NASA NASA-TP-2519 19860010877

Technical

Paper

2519

December 1985

Static Internal Performance

of a Single-Engine

Nonaxisymmetric-Nozzle

Vaned :Thrust.Reverser

Design With Thrust

Modulation Capabilities

Laurence D. Leavitt

and James R. Burley II

• _ ,_ .:A..,:_._C:.] :,, ......

NASA

Technical

Paper

2519

1985

Static Internal Performance

of a Single-Engine

Nonaxisymmetric-Nozzle

Vaned-Thrust-Reverser

Design With Thrust

Modulation Capabilities

Laurence D. Leavitt

and James R. Burley II

LangleyResearchCenter

Hampton, Virginia

National Aeronautics

and Space Administration

Scientific and Technical

Information Branch

Summary cal aircraft, the Saab-Scania Viggen and the Panavia

An investigation has been conducted at wind-off Tornado, currently use reversers for reduction in

landing ground roll. However, much of the capability

conditions in the static-test facility of the Langley of these landing-ground-roll reversers is lost by de- 16-Foot Transonic Tunnel. The tests were conducted

lays in deployment of the reverser upon touchdown

on a single-engine reverser configuration with par- and by the relatively long engine spool-up times retial and full reverse-thrust modulation capabilities, quired for maximum reverse thrust (i.e., maximum The reverser design had four ports with equal areas.

The ports were angled outboard 30° from the verti- dry power) once aircraft touchdown has occurred.

cal to impart a splay angle to the reverser exhaust (See refs. 1 and 2.)

To eliminate the spool-up time and thus make

flow. This splaying of reverser flow was intended to maximum use of thrust reversing, the engine must be

prevent impingement of exhaust flow on empennage at a high power setting during the approach prior to

surfaces and to help avoid inlet reingestion of ex- touchdown. However, the power required for flight

haust gas when the reverser is integrated into an at approach conditions is low, and a method must

actual airplane configuration. External vane boxes be developed to spoil the excess thrust generated

were located directly over each of the four ports to at high power settings. Several nozzle concepts use provide variation of reverser efflux angle from 140° to

26° (measured forward from the horizontal reference partial thrust reversing (along with thrust vectoring)

axis). The reverser model was tested with both a to achieve trim and proper approach lift-drag ratios.

Once the aircraft has touched down_ the partially butterfly-type inner door and an internal slider door

to provide area control for each individual port. In deployed reverser can be fully deployed in about

1 second to provide maximum reverse thrust nearly

addition, main nozzle throat area and vector angle at touchdown.

were varied to examine various methods of modulating thrust levels. Other model variables included Many studies on the concept of in-flight thrust

vane box configuration (four or six vanes per box), reversing have been conducted in recent years. (See

orientation of external vane boxes with respect to in- refs. 1 to 40.) Flight and wind-tunnel test proternal port walls (splay angle shims), and vane box grams on the F-94C (ref. 3), the F-100F (ref. 4),

and the F-11A airplanes (refs. 5 and 7 to 11) were sideplates. Nozzle pressure ratio was varied from 2.0

to approximately 7.0. conducted between 1956 and 1975. These early proResults of this investigation indicate that the grams demonstrated that thrust reversers integrated

into single-engine axisymmetric exhaust nozzles pro- nozzle/reverser configuration had the capability of

vided many benefits, including improved flight-path generating values of thrust ratio ranging from 0.97

control and reduced landing ground roll. However, (for the main nozzle only) to -0.70 (for full reverse

in all studies, problems were found in the handling thrust). This modulation of thrust ratio could be

obtained by varying reverser efflux angle, inner door qualities of the aircraft when the reversers were deangle, and/or main nozzle throat area. The design ployed at landing-approach (low-speed) conditions.

splay angle of 30° was not achieved. Both inner Several research programs have addressed the

door angle and slider door position provided effective central problems associated with the use of thrust remodulation of weight flow. Differential variation versers at landing-approach conditions (refs. 19 to 21,

of reverser efflux angle, inner door angle, and/or 26, 32, 33, 35 to 37, and 39). In general, thrustslider door position generated significant forces and reverser effects on airplane stability and control are

moments in addition to reverse-thrust modulation somewhat configuration dependent but may often be

capabilities. If properly integrated into an airframe, attributed to specific occurrences such as reverser

these forces and moments could augment or possibly flow blockage of the free-stream flow over tail surreplace conventional control surfaces, faces, impingement of reverser flow on empennage

surfaces, and entrainment of the free-stream flow by

Introduction reverser exhaust flow.

In addition to the stability and control problems

The design requirements for the next generation mentioned above, consideration must be given to the

of fighter aircraft may include the ability to land on possible reingestion of hot exhaust gases into the enshort or bomb-damaged runways. This capability gine inlet(s). (See ref. 34.) Reingested gases can

will probably require adding thrust-reversing capa- generate temperature and pressure distortion levels

bility to nozzle designs. The idea of thrust revers- at the compressor face which could stall the compresing after aircraft touchdown is certainly not new. sor and cause possible engine damage. Susceptibility

Commercial aircraft have been using reversers dur- to reingestion of reverser exhaust flow is a function of

ing ground roll for many years. Two European tacti- such parameters as aircraft forward airspeed, height
above the ground, inlet suction characteristics (hence figuration with thrust modulation (partial and full

local aerodynamic characteristics adjacent to the in- reverse-thrust) capabilities. The reverser design had

let), proximity of reverser port to the inlet engine four ports with equal areas. The ports were angled

weight flow, and of course, reverser effiux angle (an- outboard 30° from the vertical to impart a splay angle of the projection of reverser exhaust flow in the gle to the reverser exhaust flow. External vane boxes

normal-force/axial-force plane). The aircraft veloc- were located directly over each of the four ports to

ity at which the reingestion of reverser exhaust flow provide a variation of reverser efflux angle from 140°

occurs dictates the amount of useful reverse thrust to 26° measured forward from the horizontal referavailable for the reduction of landing ground roll. ence axis. The reverser model was tested with both a

Obviously, the longer full reverse thrust can be main- butterfly-type inner door and an internal slider door

rained during the ground roll phase, the shorter the to provide area control for each individual port. In

landing-ground-roll distance, addition, main nozzle throat area and vector angle

One method which has been shown to be very were varied to examine various methods of modulateffective in reducing the velocity at which ingestion ing thrust. Other model variables included vane box

of reverser exhaust occurs (for a given efflux angle) configuration (four or six vanes per box), orientals to splay (or cant) the reverser flow. In fact, tion of external vane boxes with respect to internal

without the ability to splay reverser flow, ingestion port walls (splay angle shims), and vane box sidespeeds can be on the order of touchdown speeds plates. Nozzle pressure ratio was varied from 2.0 to

for some configurations, as reported in reference 34. approximately 7.0.

Benefits of splaying reverser flow partially result from

increased lateral separation of the reverser jets. This Symbols

increased lateral separation tends to eliminate the All forces (with the exception of resultant gross

"fountain" effect found for many configurations with thrust) and angles are referred to the model centersplay angles of 0° once the reverser flow impinges line (body axis). A detailed discussion of the dataupon the ground, reduction and calibration procedures as well as deftReduced ground roll upon landing and improved nitions of forces, angles, and propulsion relationships

flight-path control have long been recognized as sig- used herein can be found in reference 4.

nificant advantages of thrust reversing. The problems of stability and control and engine ingestion At,main main exhaust nozzle throat area, (see

of exhaust gases are being addressed, but because fig. 2(f)), in2

of the complex nature of the nozzle requirements, F measured thrust along body axis, lbf

a reassessment of current thrust-reversing technology was warranted. This reassessment indicated that Fi ideal isentropic gross thrust,

the ability to vary reverser efflux angle to modu- _| R__,.f!___ [1_ (pa _:L_-, lbf

late thrust by a combination of main nozzle and re- Wp g'Z '_/--1 \Pt,j/

terser port area control and to splay the reverser flow

will, if properly integrated with flight and propulsion control systems, provide significant reductions FN measured normal force, lbf

in landing-field length. In addition, individual or Fr resultant gross thrust, _/F 2 + FI_ + F2y

differential variation of the nozzle variable-geometry

features (such as efflux angle and port area) could be positive for thrust vectored aft of vertical

especially useful in solving the stability and control plane and negative for thrust vectored

forward of vertical plane, lbf problems normally associated with thrust-reverser

operation at approach and landing speeds. In ad- Fr,tot individual resultant gross thrust meadition, since tail surfaces are generally sized by the sured for top left-hand reverser port only

low-speed control requirements of the aircraft, the (eq. (1)), lbf

use of propulsive controls may lead to a reduction in

the size (weight and drag) of tail surfaces or their Fy measured side force, lbf

elimination from the airframe altogether, g gravitational constant 32.174 ft/sec 2

To date, few (if any) internal performance data

are available on reverser configurations with these Ir reference length, 1.0 in.

capabilities. As a result, an investigation was con- Min t local Mach number inside nozzle

ducted at wind-off conditions in the static-test facility of the Langley 16-Foot Transonic Tunnel. The Mz measured yawing moment, in-lbf

tests were conducted on a single-engine reverser con- NPR nozzle pressure ratio, Pt,j/Pa
p local static pressure, psi sta. model station, in.

Pa ambient pressure, psi typ. typical

Pt,j jet total pressure, psi 2-D two-dimensional

R gas constant for air, 1716 ft2/sec2-°R Apparatus and Methods

Tt,j total temperature, °R

Static-Test Facility

wi ideal weight-flow rate based on nominal

throat area of 4.0 in2, lbf/sec This investigation was conducted in the statictest facility of the Langley 16-Foot Transonic Tunnel.

Wp measured weight-flow rate, lbf/sec Testing was conducted in a room with a high ceils/ ratio of specific heats, 1.3997 for air ing where the jet exhausts to atmosphere through a

large open doorway. The control room was remotely

5v geometric thrust pitch vector angle located from the test area, and a closed-circuit telemeasured from horizontal reference vision camera was used to observe the model. This

line, positive in downward direction (see facility utilized the same clean dry-air supply as that

fig. 2(f)), deg used in the 16-Foot Transonic Tunnel and a similar air-control system including valving, filters, and

0 measured reverser efftux angle,

a heat exchanger to operate the jet flow at constant

- tan-l(FN/F) for thrust directed aft stagnation temperature.

of vertical plane and 180 - tan-l(FN/F)

for thrust directed forward of vertical Single-Engine Propulsion Simulation System

plane, deg

A sketch of the single-engine air-powered nacelle

Or,Ob geometric reverser efflux angle on top model on which various nozzles were mounted is preand bottom reverser ports, respectively, sented in figure 1 with a typical nozzle configuration

measured from horizontal reference line; attached. The body shell forward of station 20.50

positive in counterclockwise direction was removed for this investigation.

on top reverser ports and in clockwise An external high-pressure air system provided a

direction on bottom reverser ports (see continuous flow of clean, dry air at a controlled temfig. 2(g)), deg perature of about 540°R and was varied up to ape resultant reverser splay angle mea- proximately 105 psi at the nozzle. This high-pressure

sured from vertical reference line, air was brought through the dolly-mounted support

tan-l(by/FN) , deg strut by six tubes, which connect to a high-pressure plenum chamber. As shown in figure 1, the air was

¢ geometric inner door angle measured then discharged perpendicularly into the model lowfrom horizontal reference line, positive in pressure plenum through eight multiholed sonic nozcounterclockwise direction on top port zles equally spaced around the high-pressure plenum.

and in clockwise direction on bottom This method was designed to minimize any forces

port (see fig. 2(a)), deg imposed by the transfer of axial momentum as the

Subscripts: air passed from the nonmetric high-pressure plenum

to the metric (mounted to the force balance) lowb bottom pressure plenum. Two flexible metal bellows were

used as seals and served to compensate for axial

l left forces caused by pressurization. The air was then

r right passed from the model low-pressure plenum (circular

in cross section) through a transition section, choke

t top plate, and instrumentation section. The transition

Abbreviations: section provided a smooth flow path for the airflow

from the round low-pressure plenum to the rectanC-D convergent-divergent gular choke plate and instrumentation section. The

conf configuration instrumentation section had a flow path width-height

ratio of 1.437 and was identical in geometry to the

dim dimension nozzle airflow entrance. The nozzles were attached to

rad. radius the instrumentation section at model station 41.13.

