NASA Technical Memorandum 4722

Design and Development of 

an F/A-18 Inlet Distortion 

Rake: A Cost and Time 

Saving Solution

October 1995

Andrew J. Yuhas, Ronald J. Ray, Richard R. Burley,

William G. Steenken, Leon Lechtenberg, and Don Thornton
NASA Technical Memorandum 4722

National Aeronautics and 

Space Administration

Office of Management

Scientific and Technical 

Information Program

1995

Andrew J. Yuhas

PRC Inc.

Edwards, California

Ronald J. Ray

Dryden Flight Research Center

Edwards, California

Richard R. Burley

NASA Lewis Research Center

Cleveland, Ohio

William G. Steenken, Leon Lechtenberg

and Don Thornton

General Electric Aircraft Engines

Evendale, Ohio

Design and Development of 

an F/A-18 Inlet Distortion 

Rake: A Cost and Time 

Saving Solution
DESIGN AND DEVELOPMENT OF AN F/A-18 INLET DISTORTION RAKE: 

A COST AND TIME SAVING SOLUTION

Andrew J. Yuhas*

PRC Inc. 

Edwards, California

Ronald J. Ray**

NASA Dryden Flight Research Center

Edwards, California

Richard R. Burley†

NASA Lewis Research Center

Cleveland, Ohio

William G. Steenken,†† Leon Lechtenberg‡

, and Don Thornton‡‡

General Electric Aircraft Engines

Evendale, Ohio

ABSTRACT

An innovative inlet total-pressure-distortion measurement rake has been designed and developed for the

F/A-18 A/B/C/D aircraft inlet. The design was conceived by NASA and General Electric Aircraft Engines

personnel. This rake has been flight qualified and flown

in the F/A-18 High Alpha Research Vehicle at NASA

Dryden Flight Research Center, Edwards, California.

The eight-legged, one-piece, wagon wheel design of

the rake was developed at a reduced cost and offered

reduced installation time compared to traditional

designs. The rake features 40 dual-measurement ports

for low- and high-frequency pressure measurements

with the high-frequency transducer mounted at the

port. This high-frequency transducer offers direct

absolute pressure measurements from low to high frequencies of interest, thereby allowing the rake to be

used during highly dynamic aircraft maneuvers. Outstanding structural characteristics are inherent to the

design through its construction and use of lightweight

materials.

NOMENCLATURE

Acronyms

AIP aerodynamic interface plane

AIR Aerospace Information Report

ALF aft looking forward

AOA angle of attack, deg

ARP Aerospace Recommended Practice

ac alternating current

CFD computational fluid dynamics

DFRC Dryden Flight Research Center, 

Edwards, California

F/A fighter–attack aircraft

FOD foreign object damage

GE General Electric

GEAE General Electric Aircraft Engines, 

Evendale, Ohio

g inertial force expressed in multiples of 

gravity

HARV High Alpha Research Vehicle

HATP High Alpha Technology Program

ID inside diameter

LeRC Lewis Research Center, Cleveland, 

Ohio

M Mach number

NASA National Aeronautics and Space 

Administration

NASTRAN NASA Structural Analysis, a finite 

element modeling program

OD outside diameter

PSI Pressure Systems Incorporated, 

Hampton, Virginia
2

PSIPP pounds force per square inch, peak to 

peak

RTV room temperature vulcanizing

SAE Society of Automotive Engineers

X experimental aircraft

Symbols

σ standard deviation

 INTRODUCTION

Designing, developing, and installing an inlet

total-pressure-distortion rake can be expensive and

time consuming. Figure 1 shows the F/A-18 High

Alpha Research Vehicle (HARV) being flown at NASA

Dryden Flight Research Center (DFRC), Edwards, California. The HARV inlet research program goals

required a low-cost, short-installation-time, and lowmaintenance solution to meet the need for a totalpressure inlet-distortion measurement system. The

HARV inlet research program evaluates the inlet

characteristics of high-performance aircraft during

stabilized and highly dynamic maneuvers at high

angles of attack.

Traditional inlet rake systems typically use eightlegged, duct-mounted cantilevered designs, such as

those used on the developmental F/A-18A1

 or inlet

guide vane leading-edge designs.2

 These designs have

commonly required extensive modifications to the aircraft or the engine. In particular, the cantilevered system, with its individual rake leg designs and

installations, can be costly, because of separate rake leg

development and testing. In addition, this system can

be time consuming because of increased structural

modifications of the aircraft and flight qualification

testing.

Figure 2 shows an inlet rake which was developed by

the NASA Lewis Research Center (LeRC), Cleveland,

Ohio, and DFRC. This rake was based on an innovative

design concept conceived by NASA and General Electric Aircraft Engines (GEAE), (Evendale, Ohio). This

approach connects all of the rake legs together at the

hub to form a one-piece, wagon wheel design. This

design simplifies rake installation and aircraft modification requirements, thus greatly reducing cost and

weight. The rake, designed and fabricated by GEAE,

EC 91 495-15

Figure 1. NASA F/A-18 HARV aircraft (preproduction aircraft number 6 modified with multiaxis thrust vectoring

paddles). 
3

has passed the needed flight qualification requirements

and has been flown on the HARV.

This paper describes the design, fabrication, installation, and qualification testing of the HARV inlet rake

system. Comparisons of cost and installation time

between this design and a previous design are made.

The paper also details the design requirements and

pressure transducer selection. All stages of flight qualification testing, from laboratory to flight test, are

described.

Use of tradenames or names of manufacturers in this

paper does not constitute an official endorsement of

such products or manufacturers, either expressed or

implied, by the National Aeronautics and Space

Administration.

INLET RAKE SYSTEM REQUIREMENTS

When considering an inlet rake design for use on the

F/A-18 HARV flight program, the basic requirements

were first established. The most important requirement

was to provide as much commonality as practical with

the planned HARV inlet wind tunnel test at the LeRC

and with previous F/A-18 inlet testing. Commonality

considerations with past and present testing included

but were not limited to instrumentation setup, rake

positioning, and probe configuration. The design also

had to allow accurate data measurements to be gathered to meet all of the HARV inlet research objectives.

