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NASA TECHNICAL. NOTE

NASA

D-5877

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NASA TN D-6178

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ACOUSTIC AND AERODYNAMIC
PERFORMANCE OF A 6-FOOT-DIAMETER
FAN FOR TURBOFAN ENGINES

III - Performance With Noise Suppressors

by Edward J. Rice, Charles E. Feiler,
and Lor en W. Acker

Lewis Research Center
Cleveland, Ohio 44155

NATIONAL AERONAUTICS AND SPACE ADMINISTRATION . WASHINGTON, D. C. • FEBRUARY 1971

TECH LIBRARY KAFB. NM

1. Report No. 2. Government Accession No.

NASA TN D-6178

4. Title and Subtitle

ACOUSTIC AND AERODYNAMIC PERFORMANCE OF A

6 -FOOT-DIAMETER FAN FOR TURBOFAN ENGINES

ni - PERFORMANCE WITH NOISE SUPPRESSORS

7. Author(s)

EdwardJ. Rice, Charles E. Feiler, and Lor en W. Acker

9. Performing Organization Name and Address

Lewis Research Center

National Aeronautics and Space Administration

Cleveland, Ohio 44135

12. Sponsoring Agency Name and Address

National Aeronautics and Space Administration
Washington, D. C. 20546

15. Supplementary Notes

3. Recipient's Catalog No.

5. Report Date

February 1971

6. Performing Organization Code

8. Performing Organization Report No.

E-5878

10. Work Unit No.
126-61

11. Contract or Grant No.

13. Type of Report and Period Covered

Technical Note

14. Sponsoring Agency Code

16. Abstract

Inlet and exhaust noise suppressors for a 6-ft- (1. 83-m-) diameter fan for a high-bypass- ratio
turbofan engine were tested. The perforated-plate-on-honeycomb suppressors provided a
much broader band noise attenuation than was predicted. Perceived noise level attenuations
obtained due to the suppressors were 13 and 12 PNdB for simulated approach and takeoff
conditions, respectively. The theory used for the design of the suppressors is discussed. In
general, the theory predicts the frequency for peak attenuation but underpredicts the peak
attenuation amplitude. For frequencies above and below peak, the observed attenuations are
more than predicted. Degradations of aerodynamic performance caused by the noise suppres-
sors were smaller than the experimental errors, which were estimated to be 2 percent.

17. Key Words (Suggested by Author(s) )

Acoustic attenuation; Acoustics; Acoustic im-
pedance; Aircraft noise; Noise reduction;
Sound transmission; Turbofan engines

18. Distribution Statement

Unclassified - unlimited

19. Security Classif. (of this report)

Unclassified

20. Security Classif. (of this page)

Unclassified

21. No. of Pages

22. Price
$3.00

For sale by the National Technical Information Service, Springfield, Virginia 22151

ACOUSTIC AND AERODYNAMIC PERFORMANCE OF A 6HF00T-

DIAMETER FAN FOR TURBOFAN ENGINES

III - PERFORMANCE WITH NOISE SUPPRESSORS

by Edward J. Rice, Charles E. Feiler, and Loren W. Acker

Lewis Research Center

SUMMARY

Inlet and exhaust noise suppressors for a 6-foot- (1. 83-m-) diameter fan for a high-
bypass-ratio turbofan engine were tested. The perforated-plate-on-honeycomb suppres-
sors provided a much broader band noise attenuation than was predicted. Perceived
noise level attenuations obtained due to the suppressors were 13 and 12 PNdB for simu-
lated approach and takeoff conditions, respectively.

The theory used for the design of the suppressors is discussed. In general, the
theory predicts the frequency for peak attenuation but underpredicts the peak attenuation
amplitude. For frequencies above and below peak, the observed attenuations are more
than predicted.

Degradations of aerodynamic performance caused by the noise suppressors were
smaller than the experimental errors, which were estimated to be 2 percent.