3
Model Description a function of the number of vanes. Four- and six-vane

configurations were tested. (See fig. 2(c).) The casNozzle geometry. The nozzle hardware down- cade vane box sideplates (fig. 2(d)) were designed to

stream of station 41.13 represented a single-engine prevent exhaust flow leakage through a slot formed

nonaxisymmetric convergent-divergent exhaust sys- between the structural support for the vanes arid the

tern with partial and full reverse-thrust capabilities, nozzle external surface. Use of the sideplates would

Sketches of the hardware tested are shown in figure 2, in effect direct all the reverser port flow through the

and photographs of full reverse-thrust configurations vanes. The reverser port tab (figs. 2(a) and 2(d))

are presented in figure 3. The reverser design con- represented a fairing that provided complete port

sisted of four ports with equal areas. The ports were closure when vanes were in a stowed position on fullangled outboard 30° from the vertical to impart a scale hardware designs. Because a region of sepadesired exhaust flow splay angle. (See fig. 2(a).) Ex- rated flow was expected on the upstream port wall

haust flow splay would direct exhaust flow away from (because of the large turning angle and sharp corner),

a single vertical-tail surface and help avoid inlet in- the upstream vane was placed slightly aft of the upgestion of exhaust flow by minimizing the fountain stream port wall in an effort to locate the vane in a

effect on the aircraft lower surface. As seen in fig- region of relatively "clean" flow. The reverser port

ure 2(a), the upstream and downstream port walls tab allows no flow to reach the upstream side of the

were parallel and were aligned 78° from the aft hor- upstream vane, and it virtually guarantees a region

izontal. External cascade vane boxes were located of separation between the corner and port tab.

directly over each of the four ports to allow variation In order to evaluate partial thrust-reversing perof reverser geometric efftux angle from 140° to 26°. formance, as would be required at landing-approach

It should be noted that when the vane boxes were lo- conditions, combinations of inner door angle were incated on each port (Splay angle = 30°), the cascade vestigated at a constant geometric efflux angle convane box sidewalls were perpendicular to the nozzle dition (t_t/Ob = 130°/130 °) with the main (normal

external surface and as a result were splayed outward forward-thrusting) exhaust nozzle flowing. Two main

only 7.6°. (See section A-A on fig. 2(a).) Because it exhaust nozzle throat areas, nominally 0.5 in2 and

was suspected that the orientation of these external 2.0 in2, and two nozzle vector angles, 5v -- 0° and

vane boxes might reduce the resultant splay angle 20°, were examined. Sketches of the main exhaust

of the individual ports, a set of splay angle shims nozzles tested are shown in figure 2(f).

(fig. 2(e)) were constructed to allow the vane box

sidewalls to be mounted parallel to the reverser port Splitter plate design. To determine the perforwalls (splayed 30°). Note that these shims did result mance of an individual reverser port/vane combinain longer port passages, tion, both force balance and weight flow measureTwo schemes for modulating port area were in- ments must be isolated from the whole reverser sysvestigated: varying the inner door angle and vary- tern. Because of model symmetry in two planes,

ing the slider door position. As shown in figure 2(a), it was believed that individual port characteristhe inner doors pivot about a longitudinal axis which tics could be obtained by making one run with

is parallel to the nozzle centerline. Values of inner both bottom ports closed (to determine normaldoor angle ranged from 60° (fully open as shown in force/axial-force characteristics) and one run with

fig. 2(a)) to 0° (fully closed for no flow through the both right-hand reverser ports closed (to determine

reverser port). The other method of modulating port side-force/axial-force characteristics). By proper

area was to vary the position of the internal slider, data-reduction techniques, isolated performance for

(See fig. 2(b).) Four slider door positions ranging the top left-hand port can be calculated. It was recfrom a fully open position to a closed position (re- ognized that the flow characteristics into a pair of reverser port area reduced by approximately 50 per- verser ports (for example, the top two ports) might

cent) were investigated. The intermediate positions, be considerably different if two of the four ports were

1/3 and 2/3, reduced the area of each individual port closed than if all ports were operating. A splitter

by approximately 17 and 33 percent, respectively, plate was utilized to alleviate this problem. Sketches

Note that the inner door remained in the fully open of the horizontal and vertical splitter plate designs

position (¢ -- 60°) when slider door positions were used during this investigation are shown in figure 4.

varied. The specific designs of these splitter plate concepts

Other reverser geometry variables were vane con- were refined on the basis of water table tests configuration (number of vanes), vane box sideplates, ducted at the University of Cincinnati, and they leaand reverser port tab. Two vane box configurations ture a corner radius on the upstream end of the splitwere tested to determine if reverser performance was ter plate. This corner radius acts as a bellmouth and

4
provides the two open reverser ports with what ap-_ the bellows when internal velocities are high and also

peared in the water table tests to be flow character: _ small differences in the forward and aft bellows spring

istics which most nearly duplicated the flow charac- constants when the bellows are pressurized. As disteristics for the reverser configuration with all ports cussed in reference 14, these bellows tares were deopen. termined by running calibration nozzles with known

performance over a range of expected normal and

Instrumentation side forces and yawing and pitching moments. The

balance data were then corrected in a manner similar

A six-component strain-gauge balance was used to that discussed in reference 14 to obtain final forces

to measure the forces and moments on the model and moments. The resultant gross thrust Fr, used in

downstream of station 20.50. (See fig. 1.) Jet the resultant thrust ratio Fr/Fi, was then determined

total pressure was measured at a fixed station in the from these corrected balance data as were the indiinstrumentation section (see fig. 1) by means of a vidual force and moment ratios FN/Fi, Fy/Fi, and

four-probe rake through the upper surface, a three- Mz/Filr. Significant differences between Fr/F i and

probe rake through the side, and a three-probe rake F/F i occur when jet-exhaust flow is directed away

through the corner. A thermocouple, also located from the axial direction. The individual force rain the instrumentation section, was used to measure tios are presented to allow a direct comparison of

jet total temperature. Flow rate of the high_pressure normal- and side-force magnitudes relative to axialair supplied to the nozzle was measured by a pair of force (thrust ratio) values.

critical flow venturis. Internal static-pressure orifices

were located on the nozzle upper flap upstream of Nozzle discharge coefficient wp/w i is the ratio

the reverser ports, in the top left-hand reverser port, of measured weight-flow rate to ideal weight-flow

and on the nozzle upper and lower flaps. Results from rate, where ideal weight-flow rate is based on jet

these static-pressure orifices (because of the limited total pressure Pt,j, jet total temperature Tt,j, and

number) were used primarily as a diagnostic tool and a nominal throat area of 4.0 in2. Since a nominal

as supporting information to the balance force and throat area (representing a scaled throat area for

moment measurements; hence, exact orifice locations an intermediate-power main nozzle) was used in the

are not defined herein. However, whenever static- computation of the discharge coefficient instead of

pressure data are plotted, the figures contain sketches actual values of throat area, variations in Wp/Wi

showing relative orifice locations, occurring as a result of configuration changes reflect

not only changes in the ability of the nozzle to pass

Data Reduction weight flow but also differences in throat area.

Data from three separate runs were used to cal- All data were recorded simultaneously on magnetic tape. Approximately 50 frames of data, taken culate the performance characteristics for a single

at a rate of 10 frames per second, were used for each port (top left-hand port). Data from complete condata point; average values were used in computa- figuration runs (all ports open) combined with data

tions. With the exception of resultant gross thrust from the horizontal (only top left and top right ports

Fr, all force data in this report are referenced to the open) and vertical (only left-side top and bottom

body axis. ports open) splitter plate runs were used to deterThe basic performance parameters used for the mine the performance for the top left-hand reverser

presentation of results were F/Fi, Fr/Fi, FN/Fi, port alone. Basic to the following analysis technique

Fy/Fi, Mz/Filr, and Wp/W i. The internal thrust is the assumption that each individual reverser port

ratio F/F i is the ratio of actual nozzle thrust (along had identical performance to the other three ports.

the body axis) to ideal nozzle thrust, where ideal Based on reverser geometry symmetry as shown in

nozzle thrust is based on measured weight-flow rate figure 2, this assumption should be valid.

and total temperature and pressure conditions in As a result of the above assumption, the thrust

the nozzle throat as defined in the symbols. The ratio F/F i (or the axial-force component) generated

balance axial-force measurement, from which actual by a single reverser port should be similar to the

nozzle thrust is subsequently obtained, is initially thrust ratio measured for either the complete configcorrected for model weight tares and balance inter- uration or for any combination of reverser ports open

actions. Although the bellows arrangement was de- or closed. In reality, however, small differences in insigned to eliminate pressure and momentum inter- dividual port geometry (port throat area, for examactions with the balance, small bellows tares on all ple) resulted in values of thrust ratio for the splitter

balance components still exist. These tares result plate runs which were not identical to those measured

from a small pressure difference between the ends of for the complete configuration. These differences in
F/F i were generally small (within 0.02). It was be- thrust ratio (axial-force component) even though all

lieved that averaging thrust ratios (at a given NPR) the flow has been turned from the axial direction.

for all three combinations of reverser ports tested On the other hand, resultant thrust of a single port,

would provide a thrust ratio more representative of which is computed by isolating the normal-, axial-,

the individual port. Ratios of normal-force and side- and side-force components, is a measure of the true

force components to ideal thrust were obtained di- efficiency of the reverser ports and is useful for assessrectly from the horizontal and vertical splitter plate ing the nozzle system performance losses associated

data, respectively. The total resultant thrust ratio with reversing the exhaust flow.

was then defined to be The individual force components for the single

port were then used to compute reverser splay angle

Fr,tot -- ] ( "_F 2 -I- (FN_ 2 + (Fy'_ 2 (1) ¢ from the equation

where ¢ = tan-1 [ --NN FY [ (5)

t,l = (rlri)+(r/ri)t 3 +(rlrd (2) The the sign absolute of thevalue splayofangle ry/r(determined N wasusedbybecause which

port is open) was not deemed important; only the

magnitude was important. The reverser effiux angle

in which (F/Fi)t, l represents the average of the 0 was determined from the following two equations.

thrust ratios obtained from runs without a splitter For configurations with exhaust directed aft of the

plate, with a horizontal splitter plate, and with a vertical plane, vertical splitter plate;

(_N.) (_.) (3) O=--tan-lFN

= F (6)

t,l t

For configurations with exhaust directed forward of

which is obtained from the horizontal splitter plate the vertical plane,

run; and

(_.) t,t=(_.) (4) O=lSO_tan -1F-NV (7)

which is obtained from the vertical splitter plate

run. Note that the above definitions require the Presentation of Results

assumption, as discussed previously, that weight flow

and flow conditions in each port are identical. The basic performance data for the current invesThe resultant thrust of a single reverser port tigation are plotted in figure 5. An index to these

Fr,tot was computed to present a more complete pic- basic data and sketches of model geometry can be

ture of actual flow conditions in the reverser ports, found in table I and figure 2, respectively.

This was necessitated by the fact that resultant Performance comparisons and summary figures

thrust, whether computed for the complete reverser are presented as follows:

or for splitter plate runs where two ports were closed, Figure

always had a net normal-force and/or side-force bal- Effect of geometric effiux angle on--

ance component equal to zero. Thus, some of the Thrust ratio .............. 6

force and flow direction information was lost. For ex- Internal static-pressure ratio ....... 7

ample, consider the complete configuration operating Effect of splay angle shims on--

with all ports open and all port variables equal. Be- Thrust ratio .............. 6

cause of the symmetry of the reverser port arrange- Internal static-pressure ratio ....... 7

ment, both the net normal- and side-force compo- Effect of differential geometric effiux angle on

nents measured by the balance are approximately thrust ratio, normal-force ratio, and

zero. The upward normal force generated by the yawing-moment ratio .......... 8

lower ports is cancelled by the downward normal Effect of inner door angle on--

force generated by the upper ports. Likewise, the Thrust ratio for 0t/0 b 2 130°/130°

symmetry in splaying the exhaust flow outboard re- and At,main = 0.0 in ......... 9

sults in zero net side force. The resultant thrust ra- Thrust ratio for Ot/O b ---- 130°/130 °

tio for this configuration is, therefore, equal to the and At,main = 2.0 in2 .......... 10

6
Nozzle discharge coefficient for zle was pitch vectored. It should be noted that Fr/Fi

Or/Oh= 130°/130 ° .......... 11 is not provided for the splitter plate configurations

Internal static-pressure ratio for with Ot/Ob -- 90°/90 ° (figs. 5(g) and 5(h)) because a

Ot/Ob = 130°/130 ° .......... 12 large change in the vertical scale would have been reEffect of slider door position on-- quired to present both Fr/F i and F/F i on the same

Thrust ratio .............. 13 figure. In each case, however, the resultant thrust

Thrust ratio and nozzle discharge coefficient, ratio is almost identical (within 0.005) to measured

summary at NPR = 3.0 ......... 14 values of Fy/F i and FN/F i for the vertical and hotInternal static-pressure ratio ....... 15 izontal splitter plate configurations, respectively.

Effect of inner door angle and slider door Configuration 2 was tested at periodic intervals

position on thrust and normal-force during the test program, and these data are preratios ................. 16 seated in figure 5(b) as a data band to indicate data

Effect of vane box configuration on repeatability. As seen in figure 5(b), the data are

thrust ratio .............. 17 scattered over a band of approximately 1 percent.

Effect of port tab on-- An average of these data was used for the comparThrust ratio .............. 18 isons presented later in the paper.

Internal static-pressure ratio ....... 19 The internal performance data shown in figure 5

Effect of vane box sideplates on exhibit characteristics typical of other convergentthrust ratio .............. 20 divergent nozzles (ref. 14), in that performance inEffect of thrust modulation on creases up to a maximum (usually occurring at the

thrust ratio .............. 21 NPR for fully expanded flow) and thereafter deEffect of main nozzle vector angle on creases throughout the NPR range. It should be

thrust ratio .............. 22 noted that a decrease in reverser performance (conComponent breakdown of resultant figurations in which exhaust is directed forward) is

flow vector for top left-hand port; distinguished by an increase in thrust ratio. RememOt/tgb = 130°/130 ° ........... 23 ber that large negative values of F/F i indicate good

Effect of geometric effiux angle and inner thrust-reverser performance.

door angle on reverser splay angle .... 24 There appear to be two exceptions to the above

Summary of geometric efflux angle and trends. For example, consider configuration 11

main nozzle throat area effects on (fig. 5(r)) in which the inner door angle is 10% Perforthrust ratio; NPR -- 3.0 ......... 25 mance of this reverse-thrust configuration generally

Summary of inner door angle and main nozzle increased (became more negative) throughout the

throat area effects on thrust ratio; range of NPR. As will be shown later, the physical

NPR = 3.0 .............. 26 throat moved upstream from the vane boxes to the

inner door. This movement of the throat resulted in a

Results and Discussion large decrease in actual port area and large changes

in the expansion ratio (ratio of exit area to throat

Basic Data area) and in the design NPR. Apparently the expanBasic data for all reverser configurations tested sion ratio change was large enough so that design

are presented in figure 5. Nozzle internal thrust ratio NPR was above 7 and was never reached.

