. Breech Cap Assembly Output Clutch Race 14, Breech Cap Chamber Output Spline Shaft Assembly 15. Pneumatic Inlet Adapter (Installation Part) 16. Ignition Connector "V" Band Coupling (Installation Part) 17. Transfer Tube Gearshaft Fig. 3 Cutaway view of cartridge-pneumatic starter

The resultant curve may look like Fig.6. The aircraft operational requirements may also dictate a maximum time-to-idle curve such as Fig.7. The significance of the curves and limits of Fig.6 is as follows. The maximum torque limit is determined by the load-carrying strength of the accessory gearbox and engine accessory power train. The -65, +59, and +130 F drag curves are representative of the typical engine plus accessory drag torque at that specific ambient temperature. At the -65 F condition, special considera-

tion is given to the rate of acceleration to light-off because of the viscous drag loading of the engine and accessories. The minimum assist or cutout speed for the starter is determined by the minimum positive torque above which the engine can accelerate itself uniformly to idle. Given the information in Figs.6 and 7, the starter designer can begin to approximate the shape of the starter output curve which will best satisfy all the requirements of minimum engine assist and time to idle. The resultant time to 3


Weight ............................................................................ 61 lbs. 6 oz. Operational Temperature Range ........................ -65 to +160 ° F. Maximum Operational Altitude ........................ 6000 feet Operational Attitude ................................................ Horizontal +10 degrees Output Rotation (when facing output spline) ................................ Counterclockwise Maximum Output Shaft Torque ............................ 680 pound feet Output Spline Shear Section ................................ 800 to 900 pound feet Pneumatic Shut-Off Actuation (output rpm) ............................................................ 2970 + 100 rpm External Disengagement at Output Spline Shaft (output rpmp) ............................................ 3000 rpm Maximum Turbine Operational Speed . .

67, 500 rpm

Cartridge Required .................................................... Air Force Type MXU-4/A or MXU-4 A Electrical Requirement for Ignition ................ 18 to 30 volts dc at 1 ampere Relief Valve Actuation Point ................................ 700 psi Gear Ratio

14 88 to 1

Pneumatic Requirement Source ........................................................................ Compressed air from MA-1A pneumatic ground cart or equivalent Maximum Pressure ............................................ 60 psia Permissible Frequency of Use Cartridge Starts per Hour ................................ 2 Minimum Time Between Starts on Aircraft (minutes) .................................... 5 Minimum Time Between Starts During Testing (minutes) .................... 30 Maximum Continuous Motoring ............................ 10 minutes Minimum Waiting Time before Removal of Spent Cartridge* ............................................ 1 minute Minimum Waiting Time before Removal of Hang Fire Cartridge* .................................... 15 minutes Oil Requirements Specification ............................................................ MIL- L-7808 Maximum Oil Capacity .................................... 275 cc(9. 3 ounces) *Assumes use of asbestos gloves

Fig. 4 Table of leading starter particulars reach engine idle (t) is derived from the following relationship. T I T

n

T

= I = I

T

e

a + I 1-

n

A

+ I

s

= Ts - (T e + T Acc )

(lb/ft)

(1)

(slug-ft2)

(2)

(1b/ft)

(3)

where T I

T

n

T

s

= net accelerating torque reflected to the starter mounting pad = total inertia reflected to the starter output shaft: l e = engine, I A = accessory, I s = starter = steady-state starter torque at the starter mounting pad: Te = engine torque, T Acc = accessory torque

4


LL

+130

I +59 LI)

z

ka.i 6-5

-65

30 40 TIME TO IDLE (SEC .1

STARTER OUTPUT SPEED IN)

Fig. 5 Starter torque versus speed envelope

Fig. 7 Engine starting requirements (time to idle) -65°F

4000

+60°F MAX ALLOWABLE STARTER TORQUE

3000

+160°F

2000

-65°F +59°F +130°F

STARTER SPEED

0.15

1500

010 009

1000

§ 0.08

2

MIN ASSIST 100 LB.-FT

0.07 0 06 0 05 004 0.03 200

ENGINE LITE- OFF = 200 RPM + 8 TO 10 SEC

300

500

1000

500

2000 3000

CHAMBER PRESSURE, PSI INERTIA ID OF ENGINE PLUS ACCESSORIES AT STARTER PAD =IXISLUG-FT.