This requirement primarily concerned instrumentation

selection. In addition, following established industry

guidelines wherever possible was desired. Considerations of cost and installation time are always important factors that ultimately constrained the design

requirements.

Commonality

Inlet instrumentation was required on the HARV to

correlate flight data with and verify test results from the

planned 9- by 15-ft wind tunnel tests scheduled at the

LeRC as part of the High Alpha Technology Program

(HATP) inlet research program. In that aspect, the

HARV instrumentation was developed to emulate the

wind tunnel set, limited primarily by the cost and complexity of modifying the full-scale vehicle. 

 Having commonality between the current flight test

rake and previous F/A-18 flight test rakes was also

desirable. In the mid-1970’s, the U.S. Navy along with

industry partners, McDonnell Douglas Aerospace

EC 93 41084-7

Figure 2. NASA and GEAE inlet pressure distortion rake mounted in HARV right inlet.
4

(St. Louis, Missouri), Northrop Aircraft Division

(Newbury Park, California), and GEAE conducted an

inlet evaluation on the second preproduction F/A-18A.

The original inlet rake design consisted of eight independent, cantilevered rakes, each spaced equiangularly,

having five measurement ports per rake located on the

centroids of five equal areas.1

 Each rake was mounted

to various structural members behind the inlet duct

wall. Because of the forces and moments acting on

each cantilevered rake, special attention was given to

the structural design and buildup. To meet the complex

inlet rake structural requirements, the bulkhead on aircraft number 2 was specifically designed to accommodate the inlet rake-mounting requirements. The cost

and time required for this modification was obviously

significant. Figure 3 shows the original rake installed in

aircraft number 2. This original inlet rake design generally followed the instrumentation guidelines established by industry. 2,3

Technical Guidelines

An Aerospace Recommended Practice titled Gas

Turbine Engine Inlet Flow Distortion Guidelines, ARP

1420, was established by the Society of Automotive

Engineers (SAE) in 1978 to ensure a consistent

approach to the development of the inlet instrumentation configuration and to provide a proven and consistent method of data analysis.3

 The SAE document

recommends that the instrumentation and analysis

methods be agreed upon among the involved parties

and remain invariant throughout the propulsion system

life cycle for all testing. This general approach was followed for the design and manufacture of the HARV

inlet rake described here. In 1983, SAE issued a more

comprehensive report, Inlet Total-Pressure-Distortion

Considerations for Gas-Turbine Engines, AIR 1419,

that provides more detailed information.2

The SAE established the “aerodynamic interface

plane” (AIP) as the location of the instrumentation

plane used to define inlet distortion and performance.

In general, the guide recommended that the AIP should

be located in a circular duct as close to the engine face

as practical. The engine face is defined by the leading

edge of the most upstream engine strut, vane, or blade

row. To be consistent with past F/A-18 testing, the

HARV AIP was required to be 4 in. in front of the

engine bullet nose. The standard also established a typical 40-port rake array for distribution of the total pressure ports at the AIP. This array consists of eight

EC 91 220-2

Figure 3. Preproduction F/A-18A aircraft number 2 cantilevered inlet rake system.
5

equiangularly spaced rakes with five ports per rake

located at the centroids of equal areas. The original

F/A-18 rake (fig. 3) followed this configuration and

was clocked (rotated clockwise aft looking forward

(ALF)) approximately 9°. Clocking the rake is often

required because of structure installation considerations. The HARV inlet flight test program goal was to

meet this arrangement. Because the original F/A-18A

rake was installed in the left inlet, the HARV rake, to

be installed in the right inlet, was required to be

clocked counterclockwise ALF approximately 9°. This

configuration would make the installation equivalent

because of symmetry.

HARV Inlet Research Objectives

Another consideration necessary for the design of the

inlet rake was its intended use during the HARV inlet

research flight test. The primary objectives of the

research were as follows: 

1. Determine whether highly dynamic aircraft

maneuvers result in a significant increase in inlet

pressure distortion levels compared to corresponding steady-flight conditions.

2. Determine whether sources other than spatial timevariant distortion lead to engine aerodynamic

instabilities during aircraft departures.

3. Assess predicted inlet distortion from computational fluid dynamics (CFD) as compared to flight

test measured levels.

These objectives have the dual requirement for gathering accurate inlet measurements during stabilized

and dynamic maneuvers, including aircraft departures.

The instrumentation setup for dynamic maneuvers

would require more attention than the setup for stabilized maneuvers.

Instrumentation

Inlet pressure distortion instrumentation typically has

the requirement for accurately measuring pressure levels at high frequencies. This requirement is typically

achieved by measuring the inlet characteristics with a

dual-probe configuration using low- and highfrequency response sensors. The low-frequency

response probe usually consists of pneumatic tubing

routed through and beyond the rake where it is connected to a highly accurate transducer. This pressure

transducer can either be absolute or differential (with

an accurate reference source). This response system

measures an accurate absolute pressure level. The

high-frequency response probe typically consists of a

miniature transducer mounted at the AIP. This response

system measures the time-dependent component of the

pressure but, normally, not an accurate absolute pressure level.

The HARV research objectives required instrumentation to measure stabilized and dynamic maneuvers.

The typical instrumentation setup described in the previous paragraph would not be adequate to meet the

demanding requirement of measuring inlet characteristics during dynamic maneuvers without introducing a

large amount of measurement uncertainty. The HARV

requires a system which would minimize the effects of

two known drawbacks of the typical instrumentation

setup that affect the ability to measure an accurate pressure level during a dynamic maneuver: pneumatic lag

and thermal zero shift. Pneumatic lag describes the

condition where the pressure signal at the AIP is

delayed in reference to time to the transducer at the end

of the tubing and, therefore, affects low-frequency

response accuracy. Thermal zero shift affects the ability of the low- and, especially, the high-frequency

response transducer to accurately measure the pressure

level at varying inlet temperature conditions. Thermal

zero shift describes the calibration shift of the zero

voltage condition experienced as a pressure-sensing

element of the transducer varies with temperature.