F/F i and discharge coefficient Wp/Wi are presented The approach configurations in which At,main =

as a function of NPR for each configuration. Resul- 2.0 (see figs. 5(nn) to 5(pp)) indicate trends which

tant thrust ratio Fr/F i is generally shown for those also appear to be different from the general case.

configurations in which values of Fr/F i differed from These data show that the general character (shape of

values of F/F i (at a constant NPR) by 0.005 or more. curve, NPR at which maximum and minimum perWhen Fr/F i is presented, the ratios of normal force formance occurs, etc.) of each performance curve

FN and side force Fy to ideal thrust Fi are usually is a function of inner door angle; however, the abprovided to allow the reader to determine the reason solute level of each curve is dependent upon the

Fr/Fi differs from F/F i. The above condition gen- percentage of the total weight flow being passed

erally resulted for configurations in which two of the through the main exhaust nozzle. For example,

reverser ports were closed (splitter plate runs); for consider configuration 33 (fig. 5(nn)) with 0t/0 b =

configurations in which model variables such as geo- 130°/130°, Ct/¢b = 60°/60 °, and At,main = 2.0 in2.

metric effiux angle, inner door angle, or slider door When compared with configuration 2 (same reverser

position were differentially configured; and for ap- vane angles and inner door angles, At,main = 0.0 in2),

proach configurations in which the main exhaust noz- it can be seen that the performance curves are very
similar in shape. The absolute levels of thrust ra- Reverser geometric effiux angle. The effects

tio indicate a considerable amount of forward thrust of reverser geometric efflux angle on thrust ratio

(positive F/Fi) is being generated by the main ex- and on internal static-pressure ratio are presented

haust nozzle (configuration 33). In fact, at NPR = in figures 6 to 8. Thrust ratio data are presented

3.75 the forward thrust generated by the main noz- for configurations in which reverser geometric efflux

zle exactly cancels the reverse thrust generated by angle was the same for all reverser ports (symmetrical

the reverser ports; hence, F/F i = 0. effiux angles) and for configurations in which reverser

Values of nozzle discharge coefficient presented geometric efflux angle was set asymmetrically to

in figure 5 range from 0.20 to approximately 1.37. generate a particular force or moment (differential

Again, recall that the ideal weight flow wi is based on efflux angles).

a constant throat area of 4.0 in2. Hence, variations Data for the configurations with symmetrical efin nozzle discharge coefficient between configurations flux angles are presented in figure 6. For the configcan reflect physical throat area changes as well as uration with no splay angle shims, values of F/F i

changes in the ability of the throat to pass weight range from -0.70 to 0.73 and depend, of course,

flow. on values of geometric effiux angle Ot/Ob. As can

Attention should also be given to configurations be seen, variation of reverser effiux angle is a very

in which F/F i (hence measured axial force) is nearly powerful thrust modulator. Thrust ratio F/F i for

equal to zero, such as for 0t/0 b = 90°/90 ° (fig. 5(f)). cascade vanes removed (configuration 13) fell beThe variation in Fr/F i (dashed line) occurring for tween performance levels for Ot/Ob = 90°/90 ° and

NPR > 5 is misleading and is a result of the change Ot/Ob -- 66°/66 °. This result was expected, since

in sign convention for Fr/F i when the sign of F/F i the upstream and downstream walls of the reverser

changes as noted in the symbols section. As shown in port were aligned 78° from the horizontal, as noted

figure 5(f), neither normal-force nor side-force ratio in figure 2(a). Peak reverser performance, whether

would indicate such a large change in the magnitude for spoiled (0 < 90°) or reverse (0 > 90°) thrust,

of the resultant thrust ratio, generally occurred near NPR -- 3 and decreased as

The splitter plate configurations were tested to NPR increased.

isolate the internal performance characteristics of the The exceptions to this were for Ot/Ob -- 90°/90 °

top left-hand port. This information provides hard- and 66°/66 °. For Ot/O b = 90°/90 °, thrust along

ware designers with more information on loads and the body axis F is equal to zero, and thus the

load directions than would otherwise be obtained variation in F/F i is very small. For the configuby testing only the complete configuration. Splitter ration with Ot/Ob = 66°/66 °, thrust ratio tended

plate data were obtained only for selected configura- to increase slightly (approximately 3 percent of Fi)

tions as indicated in table I. with increasing NPR until a performance peak ocAs discussed previously, much of the analysis of curred at an NPR of about 5.0. It should be noted

the splitter plate data was based on the fact that that these two configurations (Ot/Ob = 90°/90 ° and

the reverser configurations have geometric symmetry 66°/66 °) require less flow turning from the cascade

about the centerline vertical and horizontal planes, vanes than the other geometric reverser efflux angles

Because of this symmetry, normal-force and side- tested. This result is supported by the relative level

force values equal to zero were expected from vertical of the data obtained for configurations with the casand horizontal splitter plate runs, respectively. As cade vanes removed.

shown in figures 5(c) and 5(d), this was generally Characteristic of all the effiux angle variations exnot the case. It is believed that the nonzero values cept t_t/gb = 90°/90 ° and 66°/66 °, discussed previof normal force (vertical splitter plate, fig. 5(c)) and ously, is the relatively large impact of NPR on measide force (horizontal splitter plate, fig. 5(d)) are the sured thrust ratios. As shown in figure 6, this variresult of small differences in port geometry, ation in thrust ratio tended to increase as geometric

The measured internal performance characteris- effiux angle increased or decreased from 90°/90 ° and

tics of various reverser port configurations deter- 66°/66 °. Examination of figures 2(c) and 2(d) will

mined by the splitter plate technique are intended lend some insight as to why these trends occur. Each

to be more qualitative in nature than quantitative, passage between the turning vanes forms a discrete

single-expansion-ramp nozzle (one solid jet boundary

Performance Comparisons and one free jet boundary downstream of exhaust

Unless otherwise noted, all configurations have exit). As noted in reference 23, nozzle configurations

the four-vane external cascade boxes, port tabs in- with an external expansion surface generally have a

stalled, no cascade vane box sideplates, and no splay tendency to vector or turn the exhaust flow. The

angle shims installed, direction of the resultant vector is a function of the

8
NPR and the orientation of the external expansion Port orifice locations, represented by circle and

surface with respect to the model. At values of NPR square symbols in figure 7(b), appear to have lower

greater than required for fully expanded flow (un- pressure ratios on the upstream reverser port wall

derexpanded nozzle), the exhaust tends to be vec- than on the downstream wall. These lower pressure

tored away from the external expansion surface. For ratios indicate that the reverser flow is being acceler-

(_t/O b > 90°/90 °, a loss in reverse-thrust performance ated around the upstream port corner. As indicated

would be expected at underexpanded flow conditions by the decreasing pressure ratios, the reverser ex-

(NPR > 3.0) because the aft-facing expansion surface haust flow is continuously accelerating through the

would tend to vector exhaust flow back toward the port. However, the fact that the exhaust flow velocvertical. When effiux angle is such that the external ity is not sonic (P/Pt,j <- 0.528) at the last pressure

expansion surface is forward facing, Ot/O b < 90°/90 °, orifice indicates that the port physical throat probalosses in forward thrust would be expected at under- bly forms in the cascade vane box located on top of

expanded flow conditions. These trends can be seen each port.

in figure 6. As geometric effiux angle increases from The effect of splay angle shims on reverser port

26° to 90° or decreases from 140° to 90°, the exposed static-pressure ratios is presented in figure 7(c). As

length of the external expansion surface decreases, shown, static-pressure ratio decreased with the inand the performance losses resulting from flow turn- stallation of splay angle shims. Again, an increase

ing as discussed above would be expected to decrease, in discharge coefficient (see fig. 5) accompanied this

This trend is also exhibited by the data in figure 6. decrease in static-pressure ratio in the ports.

Addition of the splay angle shims (for shim ge- In an effort to assess the usefulness of this reometry, see fig. 2(e)) reduced levels of both reverse verser configuration for providing forces and mothrust (Ot/O b > 90°/90 °) and spoiled thrust (Ot/O b < ments for control, differential geometric effiux an90°/90 °) and had little effect on thrust ratio for gle settings for generating both normal force and

Ot/Ob = 90°/90 °, as shown in figure 6. Flow visu- yawing moment were investigated. Normal force

alization observations indicated that the splay angle was generated by setting geometric efflux angles to

shims significantly increased the outward cant of ex- Ot/0b = 90°/130 °. (See fig. 8(a).) A resulting downhaust flow from each port and correspondingly de- ward normal force was generated. As expected, levcreased the effiux angle 0 as the thrust ratio data els of reverse-thrust ratio obtained with this differindicate. This observation indicates that part of the ential effiux angle configuration were approximately

axial-force component is being traded for an increase an average of the thrust ratios obtained for symmetin the side-force component. The overall character of rical efftux angle configurations Ot/Ob= 90°/90 ° and

the curves with respect to NPR remains nearly the Ot/theatb = 130°/130 °. The normal-force contribusame, as discussed previously for the configurations tion to the resultant thrust ratio was equal to or

with no splay angle shims, greater than the axial-force contribution throughout

Internal static-pressure ratios from orifices lo- the NPR range tested.

cated in the reverser port and in the nozzle upstream The configuration tested to provide a yawing moof the reverser port are presented for symmetrical ge- ment consisted of differential efflux angle settings of

ometric efflux angle configurations without splay an- 130° on the left-hand ports and 26° on the rightgle shims in figures 7(a) and 7(b). Pressure data are hand ports (OJOr = 130°/26°). Thrust and yawingpresented at NPR = 3.0 only. The data indicate that moment ratio results in figure 8(b) show that sizable

values of static-pressure ratio decreased and internal yawing moments can be generated by differential leftMach number increased as _ approached a value of hand and right-hand geometric effiux angles. The ba78°, which corresponds to the physical reverser port sic data for this differential effiux angle configuration

wall angle as shown in figure 2(a). Examination of (fig. 5(rr)) also show that this yawing moment was

nozzle discharge coefficient data Wp/Wi presented in generated with no side force. These differential elfigure 5 indicates that this increase in internal Mach flux angle configurations demonstrate that a reverser

number in the nozzle and reverser port was accompa- which has the capability of independent control of

nied by an increase in nozzle discharge coefficient. As efftux angle and port area can provide control forces

geometric effiux angle diverged from 78°, discharge to augment or replace conventional tail surfaces or

coefficient decreased, canards.

Individual pressure taps located in the nozzle

upstream of the reverser ports (fig. 7(a)) indicated Inner door angle. The effects of inner door

that flow was being accelerated to higher speeds and, angle on thrust ratio, nozzle discharge coefficient,

hence, lower static-pressure ratios as it approached and internal static-pressure ratio are presented in

the portopening, regardless ofgeometriceffluxangle, figures 9 to 12. Data are shown at values of
At,main = 0.0 in2 (full reverse) and 2.0 in2 (landing Exhaust flow passing through the main exhaust nozapproach), zle did result in acceleration of the flow in the nozzle

As discussed previously, inner door angle was in- upstream of the ports, as shown in figure 12(a). This

tended to provide weight flow (or area) control of acceleration would be expected based solely on the

the individual reverser ports. As noted in figure 9, increase in total throat area resulting from the addihowever, inner door angle also had a large impact tion of the main exhaust nozzle.

on thrust ratio (especially for NPR < 5). Thrust ra- Another area of concern to nozzle designers would

tio increased (hence reverser performance decreased) be whether or not inner door angle variations have

as the inner door angle was changed from the open any effect on pressure distributions within the main

position (60°) to the most closed position (10°). In- nozzle passage. As shown in figure 12(d), inner

sight into the reasons for these losses in peak reverse- door angle (hence the modulation of weight flow)

thrust ratio can be gained by examination of the had no effect on internal static-pressure ratio on the

port static-pressure ratios presented in figure 12(b). nozzle flap centerline. Unfortunately, no pressures

The decreased static pressures for ¢ < 40° (NPR = were measured at locations other than the nozzle flap

3.0) indicate increased velocities in the port. These eenterline (directly behind the port, for example).

increased velocities in the port, of course, mean A summary of the effect of inner door angle on

increased velocities entering the cascade vane box. nozzle discharge coefficient is presented in figure 11.