EFFECTIVE GRAIN BURNING THICKNESS = 119 IN.

2

2

EFFECTIVE GRAIN BURNING SURFACE IA, I= 129 8 IN.

Fig. 6 Engine starting requirements (torque versus speed)

A

t

= EFFECTIVE NOZZLE AREA

Fig. 8 MXU-4/A Cartridge burn rate and K n versus chamber pressure and grain temperature (fixed nozzle system) (Courtesy of Olin. Mathieson Chemical Corporation) a = acceleration reflected to the starter mounting pad N = starter output speed t = time to idle Then, from equation (1): dn

T n = I T &t

,r

idle

t

- IT

(lb/ft)

( 4)

(sec)

(5)

0

The actual calculation of time to idle using equation (5) is performed by taking AN increments, using the average torque for each increment, and determining the At for each increment; i.e.: t

I x 2/t/60x (4h.RPM) Tnet

time to idle (t) =1 0 t

(sec)

(6)

(sec) (7)

The same method of calculation is used to determine the engine speed at cartridge burnout.

B Starter Cartridge Input Energy Available The energy source for self-sufficient starting of an aircraft utilizing a cartridge-pneumatic starter is a starter cartridge such as the MXU-4/A or MXU-4A/A. The primary constituent of these cartridges is an 8-lb grain composed of pellets of an ammonium nitrate base propellant homogenously suspended in an appropriate combustible binder. The design of the grain is such that the burn rate, or rate of consumption of the grain, produces a relatively constant mass rate of gas flow having a characteristic temperature of approximately 2400 R, and composed of CH4, CO 2 , H 2 , H2 0, N 2, and Na CO 2 3' It is also characteristic of ammonium nitrate base propellants to exhibit a burn rate which is directly proportional to a function of the conditioned temperature of the grain and the operating chamber pressure. Graphical representations of cartridge chamber pressure versus time, burn rate and K versus grain temperature and n chamber pressure for the typical MXU-4/A cartridge are shown in Figs.8 and 9. The term K is the n

5


1500

1000

500

0 a_

1500

1 ,1

Lailemarmos

1500

II 111 \ t

I

V)

Er)

cc

100 0

a_ re

Fig. 10 Cartridge burn characteristics (fixed-nozzle system)

k eililiMMEMIN

rlillii iirillirili \\ IW= -- 1ENE1MIN -N

50

500

ril 1

0

—65°F ■•■MI■LUIll 8

I2

I6

2

28

THE,SECONDS

Fig. 9 MXU-4A Cartridge chamber pressure versus time (fixed-nozzle system) (Courtesy of Olin Mathieson Chemical Corporation) ratio of nozzle-throat area to the burning surface of the cartridge grain, and therefore is useful in predicting chamber pressures for a particular nozzle system. It is evident from Figs.8 and 9 that, for a fixed-nozzle size, the increased burn rate of a +160 F cartridge will result in an increased chamber pressure, which in turn results in an additional increase in the burn rate of the grain until stabilization is reached. Typical chamber pressure versus time curves for cartridge temperatures of -65 and +160 F, with a fixed-nozzle size, are shown in Fig.10. However, since the output torque of the starter is a function of the pressure ratio across the nozzle, the magnitude of this chamber pressure variation with temperature produces excessive torque output at the high ambient temperatures required. The starter described in Section 2 therefore utilizes a highly pressure-sensitive valve, which produces a variable-nozzle bypass orifice. The resultant increase in effective nozzle area reduces the chamber pressure below the fixed-nozzle condition, such that the pressure difference between the hot and cold cartridge is minimized. The maximum starter output torque and, subsequently, the impact torques sustained by the engine gearbox can then be controlled within a more desirable range. Fig.11 shows the typical chamber pressure versus time curves for cartridge tempera-