Thus, the requirement was to develop an instrumentation setup that would allow for accurate measurement

of the pressure level and the time-dependent component of the pressure during highly dynamic maneuvers.

The requirements for pressure and temperature

ranges were determined by the flight conditions where

research testing would take place. The HARV research

occurred between an altitude from 15,000 to 40,000 ft

and below Mach 0.9 (fig. 4). The necessary instrumentation pressure range was determined to be 2 to 16 psia.

The temperature was from 395 to 618 °R.

Other instrumentation considerations outlined by

industry through the SAE3

 include that “the frequency

response characteristic of the probe and transducer

combination should be determined with reference to

system accuracy requirements.” GEAE determined that

the highest frequency of interest for the F404-GE-400

engine was 105 Hz. NASA chose to increase the highest frequency of interest to 250 Hz. Industry requirements for the highest frequency of interest vary. NASA
6

required the higher range to allow the HARV inlet

research database to be used by all interested industry

customers. The instrumentation accuracy requirement

needed to meet or exceed the original F/A-18A flight

test. This testing called for the following system accuracy as a percent of reading (2σ): 3.2 percent at 2 psia,

1.3 percent at 5 psia, and 1 percent at 32 psia.1

A final requirement stated for the instrumentation

was ease of maintainability. Transducers needed to be

accessible without engine removal. The ability to

replace the transducer without removing the engine

was also desirable. Additionally, the transducer or

probe configuration should provide the sensing element

of the transducer with protection from foreign object

damage (FOD). 

Aerodynamic and Structural Requirements

Figure 4 shows the aerodynamic design flight envelope. This envelope coincides with the normal operating envelope of the HARV aircraft and was chosen to

allow unrestricted flight with the inlet rake installed.

Inlet research test points were primarily focused at the

low-speed portion of the envelope between Mach 0.3

and Mach 0.4. The worse-case dynamic pressure condition was at a Mach 0.7 at sea level conditions where

the freestream total pressure was 20.4 psia, and the hot

day total temperature was 618 °R.

Requirements for the inlet rake design included additional aerodynamic and structural considerations.

Designing the rake legs and center hub with an aerodynamic shape that minimized airflow disturbance and

with a blockage factor as small as possible was desirable. The HARV blockage goal was to be equal to or

less than the previous F/A-18A inlet rake design,

which had a flow blockage of less than 8 percent.1

Structural requirements included meeting the worsecase pressure loads. These loads include an inlet hammershock overpressure of 20 psi maximum caused by

an engine surge. The HARV structural load limits with

the thrust vectoring vanes installed are 5.4 normal g

loads and 2.0 lateral g loads. Of particular concern was

the requirement to meet dynamic structural requirements as outlined in DFRC document “Process Specification 21-2, Environmental Testing Electronic and

Electromechanical Equipment.”* In particular, the rake

had to be designed stiff enough so that it did not exceed

stress limits at its predominant structural frequencies

*

NASA Dryden internal document, “Process Specification 21-2,

Environmental Testing Electronic and Electromechanical Equipment.” Original released on May, 1968 with current updates until

Apr. 1989.

Figure 4. HARV inlet rake design flight envelope.

.4 .6 


Mach number

0 .2 .8 1.0

10

20

30

40

50 x 103

Altitude,


ft

940086

450 kts
7

within the airframe and engine operating range. NASA

also wanted a number of structural materials considered for the rake including composite materials. Use of

composites could translate into a lighter, more aerodynamic design along with structural tailoring of the rake

arms as compared to traditional designs.

Cost and Time Considerations

Two of the critical requirements that had to be met

for the HARV inlet rake design were low cost and minimal installation time. An evaluation of the rake used in

the original F/A-18 inlet compatibility program indicated costs in excess of $1.5 million and installation

time on the order of 1 year or more. Driving both these

factors were the complexity of using eight cantilever

rakes. Those designs had to be developed, tested, and

installed independently. Installation would require the

aft portion of the inlet duct to be extensively reinforced. Thus, this evaluation quickly revealed that this

approach was not viable for the HARV project. A literature search of past rake designs did not provide a viable alternative approach.4,5

During an early design conception meeting, NASA

and GEAE personnel conceived an alternative

approach in which the eight rake legs would be joined

at the center of the inlet with a hub similar to that of a

wagon wheel. The rake would thus be one piece and

would only have to consider shear loads (no moments)

at its attachment points. This alternative would greatly

simplify the structural installation requirements. Also,

it was envisioned that the rake would slip up the back

of the inlet ahead of the engine and attach to the main

bulkhead near this location. 

Similar wagon wheel rake designs of the past differ

from the NASA and GEAE approach. Previous designs

typically used a modified engine bullet nose which

acted as the hub. The NASA and GEAE concept would

be self-supported with no physical contact with the

engine. This design would minimize vibrational and

force transfer from the engine. The NASA and GEAE

design would greatly simplify installation and aircraft

modifications and significantly reduce aircraft

down time. GEAE has successfully designed, developed and built one prototype and two flight-worthy

rakes for less than $500,000. One flight-worthy rake

had an entire set of high-frequency response transducers included in the cost. The design details and specifications associated with the delivered rake along with

its flight qualification testing are described in the

HARV Inlet Rake System and Flight Qualification

Testing sections.

One significant advantage of the new rake design

includes its transportability to other aircraft which use

F404-GE-400 engines, such as other F/A-18, X-29, and

X-31 aircraft. GEAE is currently under contract to

determine the feasibility of extending the current

design to supersonic flight conditions and to scale this

design to larger inlet diameters.