INTRODUCTION

Current four-ergine aircraft using ducted fans of the 707/DC-8 class produce about
120 perceived noise decibels (PNdB) at standard measuring points during both takeoff
and landing. The NASA Quiet Engine Program was initiated to demonstrate the feasi-
bility of reducing this noise 15 PNdB by engine design changes and perhaps another
10 PNdB by the use of absorptive duct liners. In support of these goals, a facility
described in reference 1 has been built for determining the performance and noise char-
acteristics of full-scale fans. Performance and noise data without acoustic treatment
are reported in reference 2 for the first 6-foot- (1.83-m-) diameter fan tested.

This report describes the noise reduction attained by the use of acoustic liners in

the inlet and exhaust ducts of the fan. The use of acoustic treatment to suppress engine
noise has recently received considerable attention. Numerous experimental and analyti-
cal studies have been performed to develop liner design methods (e. g. , refs. 3 and 4).

The liners used in this study were of a perforated-plate-on-honeycomb construc-
tion. The basis for their design is discussed and a comparison of experimental and
theoretical results is presented. Estimates are made of the flyover noise characteris-
tics (without core engine noise) of an airplane having this fan and treatment on four
engines.

SYMBOLS

S/A

6
V

P
a

max

backing depth of liner resonators, ft (m)

speed of sound, ft/sec (m/sec)

circular-duct diameter or annular-duct height, ft (m)

perforated- plate hole diameter, ft (m)

maximum possible sound power attenuation for a given -q and L/D, dB

frequency, Hz

length of acoustic treatment, ft (m)

average steady flow Mach number

2 2

acoustic pressure, Ibf/ft (N/m )

ratio of acoustic treatment area to duct cross-section area

perforated- plate sheet thickness, ft (m)

normal gas velocity at a lined wall, ft/sec (m/sec)

orifice gas velocity in Helmholtz resonator, ft/sec (m/sec)

orifice-end correction (see eq. (7)), ft (m)

frequency parameter, Df/c

specific acoustic resistance (see eqs. (2) and (3))

nonlinear specific acoustic resistance (see eq. (4))

sound wavelength, ft (m)

2 2

gas kinematic viscosity, ft /sec (m /sec)

gas density, Ibm/ft (kg/m )

open-area ratio (orifice area to wall area)

specific acoustic reactance (see eq. (2))
angular frequency

APPARATUS AND PROCEDURE
Fan Description

A cutaway view of the fan is shown in figure 1. The single-stage fan has a large
rotor-stator spacing (3.6 rotor chords) and is without inlet guide vanes. The fan is
driven by electric motors through the shaft shown emerging from the fan inlet.

The detailed design of the fan is presented in reference 1. The following fan
parameters are given here for convenience. Aerodynamic parameters are corrected
to standard sea-level atmosphere of 518. 7° R (288. 2 K) and 2116. 2 pounds per square
foot (1.013x10^ N/m^):

Rotor tip diameter, in. (m) 71.81 (1.8240)

Stator tip diameter, in. (m) 67.94 (1.7257)

Rotor tip speed at 3533 rpm (cruise design corrected value),

ft/sec (m/sec) 1107 (337.4)

Design stagnation pressure ratio 1.5

Design weight flow (corrected value), Ibm/sec (kg/sec) 873 (396)

Rotor hub-tip ratio (inflow face) 0. 50

Stator hub-tip ratio 0.59

Rotor to stator spacing (trailing to leading edges), in. (cm) 20 (50.8)

Number of blades:

Rotor 53

Stator 112

Chord length, in. (cm):

Rotor 5. 5 (13. 97)

Stator 2.69 (6. 83)

Noise Suppressor Construction

The inlet and exhaust noise suppressors can be seen in figure 1. The inlet suppres-
sor consists of a lined outer cowl and three splitter rings with acoustic lining on both
sides. The exhaust suppressor has only the lined outer cowl and centerbody.

>(^

r Centerbody

Exhaust cowl

Rotor-stator support ring

Splitter rings
(acoustic
lining only) -v

Fan drive shaft

Perforated plate 7

Partitioned cavity

C0-l(XB6-02

Figure 1, - Cutaway view of fan and suppressor assembly.

-58 in. (1.473 m)

x^///^//////j'/////////'j'/////j'y////^j'///y///y////'/ //////////////////A \ ^ _

-34 in. (0.864 m) - ^^ -I

2&5in. ]
(0.724 m)

20 in.
(0.508 m)

12 in.
(0.305 m)'

36.9 in.
(0.937 m)

Location

Open-area
ratio,

Perforated-plate
hole diameter, d

in.