tures of -65 and +160 F while using a valve to produce a variable orifice. Fig.11 also shows that, with the variable orifice design, the reduced cartridge pressure results in a slower burn rate and correspondingly longer cartridge burn time for the hot-day conditions. This resultant characteristic is beneficial in starting engines which are slow to light-off. The available energy, or available adiabatic head (H ad ), produced by the cartridge may be calculated from published values of the cartridge specific heat ratio (k), specific gas constant (R), the flame temperature (T p ), and the appropriate design operating pressure (Po) and exhaust exit pressure (Pe). Thus H

ad =

T [ 1-(pe ) Po

K R K-1

(ft)

(8)

For the typical MXU-4A and MXU-4A/A cartridge gas:

R

= 81

K =

1.27

T =2400 ° R To simplify equation (8), let K-1

(9)

Y = 1-( Pe) K Po

Then, from Fig.12, the value of Y can be determined for any Pe/Po with K = 1.2 to 1.4. Therefore, equation (8) may be expressed as H

ad =

K K-1

RT,Y

(ft)

(10)

The spouting velocity (Co), or maximum theoretical gas velocity which can be obtained at a nozzle exit from the available energy, can then be determined from


825 750

16

TIME - SECONDS

21

Fig. 11 Cartridge burn characteristics (variable-nozzle system)

05 P E

1/2 C o = ( 2g H ad )

or (2g

k RT p Y) k-1

0

1/2

01

(ft/sec)(11) 005

The energy available from the cartridge may also be calculated in terms of gas horsepower (GHP) from the mass rate of cartridge gas flow (W f ) and the adiabatic head (H ad ), where W f is a function of the grain burn rate (r b ) at the desired operating pressure and temperature, the grain burning area (A s ), and the grain density (p ). Thus W

f

001

0005

r,

from Fig.9 (12) A s = 129.8 in 2 (typical) jOl p = .053 #/in 3

= rb As

0001

4

0

w H GHP - 5f50ad

(HP)

(13)

Approximate values of the above properties for a typical MXU-4A cartridge at nominal operating conditions of +60 F, 865 psia, and exhaust pressure of 18 psia are H

C

ad :Z.- 506 x 10 3 FT 3 o C:= 5.7 x 10 FT/sec

f GHP

.52 1,13 m /sec

481 HP

Fig. 12 Pressure ratio versus Y diverging nozzles requires that the Mach number (M) at the nozzle throat must be unity, and that the pressure ratios P e /P o must be less than the critical pressure ratio P*/P o required for M = 1 at the throat, where P* is the pressure at the throat, and the critical pressure ratio P*/P o is defined by *

P (

Po

Starter Nozzles and Nozzle Energy Conversion The first stage in the transformation of the heat energy produced by the cartridge to usable work is the conversion to kinetic energy by a nozzle system. The starter accomplishes this energy conversion by expansion of the gas through supersonic converging-diverging nozzles. This type of nozzle is used because of its characteristically high efficiency for high pressure ratios P o /P e . Before defining the energy conversion by the nozzles, it should be recalled that supersonic isentropic flow of a gas through convergingC

k

2 k+1

) K -1

(14)

It should also be noted that the actual design value of P o must not only be consistent with the requirements for supersonic isentropic flow, but must also be of a sufficient magnitude to produce the required starter performance with the final nozzle and turbine design incorporated in the starter. Experience has shown that values of P o in the range of 800 to 1000 psig are generally sufficient. The conditions of the gas at the nozzle exit may be determined from the following pressure, temperature, and density relationships for isentropic flow:

7


Po

7

or (Po ) 15;

k -1 k = To Tr— -e (15)

= .Po 'doe

e

Co = ( 2°48- 6 )

12

/

(ft/sec)

The actual nozzle exit velocity is, therefore:

( 1 8)