HARV INLET RAKE SYSTEM

This section describes the HARV Inlet Rake System.

The design and development of the HARV inlet rake

system are detailed in the Rake Description, Pressure

Transducer Selection, and Rake Fabrication subsections. Placement of the rake into the HARV is

described in the Rake Installation subsection. A summary description of the rake system is described in the

Operating Principles subsection.

Rake Description

The HARV rake is similar to a wagon wheel with the

streamlined centerbody acting as the hub, the eight

aerodynamic rake bodies being the spokes, and the

inlet duct being the rim. The load-bearing structure is a

welded Inconel 625* unit that joins the rake bodies and

the central hub into a single piece that is supported by

integral footpads and bolted to the aircraft inlet duct

flange (fig. 5). Each of the eight rake legs contain five

ports located on the centroids of five equal areas of the

flow area. The ports are aligned within 2° of the anticipated steady-flow streamlines. The innermost port is

the only one that had to be angled (5.5°) with respect to

the rake body. All others were already within 2° of the

flow angle.

The central hub also contains an isolated metal

damper ring potted in the polyurethane centerbody

(fig. 6). This configuration allows the damping material

to dissipate vibration energy more effectively than an

all-metal body. The same polyurethane material also

forms the streamlining of both the centerbody and the

trailing edges of the rake bodies. The rake bodies

(or spokes) of the structure are made by forming sheet

metal into the leading edge and sides of the airfoil 

*Inconel 625 is a registered trademark of Huntington Alloy Products Division, International Nickel Company, Huntington Beach,

West Virginia.
8

(a) Front and side views.

Figure 5. Inlet distortion rake.

D

940096

D

C

C

Front view of rake

 Section B

 Section A

Side view of rake

Engine bullet nose

Metal damper ring
9

(b) Sections A through D.

Figure 5. Concluded.

Lead-out tubes

Measurement port


Section A

Mounting installation


Section B

Footpad


Section C

Rake body cross-section


Section D

K-seal

Engine
 front frame

Epoxy filler

Fuselage
 bulkhead

Shim

Footpad

Stagnation shield

Shrink tube

High-frequency
 response transducer

Low-frequency response
 pneumatic tubing

Metal rake body

Elastomer filler

Lead-out tubes

Bolt holes

940173
10

shape. The sheet metal is left open at the trailing edge.

This configuration allows the installation of the sensor

and leadout tubes. 

Flexane 94* is the polyurethane elastomer used as

the potting and flow contour material. This 2 part mix,

room temperature curing, pourable urethane bonds

well to metal, meets the 618 °R maximum temperature

requirement for this application, and has good vibration-damping properties. The rake assembly weighs

approximately 15 lb.

The measurement plane of the sensors is located 4 in.

in front of the bullet nose of the engine with the rakes

oriented 45° from each other starting at 9° counterclockwise off top center (aft looking forward (ALF))

for the right-hand inlet. This orientation satisfies the

mounting accessibility and correspondence to previous

test vehicle requirements. The rake sensors are shielded

total-pressure-measuring sensors consisting of an

impact-mounted high-frequency-response pressure

transducer and a 1/16-in.-diameter, steady-state pressure tube. The stagnation shield configuration was

tested to show its ability to measure the true total input

pressure at varying flow angles. This configuration

*

Flexane 94 is a registered trademark of Devcon Corporation, Danvers, Massachusetts.

allows the sensors to read true pressure levels at yaw

angles from ±25° and pitch angles of 15° and –25°.

Positive angle is toward the engine centerline. Maximum flow blockage is less than 8 percent of flowpath at

the maximum thickness of the rake. This maximum

thickness occurs 1.5 in. behind the AIP.

A composite material was not chosen for the rake

because of increased design costs. Instead, the selection of the Inconel 625 frame with the bonded elastomer was an adequate compromise. The elastomer

allowed the overall weight to be reduced and the rake

struts to be aerodynamically shaped more easily than

an all-metal body. In addition, the elastomer acted as

an excellent damping material.

Pressure Transducer Selection and Installation

The NASA and GEAE team decided to use the same

dual-probe configuration with low- and high-frequency

response probes which had been used during the previous F/A-18A flight test for the HARV inlet rake. The

low-frequency response probe would use 1/16-in. outside diameter (OD) tubing routed through the rake into

the engine bay. A differential transducer with a reference pressure was selected. The DFRC has had a considerable amount of experience using ESP-3205BSL

Figure 6. Floating damper ring.
11

differential transducers (Pressure Systems Incorporated

(PSI), Hampton, Virginia). This transducer unit (fig. 7)

was thermally stabilized to increase accuracy by minimizing thermal zero drift. This stabilization was

accomplished by wrapping the transducer unit in a

temperature-controlled thermal blanket. Another feature of the transducer, used to increase accuracy, was

its ability to perform in-flight calibrations. This inflight calibration allows for any calibration bias error to

be removed during postflight data processing. The

in-flight calibration is accomplished by applying the

reference pressure to both sides of the differential

transducer.

The right-hand ALF engine bay pressure was

selected as the reference pressure for the

ESP-3205BSL transducers. A Sonix PS1019 transducer (Pressure System Incorporated, Hampton, Virginia) was selected to measure the reference pressure.

This pressure was chosen after careful consideration of

a number of possible reference sources. Other possibilities included inlet duct throat wall static pressure, an

inlet rake total-pressure port, and a pressurized tank.

The inlet duct throat wall static pressure and inlet rake

total-pressure ports were eliminated because of their

unsteady pressure levels caused by inlet airflow variation. The pressurized tank was eliminated because of

increased maintainability requirements. The engine

bay pressure was finally chosen because of its steady

rate of change with altitude and its ease of installation

on the aircraft. This pressure is also relatively unaffected by angle of attack. 