Inlet
Exhaust

0.025
.08

0.032
.050

0.81
1.27

Surface

Honeycomb
backing depth

in.

1
2
3

0.88
.20
.68

2.24

.51

1.73

34.03 in.
(0. 864 m)

Flow

/"I

19.75 in. '
(0. 502 m)

!in. (2.235 m)

Figure 2. - Suppressor dimensions and materials. All perforated-plate sheet, 0.02-inch- (0. 51-mm-) thick
aluminum; all honeycomb, 3/8 inch (0.95 cm) hexagonal.

The suppressor dimensions and the materials used in the acoustic liners are shown
in figure 2. The liner is constructed with a perforated aluminum sheet bonded to a
honeycomb backing. All facing materials are 0. 020- inch- (0. 51-mm-) thick perforated
aluminum sheet metal. The three surface treatments indicated on the inlet differ only
in the thickness of the honeycomb backing material.

A cross section of an inlet splitter ring is shown in figure 3, where the two different
honeycomb thicknesses are apparent. This construction was used to broaden the fre-
quency range of noise suppression. Each passage in the inlet was thus bounded by two
surfaces tuned to two different frequencies.

The construction of the noise suppressors followed the technique outlined in refer-
ence 3 (p. 63). A modified epoxy adhesive film, supported by a synthetic fabric carrier,
is inserted between the honeycomb core and the septum of the splitter ring. The bond is
then made by oven curing the above assembly. An adhesive film without a fabric carrier
is then applied to the honeycomb. Heat is applied to the adhesive and it coagulates into a
bead along the honeycomb edges. The perforated sheet- metal facing plate is then ap-
plied to the honeycomb and bonded by oven curing.

Ill I nil

Figure 4. - Inlet duct with splitter rings.

|C-70-1957
Figure 3. - Cross section of inlet splitter ring.

The assembled inlet noise suppressor is shown in figure 4. The fan drive shaft
enters through the center, reducing the size of this passage.

Data for four configurations are reported herein. These result from combinations
of two inlet lengths and two exhaust nozzle areas. The length of the inlet is the distance
from the bellmouth flange to the front ec^e of the rotor-stator support ring (see fig. 1).
The leading edge of the rotor at the root lies 5. 2 inches (0. 132 m) behind this support-
ring front edge. The short inlet is 60. 5 inches (1. 537 m) and the long inlet is 101. 5

2
inches (2. 578 m). The standard nozzle area is 1895 square inches (1.223 m ), which is

97 percent of the design value. The 10-percent-oversized nozzle has an area of 2150

2
square inches (1. 387 m ) and is 110 percent of the design value.

Noise Suppressor Design

The noise suppressor design philosophy is contained in the answers to three ques-
tions. First, how much and in what configuration can the noise-absorbing material be

applied to the fan ducts without excessive losses due to flow blockage? Second, what
linii^ material impedance will result in the maximum noise attenuation within the im-
posed geometry restrictions? Finally, what wall construction will provide the optimum
wall impedance at the frequency of the maximum noise?

For an estimate of the necessary noise suppressor treatment, the theory of refer-
ence 5 is used. This theory describes the propagation of an initially plane pressure
wave in a cylindrical duct without mean flow. F^ure 5 (ref. 5) relates maximum pos-
sible sound power attenuation to the duct length-to-diameter ratio L/D and frequency
parameter tj where

(1)

Although figure 5 was derived from the circular-duct theory, an estimate for the attenua-
tion in annular ducts can be made by using one-half the ordinate. This factor results
from the consideration of duct-lined area to cross-sectional area ratio. Eqiiation (8)
(p. 17) shows the sound power attenuation to be proportional to this ratio (an approxi-

lOOr—

- 60

Duct length-to
diameter ratio,
l/D

-- 2
3

I
.2

I .1.1 L

.6 .8 1 2

Frequency parameter, 17

Figure 5. - Dependence of maximum possible sound power attenuation on frequency
parameter (tj) for various duct length-to-diameter ratios.

mation). This ratio is twice as large for a circular duct as for an annular duct when D
is interpreted as the distance between lined surfaces in the annulus.