To summarize the conversion of the heat energy of the gas to kinetic energy, it can be stated that the nozzles transform the lowvelocity, high-temperature, high-pressure gas produced by the cartridge to a lower static temperature, lower static pressure gas having a supersonic velocity and the same mass rate of flow. The nozzle exit W can be determined from f the relationships for isentropic flow at the critical nozzle area A*. t' f

C*

= Jo*

C*

A*

(KgRT*) 1/2 •

..e

e Fi. 7,

(KgR ) 1/2 T o 1/2

P

(T-71)

(19)

A* = Kd A t * K d = Nozzle discharge coeff = .96 Therefore: Wf =

A* P o (Kg) 1/2 )K+1 (RT 0 )1/2 "' T7=1)

K+1

(FTC/sec)

2(K-1)

(21)

k 2 I For the typical cartridge gas K

1.27

Max.

3.78

Therefore W f = A* P o max (lbm/sec) (22) (RT 0 )

(17)

The actual nozzle exit velocity, however, is dependent on the nozzle velocity coefficient (k v ), which is the ratio of the actual exit velocity (V ) to the theoretical exit velocity with isen1 tropic flow and the same exit pressure P e . The value of k v for supersonic nozzles having straight axes has been found from experience to generally equal 0.96.

V1 = .96 C o (Ft/Sec)

12

(Kg) / K-+1N --

(16)

where p is the gas density. As noted in section B, the maximum theoretical exit velocity of the gas at the nozzle exit may be determined from

W

le max =

(lb m/sec) (20)

For pressure ratio (Po/Pe) greater than the critical ratio, the foregoing Wf equation can be simplified as follows by using the maximum compressibility factor ( y max), where

D Starter Turbine Rotor and Turbine Energy Conversion The second stage in the transformation of the cartridge gas energy is the conversion of the kinetic energy of the gas leaving the nozzle system to usable shaft power by a means of a singlestage impulse turbine. The absolute blade speed used for the design of a starter turbine is not necessarily the speed at which maximum efficiency is obtained, because the starter must be designed to yield a maximum output torque at the most critical starter output speed. The design speed used for the blading design, and therefore the entire starter system design, is therefore the one corresponding to the starter output speed at which the maximum engine drag torque must be exceeded. The actual turbine blading utilized by a cartridge-pneumatic starter must also be designed on the basis of a compromise between the most practical turbine efficiency for both the cartridge and pneumatic-energy sources. The reason for this is that the mass rate of gas flow from the pneumatic-energy source is generally greater than that produced by the cartridge, and the size and weight requirements of the starter assembly necessitate the use of a cascade of converging nozzles for most efficient energy conversion of the pneumatic gases. Therefore, the nozzle exit conditions and velocities are significantly different for the two energy sources. Let us,begin the analysis of energy conversion by an impulse turbine by defining an impulse force as one which results from a decrease in magnitude or change in direction of the tangential velocity of a fluid. For the ideal impulse blade shown in Fig.13, if the absolute tangential velocity of the blade is U and the absolute velocity of the fluid leaving the nozzles and entering the blade is V 1 , then the relative velocity of the fluid entering the blade (W 1 ) is, vectorially:

8


VI

W =v - U 1

I

U

VZ W 2 =-W I =U - V 1

Fig. 13 Schematic of ideal impulse blade Fig. 14 Velocity diagram for simple impulse turbine W V=1 --.- U

(23)

1

Using similar notation, and assuming frictionless contact between the fluid and the blade, it can be seen that the relative velocity of the fluid leaving the blade (W 2 ) must be equal in magnitude but opposite in direction to W. Therefore =--.