Vane turning losses resulting from separation on the The modulation of weight flow was the primary runeleeward side of each vane would be expected to in- tion for the variable-angle inner door and, as shown

crease as velocities in the port increased. (Individual in figure 11, a smooth and effective modulation of

vanes are at a high angle of attack with respect to weight flow results from varying inner door angle. In

the port flow.) combination with the variation of main nozzle throat

The performance data for the inner door angles area, inner door angle was shown to be capable of

of 10° and 20° (fig. 9) have, in general, different providing a wide range in values of nozzle discharge

characteristics than the performance data for the coefficient. It should again be stressed that ideal

larger inner door angles, ¢ > 20°. Examination of weight flow rate wi is based on a nominal constant throat area of 4.0 in2.

port static pressures (fig. 12(b)) indicates supersonic

flow (P/Pt,j <-0.528) in the ports for these two inner

door angle configurations. The thrust-reverser throat Slider door position. The slider door shown

(Min t = 1.0)is apparently formed between the inner in figure 2(b) provides an alternative method of

door and the port passage walls when the inner doors port area control to the inner door. As shown in

figure 13, slider door position did not have as large are nearly closed (¢ = 10° or 20°); hence, large vaneturning losses would be expected in the presence of an effect on thrust ratio as inner door angle had

this supersonic flow. for Ot/Ob = 130°/130 °. (See fig. 10.) An even smaller effect on thrust ratio can be noted for the

The effect of inner door angle on thrust ratio for reverser configuration with Ot/O b : 90°/90 °. A

the reverser in an approach configuration (At,main = summary of slider door position effects on thrust

2.0 in2) is shown in figure 10. Inner door angle ratio and nozzle discharge coefficient is presented in

variations resulted in large changes in thrust for both figure 14. In general, reverser performance tended to

0° and 20° vectoring configurations. The inner doors decrease (become less negative) as the slider door was

are obviously very effective in modulating weight flow closed. The closed-slider-door position resulted in a

between the main nozzle and reverser ports. 50-percent reduction in port area. As seen, discharge

Internal port static-pressure ratios for the ap- coefficient for the slider door in the closed position

proach configuration at 5v : 0° are shown in fig- was approximately one-half that for the slider door

ure 12(c) and indicate, as discussed previously for in the open position.

At,main = 0.0 in2, that port static pressures de- Reverser upper flap and port static pressures are

creased as inner door angle decreased. Port pressures presented in figure 15. As was the case any time

for the approach configurations were essentially the throat area was reduced, static pressures on the nozsame on the downstream wall and slightly lower on zle flap upstream of the port increased, and the corthe upstream wall for ¢ > 20° than for the'full re- responding internal Mach number decreased. Presverse configuration, At,main = 0.0 in2 (fig. 12(b)). sure ratio trends in the port (fig. 15(5)) were, howPort pressures for those configurations in which inner ever, dependent upon orifice location. The orifices

door angle was < 20° did not generally change, prob- located directly upstream of the slider door (circle

ably because flow is becoming supersonic upstream of symbols) show increased pressure ratios as the slider

the pressure measurements, as discussed previously, door is closed. In fact, in the closed position, the

10
orifices represented by circle symbols indicate that Port tab. The effect of the port tab (see fig. 2(d))

total pressure is being measured. The only other gen- on thrust ratio and reverser static pressures is preeral trend that can be noted is that when the slider sented in figures 18 and 19, respectively. As seen in

door is closed and the exhaust flow passes between figure 18, removal of the reverser port tab had little

the outboard wall of the reverser port and the inner or no effect on thrust ratio for values of NPR < 3.0.

door, pressures measured by the remaining orifices However, for NPR > 3.0, reverse-thrust performance

were generally lower than for any other slider door was significantly higher (more negative) for the conposition. It is obvious that the flow is being accel- figuration with the port tab removed.

crated in the outboard portion of the reverser port Reasons for this higher performance can best be

because the inboard passage is blocked, explained by referring to the sketch of the vane box

Because the slider door method of weight-flow and port tab in figure 2(d). With the port tab incontrol affected thrust ratio much less than did inner stalled, the vanes form three distinct nozzle passages.

door angle for a similar variation in discharge coeffi- As discussed previously, all three passages have an

cient, it in all probability would be considered a more aft-facing external expansion surface which is somedesirable method of reverser port area control, what larger for the effiux angle of 130° than for any

other efflux angles except 140° and 26°. As noted in

Asymmetric port area modulation. As discussed reference 23, nozzle configurations with an external

previously, asymmetric effiux angle configurations expansion surface generally have a tendency to vecwere capable of generating substantial normal-force tor exhaust flow away from the external expansion

and yawing-moment components. Two methods of surface once the NPR is above that required for fully

modulating the magnitudes of these forces and mo- expanded flow on the expansion surface. Reversements are (1) the use of differential inner door an- thrust losses (exhaust flow turned more aft) would,

gle and (2) the use of differential slider door posi- therefore, be expected for the configurations with the

tion. The effects of differential inner door angle and geometric effiux angles of 130° at the higher NPR

slider door position on thrust and normal-force ratios values, for which exhaust flow is underexpanded.

for an asymmetric geometric efi_ux angle configura- With the port tab removed, an additional noztion (Ot/Ob = 90°/130 °) are presented in figure 16. zle passage is created. This passage would exhibit

Both methods of port area modulation provided large external expansion surface flow characteristics simincreases in normal-force ratio (increases on the or- ilar to those discussed above, except that the most

der of 70 to 115 percent) over the symmetrical inner upstream vane provides an external expansion surdoor/slider door configuration (data represented by face which is forward facing. The external expansion

circles). Values of normal force were approximately surface for this passage should help turn the exhaust

3 to 4 times larger than the axial-force component flow upstream (in the reverse-thrust direction) at the

(F/Fi)" higher NPR values. The result is a net decrease in

It appears that by independent actuation of ge- the underexpanded flow turning associated with the

ometric efflux angle, inner door angle, and/or slider aft-facing expansion surfaces of the aft three passages

door position, the potential for tailoring individual (and a net increase in reverse-thrust performance, as

forces and moments generated by the propulsion sys- shown in fig. 18).

tem would be nearly limitless. Such a capability The static pressures on the flap upstream of the

could be used for low-speed control augmentation reverser ports and in the reverser ports (fig. 19)were

during landing approach, reduced when the port tab was removed. These reductions in static pressure are indicative of an inCascade vane box configuration. The effect of cos- crease in throat area. Examination of discharge cocode vane box configuration on thrust ratio for the efficients presented in figures 5(b) and 5(hh) confirms

reverser configuration with Ot/Ob= 130°/130 ° is pre- this.

sented in figure 17. As shown, the six-vane configuration provided higher levels of reverse thrust than the Cascade vane box sideplates. The effect of cosfour-vane configuration throughout the NPR range code vane box sideplates on thrust ratio is presented

tested. The improvement in reverse-thrust perfor- in figure 20. The reverser configuration with the vane

mance is believed to be a direct result of decreas- box sideplates installed provided higher levels of reing the spacing between the vanes. Decreasing vane verse thrust than those for the same configuration

spacing reduces the length of the external expansion without sideplates. The need for the sideplate was

surface available for the exhaust flow to act upon identified by means of flow visualization techniques.

and thus reduces the adverse effect on reverser per- It was noted during the test that high-energy flow

formance discussed previously, was leaking out of the slot formed between the noz11
zle boattail surface and the structural support for between F/F i for the configuration with 5v = 0°

the vanes. (See fig. 2(d).) This leakage did not con- and Fr/F i for the configuration with 5v = 20° are

tribute to the reverse-thrust component because it a measure of the internal performance losses assowas directed to the side. When the sideplates were ciated with turning (pitch thrust vectoring) the exadded, all the exhaust flow was then directed through haust flow. As can be seen, these losses were less

the vanes and therefore contributed to the reverse- than 1 percent for the forward-thrust-mode (reverserthrust component, ports-closed) configuration (fig. 22(a)).

Examination of nozzle discharge coefficient data Increases in inner door angle resulted in decreases

for these two configurations (figs. 5(5) and 5(ii)) in thrust ratio (as well as resultant thrust ratio), as

shows that there is a 3- to 4-percent reduction in noted previously, since a larger percentage of the

discharge coefficient resulting from the addition of exhaust flow is being directed out of the reverser

sideplates. This observation lends further credibility ports.

to the hypothesis that the physical throat of many

of these reverser configurations (especially when the Splay angle. One of the primary objectives of

inner doors are fully open) is located in the vane box. the present reverser design was to provide some outboard cant (splay) of exhaust flow. A splay angle

Main exhaust nozzle throat area. Under actual of 30° was originally desired to help prevent hot gas

operating conditions, modulation of thrust ratio will impingement on tail surfaces and reingestion of exin all probability be accompanied by simultaneous haust flow by the inlet on a fighter configuration. A

variation of reverser port area and main exhaust component breakdown of the various forces measured

nozzle area. Port area will open as main nozzle area and the relationships used to compute reverser splay

is closed. Thus, effective throat area will be held angle are illustrated in figure 23 for configuration 2

constant to prevent disruptions in engine operation (Ot/Ob ---- 130°/130°, _)t/¢b = 60°/60 °, slider door

resulting in stall or overspeed, open). Data are presented at an NPR of 3.0. Note

The effect of varying main exhaust nozzle throat that the signs of all forces presented were reversed in

area on thrust ratio is presented in figure 21. For an effort to provide graphical representation of the

all configurations in which the reverser ports were reverser flow direction instead of the reaction thrust

open, geometric efflux angle was held constant at vectors. As discussed in the "Data Reduction" sec0t/0 b = 130°/130 ° and the inner door and slider tion, this analysis was based on measurements made

doors were fully open. during splitter plate runs as well as complete conAs shown in figure 21, values of thrust ratio figuration runs. Splay angle was computed to be

ranged from 0.97 for the main nozzle only to -0.55 10.8°, a value considerably below the desired splay

for the reversers only. Depending on main nozzle angle of 30°. The measured emux angle of 135.5° inthroat area, thrust ratio varied from 0.07 to -0.55 dicates that reverser flow for the configuration with

with reversers operating. As certainly would be ex- a geometric effiux angle of 130° was slightly overpected, decreasing the main nozzle throat area re- turned. This overturning of the exhaust flow may be

sulted in a greater percentage of the exhaust flow the result of the external-expansion-ramp characterbeing directed to the reverser ports and, hence, an istic of each individual nozzle passage operating at

increase in reverser performance (larger negative val- overexpanded flow conditions. External-expansionues of F/Fi). ramp nozzles operating at values of NPR below those

for fully expanded flow on the external surface often

Main nozzle pitch vector angle. The effect tend to vector exhaust flow toward the external exof main nozzle pitch vector angle on thrust ra- pansion surface (ref. 23). Hence for the reverser contio is presented in figure 22 for forward-thrust- figuration tested, the flow is turned even more than

mode (reverser-ports-closed) and approach (com- the geometric effiux angle.

bined main nozzle/reverser) configurations. Data are The splitter plate data technique was used to depresented for several values of inner door angle. With termine the effects of geometric efflux angle and inner

the main exhaust nozzle alone (reverser ports closed), door angle on reverser splay angle, and the results are

significant differences between thrust ratio and resul- presented in figure 24. In all cases, splay angle was

tant thrust ratio, shown by the dashed line, occur for less than the desired value of 30°. In fact, the reverser

the configuration with 5v -- 20°. These differences are with a geometric effiux angle of 130° provided splay

a function of the resultant thrust-vector angle and oc- angles less than 12° throughout the NPR range. One

cur because jet exhaust flow is directed away from the reason the desired splay angles were not attained is

axial direction. For the forward-thrust-mode configu- believed to be the vane box orientation (in the splay

ration at 5v = 0°; F/Fi is equal to Fr/F i. Differences direction) with respect to the reverser port wall angle
as shown in figure 2(a). The port wails were aligned that Ot/Ob = 0°/0 ° represents the reverser-portsat 30° from the vertical; however, the cascade vane closed (or forward-thrust) configuration. The effects

box sidewalls were only canted 7.6°. Flow is being of main nozzle throat area on F/F i were large for an

directed back toward the vertical by the vane box efflux angle of 130° and, as indicated by the dashed

sidewalls. Had the sidewalls been aligned parallel to lines, would be expected to decrease as efftux angle

the port walls, larger splay angles would probably decreased. This decrease is expected, since a decrease

have resulted. This hypothesis is supported by the in efflux angle results in reverser flow being directed

data shown in figure 24 for the configuration with more toward the aft direction coincident with the

Ot/Ob = 130°/130 ° and cascade vane box sideplates main exhaust nozzle flow.

installed and is discussed later. With no cascade vane A summary of inner door angle and main nozbox sideplates, splay angle increased as geometric ef- zle throat area effects on thrust ratio for Ot/Ob =

flux angle decreased from 130° to 26°. Comparison of 130°/130 ° is shown in figure 26 at NPR = 3.0. Again

the individual normal- and side-force components for a very large variation in thrust ratio can be generthe efflux angle of 130° (figs. 5(c) and 5(d)) and the ated by combined use of inner door angle and main

effiux angle of 26° (figs. 5(1) and 5(m)) points out the nozzle throat area. Thrust ratios ranging from 0.58

interdependency of splay angle on effiux angle. The to -0.55 were generated at a fixed geometric efflux

side-force components for both configurations were angle. The effect of nozzle throat area on F/F i innearly equal (figs. 5(c) and 5(1)), but since the geo- creased as inner door angle decreased. This result

metric efl=luxangle of 26° turned flow back toward the would be expected on the basis of the relatively large

horizontal more than the 130° angle, normal-force impact of inner door angle (for ¢ < 20°) on thrust

component for the efflux angle of 26° was consider- ratio, as noted previously. (See fig. 9.)

ably smaller (figs. 5(d)and 5(m)); hence, the resul- Summary figures 25 and 26 indicate that the

tant splay angle was significantly larger, nozzle reverser design of the present investigation