The curves of figure 5 can be used to give a quick estimate of the performance of a
proposed liner configuration. For the following estimates a frequency of 3500 hertz
(X = 3. 84 in. , 9. 75 cm) is assumed. First, consider just lining the 6-foot- (1. 83-m-)
diameter inlet for a length of 6 feet (1. 83 m). The duct has an L/D of 1 and an ry of
18. 75. By extrapolating in figure 5 and applying the factor of 1/2 to the ordinate, an
attenuation of less than -1 decibel is obtained. Clearly, splitter rings or lined struts
must be used to increase the effective L/D and decrease rj. Next, consider three
splitter rings of 3-foot (0. 91-m) length in the fan inlet. Now D « 8 inches (20. 3 cm),
L/D « 4, and 77 ~ 2.1. From figure 5 the maximum possible sound power attenuation
is -19 decibels (the ordinate in fig. 5 multiplied by 1/2 as previously discussed). If this
attenuation could be achieved at 3500 hertz by the proper wall construction, it would be
sufficient for the purposes of this experiment. More noise attenuation could be obtained
by increasing the number of splitter rings, but three was considered as a reasonable
compromise between noise attenuation and flow losses.

The spacings of the three inlet splitter rings are shown in figure 2. Due to the en-
largement of the hub near the rotor, the inner rings had to be made shorter than the
outer rings. The spacing between splitters was thus adjusted by using figure 5 to pro-
vide approximately the same maximum sound power attenuation (-dB_,„^) for all pas-
sages.

Once the overall geometry of the suppressor is decided, the optimum wall imped-
ance can be determined from figure 6. These curves, obtained from reference 5, show
the relation between L/D, 77, and wall impedance required to produce the attenuations of
figure 5. Continuing the previous example with L/D « 4 and 77 « 2. 1, the wall specific
resistance should be 2.4 and reactance should be -1.25.

The optimum wall impedance can then be corrected for a finite steady flow by
dividing it by 1 + M according to reference 6. The Mach number is considered posi-
tive for the exhaust duct and negative for the inlet, where the sound propagation opposes
the steady gas flow.

Thus for the same L/D and 77 the inlet must have a larger resistance and more
negative reactance than the exhaust duct. In reference 6 it was shown that the curves in
figure 5 were not altered by a steady, uniform gas flow.

Finally, when the required wall impedance is known at a particular frequency, the
wall construction must be specified to produce the desired impedance. The following
relations are given to relate acoustic impedance to wall construction for a perforated
plate mounted over a backing cavity. The equations, or in some cases the data from
which the equations were determined, are available in the literature and are reviewed
in reference 7.

10(—
8-
6-

•Constant duct length-to-
diameter ratio, L/D
Constant frequency, tj

L/D = 5 -.

T)=10

1,1,1
.6 .8 1
Wall reactance,

1
2

, 1

1 1 1

8 10

Figure 6. - Locus of maximum sound power attenuation in wall impedance plane.

The specific acoustic impedance is defined as

pcV,

= + IX

(2)

The resistance is given by

e =

Uit)

1 +-\ + e

CTC

(3)

where

'NL

-^- (1 -1- 6.67 M)

CTC

(4)

The first term in equation (3) is the linear resistance of a Helmholtz resonator array
due to viscous dissipation in the oscillatory boundary layers at the walls and in the ori-

fice In the absence of steady grazing flow and for very small sound levels, this is the
entire acoustic resistance of the wall.

The second term in equation (4) is an empirical expression to account for the in-
crease in acoustic resistance of the wall due to grazing flow. The peak orifice velocity
accounts for the nonlinear acoustic resistance due to finite pressure amplitude, and is
related to the pressure amplitude by

V„|= ^ (5)

The specific acoustic reactance of the wall can be expressed as

X^ ^^^^^> -cot(^) (6)

ac \ c /

where 5 is the orifice end correction and is given by

g ^ 0.85d(l - 0.7\/^) ^7)

1 + 305 M^

When the liner geometry, flow conditions, and noise spectrum are given, equations (3)
to (5) must be solved by iteration for 6. The reactance is obtained directly from equa-
tions (6) and (7).