W2

--a

"'

U

W

1

--0.•

(24)

V1

Then, since W2 is opposite in direction to Wi and V 1 , it can be seen that the absolute velocity of the fluid leaving the blade (V2) must be V2 = W 2 U ===-1.W1.4-11. U = 2U

V1

(25)

This brief analysis of the ideal impulse blade clearly shows that the effective tangential velocity of the fluid entering the blade (V 1U ) is decreased, resulting in an impulse force. The force produced by the gas is derived from Newton's laws of motion, which state that momentum is the product of the mass of a body and its velocity and that the rate of change of momentum is equal to the sum of the external forces acting on the body. Therefore F

= d dt

(M) = d

( E—

dt

g

(1bf)

(26)

The impulse force produced by the tangential velocities of a gas entering and leaving a turbine blade passage may therefore be defined as F

i =

dt

W

)(

V

1U

V2U)

(lb f )

(27)

Now by noting that the mass rate of fluid flow (W f ) is equal to dW/dt, and that torque about a fixed axis is equal to the product of the force and the distance from the axis to the point at which the fluid enters the blade (r), the turbine torque produced by the fluid may be defined as

/7/ = r ( dw

dt /

( V 1U - V2U (lb/ft)

T =

(28)

(V1U - v2u)

The analysis of a real impulse turbine is facilitated by a velocity diagram such as the one shown in Fig.14, where, in addition to the nomenclature previously defined: a = angle of absolute fluid velocity V 1 leaving nozzle 0 = angle of relative fluid velocity W1 entering blades y = angle of relative fluid velocity W 2 leaving blades 8 = angle of absolute fluid velocity V2 leaving blades V ia and V 2a = resultant axial components of absolute fluid velocities V 1 and V 2 . Using equation (28) and the relationships V 1U

= V1 Cos oCand

V2u = V 2 CoSS

derived from the velocity diagram of Fig.14, the value of T for a real impulse turbine may be bxpressed as

=

W f r (V 1 Cos0C- V2 Cos g

)

(lb/ft)

(29)

Since the relationships derived from the analysis of the ideal impulse blade assume steady flow and no blade passage losses, it is only necessary to apply suitable design coefficients for utilization of those relationships for practical turbine analysis and design. These design coefficients, which account for the difference between the relative fluid entrance velocity W i and the relative fluid exit velocity cannot be calculated because of the numerous W 2' indeterminable factors which effect the blade passage losses. However, from the interpretation 9


torque, which occurs at zero rotor rpm, may then be derived from equation (32) by equating k p k i = kb and setting U = 0. Therefore

2

4

6

8

Ts

L0

= w f rv,

+ k b k d Sin

(Cos

U/v i

and evaluation of extensive testing, a satisfactory correlation between predicted and actual results has been obtained by assuming that the significant losses occurring in single-stage impulse turbines are represented by a profile-loss coefficient (k p ), and an incidence-loss coefficient (ki). The coefficient k p accounts for the losses due to fluid turbulence, friction, deflection of the fluid within the passage, curvature of the blade profile, and the deviation of the actual fluid exit angle y from the geometric blade exit angle y'. The coefficient k i accounts for the losses due to turbulence caused by the difference between the entrance angle 13 of the relative fluid velocity W i and the geometric blade entrance angle Another coefficient, which is required for partial-admission systems only, is the partialadmission factor kd. The value of this coefficient is a function of the blade profile width (c), the geometric blade exit angle (y 1 ), the thickness of the blade trailing edge (t e ), and the arc length of admission (e), and may be calculated from

-

E

yo.

2

-0-t e (30)

e

(ft/sec)

(31)

Using the relationship of equation (31) and the identities V 2 Cos

S

= U

-

W2 Cos r and W 1 = V1 SinSin/g

from Fig.15, equation (29) may be expressed as

T.