The reference pressure range of the Sonix PS1019

transducer (0.4 to 19 psia) met the requirement of 2 to

16 psia. The range of the differential pressure transducer unit was sized to provide the smallest range

about the reference pressure, while allowing for the

expected differential between the engine bay pressure

and the highly distorted total pressures at the inlet rake.

A range of ±5 psid was selected. This minimization of

the differential pressure range allowed for increased

resolution. Both the differential and reference pressure

transducers were readily accessible. The differential

transducer was located in the missile bay under the

right wing ALF at the fuselage. The reference pressure

transducer was located in the right-hand ALF landing

gear wheel well. The differential and reference pressure transducer low-frequency response setups met the

system accuracy requirements during stabilized aircraft

maneuvers. This setup should not be depended upon

during highly dynamic aircraft maneuvers. Instead, the

high-frequency response instrumentation setup should

be used.

Figure 7. Installed HARV low-response differential transducer setup.
12

Each high-frequency response probe used an

XCS-27L-093-20A temperature-compensated pressure transducer (Kulite Semiconductor Products, Inc.,

Leona, New Jersey). This transducer has a 0.093-in.

diameter by 0.50 in. length with an absolute pressure

range from 0 to 20 psia. The frequency response of this

transducer met the 250 Hz requirement. Forty of these

transducers were mounted at the rake measurement

ports in close proximity to the low-frequency response

pneumatic tubing (fig. 8). Each high-frequency

response transducer had a protective FOD screen over

the sensing element. This transducer was selected

because of its ability to minimize thermal zero drift

through passive temperature compensation. To further

increase the accuracy of the transducer measurement, a

series of pressure calibrations were performed over the

entire required pressure and temperature range, up to

20 psia and at 395, 425, 460, 535, and 610 °R. Along

with the measured engine inlet temperature, these calibrations would allow for any remaining zero thermal

drift to be removed during postflight data processing.

The differential transducer low-frequency response

measurements will be used to verify the pressure levels

of the temperature-compensated pressure transducer at

stabilized conditions. This high-frequency response

transducer setup allows for the accurate measurement

of high-frequency pressure levels during highly

dynamic maneuvers and meets the system accuracy

requirements.

The temperature-compensated pressure transducers

were installed in the inlet rake 0.125-in. OD carrier

tubes. These tubes have a 0.099-in. ID counterbore to

receive the transducers. Installation was accomplished

by feeding the transducer electrical leads through the

0.125-in. tube from the sensor end, coating the back of

the transducer with room temperature vulcanizing

(RTV) silicone adhesive, and inserting the transducer

into the counterbore. Next, the transducer was covered

with heat-shrinkable tubing, and the silicone adhesive

was allowed to cure. Then, the shrink tube was

trimmed flush with the transducer face. The lead exit

was also sealed with the silicone adhesive and shrinkable tubing. This arrangement provided a secure

mounting for the transducers but still allowed removal

and replacement while the rake was still in the aircraft

in case of transducer failure.

Rake Fabrication

Fabrication of the rake unit began with forming the

sheet metal rake body airfoils, drilling them for the

sensor tubes, and welding them to a machined hub. The

Figure 8. HARV inlet rake sensor setup with high-response, temperature-compensated pressure transducer

installed.
13

footpads were then machined, formed to the duct

radius, and welded to the rake bodies. After stress

relieving, the welds were inspected using the florescent

penetrant inspection technique. (This nondestructive

inspection technique detects cracks by applying a penetrating fluorescent solvent to the material.) Sensor

tubes and stagnation shields were then brazed in place

with a gold and nickel braze. The bend radii in the

leadout tubes were kept as large as possible to facilitate

installation of the transducer leads later in the assembly

(fig. 9). A damper ring was then installed around the

hub and rake struts. Welds and brazes were reinspected

after the metal fabrication portion of the assembly concluded. Strain gages were applied to the hub and rake

bodies in the areas that would later be covered with

polyurethane elastomer (fig. 10). The rake bodies were

then filled and the trailing edges of the rakes were

formed with the elastomer. The center hub form was

then cast in place completing the airfoil blending of the

rakes and centerbody. Figure 11 shows the completed

rake assembly.

Rake Installation

Installation of the rake on the aircraft was accomplished after the right-hand engine (ALF) had been

removed. The minimal airframe modifications that

were required consisted of 1/4-in. bolt holes being

drilled in the inlet duct flange aft of the rear bulkhead.

These 16 holes were located by trial positioning the

rake in the inlet duct, confirming that none of the bolt

holes would interfere with any bulkhead bracing, and

transfer-punching the hole locations through the rakemounting footpads. Then, the rake assembly was

removed, and the holes were drilled in the flange using

a hand-held drill. Backup washers were then placed at

each hole on the outside of the duct and expoxied in

place to provide a solid and flat surface for seating the

nuts of the rake-mounting bolts (fig. 5). This seating

was done by inserting a nylon mandrel through the

washer and rake-mounting holes in the duct to position

the washers on an aluminum-filled epoxy base that

filled the gap under the washer. These mandrels were

removed after the epoxy had cured overnight. 

EC 92 5192-3

Figure 9. Rake leg instrumentation tube leadouts.
14

S 3405-2

Figure 10. Stain-gaged hub of distortion rake.

S 825-3

Figure 11. Completed prototype distortion rake with polyurethane elastomer rake leg fairings and centerbody.
15

Testing showed that this configuration could withstand over 200 in.-lb torque applied to the mounting

bolts without deforming the washer, epoxy, or flange.

The originally planned rework to the exit lip of the inlet

duct was eliminated to retain as much of the original

integrity of the flange as possible. To accomplish this

goal, the leadout tubes of the rakes were reformed to

exit the duct just behind the lip, and a support shim was

made from polyurethane elastomer to support the tubes

under the K-seal. This seal around the inlet duct exit

provides the transition from the inlet duct to the engine

front-frame lip. The K-seal was contoured to allow the

rake tubes to pass freely under it (fig. 12).