The converse of the problem of calculating acoustic impedance is of most interest
for liner design. Given the optimum specific acoustic impedance at a certain frequency,
the steady flow velocity, and the noise spectrum, what are the required values of u, t,
d, and b? The wall geometry can be determined from equations (3) to (7) when the fol-
lowii^ two conditions are given: First, the peak overall sound pressure level is used in
equation (5). This approximation accounts for the nonlinear effect of the noise spectrum
upon the resistance at the frequency in question. In reference 7 this approach was used
with some success to correlate the results of a two-frequency resistance experiment.
The second condition is that the facing sheet thickness and hole diameter must be speci-
fied. The facing sheet thickness would probably be specified by material strength con-
siderations. The hole diameter should be at least as large and preferably larger than
the thickness if the holes are to be punched. The hole diameter d is a very weak
parameter when high flow velocity and sound pressure level exist in the suppressor.
The linear resistance (first term in eq. (3)), in which hole diameter appears, is usually
much smaller than the nonlinear resistance. The hole diameter appears in the reac-

tance through the orifice end correction, which approaches zero as flow velocity is in-
creased (see eq. (7)).

When thickness t and hole diameter d are specified, the open-area ratio cr can
be obtained by combining equations (3) to (5). The backing distance b is then obtained
by combining equations (6) and (7).

Instrumentation

The instrumentation to measure the aerodynamic performance of the fan was re-
ported in detail in reference 1. Some important aspects of the noise measuring appara-
tus and the test site are mentioned here for convenience.

The fan center is located 19 feet (5. 79 m) above the ground. The 1/2-inch (1. 27-cm)
condenser microphones are at the same elevation as the fan center.

The microphone locations are shown in figure 7. The microphones are located at
10° increments from 10° to 160° with 0° being the fan axis at the inlet. The 70° to 160°

Main Drive
Motor Building-,

CD-10755-11

Figure 7. - Overall plan of test facility with microphone Ixations.

microphones are located on a 100-foot (30. 48-m) radius measured from the center of
the fan rotor. The 10° to 60° microphones are on a line which is perpendicular to the
fan axis and are located 31. 5 feet (9.6 m) in front of the fan.

The area between the fan and the microphones is asphalt surfaced. The face of the
drive motor building is lined with 6-inch- (15. 24-cm-) thick polyurethane ether open-
cell foam.

Test Procedure and Data Reduction

The techniques for acquisition and reduction of acoustic data are presented in ref-
erence 1. Some aspects of the procedure are given here for convenience.

The tests were run from low speed to high speed and then back down to low speed.
One set of data was obtained at each speed on the way up and two sets on the way down.
Each of the three data samples is from 1 to 3 minutes duration. Each microphone out-
put was recorded on magnetic tape. Acoustic data were not taken if the wind speed ex-
ceeded 13 knots (6. 7 m/sec) or if there was any precipitation.

Cable loss corrections were applied to the data. The largest correction was 3 deci-
bles at 10 000 hertz. The data were corrected to a 100-foot (30. 48-m) radius for
those microphones which were not at 100 feet (30. 48 m), that is, those located at 10° to
60°, using the inverse squared distance rule. A 1/3-octave band analysis from 50 to
10 000 hertz was performed on all the data. The sound pressure levels were corrected
to standard-day conditions (70 percent relative humidity; 59° F, 288. 2 K) using the
methods of reference 8.

No site calibration corrections were applied to the acoustic data. In reference 1 a
test of the site showed that the nearby building with foam lining did not affect the data
above 500 hertz. However, at least a ground reflection effect, which has not been ac-
counted for, remains.

RESULTS AND DISCUSSION
Acoustic Data

The acoustic data in the form of sound pressure level frequency spectra (in dB ref-
_4
erenced to 2x10 microbar) are presented in the appendix in figures 12 to 75. The

sound pressure levels have been corrected to a 100-foot (30. 48-m) radius and to

standard-day conditions, as described in the previous section. Each figure presents the

results of four fan speeds at a particular microphone location.

All the data presented herein are for acoustically treated inlet and exhaust ducts.
To obtain the noise suppressor performance at a particular speed and polar angle, these
data must be compared to the corresponding hard-wall data of reference 2. The per-
formance of the noise suppressors on the basis of overall acoustic power is discussed
in the next two sections.