=

WfrV

( Cos

kp k i k d Sin

64. Cos Sin /9 (lb/ft)

- u (32)

A simplified equation for the turbine stall torque T s , or the maximum obtainable turbine

(33)

For practical analysis of starter turbine systems, it is a common practice to account for all the losses which occur in the nozzle-turbine system and the fixed geometric parameters by means of a stall torque coefficient, designated as tau ( From equation (33) let

T max = Cos a '

kb k d Sin of Cosr (34) Sinif

The equation for turbine stall torque may then be expressed as

Ts

=

'

rmax

W

fV1 r

(lb/ft) (35)

Also, since the turbine torque will be equal to zero when the tangential blade velocity is equal to tangential velocity of the gas entering the blade, V 1 Cos a, the value of T at the corresponding rotor velocity must equal zero. However, in the actual turbine system, the maximum tangential velocity of the gas will never equal V 1 Cosa, but will incur some losses due to the deviation of the angle of the gas from the geometric angle of the nozzle. Therefore U

Therefore, recalling that the effect of the blade losses is to decrease the magnitude of W2, the actual value of W 2 is defined by W2 = k p k i kd

r

(lb/ft)

Fig. 15 Typical torque coefficient versus 11/V 1 (cartridge mode of operation)

kd = 1

Cos

Sin/

max

= KV Cos c<.. 1

or ( U 1 ) max

= K Cosc.C.

(36)

Empirically it has been found that for the starter turbine systems, which have a nozzle angle (a) in the range of 17 to 20 deg, the maximum value of (U/V 1 ) max is approximately equal to 0.90, which is also the speed ratio at which 'C equals zero. Therefore, since turbine torque versus speed is very nearly a linear function, a curve of T versus U/V for a typical starter turbine system 1 may be derived as shown in Fig.15. The simple expression of torque in equation (35) is very useful in the evaluation of modifications to an existing unit configuration and for predicting unit performance with intermediate or new cartridge pressure levels. With a known fixed-nozzle diameter, the value of W f and V 1 can be determined from the measured breech pressure. The 'C factor can then be determined from the

10


measured starter output torque, or the torque can be predicted by a previously determined "t" versus U/V relationship. 1 It is apparent that small-diameter impulse turbines must operate at extremely high rotative speeds for most efficient energy conversion of the high-energy gas produced by a cartridge. Therefore, since the torque-speed characteristics of the turbine will almost never match the torquespeed characteristics required for starter output, a gear ratio must be employed between the turbine and the starter output shaft. The value of this gear ratio must be a compromise between the gear ratios which will produce the most desirable starter torque-speed curves for both the cartridge and pneumatic modes of starting. The reason for this compromise is that, because the stall torques resulting from a typical Air Force MA-1A type ground cart pneumatic-energy source are normally lower than those resulting from a cartridge energy source, the required slope of the torque-speed curve for the pneumatic mode must be much less than that of the cartridge mode in order to exceed the engine drag torque at all output speeds below the required minimum starter assist speed. The gear ratios actually used for starter systems range from 13:1 to 22:1. Utilizing the required gear ratio (G) and the corresponding efficiency (% ) of the gearbox G required, the actual starter output torque may be determined from

T

starter

=

TWf

r

via

(lb/ft)

(37)

g

Typical torque versus speed and horsepower versus speed curves for a starter operating in the cartridge mode at ambient temperature are shown in Fig.16. E Hot-Gas Handling and Transfer System Special attention must be given to the selection of the materials used for the starter components carrying the cartridge gases, in order to provide: 1 Sufficient pressure vessel strength at maximum temperature. 2 Minimum surface temperatures. 3 Minimum localized fatigue stresses due to high-temperature cycling. 4 Minimum erosion by the cartridge gas. 5 Minimum high-temperature corrosion by the cartridge gas residues. The problems encountered by the designer and metallurgist in achieving the desired endurance of the components subjected to contact with the hightemperature cartridge gas (1900+ F) have been formidable.