With cascade vane box sideplates installed on the has the capability of generat!ng any required value

reverser configuration with an efftux angle of 130°, of thrust ratio from 0.97 (for main nozzles only) to

a further reduction in splay angle was measured, as almost -0.70 (for full reverse). This variation of

can be seen in figure 24. As discussed previously, thrust ratio can be obtained by various combinations

cascade vane box sideplates reduce lateral leakage of geometric efflux angle, inner door angle, and main

and direct more flow over the vanes. Both normal nozzle throat area.

force and axial force increase relative to side force

for the sideplates-installed configuration (compare Conclusiolls

figs. 5(c) and 5(d) with 5(jj) and 5(kk)); hence,

the reverser configuration with sideplates produces An investigation has been conducted at wind-off

a smaller splay angle than the reverser configuration conditions in the static-test facility of the Langley

without sideplates. 16-Foot Transonic Tunnel. The tests were conducted

Variation of inner door angle for the reverser with on a single-engine reverser configuration with par-

_t/Ob = 130°/130 ° and had little or no effect on tial (thrust modulation) and full reverse-thrust casplay angle for the inner door angles presented. (See pabilities. The reverser design had four ports with

fig. 24.) equal areas. The ports were angled outboard 30°

Results of the splay angle analysis highlight the from the vertical to impart a splay angle to the reinterrelationship between efftux angle and splay an- verser exhaust flow. This splay angle was intended

gle. In order to meet both a splay angle and an efflux to prevent impingement of exhaust flow on empenangle requirement simultaneously, exhaust flow must nage surfaces and to help avoid inlet reingestion of

be overturned more than the desired angles in both exhaust gas when the reverser is integrated into an

directions, actual airplane configuration. External vane boxes

were located directly over each of the four ports to

Summary Figures provide variation of reverser effiux angle from 26° to

140° (measured forward from the horizontal reference

A summary plot of geometric efflux angle and axis). The reverser model was tested with both a

main nozzle throat area effects on thrust ratio for butterfly-type inner door and an internal slider door

an inner door angle setting of 60°/60 ° is presented to provide area control for each individual port. In

in figure 25 at NPR -- 3.0. Combined use of efflux addition, main nozzle throat area and vector angle

angle and main nozzle throat area provides very were varied to examine various methods of modueffective modulation of thrust ratio. Values of F/F i lating thrust levels. Other model variables included

ranged from 0.97 to approximately -0.70. Note vane box configuration (four or six vanes per box),

13
orientation of external vane boxes with respect to in- be used to augment or possibly replace conventional

ternal port walls (splay angle shims), and vane box control surfaces.

sideplates. Nozzle pressure ratio was varied from 2.0 9. The six-vane reverser configuration provided

to approximately 7.0. Results of this study indicate higher levels of reverse thrust than the four-vane

the following conclusions: configuration.

1. For a nonaxisymmetric nozzle with an integrated thrust reverser, use of reverser efftux angle, in- NASA Langley Research Center

ner door angle, and main nozzle throat area provides Hampton, VA 23665-5225

effective control of thrust ratio (magnitude of forward September 10, 1985

or reverse thrust). For the configuration tested, values of thrust ratio ranged from 0.97 for the main

nozzle only to almost -0.70 for full reverse thrust at

a typical operating nozzle pressure ratio of 3.0. References

2. Orientation of cascade vane box with respect 1. Lorincz, Dale J.; Chiarelli, Charles; and Hunt,

to reverser port walls appears to have a large effect on Brian L.: Effect of In-Flight Thrust Reverser Deployreverser exhaust splay angle. For the configuration ment on Tactical Aircraft Stability and Control. AIAA

tested, splay angle values less than 12° were mea- Paper No. 81-1446,July 1981.

sured for the full reverse-thrust configurations. The 2. Blackman, J. P.; and Eigenmann, M. F.: Axisymmetric

goal of a splay angle of 30° for each port was not met, Approach and Landing Thrust Reversers. AIAA Paper

partially, because the vane box sidewalls redirected No. 81-1650, Aug. 1981.

reverser exhaust flow (exiting the reverser ports at 3. Anderson, Seth B.; Cooper, George E.; and Faye,

a splay angle of approximately 30°) back toward the Alan E., Jr.: FlightMeasurementsof the Effect of a Convertical, trollableThrust Reverser on the Flight Characteristics of

3. Flow visualization observations and measured a Single-EngineJet Airplane. NASA MEMO 4- 26-59A,

performance data indicated that addition of the splay 1959.

4. Kelly, Mark W.; Greif, Richard K.; and Tolhurst, angle shims to align cascade vane boxes with reverser

William H., Jr.: Full-Scale Wind-Tunnel Tests of a

port walls significantly increased the reverser flow

Swept-Wing Airplane With a Cascade-TypeThrust Resplay angle for each port but correspondingly de- vcrser. NASA TN D-311, 1960.

creased the effiux angle. 5. Simpson, W. R.; Covey, M. W.; Palmer, D. F.; and

4. The interrelationship between reverser exhaust Hewett, M. D.: Navy Evaluation of F-11A In-Flight

effiux angle and splay angle dictates that reverser Thrust Control System. SA-75R-75, U.S. Naval Air

exhaust flow must be overturned in both the effiux Test Center, Dec. 15, 1975. (Available from DTIC as

and splay angle directions in order to meet the efflux AD A019 954.)

and splay angle goals simultaneously. 6. Steffen, Fred W.; and McArdle, Jack G.: Performance

5. Variable inner doors provide a smooth and Characteristics of Cylindrical Target-TypeThrust Reeffective modulation of weight flow on both land- versers. NACA RM E55129, 1956.

ing (full reverse-thrust) and approach (combined 7. Swihart, John M.: Effect of Target-Type Thrust Reverser

reverser/main nozzle) configurations. In addition, on Transonic Aerodynamic Characteristics of a SingleEngine Fighter Model. NACA RM L57J16, 1958.

they had a large impact on thrust ratio. As inner 8. Weiss, D. C.; and McGuigan, W. M.: Inflight Thrust

doors were moved from the fully open to the most Control for Fighter Aircraft. AIAA Paper No. 70-513,

closed position, losses in reverse thrust performance Mar. 1970.

occurred. 9. Linderman, D. L.; and Mount, J. S.: Developmentof an

6. Inner door angle (hence variation of weight In-Flight Thrust Reverser for Tactical/Attack Aircraft.

flow passing through reverser ports) had no effect AIAA Paper No. 70-699,June 1970.

on static pressures measured along the main nozzle 10. Maiden, Donald L.; and Mercer, Charles E.: Perforflap centerline, mance Characteristics of a Single-Engine Fighter Model

7. A variable internal slider door is probably a Fitted With an In-Flight Thrust Reverser. NASA TN

more desirable device for port area control than D-6460, 1971.

11. Mercer, Charles E.; and Maiden, Donald L.: Effects of a variable inner door because varying slider door

an In-Flight Thrust Reverser on the Stability and Conposition had a much smaller impact on thrust ratio trol Characteristics of a Single-Engine Fighter Airplane

for a given change in nozzle discharge coefficient. Model. NASA TN D-6886, 1972.

8. Differential variation of geometric effiux angle, 12. Goetz, Gerald F.; Young, John H.; and Palcza,

inner door angle, and/or slider door position gen- J. Lawrence: A Two-DimensionMAirframe Integrated

erated significant forces and moments which could Nozzle Design with Inflight Thrust Vectoring and Re14
versing Capabilities for Advanced Fighter Aircraft. 26. Carson, George T., Jr.; Capone, Francis J.; and MaAIAA Paper No. 76-626, July 1976. son, Mary L.: Aeropropulsive Characteristics of Nonaxi13. Willard, C. M.; Capone, F. J.; Konarski, M.; and symmetric-Nozzle Thrust Reversers at Mach Numbers

Stevens, H. L.: Static Performance of Vectoring/ From 0 to 1.20. NASA TP-2306, 1984.

Reversing Non-Axisymmetric Nozzles. AIAA Paper 27. Leavitt, Laurence D.; and Re, Richard J.: Static InterNo. 77-840, July 1977. nal Performance Characteristics of Two Thrust-Reverser

14. Capone, Francis J.: Static Performance of Five Twin- 28. Rowe, R. Kevin; Dusa, D. J.; and Leavitt,

Engine Nonaxisymmetric Nozzles With Vectoring and Re- Laurence D.: Static Internal Performance Evaluation of

versing Capability. NASA TP-1224, 1978. Several Thrust Reversing Concepts for 2D-CD Nozzles.

15. Hiley, P. E.; Kitzmiller, D. E.; and Willard, C. M.: In- AIAA Paper No. 84-1174, June 1984.

stalled Performance of Vectoring/Reversing Nonaxisym- 29. Obye, R. C.; and Hakim, A. D.: Axisymmetric Approach

metric Nozzles. J. Aircr., vol. 16, no. 8, Aug. 1979, and Landing Thrust Reverser Concepts: Hot Flow Test

pp. 532-538. Results. AIAA Paper No. 84-1176, June 1984.

16. Capone, Francis J.; Gowadia, Nashir S.; and Wooten, 30. McLafferty, George H.; and Peterson, Jeffrey L.: Results

W. H.: Performance Characteristics of Nonaxisymmet- of Tests of a Rectangular Vectoring/Reversing Nozzle on

ric Nozzles Installed on the F-18 Aircraft. AIAA Paper an F100 Engine. AIAA Paper No. 83-1285, June 1983.

No. 79-0101, Jan. 1979. 31. Nelson, B. D.; and Nicolai, L. M.: Application of Multi17. Petit, John E.; and Capone, Francis J.: Performance Function Nozzles to Advanced Fighters. AIAA Paper

Characteristics of a Wedge Nozzle Installed on an F- No. 81-2618, Dec. 1981.

18 Propulsion Wind Tunnel Model. AIAA Paper No. 32. Banks, Daniel W.; and Paulson, .John W., Jr.: Aero79-1164, June 1979. dynamic Characteristics and Predicted Landing Perfor18. Laughrey, J. A.; Drape, D. J.; and Hiley, P. E.: Per- mance of an F-15 Fighter With Two Thrust-Reverser

formance Evaluation of an Air Vehicle Utilizing Non- Configurations. NASA TP-2466, 1985.

axisymmetric Nozzles. AIAA Paper No. 79-1811, Aug. 33. Glezer, A.; Hughes, R. V.; and Hunt, B. L.: Thrust

1979. Reverser Effects on the Tail Surface Aerodynamics of

an F-18 Type Configuration. AIAA Paper No. 83-1860, 19. Capone, Francis J.; and Berrier, Bobby L.: Investigation

of Axisymmetric and Nonaxisymmetric Nozzles Installed July 1983. 34. Amin, N. F.; and Richards, C. J.: Thrust Reverser

on a 0.10-Scale F-18 Prototype Airplane Model. NASA Exhaust Plume Reingestion Tests for a STOL Fighter

TP-1638, 1980. Model. AIAA Paper No. 83-1229, June 1983.

20. Bare, E. Ann; Berrier, Bobby L.; and Capone, 35. Chiarelli, Charles; Lorincz, Dale; and Hunt, Brian: Francis J.: Effects of Simulated In-Flight Thrust Revers- Thrust Reverser Induced Flow Interference on Tactical

ing on Vertical-Tail Loads of F-18 and F-15 Airplane Aircraft Stability and Control. AIAA Paper No. 82- Models. NASA TP-1890, 1981. 1133, June 1982.

21. Capone, Francis J.; Re, Richard J.; and Bare, E. 36. Chiarelli, Charles; and Compton, Michael: Wind TunAnn: Thrust Reversing Effects on Twin-Engine Aircraft nel Evaluation of Tactical Aircraft Stability and Control

Having Nonaxisymmetric Nozzles. AIAA Paper No. as Affected by Nozzle Thrust Reverser Parameter Vari81-2639, Dec. 1981. ations. AIAA Paper No. 83-1228, June 1983.

22. Banks, D. W.; Quinto, P. F.; and Paulson, J. W., Jr.: 37. Burger, Karl Heinz: In-Flight Short Field Landing InThrust-Induced Effects on Low-Speed Aerodynamics of vestigations on a Combat Aircraft With Thrust ReFighter Aircraft. AIAA Paper No. 81-2612, Dec. 1981. verser. AIAA Paper No. 83-2693, Nov. 1983.

23. Re, Richard J.; and Berrier, Bobby L.: Static Inter- 38. Miller, E. H.: Performance of a Forward Swept Wing

nal Performance of Single Expansion-Ramp Nozzles With Fighter Utilizing Thrust Vectoring. AIAA Paper No.

Thrust Vectoring and Reversing. NASA TP-1962, 1982. 83-2482, Oct. 1983.

24. Hiley, P. E.; and Bowers, D. L.: Advanced Nozzle 39. Eigenmann, M. F.; Kitzmiller, D. E.; and Hakim, A. D.:

Integration for Supersonic Strike Fighter Application. Axisymmetric Approach and Landing Thrust Reverser

AIAA Paper No. 81-1441, July 1981. Concepts: In-Ground Effects Wind Tunnel Test Results.

25. Re, Richard J.; and Leavitt, Laurence D.: Static Internal AIAA Paper No. 84-1215, June 1984.

Performance Including Thrust Vectoring and Reversing of 40. Bittrick, W. C.: Axisymmetric Approach and Landing

Two-Dimensional Convergent-Divergent Nozzles. NASA Thrust Reversers for Single Engine Fighters. AIAA

TP-2253, 1984. Paper No. 84-1214, June 1984.