Data for four configurations (10 to 13) are referred to in figures 12 to 75. As seen
in the figure titles, these configurations differ in inlet length and nozzle area. These
configurations have been discussed in the apparatus section of this report.

Acoustic Power Comparisons - Hard and Soft Ducts

Several acoustic power spectra are presented in figures 8 and 9. The inlet and ex-
haust powers are determined by summation of the power in the front and rear hemi-
spheres. The 90 and 60 percent speeds are representative of takeoff and landing engine
speeds. On each figure data for long and short inlets are shown with and without
acoustic treatment. Figures 8(a) to (d) are for the standard nozzle, while figure 9 also
contains the data for the lO-percent-oversized nozzle.

The power spectra for the inlet at 60 percent speed are shown in figure 8(a). About
a 1-decibel difference between the long and short hard-cowl data can be seen over the
entire frequency spectrum. Similar shifts of up to 3 decibels can be seen in later fig-
ures. These apparent shifts in the data have not as yet been explained. They may come
from an insufficient averaging time in the sound pressure level averaging circuits (about
1. 5 sec), or possibly from some systematic error. Small differences in acoustic power
spectra, such as in figure 8(a), may not be significant. However, several significant
points about the power spectra should be noted. A large sound power attenuation (18 dB
at the blade passage frequency) was obtained with the short treated inlet over a wide fre-
quency range. Only at the blade passage frequency did the long treated inlet provide a
significant additional attenuation. An increased attenuation was observed between 2500
and 6300 hertz, but it was small and possibly not significant. Doubling the amount of
acoustic treatment without further reduction of sound power level strongly suggests that
a noise floor has been reached. This implies that the noise measured in the far field in
front of the fan is not dominated by direct radiation out of the inlet. The source of this
noise floor has not been determined. It may be caused by radiation out of the fan shell,
a flanking ground path, or even electric drive motor noise.

Figure 8(b) shows the rear hemisphere power spectra at 60 percent speed. Again a
significant noise attenuation was obtained over a wide frequency range with the short in-
let suppressor. The long inlet suppressor should not cause a further reduction in the
rear-end noise. In this case a sUght increase in noise was observed which may not be
significant.

140

135

130

125

120
3 115

llOl — L

~ 140

O Short inlet
n Long inlet

Open symbols denote hard cowls
Solid symbols denote acoustic
treatment

1 1 .1 U..1 iJ _
(a) Front hemisphere, 60 percent speed.

I 1 ij.ll

.4 .6 .8 1

Frequency, kHz

j__ I I Lj_bJ

4 6 8 10

(b) Rear hemisphere, 60 percent speed.
Figures. - Acoustic power spectra. Standard nozzle.

155

150

145

140

135

130

125

120

150i—

145

140

135

130

125

O Short inlet
D Long inlet

Open symlMis denote hard cowls
Solid symbols denote acoustic
treatment

I I I 1 I I I
(c) Front hemisphere, 90 percent speed.

I I I

I I I I I
04 .06.08.1

I I I I
.4 .6 .8 1
Frequency, kHz

(d) Rear hemisphere, 90 percent speed.
Figures. -Concluded.

The inlet sound power levels at 90 percent speed are shown in figure 8(c). The
situation is much like that of figure 8(a), The short suppressor greatly reduces the
broad band and discrete tone noises. Additional suppressor length reduces only the dis-
crete tones which were above the background level. With the exception of the discrete
tones, a noise floor had already been reached with the short suppressor over most of the
noise spectrum.

A comparison of the hard-cowl and suppressed data of figures 8(c) and (d) shows
that the blade passage frequencies (fundamental and second harmonic) lie in different
1/3-octave bands. The cause of this difference was the considerably different ambient
temperatures at which the data were taken and altering the rotational speed of the fan to
maintain constant corrected speed.

In figure 8(d) the rear hemisphere noise at 90 percent speed can be seen. Again the
longer inlet suppressor yields no additional noise reduction, which is as expected. The
noise data obtained with acoustic treatment appear to identify two noise sources. The
low-frequency peak (125 Hz) and a steady decrease in power with frequency (to 2000 Hz)
is typical of the externally generated jet noise. Above 2000 hertz the noise is probably
associated with the fan. Additional aft- end suppressors probably would not result in a
gain in attenuation below 2000 hertz .