The rake was placed in the duct. An even fit for each

rake strut was achieved by placing shims between the

rake metal footpad and the inlet duct wall. The rake

assembly was then installed in its final position. Room

temperature vulcanizing silicone adhesive was placed

on each side of the fiberglass and polyurethane elastomer footpads. The metal footpads were then bolted in

place, and the silicone adhesive was cured. This configuration gave the assembly a firm, elastomer-damped

mounting.

The K-seal was then bolted in place on the aft duct

flange. The electrical leads for the pressure transducers

and the steady-state pressure tubes were routed to their

respective connector locations. The rework to the

airframe, installation of the rake assembly, modification to the K-seal, and installation of the seal ready for

lead routing was accomplished in less than 3 days.

Operating Principles

Single-piece construction of the HARV rake assembly resulted in a redundant structure that allows each

spoke of the wheel configuration to share the load of

the structure through the hub and transfer it to the airframe duct as shear loads to the bolts in the duct flange.

These rakes are fastened at the footpad on the duct

wall, allowing the duct flange to support the rake without inducing any bending load in the sheet metal wall

of the duct.

Vibratory energy within the rake struts is damped as

the flex and twist motion of the open-backed rakes put

the polyurethane elastomer in shear. The elastomer is

bonded to the sheet metal. Energy at the hub from the

rake struts is dissipated using the action of the damping

ring imbedded within the elastomer.

FLIGHT QUALIFICATION TESTING

Flight qualification testing consisted of three phases:

laboratory, ground testing, and flight testing. The laboratory phase determines the baseline structural and

Figure 12. Instrumentation tubes passing under K-seal at exit lip of HARV inlet.
16

vibrational characteristics of the rake. This phase consists of rake modeling using NASA Structural Analysis

(NASTRAN) along with Zonic ping (Zonic Corporation, Milford, Ohio) and vibrational shake table

(Unholtz-Dickie Corporation, Wallingford, Connecticut) testing on a prototype rake. Results of this phase

are compared to the environmental conditions that the

rake will experience in operation. Then, baseline

results from the laboratory tests are used for comparison with the results from the installed ground testing

and flight testing phases. Each phase is described in the

following subsections.

Laboratory Testing

Computer modeling of the rake structure using the

NASTRAN program was used to identify the possible

vibration modes and the approximate frequencies of

those modes. The computer model identified 15 potential vibrational modes between 0 to 7000 Hz. Most of

these modes could not be driven to significant levels

during subsequent laboratory, ground, or flight testing.

The laboratory testing used a prototype rake instrumented with strain gages. This rake was installed in an

inlet test fixture manufactured from a modified F/A-18

aircraft inlet. This test phase consisted of ping tests and

vibration table testing. Objectives of vibration testing

with the instrumented prototype were to:

1. Measure the vibration characteristics of the rake

with axial and transverse excitation.

2. Determine the rake stress distribution at an established maximum rake stress level.

3. Excite the rake randomly at the flight qualification

excitation level for a specified period without damaging the rake.

4. Excite the rake sinusoidally at its resonant frequency at a specified excitation level to empirically

define the rake endurance limit.

5. Select the critical strain gage locations for ground

test monitoring based on the strain-gage-derived

stress distribution.

Twenty-eight dynamic strain gages were applied to

two (of eight) rake struts. That is, 14 each on 2 struts

spaced 90° with respect to each other and circumferentially oriented in the inlet so that one of the

instrumented struts would be exposed to maximum excitation during the transverse vibration test. In

addition, accelerometers were mounted on the inlet test

fixture and on the vibration table. This instrumentation

would verify that there was no displacement amplification caused by resonance of the test fixture. This instrumentation was also used to verify the predicted

damping characteristics of the rake structure.

Strain gage excitation and signal conditioning were

accomplished by a GEAE designed and fabricated constant current excitation, alternating current (ac) coupled amplification system. Data were recorded on an

X-Y plotter to show overall strain levels as a function

of frequency. The instrumented inlet test fixture and

rake assembly was mounted on a 17,500 lb capacity

vibration table.

The ping test uses an instrumented hammer for

applying a force to the rake structure and measures the

resultant motion of the rake. This method identified

four primary modes at which the rake might be driven.

These modes were axial, torsion, strut bending, and

second axial.

Vibration table testing consisted of three procedures:

axial-sinusoidal, transverse-sinusoidal, and transverserandom excitation. The axial- and transverse-sinusoidal

vibration test procedures required several steps. First,

the rake was excited axially with respect to the engine

centerline. Next, the rake was excited transversely

using sinusoidal-driving forces over the frequency

from 10 to 2000 Hz at 1-g excitation to locate the resonant frequencies of the rake. Then, the axial and transverse stress distributions were obtained by dwelling at

the resonant points and increasing the input acceleration levels until the established maximum stress level

was reached on one of the strain gages. Subsequently,

the rake was exposed to transverse-random excitation

(nonsinusoidal) over the frequency range of 15 to

2000 Hz at 18.4-g excitation (flight qualification level).

Finally, the rake was vibrated at resonance at an excitation level necessary to verify the rake endurance limit. 

Rake maximum stress levels were established by taking a number of factors into consideration. These factors included the material properties of the rake

structure, placement and number of strain gages

installed, stress concentrations, and safety margin. The

stress limits of 30,000 and 50,000 lb/in2 peak-to-peak

(PSIPP), dependent on strain gage location, were used. 

The axial-input frequency resonant sweep test

showed an axial mode at 147 Hz with 1-g excitation.