15

TABLE I. INDEX TO BASIC DATA

Basic data figure for

configuration withGeometric efflux Inner door

angle, Ot/_b, angle, Ct/¢b, Slider door Vane box Vane box Splay angle No Vertical Horizontal

Conf deg deg position configuration Port tab sideplates shims AtTmain, in2 5v, deg splitter splitter splitter

1 140/140 60/60 Open Four vanes On Off Off 0.0 N/A 5(a)

2 130/130 5(5) 5(c) 5(d)

3 110/110 5(e)

4 90/90 5(f) 5(g) 5(h)

5 66/66 5(i)

6 46/46 50)

7 26/26 5(k) 5(1) 5(m)

8 130/130 40340 5(n)

9 30/30 5(n)

10 20/20 5(0) 5(p) 5(q)

11 10/10 5(r)

12 ,[ Off/Off 5(s) 5(t) 5(u)

13 Off/Off 60/60 5(v) 5(w)

14 130/130 Closed 5(x) 5(y) 5(z)

15 130/130 2/3 5(a_)

16 130/130 1/3 5(aa)

17 90/90 Closed 5(bb) 5(cc)

18 90/130 60310 Open 5(dd)

19 90/130 60/60 Open 5(ee)

20 90/130 | Open/Closed .. i _ 5(if)

2122 130/130 130 l Open Open Four Six vanes vanes Off .. ._ ] 5(hh) 5(gg)
TABLE I. Concluded

Basic data figure for

configuration withGeometric effiux Inner door

angle, Ot/Ob, angle, Ct/¢b, Slider door Vane box Vane box Splay angle No Vertical Horizontal

Conf deg deg position configuration Port tab sideplates shims At,main, in 2 6v, deg splitter splitter splitter

23 130/130 60/60 Open Four vanes On On Off 0.0 N/A 5(ii) 5(jj) 5(kk)

24 130/130 I Off On 5(b)

25 110/110 5(e)

26 90/90 5(1!/

27 66/66 I 5(i)

28 46/46 5(j)

29 26/26 .. 5(k)

30 130/130 60/60-- Off .5 "; 5(mm)

31 20/20 .5 5(mm)

32 60/20 .5 5(mm)

33 60/60 2.0 5(nn)

34 30/30 i 5(nn)

35 20/20 5(00)

36 10/10 : 5(00)

37 20/10 5(00)

3s 30/30 2o 5(pp) l

39 20/20 / i 20 5(pp)

40 10/10 1 L . 20 5(pp)

41 Closed 2.0 0 5(q )

42 *Blank/Blank Closed N/A N/A N/A N/A N/A 2.0 20 5(qq)

43 t130/126 60/60 Open Four vanes On Off Off .0 N/A 5(rr)

* Reverser ports closed.

t Geometric efflux angle on left and right sides of mode! instead of top and bottom.
Sfa. 0 Sta. 20.50 Sta. 41.13

High-pressure plenum seal

(metal bellows) plate Total - pressure

8J7 rod. -Low pressure plenum probes

14°

7

Total-temperature

probe

Airflc Instrumentation

-8 equally spaced section

50 nozzles exiting

o, cho , rodia,, 0

40 percent

chord

5-percent thickness ratio paraliel_-_.__ _

to model centerline, 50 Typical section ahead Typical section in 20.O-in.chord at model f chord 45 °

centerline ,/_/_ ,_ of transition transition section

/

_"'_"_'_"_'_--_--To taI- pressure

Total-temperature _=,=___:_ probes

probe_ ;

Typical section in instrumentation section

Figure 1. Sketch of air-powered nacelle model with typical reverser configuration installed. All dimensions are

in inches unless otherwise noted.
Porttab "_ Sidewall

4.00 l-- ]2"0° "_

3.12

4.70 > / I I

< 7, 6i

< 8.72 >

(a) Details of model geometry.

Figure 2. Model sketches of single-engine reverser. All dimensions in inches unless otherwise noted.

20
vanes

Inner door

SectionA-A

m IFI

E,de9 Dim 'D' Dim 'E'

-- " I .74 !02 .45 .27

•-'(-1.48 -- _ 122 .30 .21

142 .29 .ll

-'('_ 2.08--'_')" 152 .22 .04

ViewB-B SectionC-C

(a) Concluded.

Figure 2. Continued.

21
Position Dim'F' Dim'G'

Open Open .64 .46

1/3 .87 .46

1/3 2/3 1.10 .50

'F' Closed 1.33 ..53

2/3

Closed

Structural support .I0

(b) Slider door positions.

Figure 2. Continued.

22
Six-vaneconfiguration

I

.273_ I <

-_--. 59---_-

5equalspaces --_

Four-vaneconfiguration

455 _. "_r]

_ 3 equal spaces _

(c) Vane box configurations.

Figure 2. Continued.

23
Externalvanebox Vaneboxsideplates

Upstreamport

wall Porttab

(d) Port tab and vane box sideplates.

Figure 2. Continued.

24
Splayangleshims

Externalvanes

Inner door

(e) Splay angle shims.

Figure 2. Continued.

25
3.59

Typ.

Sta. Sta.

47.45 47.45

At main• 2.0 in2, 5v • 0° At,main" 2.0 in2, 0v • 20°

.59

1.00 I

i_i o°

i

_1.70

I • 3.74 _ 3.62 >

3.64 _-

(f) 2-D C-D main exhaust nozzle geometry.

Figure 2. Continued.

26
12°

Nozzleexternal

surface

port et

walls \

Model _ "_[

centerline ,_ Externalvane /

boxes

(g) Definition of top and bottom efflux angles.

Figure 2. Concluded.

27
L-85-144

(a) Four-vane cascade box.

Figure 3. Reverser configurations.

28
L-85-145

(b) Six-vane cascade box.

Figure 3. Continued.

29
ii!!i _!

ilili_i!ii_

iiii_iiiii

i%_i!ii! _,

@

L-85-146

(c) Splay angle shims installed.

Figure 3. Continued.

30
(d) Inner doors.

Figure 3. Concluded.

31
Right-side ports closed

Verticalsplitter

plate

Figure 4. Orientation of horizontal and vertical splitter plates.

32
8t/eb=140°/140°, (bt/Ctb=60°160°, sliderdooropen,At, main=0.0 in2

-.4

-.5

F/Fi

".6

-.7

-.8

.58

•64

Wp/Wi

•60

•56

1 3 5 7 9

NPR

(a) Configuration 1; no splitter plate.

Figure 5. Variation of static internal performance characteristics with nozzle pressure ratio. Lines with no

symbols indicate resultant thrust ratio Fr/F i. Unless otherwise noted, all configurations have four-vane

external cascade boxes, port tabs installed, no vane box sideplates, and no splay angle shims.

33
Ot/Ob- 130°/130°, @t/@b- 60°/60°, sliderdooropen,At, main=0.0 in2

Conf2, withoutsplayangleshims Conf24,withsplayangleshims -.3 -.3

.92 i.12

.88 1.08

Wp/Wi Wp/Wi

.84 1.04

•80 1.O0

1 3 5 7 0 I 3 5 7 9

NPR NPR

(b) Configurations 2 and 24; no splitter plate.

Figure 5. Continued.

34
=0.0in2

OtlOb--130°1130°, St/$b,,60°160°, sliderdooropen,At, main

-.3 .1

-.4 0

-.5 -.1

F/Fi FN/Fi

-.6 -.2

.52 .2

.48 . I

Wp/Wi Fy/Fi

.44 0

.40 -.I

I 3 5 7 9 1 3 5 7 9

NPR NPR

(c) Configuration 2; vertical splitter plate.

Figure 5. Continued.

35
8t/Sb=130°/130°, @tJ@b=50°/5(P,sliderdooropen,At,main=0.0 in2

-.4 -.4

.48 .1

Wp/Wi .44 Fy/Fi 0

.40 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(d) Configuration 2; horizontal splitter plate.

Figure 5. Continued.

36
Ot/Ob=110°/110°, @t/@b- 60°/60°, sliderdooropen,At, main"0.0 in2

Conf3, withoutsplayangleshims Conf25, withsplayangleshims 0 0

I.20 I.32"

1.16 1.28

Wp/Wi Wp/Wi

1.12 1.24

1.08 t. 20

1 3 5 7 9 1 3 5 7 9

NPR NPR

(e) Configurations 3 and 25; no splitter plate.

Figure 5. Continued.

37
Ot/O=b90°19°0, @t/@b--60°16°,0 sliderdooropen,At, main=0.0 in2

.2 .2

.1 .1

F/Fi 0 FN/Fi 0

-.l -o].

1.24 .1

 p/Wi Fyi/F

1.20 0

1.16 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(f) Configuration 4; no splitter plate.

Figure 5. Continued.

38
--0.0in2

8t/Sb=90°/90°, _#t/(#b- 60°/60°, sliderdooropen,At, main

.2 .2

.1 .1

F/Fi 0 FN/Fi 0

.72 .3

.68 .2

Wp/Wi Fy/Fi

.64 .I

.60 0

1 3 5 7 9 1 3 5 7 9

NPR NPR

(g) Configuration 4; vertical splitter plate.

Figure 5. Continued.

39
Ot/Ob--90o190o, @t/@=b60°/60°, sliderdooropen,At, main=0.0 in2

.1 -.6

.72 .2

.68 .1

Wp/Wi Fy/Fi

.64 0

.60 -. 1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(h) Configuration 4; horizontal splitter plate.

Figure 5. Continued.

40
Ot/Ob=66°/66°, @t/@b=60°/60°, slider dooropen,At, main=0.0 in2

Conf.5,withoutsplayangleshims Conf27,withsplayangleshims

.4 .4

.2 .2

F/Fi F/Fi

.1 .1

0 0

1.28 i.36

]24 I,32

Wp/Wi Wp/Wi

1.20 I.28

1.16 1.24

l 3 5 7 9 1 3 5 7 9

NPR NPR

(i) Configurations 5 and 27; no splitter plate.

Figure 5. Continued.

41
--0.0in2

OtlOb=z16°/4°,6@t/@b--60°[60°, sliderdooropen,At, main

Conf6, withoutsplayangleshims Conf28,withsplayangleshims

.6 .6

.4 .4

F/Fi F/Fi

.3 .3

I.20 I.32

1.16 1.28

WplWi Wp/Wi

1.12 1.24

1.08 1.20

1 3 5 7 9 1 3 5 7 9

NPR NPR

(j) Configurations 6 and 28; no splitter plate.

Figure 5. Continued.

42
--0.0 in2

Ot/Sb=26°/26°, _t/_b- 60°/60°, sliderdooropen,At,main

Conf7; withoutsplayangleshims Conf29,withsplayangleshims

.8 .8

.6 .6

F/Fi F/Fi

.5 .5

.4 .4

•96 ]..]2

•92 1.08

Wp/wi Wp/wi

•88 1.04

•84 1.00

1 3 5 7 9 1 3 5 7 9

NPR NPR

(k) Configurations 7 and 29; no splitter plate•

Figure 5. Continued.

43
8tlSb=26°/26°, @tl@--b60°/60°, sliderdooropen,At, main--0.0 in2

.8 .2

.7 .1

F/Fi .6 FN/Fi 0

•52 .2

.48 .i

Wp/Wi Fy/Fi

.44 0

.ZlO -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(l) Configuration 7; vertical splitter plate.

Figure 5. Continued.

44
et/Ob=26°126°, _#t/_bo60°/60°, sliderdooropen,At,main"0.0 in2

.9 -.1

.52 .i

•48 0

Wp/Wi Fy/Fi

.44 -.1

.40 -.2

1 3 5 7 9 l 3 5 7 9

NPR NPR

(m) Configuration 7; horizontal splitter plate.

Figure 5. Continued.

45
Ot/Ob--130°/130°, sliderdooropen,At,main=0.0 in2

Conf8, St/_b--40°/40° Conf9,St/$b=30°/31P

-°2 -°2

•92 .80

•88 .76

Wp/Wi Wp/Wi

•84 .72

.80 .68

1 3 5 7 9 1 3 5 7 9

NPR NPR

(n) Configurations 8 and 9; no splitter plate.

Figure 5. Continued.

46
--0.0 in2

8t/Sb=130°/130°, _t/_b=20°/20°, sliderdooropen,At, main

-.1

-.2

F/Fi

-.3

-.4

-.5

•60

•56

Wp/Wi

•52

•48

1 3 5 7 9

NPR

(o) Configuration 10; no splitter plate.

Figure 5. Continued.

47
OtlOb=130°1130°, _tl(Pb=20°/20°, sliderdooropen,At,main=0.0in2

F/Fi -.5 FN/Fi 0

.36 .2

.32 .1

Wp/Wi Fy/Fi

.28 0

.24 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(p) Configuration 10; vertical splitter plate.

Figure 5. Continued.

48
Ot/O=b 130°1130°, _t/_b--20°120°, sliderdooropen,At,main"0.0 inz

-.5 -.5

F/Fi FN/Fi

-.6 -.6

.32 .1

Wp/Wi .28 Fy/Fi 0

.24 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(q) Configuration 10; horizontal splitter plate.

Figure 5. Continued.

49
--0.0 in2

Ot/eb=130°/130°, @t/@b--10°/10°, sliderdooropen,At, main

-.1

-.2

F/Fi

-.3

-.4

-.5

•32

.28

Wp/Wi

•24

•20

1 3 5 7 9

NPR

(r) Configuration 11; no splitter plate.