600 CO CO

500

w 400 ce

300 200 100

120

HP VS SPEED

%EMMEN= • •

100 80

■■11E 60 VS. SPEED 1111■ 3RQUE 40

F■IIIIM11■■ 20 1000

2000

3000

0 4000

STARTER OUTPUT SPEED — RPM

Fig. 16 Typical starter torque and starter horsepower

versus output speed (cartridge mode of operation) The first problem encountered was that of material cracking due to thermal cycling stresses. The selection of materials used to confront this problem was accomplished mostly by experimental testing with the high-temperature nickel and cobalt-base alloys. Test components were made from both investment cast and wrought materials. In most of the applications it was found that the cobalt-base alloys exhibited a slightly higher resistance to thermal cracking, and that the wrought material provided a longer endurance life. It was also found that the resistance to thermal cracking was greatly increased by the incorporation of the most practical gradual transition between thin and thick cross sections, and by utilization of the most generous radii practical, at all internal corners. Weld joints were found to be particularly prone to cracking in the heat-affected area of the parent material, which was minimized by using equal cross sections for the components at the weld joint, whenever possible. High-temperature corrosion and/or erosion potential exists in all areas exposed to the flow of the gases. Here again, the wrought cobalt-base materials exhibited the better resistance. It has been found that the cartridge ignition gases contain sodium and potassium salts along with various degrees of sulfur compounds, depending upon the particular ignition design. These elements, in combination with a reducing atmosphere, react with the metals at high temperature to form nickel and chromium sulfides. Also, the chromium reacts with the carbon in the reducing atmosphere to form chromium carbides. Both of these reactions cause the chromium to be depleted as a corrosion-resistant element and allow the nickel or cobalt to later oxidize more rapidly. It is generally found that the nickel alloys are impervious to this corrosion below 1200 F and the cobalt alloys are good for approximately 200 deg F higher. Diffusion coatings are helpful in resisting this corro11


I BREECH CHAMBER ASSEMBLY SWITCH

BREECH CHAMBER PRESSURE VESSEL INTERNAL GAS SHIELD INTERMEDIATE RADIATION SHIELD GAS FLOW LINER TUBE

INSULATION INSULATION 8 RADIATION GAS FLOW

Fig. 18 Overspeed protection which diverts gas before nozzle

INTERNAL (FREE FLOATING) GAS TUBE PRESSURE VESSEL TUBE OUTER SKIN III NOZZLE BLOCK ASSEMBLY

w‘s mmvi

s‘‘‘m ex

r

GAS FLOW ►

OUTER SKIN s‘ .

x x

IN

IN S U L A T I 0 N PRESSURE VESSEL RADIATION SHIELD AND INSULATION FLOW LINER

SENSING ARRANGEMENT AND GEARBOX APPROACH 2

Fig. 17 Schematics of hot-gas transfer components

Fig. 19 Overspeed protection which diverts gas after nozzle

sion, and in some areas of the starter where corrosion was particularly bad, a chromized diffusion coating has been used with success. The basic design approach used for the starter described in section 2 was to use a material which exhibited high corrosion resistance and good high-temperature strength, as an internal gas shield and/or gas-transfer member, wherever possible. These internal shield and/or transfer members are only required to direct the main flow of cartridge gases, and are not required to withstand the stresses produced by the gas pressure. Therefore, these members are designed to be free-floating wherever possible, thus minimizing the stresses due to thermal expansion and internal pressure. The structural members, which must actually withstand the gas pressures, are protected from the high-temperature cycling by the use of insulation materials, air gaps, and/or shielding between it and the gas-carrying member. From Fig.17 it can be seen how this design approach is applied to each area of the starter which transfers the hot gases.

diameter required for sufficient output torque, it is not practical to design a turbine rotor which can structurally withstand a free-running condition at tip speeds corresponding to rotor speeds near 160,000 rpm. It is therefore necessary to provide a reliable means of limiting the maximum rotor speed. There are several methods of providing turbine overspeed protection in a cartridge starter. A few of the devices which have been utilized are: 1 Diversion of the hot-gas flow so that it does not pass through the nozzle when cutout speed is attained. 2 Diversion of the hot gas between the nozzle exit and the turbine inlet, either partially or completely, when cutout speed is attained. 3 Modification of the single-stage turbine output torque so that the turbine is aerodynamically or inherently speed limited within a safe design speed. Each of these methods of speed limiting has been used with varying degrees of success, and there are advantages and disadvantages to each. Method 1. Fig.18 shows an arrangement whereby the gas is diverted before reaching the nozzle. In this case, it is necessary to have some type of speed or acceleration sensing device with a cutout or bypass arrangement in the cir-