The input acceleration level was then increased until

the maximum stress level of 30,000 PSIPP was

reached. The maximum strain gage outputs were
17

located on the leading edge and trailing edge of the

rake foil at the hub. A decrease in the axial mode frequency to 129 Hz at 3.8-g excitation was observed. The

maximum response of 30,000 PSIPP at 3.8-g axial

input meets the damping and structural design requirements of the rake. The lower resonance frequency

(129 Hz) at the higher vibration level is a result of the

soft mount interaction of the fixture and the

elastomer-damping configuration. The soft mount of

the test fixture is not representative of a full aircraft

inlet. This soft mount was improved by adding bags of

lead shot to the fixture. The frequency shift of the rake

and the lower 1/rev crossover on the fan Campbell diagram (fig. 13) at the higher excitation level (147 Hz) is

not critical to the highly damped rake structure. 

Figure 13 shows the various vibration mode frequencies of the rake (axial, torsion, strut bending, and second axial) and lines representing multiples of fan rotor

speed frequencies (1/rev, 2/rev) are superimposed

across the entire fan rotor speed range. Any intersection allows for fan rotor speeds to be identified where

rake resonance frequencies are likely to occur. For

example, the rake axial mode resonance (147 Hz) is

likely to occur at a fan rotor speed of approximately

8500 rpm. This low fan rotor speed means less energy

which can be imparted onto the rake. Experience has

shown that axial excitation has never been a problem

for an inlet rake installed directly in front of an engine.

The transverse-sinusoidal input sweep showed a resonant frequency of 123 Hz (a combination of rake axial

mode and fixture resonance) and a strut bending mode

of 240 Hz with the stress distribution remaining constant with increasing input levels. A transverse excitation of 7.8 g was reached at the maximum stress level

of 30,000 PSIPP. 

The transverse-random test provided a nonsinusoidal

input to test the hardware for interaction and structural

flaws caused by 18.4-g random input (flight qualification level) in the 15 to 2000 Hz frequency band. The

test had to be performed for a minimum of 20 min to

meet the flight qualification requirements. This test was

performed for 26 min. Stress level and distribution data

from the nonsinusoidal excitation are not related to the

transverse sinusoidal sweep data. During the random

tests, strain gages located on the leading and trailing

edges of the rake foil near the footpad provided the

highest relative output signals. Spectral data from the

random excitation again indicated resonant frequencies

of 129 Hz (rake axial with fixture resonance) and

240 Hz (strut bending). Based on the data from the

vibration table tests, 12 strain gages at the leading and

trailing edges of the airfoil at both the hub and the

Figure 13. Campbell diagram showing rake vibration responses.

800

700

600

500

400

300

200

100

0

Frequency,


Hz

4 5 6 7 8 9


rpm x 1000

8500

11 12 14 13

147

Idle Takeoff Fan rotor speeds where rake


 resonance frequencies are


 likely to occur

1/rev

2/rev

3/rev

4/rev

Strut bending

Torsion


Axial


Axial with fixture resonance

Second axial 

625 Hz

262 Hz

123 Hz

162 Hz

940095

10
18

outer ends of the rake strut were selected to be monitored during the ground and flight test program. This

test resulted in no structural damage to the rake.

Empirical determination of the structural endurance

limit was attempted by exciting the rake at its resonant

frequency at the required 10-g, transverse-sinusoidal

level. The 10-g, transverse excitation showed a stress

level of 41,000 PSIPP. The 10-g input was to be held

until the rake structure failed or reached 107

 vibration

cycles. The duration would define the endurance limit.

The resonant frequency changed from 122 to 117 Hz

during the test. This change probably resulted from the

heating effect on modulus of the elastomer potting

material. The 10-g input level was extremely severe on

the inlet test fixture. The test was shut down after 2 hr

of dwell time at 10 g to avoid catastrophic damage to

the test fixture. Empirical determination of the rake

endurance limit could not be achieved because the

high-damping characteristic of the rake made it significantly more durable than the inlet test fixture. The inlet

test fixture could not tolerate excitation levels necessary to damage the rake. The rake system showed no

signs of structural damage for the duration of the test

and was confirmed with a posttest X-ray and visual

inspection of the rake hardware.

The results of the vibration shake table test have been

compared to the NASTRAN computer model predictions and to the ping test results. These dynamic analysis results are listed next.

The rake vibration testing confirmed the analytical

design calculations. The laboratory results from the

ping and vibration table tests formed the baseline predominant frequencies for comparison to ground and

flight testing. A frequency shift limit of ±20 percent

was determined to be reasonable. Shifts greater than

this limit indicate either a damaged or failed structural

component. The stress limits of 30,000 and

50,000 PSIPP were satisfactory for use during ground

and flight testing. Use of these limits depended on the

location of the strain gage. 

Ground Testing

A ping test was performed on the inlet rake installed

in the F/A-18 HARV while in the hangar. This test was

used to verify the results gathered during laboratory

testing. Frequencies showed good agreement with the

past laboratory ping test results.

The installed rake was ground tested on the F/A-18

HARV with the aircraft tied down. Twelve strain gages

were monitored while the right-hand engine was operated through its full range with a slow acceleration

from idle power to full maximum afterburning and a

slow deceleration to idle power. This procedure

allowed predominant frequencies to be identified over

the entire fan rotor speed range. Stress levels and spectral data were taken during the test. 

The spectra showed predominant frequencies of 153

and 162 Hz corresponding to axial and torsional frequencies and a broad band about 250 Hz at a lower

level corresponding to strut bending modes. These predominant frequencies showed consistent agreement

with the laboratory baseline frequencies. Maximum

stress levels during the test were approximately 10 percent of limit. Again, the strain gages showing the greatest stresses were located at the leading and trailing

edges of the airfoil, at the hub, and at the outer end of

the rake strut.

Flight Testing

The four most active strain gages were located at the

leading and trailing edges of the airfoil, at the hub, and

at the outer end of the rake strut, and were selected for

monitoring during rake qualification flights. Two were

monitored by telemetry in real time, and two were

recorded on the aircraft for postflight analysis.