Figure 5. Continued.

5O
=0.0 in2

et/eb= 130°/130°, _t/_b--off/off, sliderdooropen,At, main

-.2

-.3

F/F. I

-.5

-.5

.95

•92

wp/wi

•88

• 84

1 3 5 7 9

NPR

(s) Configuration 12; no splitter plate.

Figure 5. Continued.

51
8tleb--130°1130°, _tl_b--off/off, sliderdooropen,At,main=0.0 in2

-.4 .2

.52 .2

.48 . l

Wp/Wi Fy/Fi

.44 0

._ -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(t) Configuration 12; vertical splitter plate.

Figure 5. Continued.

52
et/Sb=130°/130°, St/_b=off/off, sliderdooropen,At, main=0.0 in2

-4 -4

-6 -5

F/Fi FNIFi

-7 -7

•48 1

Wp/Wi .44 Fy/Fi 0

._ -.1

1 3 5 7 9 I 3 5 7 9

NPR NPR

(u) Configuration 12; horizontal splitter plate.

Figure 5. Continued.

53
Ot/eb--off/off, _t/_b--60°/60°, sliderdooropen,At, main=0.0 in2

.3

.2

.1

F/Fi

0

-.1

1.36

I.32

Wp/Wi

1.28

1.24

1 3 5 7 9

NPR

(v) Configuration 13; no splitter plate.

Figure 5. Continued.

54
--0.0in2

Ot/Ob--off/off,_t/t//b=60°/60°, sliderdooropen,At, main

.6 .1

.4 -.1

F/Fi FN/Fi

.3 -.2

.76 .4

•72 .3

Wp/Wi Fy/Fi

.68 .2

.64 .l

1 3 5 7 9 1 3 5 7 9

NPR NPR

(w) Configuration 13; vertical splitter plate.

Figure 5. Continued.

55
etleb--]30°1130°, CtlCb--60°160°, slider doorclosed,At,main- 0.0in2

-.2

-3

-4

F/F.

I

-.5

-6

52

48

WplWi

44

4O

1 3 5 7 9

NPR

(x) Configuration 14; no splitter plate.

Figure 5. Continued.

5O
Ot/Ob=130°/130°, _t/_b=60°/60°, sliderdoorclosed,At, main=0.0 in2

•32 .2

.28 .1

Wp/Wi Fy/Fi

.24 0

.20 -. 1

1 3 5 7 9 1 3 .5 7 9

NPR NPR

(y) Configuration 14; vertical splitter plate.

Figure 5. Continued.

67
Ot/eb--130°/130°, Ct/$b--60°/60°, slider doorclosed,At, main--0.0 in2

Wp/Wi .24 Fy/Fi 0

- 1

"201 3 5 7 q " 1 3 5 7 9

NPR NPR

(z) Configuration 14; horizontal splitter plate.

Figure 5. Continued.

58
Ot/Ob=130°/130°, _t/Ob=60°/60°, At,main=0.0in2

Conf1.5,slider door2/3 position Conf16, slider doori/3 position

-.2 -.2

•72 .80

•68 .76

WplWi Wp/Wi

•64 .72

•60 .68

1 3 5 l 9 1 3 5 7 9

NPR NPR

(aa) Configurations 15 and 16; no splitter plate.

Figure 5. Continued.

59
Ot/Ob=90°/90°, _t/g/b--60°/60°, sliderdoorclosed,At, main=0.0 in2

.1

0

F/Fi

".1

-.2

-.3

•64

.60

WplWi

•.56

.52

1 3 5 7 9

NPR

(bb) Configuration 17; no splitter plate.

Figure 5. Continued.

6O
8t/Sb-"90°190°, @t/@b=60°/60°, sliderdoorclosed,At, main=0.0 in2

F/Fi 0 FN/Fi -.7

35 .1

32 0

Wp/wi Fy]Fi

28 -,]

24 -.2

1 3 5 7 9 i 3 5 7 0

NPR NPR

(cc) Configuration 17; horizontal splitter plate.

Figure 5. Continued.

61
et/Sb--90°/130°, ,_t/_b--60°/10°, sliderdooropen,At,main=0.0 in2

1.00 .2

.96 .I

Wpwl i FylFi

•92 0

.88 -.I

1 3 5 7 9 1 3 5 7 9

NPR NPR

(dd) Configuration 18; no splitter plate.

Figure 5. Continued.

62
8t/Sb=90°1130°, _#t/_b--60°/60°, sliderdooropen,At,main=0.0in2

0 0

F/Fi -.2 FN/Fi -.2

1.12 .2

1.08 .1

Wp/Wi Fy/Fi

1.04 0

1.00 -. 1

1 3 5 7 9 i 3 5 7 9

NPR NPR

(ee) Configuration 19; no splitter plate.

Figure 5. Continued.

63
et/Sb--90°/130°, _t/_b =60°/60°, slider dooropenontopandclosedon bottom,At, main=0.0 in2

0 0

.92 .1

Wp/Wi .88 Fy/Fi 0

.84 -.i

1 3 5 7 9 1 3 5 7 9

NPR NPR

(if) Configuration 20; no splitter plate.

Figure 5. Continued.

64
8t/eb=130°1130°, _t/_b=60°/60°, sliderdooropen,At,main=0.0 in2,

Six-vane configuration

-.3

FIFi

.76

Wp/Wi

•72

.68

1 3 5 7 9

NPR

(gg) Configuration 21; no splitter plate.

Figure 5. Continued.

65
et/eb=130°1130°, 0t/0b=60°/60°, sliderdooropen,At, main=0.0 in2,

porttabremoved

-.2

-.3

F/Fi

-._

-.5

-.6

I. 12

1.08

Wp/Wi

i. 04

1.O0

I 3 5 7 9

NPR

(hh) Configuration 22; no splitter plate.

Figure 5. Continued.

66
=0.0 in2,

et/Ob=130°/130°, _Dt/_b--60°/60°, sliderdooropen,At, main

vaneboxsideplatesinstalled

-.2

-.3

F/Fi

-._

-..5

-.6

•92

•88

Wp/Wi

•84

•80

1 3 5 7 9

NPR

(ii) Configuration 23; no splitter plate.

Figure 5. Continued.

67
8t/Sb=130°/130°, _tl_b=60°/60°, sliderdooropen,At, main=0.0in2, vaneboxsideplatesinstalled

-.3 .2

-.5 0

F/Fi FN/Fi

-.6 -.1

.48 .2

.44 .1

Wp/Wi Fy/Fi

.40 0

•36 -. 1

] 3 5 7 9 1 3 5 7 9

NPR NPR

(jj) Configuration 23; vertical splitter plate.

Figure 5. Continued.

68
et/Sb--130°/130°, _t/_b =60°/60°, slider dooropen,At, main= 0.0 in2, vaneboxsideplatesinstalled

Wp/Wi .40 Fy/Fi 0

.36 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(kk) Configuration 23; horizontal splitter plate.

Figure 5. Continued.

69
et/ab--90°190°, t#tl_b--60°/60°, sliderdooropen,At, main"0.0 in2, withsplayangleshims

.2 .2

F/Fi 0 FN/Fi 0

1.36 .2

1.32 .1

Wp/wi Fy/Fi

1.28 0

1.24 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(ll) Configuration 26; no splitter plate.

Figure 5. Continued.

7O
erieb- 130°/130°, sliderdooropen,At, main=0.5 in2

Conf31, _t/_b=20°120° Conf30, _tlCb=60°/60° Conf32, Ct/¢b=60°/20°

.1 0 0

0 -.1 -.1

F/Fi -.1 F/Fi -.2 F/Fi -.2

-.2 -.3 -.3

-.3 -.4 -.4

.72 1.00 .92

•68 .96 .88

Wp/Wi Wp/Wi Wp/Wi

•64 .92 :84

.60 .88 .80

1 3 5 7 9 1 3 5 7 9 1 3 5 7 9

NPR NPR NPR

(ram) Configurations 30, 31, and 32; no splitter plate.

Figure 5. Continued.

..,1

i.,.t
et/Ob=130°/130°, sliderdooropen,At,main--2.0in2

Conf34,@t/@b--30°/30° Conf33,@t/@b--60°/60°

.2 .2

F/Fi 0 F/Fi 0

1.28 1.44

1.24 1.40

Wp/Wi Wp/Wi

1.20 I.36

i. 16 I. 32

1 3 5 7 9 1 3 5 7 9

NPR NPR

(nn) Configurations 33 and 34; no splitter plate.

Figure 5. Continued.

72
8t/Sb=130°/130°, slider dooropen,At, main=2.0 in2

Conf35,t#t/_b=21P/20° Conf37,@t/t#b=2(P/10° Conf35, ¢,t/_b=10°110°

.4 .4 .6

.3 .3 .5

F/Fi .2 F/Fi .2 F/Fi .4

.I .I .3

0 0 .2

i. 08 .92 .76

1.04 .88 .72

Wp/Wi Wp/Wi Wp/Wi

1.00 .84 .68

.96 .80 .64

I 3 5 7 9 1 3 5 7 9 I 3 5 7 9

NPR NPR NPR

(oo) Configurations 35, 36, and 37; no splitter plate.

Figure 5. Continued.
et/Sb=130°/130°, slider dooropen,At, main=2.0 in2, 8v=20°

C0nf40, $t/¢b=10°/10° Conf39, ¢t/¢b=20°/20° C0nf38, ¢t/¢b- 30°/30°

.6 .4 .4

.3

.5 .3

F/Fi .4 F/Fi .2 F/Fi .2

.3 .I .I

.2 0 0

•72 1.08 1.24

•68 I.04 i.20

Wp/Wi Wp/wi Wp/Wi

•64 I.O0 i.16

•60 .96 i. 12

1 3 5 7 9 1 3 5 7 9 1 3 5 7 9

NPR NPR NPR

(pp) Configurations 38, 39, and 40; no splitter plate.

Figure 5. Continued.
At, main=2.0 in2, reverserportsclosed

Conf41,6v--0° Conf42,6v;20°

i. O0 I. O0

•96 .96

.92 .92

F/Fi F/Fi

•88 .88

Wp/Wi .44 Wp/Wi .44

1 3 5 7 9 I 3 5 7 9

NPR NPR

(qq) Configurations 41 and 42; no splitter plate.

Figure 5. Continued.

75
el/e r- 130°/26°, St/$b- 60°160°, sliderdooropen,At,main=0.0 in2

Note:130° reverservanesinstalledon leftsideand26° reverservanesinstalledon right side

.2 1.0

.I .9

F/Fi 0 Mz/Fil'r .8

.96 .2

.92 .1

Wp/Wi Fy/Fi

•88 0

.84 -.1

1 3 5 7 9 1 3 5 7 9

NPR NPR

(rr) Configuration 43; no splitter plate.

Figure 5. Concluded.

76
etle b NosplayConf angleshims WithsplayConf angleshims

0 140°/140° 1

E] 130°/130° 2 24

0 110°/110° 3 25

e_ 90°/go° 4 26

IX 66°/66° 5 27

{3 460146° 6 28

(2) 260126° 7 29

0 off/off 13

•8 Nosplayangleshims .8

.6 .6

.4 .4

.2 .2

F/Fi F/Fi

0 0

-.8

1 3 5 7 9 1 3 5 7 9

NPR NPR

Figure 6. Effect of geometric eftiux angle on thrust ratio. Ct/¢b = 60°/60°; slider door open;

At,main = 0.0 in2.

77

i
\

1.0 .4

P/Pt,j Mint

.8 .2

.6 0

0 20 40 60 80 i00 120 140 160

O,deg

(a) Upper flap pressures and Mach numbers; splay angle shims off.

Figure 7. Effect of geometric ei_lux angle and splay angle shims on reverser static-pressure ratios and internal

Mach numbers. Ct/¢b = 60°/60°; slider door open; At,main = 0.0 in2; NPR = 3.0.

78
1.0

.8

Leftupstreamportwall

P/Pt,j .4"6

.2

1.0

.8

P/Pt,j Leftdownstreamportwall

.4

.2

0 20 40 60 80 I00 120 140 160

8, deg

(b) Port pressures; splay angle shims off.

Figure 7. Continued.

79
1.0

Leftupstreamportwall

.8

P/Pt,j _l

.4

.2

1.0

Leftdownstreamportwall

.8

P/Pt,j

.6

.4

.2

off on

Splayangleshim

(c) Port pressures; Ot/O b = 130°/130°.

Figure 7. Concluded.

80
et/eb Conf

0 130°/130° 2

------ f-I 90°/130° 19

90°/90° 4

.1 .1

0 0

-.2 -.2

F/Fi FN/Fi

-.3 -.3

-.6 -.6

1 3 5 7 9 I 3 5 7 9

NPR NPR

(a) Differential efflux angles for generating normal force.

Figure 8. Effect of differential geometric effiux angles on thrust ratio, normal-force ratio, and yawing-moment

ratio. Ct/¢b = 60°//60°; slider door open; At,main = 0.0 in2. Line with no symbols indicates resultant

thrust ratio Fr/ Fi.

81
_/e r Conf

0 130°/130° 2

[] 130°126o 43

26°/26° 7

.8 1.0

.2 .4

F/Fi MZIFIZ: r

0 .2

-.8 -.6

1 3 5 7 9 1 3 5 7 9

NPR NPR

(b) Differential effiux angles for generating yawing moment.

Figure 8. Concluded.