F Speed Control and Safety Devices Because of the high spouting velocity (C o ) of the cartridge gas (5700 fps) and the rotor


FAN

0 0 CARTRIDGE RUBBER SEAL REQUIRED LEAKAGE GAP

ELECTRIC CONNECTOR AND BREECH LOCK

Fig. 21 Breech locking design REENTRY DUCT

Fig. 20 Aerodynamic speed control

cuitry. The sensing device may include a flyball governor system in conjunction with a switch, squib, or solenoid, and the necessary electrical circuitry. The governor system may be mechanically connected to a bypass valve in the hot-gas circuit. Method 2. Fig.19 shows the arrangement of the cutout device which diverts the gas from completely going through the turbine wheel downstream of the cartridge nozzle system. This approach to speed control has a sensing arrangement which may either sense the output speed, the output acceleration, or the output torque of the starter. There have been two schemes of control mechanisms with this type of design approach. The first would be to place some type of baffle plate between the nozzle and the turbine, so that the gas coming from the nozzle would be deflected and prevented from going through the turbine. The second would be to vary the position of the turbine with respect to the nozzles, such that the blades could be moved out of the nozzle gas flow. Method 3. Fig.20 is a diagram showing two approaches to aerodynamic or inherent turbine speed control. The first arrangement is that of using a torque-absorbing fan as a part of the rotating single-stage turbine assembly. The second approach shown is that of staging the turbine wheel, thus producing a torque curve which in itself goes to zero torque within a safe turbine yield speed. There are some basic efficiency losses in both systems; however, in the fan approach, the losses in the critical starter operating speeds can be kept at a minimum.

G Breech Electrical Interlock Fig.21 shows a typical starter cartridge breech assembly. It includes the basic pressure vessel in which the cartridge is burned as well as a joint at which the pressure vessel separates for the purpose of installing the cartridge. The breech must also include a locking device for indexing and locking the two halves of the breech after the cartridge has been installed. This locking device includes the contact for the electrical circuit which is used to ignite the cartridge. The breech handle which carries part of the electrical circuitry is designed so that the breech assembly has to be closed and locked before the breech electrical circuit can be completed. This is to prevent any inadvertent firing of the cartridge before the breech is completely closed and locked. H Overpressure Protection Overpressure protection is a significant design requirement for the safe handling and utilization of the cartridge gas in a starter system. There are two basic approaches to overpressure protection. One design is to use a rupture diaphragm. The second is to use a spring-loaded plunger. Both systems have been used satisfactorily. 4 SUMMARY In summary, the design of the cartridgepneumatic starter begins when the prime contractor combines the engine starting requirements with the accessory loading and aircraft performance requirements to establish the starter performance limits. The starter manufacturer first analyzes the airframe requirements and approximates the starter output torque versus speed curve which will satisfy a particular application. The size and type of gas generator is determined to supply the required rate of gas flow (namely, Wf, C o , and so forth) and the total propellant burn time. The

13


starter nozzles and turbine rotor are designed to convert the gas energy to mechanical energy within an economical weight and package size. A compromise is made in the nozzle and turbine to provide both cartridge and pneumatic-mode operation. Special design considerations unique to the cartridge starter are given to the handling of the high-temperature and pressure cartridge gas to arrive at reasonable endurance life. Safety features must be included to provide safe no-load operation and overpressure relief.

In the early 1950's the U.S. Air Force haa its first jet engine cartridge starter operational on the B-57 aircraft. Beginning in the late 1950 1 s, cartridge-pneumatic starters were developed and have been installed on operational F-100, F-101, F-105, B-52, KC/C-135, F-4C, and F-111A aircraft. It is evident from this lengthy list of U.S. Air Force applications that the cartridgepneumatic starter has proven itself to be a reliable means of providing self-sufficient jet engine starting.

14


Cartridge Pneumatic Starter

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