Qualification runs for the rake consisted of flight

points to give maximum unsteady loads (an angle of

attack of 60° at an altitude of 20,000 ft), temperature

and pressure (Mach 0.7 on the deck), and combination

of temperature, pressure, and unsteady loads (Mach 0.9

at an altitude of 18,000 ft) within the HARV flight

envelope. The latter two points were at the limits of the

HARV flight envelope. The first point was flown to a

high-angle-of-attack condition. The latter two points

were obtained at the maximum maneuver-loading

limit.

Mode

NASTRAN

model, Hz

Ping 

test, Hz

Vibration 

table, Hz

Axial 140 155 147

Torsion 147 162 N/A

Axial with 

fixture 

resonance

N/A N/A 123–129

Strut bending 250 236–262 240

Second axial 594 625 N/A
19

The spectra of the rake vibrations remained consistent with the ground and laboratory tests. The stress

levels observed were 26 percent of limits during highangle-of-attack flight and less than 10 percent of limits

during maximum maneuver-loading flight. The maximum stress limits observed were 30 percent of limits

during aircraft takeoff. Based on the laboratory,

ground, and flight test results, the rake is now fully

cleared for conducting flight research within the entire

HARV flight envelope with no restrictions.

CONCLUDING REMARKS

An improved cost and installation-time saving inlet

distortion pressure rake was successfully designed,

built, and validated for flight testing on the F/A-18

HARV research aircraft. The innovative design consists

of a one-piece, wagon wheel approach that resulted in

ease of installation with minimal aircraft modifications.

Design advantages include lightweight, high strength,

low structural resonance, low flow blockage, and easy

transducer removal and replacement. A prototype of

the new rake was environmentally tested in a laboratory where it passed all vibration structural requirements. Ground test verified the expected frequency

response predicted from laboratory test. Flight qualification was completed and the rake is now cleared for

flight testing on the HARV aircraft. All stress levels

observed during ground and flight qualification were

less then 30 percent of limits.

The simple design approach and creative use of lightweight materials such as elastomers resulted in a superior rake design at a significantly lower cost than previous design approaches. The rake has promise of ease of

transportability to other F404-GE-400 equipped aircraft. Further design efforts are expected to allow the

design flight envelope to be expanded to supersonic

conditions and allow scaling to other aircraft inlets.

The cost and time saving solution to the inlet rake

design will allow the F/A-18 HARV to conduct

important inlet research at flight conditions never

before explored.

ACKNOWLEDGMENT

The DFRC, LeRC, and GEAE team members wish to

recognize posthumously the creative energy brought by

co-author Leon Lechtenberg to this project. It is unfortunate that his tragic and untimely death did not allow

him to see the fruit of his labor that resulted in successful ground and flight clearance of this novel rake

design.

REFERENCES

1

Amin, N.F. and D.J. Hollweger, “F/A-18A Inlet/Engine

Compatibility Flight Test Results,” AIAA-81-1393,

July, 1981.

2

Inlet Total-Pressure-Distortion Considerations for GasTurbine Engines, Society of Automotive Engineers

Aerospace Information Report, AIR 1419, May 1983.

3

Gas Turbine Engine Inlet Flow Distortion Guidelines,

Society of Automotive Engineers Aerospace Recom-mended

Practice, ARP 1420, Mar. 1978.

4

Holzman, Jon K. and Gordon A. Payne, Design and Flight

Testing of a Nullable Compressor Face Rake, NASA TN

D-7162, Jan. 1973.

5

Farr, A.P. and G.A. Schumacher, “System For Evaluation

of F-15 Inlet Dynamic Distortion,” Symposium on

Instrumentation for Airbreathing Propulsion, MCAIR

72-043, Sept. 1972.
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NSN 7540-01-280-5500 Standard Form 298 (Rev. 2-89)

Prescribed by ANSI Std. Z39-18

298-102

Design and Development of an F/A-18 Inlet Distortion Rake: A Cost and

Time Saving Solution

WU 505-68

Andrew J. Yuhas, Ronald J. Ray, Richard R. Burley, William G. Steenken,

Leon Lechtenberg, and Don Thornton

NASA Dryden Flight Research Center

P.O. Box 273

Edwards, California 93523-0273

H-2078

National Aeronautics and Space Administration

Washington, DC 20546-0001 NASA TM-4722

An innovative inlet total-pressure-distortion measurement rake has been designed and developed for the

F/A-18 A/B/C/D aircraft inlet. The design was conceived by NASA and General Electric Aircraft Engines

personnel. This rake has been flight qualified and flown in the F/A-18 High Alpha Research Vehicle at NASA

Dryden Flight Research Center, Edwards, California. The eight-legged, one-piece, wagon wheel design of the

rake was developed at a reduced cost and offered reduced installation time compared to traditional designs.

The rake features 40 dual-measurement ports for low- and high-frequency pressure measurements with the

high-frequency transducer mounted at the port. This high-frequency transducer offers direct absolute pressure

measurements from low to high frequencies of interest, thereby allowing the rake to be used during highly

dynamic aircraft maneuvers. Outstanding structural characteristics are inherent to the design through its

construction and use of lightweight materials.

F-18 aircraft; Flow measurement; In-flight monitoring; Inlet flow; Pressure

measurement; Pressure sensors A03

21

Unclassified Unclassified Unclassified Unlimited

October 1995 Technical Memorandum

Available from the NASA Center for AeroSpace Information, 800 Elkridge Landing Road, 

Linthicum Heights, MD 21090; (301)621-0390

Presented as AIAA 94-2132 at the 7th Biennial AIAA Flight Test Conference, Colorado Springs, Colorado, June 20–23, 1994. A.J. Yuhas, PRC Inc.,

Edwards, California; R.J. Ray, Dryden Flight Research Center, Edwards, California; R.R. Burley, NASA Lewis Research Center, Cleveland, Ohio;

W.G. Steenken, L. Lechtenberg, and D. Thornton, General Electric Aircraft Engines, Evendale, Ohio.

Unclassified—Unlimited

Subject Category 02