82
0

St/_b Conf

0 600160° 2

-. 1 IZI 400140° 8

0 300130° 9

r, 200120° 10

r,, 10o/10° 11

-.2 1:3Off/Off 12

F/F.

I

-.3

-4

-5

-.6

1 3 5 7 9

NPR

Figure 9. Effect of inner door angle on thrust ratio. _t/8 b = 130°/130°; slider doer open; At,main -- 0.0 in2.

83
_#t/@b 5v-O °,Conf- 5v-20 °,Conf0 60°/60° 33

U 30°130° 34 38

--- 0 20°/20° 35 39

----- •z_ 10°110° 36 40

5:0 ° IX 20°110° 37 6:20 °

V V

.6 .6

.3 .3

F/Fi F/Fi

.2 .2

0 0

-.I -.I

I 3 5 7 9 1 3 5 7 9

NPR NPR

Figure 10. Effect of inner door angle on thrust ratio for approach configuration. _/0b = 130°/130°; slider

door open; At,main -- 2.0 in2. Lines with no symbols indicate resultant thrust ratio Fr/F i.

84
At, main 5v, deg

O 0 -

O 2.0 0

0 2.0 20

1.6

!.4

1.2

1.0

Wp/Wi

.8

.5

.4

.2

0

10 20 30 40 50 60 off

_, deg

Figure 11. Summary of the effect of inner door angle on nozzle discharge coefficient for various reverser

configurations. Ot/Ob= 130°/130°; slider door open; NPR = 3.0.

85
/

.7 0

FlaggedsymbolsrepresentvaluesofMint

1.0 .4

.9 .3

P/Ptj, Mint

.8 .2

At, main

.7 5v.0o .1

.6 0

CIosed l0 20 30 40 50 60 Off

@,deg

(a) Upper flap pressures and Mach numbers.

Figure 12. Effect of inner door angle on reverser static-pressure ratios and internal Mach numbers. Ot/Sb=

130°/130°; slider door open; NPR = 3.0.

86
1.0

•8 Leftupstreamportwall

P/Ptj,

.5

.4

.2

1.0

•8 Leftdownstreamportwall

p/Pt,j

I

, i/°.

.4

.2

10 20 30 40 50 50 Off

_, deg

(b) Port pressures; At,main : 0.0 in2.

Figure 12. Continued.

87
1.0

Left upstreamportwall

.8

P/Pt,j .6.4

.2

1.0

Leftdownstreamportwall

.8

p/pt,J .6

.4

.2

10 20 30 40 50 60

@,deg

(c) Port pressures; At,main = 2.0 in2; 6v = 0°.

Figure 12. Continued.

88
1.0

.8

P/Pt,j

.6

.4

.2

0

CIosed 10 20 30 40 50 60

_, deg

(d) Main exhaust nozzle upper flap pressures; At,main = 2.0 in2;5v = 0°.

Figure 12. Concluded.

89
Sliderdoor etleb-- 13°°113°° etleb= 9°°19°°

position conf Conf

O open 2 4

(3 1/3 16

213 15

z,, closed 14 17

-.] .3

-.2 .2

F/Fi -.4 F/Fi 0

-.7 -.3

1 3 5 7 9 1 3 5 7 9

NPR NPR

Figure 13. Effect of slider door position on thrust ratio. Ct/¢b = 60°/60°; At,main = 0.0 in2.

90
)

-.5 .8 ,

I

F Wp r/F. /w,

I " I ;I

I

-6 6 i,

;i

rl

I

!

-.7 .4 !

Closed 2/3 1/3 Open Closed 2/3 113 Open

Sliderdoorposition Sliderdoorposition

Figure 14. Summary of the effects of slider door position on thrust ratio and nozzle discharge coefficient.

Ot/Ob = 130°/130°; Ct/¢b : 60°/60°; At,main = 0.0 in2; NPR = 3.0.
1.0 .4

7

.9 .3 -- \

P/Pt,j Mint .,o n"

.8 .2

"/

.7 .1

.6 0

Closed 213 I/3 Open

Sliderdoorposition

(a) Upper flap pressures and Mach numbers.

Figure 15. Effect of slider door position on reverser static-pressure ratios and internal Mach numbers.

Ot/Ob: 130°/130°; Ct/¢b : 607/60°; At,main = 0.0 in2; NPR = 3.0.

92
1.0

Left upstreamportwall .8

P/Ptj,

I

.5

.4

.2

1.0

•8 Leftdownstreamportwall

P/Ptj ,

I

i/o

.4

.2

Closed 2/3 1/3 Open

Sliderdoorposition

(b) Port pressures.

Figure 15. Concluded.

93
Sliderdoor

_t/_b position Conf

0 60°/60° open 19

13 60°/10° open 18

0 600160° open/closed 20

0 0

-.1 -.1

-.2 -.2

F/Fi FN/Fi

-.3 -.3

-.4 -.4

-.5 -.5

I 3 5 7 9 I 3 5 7 9

NPR NPR

Figure 16. Effect of inner door angle and slider door position on thrust and normal-force ratios. 8j0 b =

90°/130°; At,main = 0.0 in2. Lines with no symbols indicate resultant thrust ratio Fr/F i.
0

-.1

-.2 1:3

F/F.

I

-.3

-.4

-.5

-.6

1 3 5 7 9

NPR

Figure 17. Effect of vane box configuration on thrust ratio. Ot/_b = 130°/130°; Ct/¢b = 60°/60°; slider door

open; At,main -- 0.0 in2.

95
-.1

0

-.2 0

-3

F/F.

I

-.4

-5

-.6

-7

1 3 5 7 9

NPR

Figure 18. Effect of port tab on thrust ratio. Ot/t_b : 130°/130°; _bt/¢b : 60°/60°; slider door open;

At,main -- 0.0 in2.

96
Flaggedsymbolsindicate

valuesof Mi nt

1.0 .4

.9 3 -L

P/Pt,j Mint .,o .

.8 .2 u"

.7 .1 _ /

.6 0

On Off

Porttab

(a) Upper flap pressures and Mach numbers.

Figure 19. Effect of port tab on reverser static-pressure ratio and internal Mach number. 0t]Ob = 130°/130°;

Ct/¢b = 60°/60°; slider door open; At,main = 0.0 in2; NPR = 3.0.

97
1.0

Leftupstreamportwall

.8

p/Pt,j

.5

.4

.2

1.0

Leftdownstreamportwall

.8

P/Pt,j

.6

.4

.2

On Off

Porttab

(b) Port pressures.

Figure 19. Concluded.

98
-.1

-2 0

E3

-3

F/F.

I

-4

-.5

-.6

-7

I 3 5 7 9

NPR

Figure 20. Effect of vane box sideplates on thrust ratio. _t/_ b ---- 130°/130°; Ct/_b ---- 60°/60°; slider door

open; At,main = 0.0 in2.

g9
1.0

.8

6

• At,main' in2 Ot/Ob Conf

O 2.0 closed/closed 41

IZI 2.0 130°/130° 33

.4 O 0.5 J, 30

z_ 0.0 T 2

F/F.I

.2

-.2

-.4

.6

1 3 5 7 9

NPR

Figure 21. Effect of thrust modulation on thrust ratio. Ct/¢b = 60°/60°; slider door open when thrust reverser

was in operation; 6v = 0°.

100
8v, deg Conf 8v, deg Conf

0 0 41 0 0 36

------CI 20 42 -----El 20 40

1.00 .7

.96 .6

.92 .5

F/Fi F/Fi

•88 .4

•84 .3

•80 .2

1 3 5 7 9 1 3 5 7 9

NPR NPR

(a) Reverser ports closed. (b) Ct/tb5 = 10°/10°.

Figure 22. Effect of main nozzle vector angle on thrust ratio for forward mode and approach configurations.

Ot/Ob = 130°/130°; slider door open when reverser was in operation; At,main = 2.0 in2. Lines with no

symbols indicate resultant thrust ratio Fr/F i.
5v, deg Conf 5v, deg Conf

0 0 35 0 0 34

O 20 39 ------ O 20 38

.4 .4

.3 .3

.2 .2

F/Fi F/Fi

.1 .1

0 0

-.I -.I

1 3 5 7 9 1 3 5 7 9

NPR NPR

(c) ¢_/¢b = 20°/20% (d) CtlCb = 30°/30°.

Figure 22. Concluded.
-%

\ I \ \

\1 \

FN/F NOTE:All signshavebeenreversedto allowpresentationof flow

\ \ \ diredJoninsteadofthe reactionforce.

\

\ \

F/Fi _ \ \ \ \

I \ \ \ \ Fr,tot/Fi = / (F/Fi)2+ (FN/Fi)2 + (Fy/Fi)2 \ \

\ \ 7 1.56)2 (.55)2 + (-. 105)2

Fkr,tot/iF \ \ = +

, _'.... .55 = .792

f "' I

I _,._=tan-1 Fy/Fi = tan4-.!05 80 I _ =lO.

, I

\

\

\

\

\

\

\

\

IX / _" Fy/Fi

!80o-tan-] FN'Fi : !80o- tan-] _=_ =135.5°

I _ _ O_

\ F/Fi

Figure 23. Component breakdown of resultant flow vector for top left-hand port. Conf 2; Ot/t_b = 130°/130°;

_., Ct/_bb = 60°/60°; slider door open; At,main = 0.0 in2; NPR = 3.0. Data obtained from splitter plate runs.
et/Ob VaneboxsideplatesConf

0 130°1130° no 2

0 90°/90° I 4

40 0 26°126° f 7 40

Z_ 130°1130° yes 23 Ot/Ob=130°1130°, slider dooropen,NPR=3.0

30 30

_p,deg _#,deg

20 20

Conf12

i0 i0

0

1 3 5 7 9 0 20 40 60 Off

NPR _t and¢b' deg

Figure 24. Effect of geometric efflux angle and inner door angle on reverser splay angle. Slider door open.
Four-vaneconfiguration,porttabon, novaneboxsideplatesno, splayangleshims,5v=0°

1.0

.8

.6

.4

.2

FIFi 0

-.2

-.4

-.6

-.8

-i.0

0 20 40 60 80 I00 120 140 160

0t and8b, deg

Figure 25. Summary of geometric efflux angle and main nozzle throat area effects on thrust ratio. _/_b =

60°/60°; slider door open; NPR = 3.0.

105
Four-vane configuration,porttab on, novaneboxsideplates,nosplayangleshims, 8v=0°

1.0

.8

.5

.4

.2

F/Fi

0

-.2

-.4

-.5

-.8

-1.0

0 10 20 30 40 50 50 70

St andSb, deg

Figure 26. Summary of inner door angle and main nozzle throat area effects on thrust ratio. Ot//{)b = 130°/130°;

slider door open; NPR -- 3.0.

106



1. Report No. ] 2. Government Accession No. 3. Recipient's Catalog No.

NASA TP-2519 ]

4. Title and Subtitle 5. Report Date

Static Internal Performance of a Single-Engine December 1985

Nonaxisymmetric-Nozzle Vaned-Thrust-Reverser Design 6. PerformingOrganizationCode

With Thrust Modulation Capabilities 505-40-90-01

7. Author(s)

Laurence D. Leavitt and James R. Burley II s. PerformingOrganizationReportNo.

L-15991

9. Performing Organization Name and Address 10. Work Unit No.

NASA Langley Research Center

Hampton, VA 23665-5225 11. Contract or Grant No.

13. Type of Report and Period Covered

12. Sponsoring Agency Name and Address Technical Paper

National Aeronautics and Space Administration

Washington, DC 20546-0001 14. SponsoringAgencyCode

15. Supplementary Notes

16. Abstract

An investigation has been conducted at wind-off conditions in the static-test facility of the Langley 16-Foot

Transonic Tunnel. The tests were conducted on a single-engine reverser configuration with partial and

full reverse-thrust modulation capabilities. The reverser design had four ports with equal areas. These

ports were angled outboard 30° from the vertical to impart a splay angle to the reverser exhaust flow.

This splaying of reverser flow was intended to prevent impingement of exhaust flow on empennage surfaces

and to help avoid inlet reingestion of exhaust gas when the reverser is integrated into an actual airplane

configuration. External vane boxes were located directly over each of the four ports to provide variation of

reverser efflux angle from 140° to 26° (measured forward from the horizontal reference axis). The reverser

model was tested with both a butterfly-type inner door and an internal slider door to provide area control

for each individual port. In addition, main nozzle throat area and vector angle were varied to examine

various methods of modulating thrust levels. Other model variables included vane box configuration (four

or six vanes per box), orientation of external vane boxes with respect to internal port walls (splay angle

shims), and vane box sideplates. Nozzle pressure ratio was varied from 2.0 to approximately 7.0.

17. Key Words (Suggested by Authors(s)) 18. Distribution Statement

Nonaxisymmetric nozzles Unclassified--Unlimited

Thrust reversing

Internal performance

External vanes

Subject Category 02

19. Security Classif.(of this report) 20. Security Classif.(of this page) 21_ No. of Pages 22. Price

Unclassified Unclassified 107 A06

For sale by the National Technical Information Service, Springfield, Virginia 22161

NASA-Langley, 198

National Aeronauticsand

Space Administration BULKRATE

Code NIT-4 POSTAGE & FEES PAID

NASA Washington,DC

Washington, D.C. Permit No. G-27

20546-0001

Official Business

Penalty for Private Use, $300

If Undeliverable (Section 158

POSTMASTER:

Postal Manual) Do Not Return