In figure 9 the curves are all for the short inlet configurations, both treated and
hard. Two different exhaust nozzles are shown, however, which yield considerably dif-
ferent noise spectra with hard cowls. The oversized nozzle (+10 percent) configuration

155

150

145

140

■^ 135

130

125

▲

Config-
uration

2
10

1
13

Nozzle

Standard
+10 Percent

Open symtjois denote hard cowls f^
Solid symbols denote acoustic /
treatment

Frequency, kHz
Figure 9. -Acoustic power spectra, front hemispttere, 90 percent speed, short

inlet.

is rich in multiple pure tones (ref. 2) at 1250 and 1600 hertz and deficient in the blade
passage frequency (and harmonics) in comiparison with the standard nozzle configura-
tion. The multiple pure tones are seen to be completely removed by the inlet sup-
pressor.

The argument for a noise floor beii^ reached with the short suppressor is also
reinforced by the evidence in figure 9. The noise spectra after suppression are almost
identical for the two nozzle configurations, although without suppression they are con-
siderably different.

Acoustic Power Attenuations - Experimental and Theoretical

The experimental acoustic power attenuations for the exhaust and short-inlet sup-
pressors at 60 and 90 percent speed are shown in figures 10(a) to (d). These attenua-
tions are the differences between the hard- and soft-cowl, short- inlet curves of figure 8.
As discussed previously when comparing hard- and soft-cowl data, the 90-percent-speed
data show a shift in the 1/3-octave band in which the discrete tones are found. For fig-
ures 10(b) and (d) the 1/3-octave bands in which the discrete tones are found and the
band above these have been reversed for the hard- cowl data.

Theoretical predictions of suppressor performance are also shown in figure 10.
The attenuation theory was based upon the model of references 5 and 6, except that rec-
tangular instead of circular geometry was used. The model assumes that a plane pres-
sure wave enters the duct, traveling in the direction of the duct axis. A uniform, steady
flow field was assumed without boundary layers or velocity gradients over the duct. The
wall impedance was calculated according to equations (3) to (7). Estimates of the overall
sound pressure level within the ducts were obtained from the hard-cowl, far-field data.
This pressure was used in equation (5) to define the nonlinear resistance.

The calculation method based on references 5 and 6 provides attenuations for semi-
infinite rectangular ducts with the same impedance on both walls. For the inlet suppres-
sor, each passage has different materials and thus different impedances on opposing
walls. To handle this situation, a result from the approximate theory of Morse (ref. 9)
was used. The sound power attenuation in a duct is approximated by

-4.34 e -
A

e'.x'

AdB « i^ (8)

where S/A is the ratio of treated surface area to duct cross- sectional area. When put
in terms of S/A, equation (8) provides an approximation for either circular or rectangu-

(a) Inlet, 50 percent speed.

Frequency, kHz

(b) Exhaust duct, 60 percent speed.

Figure 10. - Comparison of theoretical and experimental sound power attenuation.

lar ducts. For the case of walls of different impedance on two sides of a duct, the fol-
lowing procedure was used: The attenuation calculations were performed as if just one
of the materials was present on both walls of the duct. The calculations were repeated
with the second material. According to equation (8) the attenuation is approximately
proportional to the area of acoustic treatment. The attenuation for each material was
thus weighted by its proportion of the total area of acoustic treatment. The total attenu-
ation was then the sum of these weighted attenuations.

When several lined passages are involved, such as in the fan inlet suppressor, a
uniform sound power flux was assumed at the duct inlets. In cases where a radial sound

16 —

—

\ o-<l

/
/
/

Y\.

—

ii\ ^\

\ \ /

/ J

\ V

^I^-'"!^

-Y\^ 1 1

\ ^

1 < 1 1 1 , 1 r

12 —

4 —

.5 .8 1 2

Frequency, kHz

(d) Exhaust, 90 percent speed.

Figure 10. - Concluded.

intensity profile is known to exist, this assumption should be changed.

Figures 10(a) to (c) have several characteristics in common. For frequencies both
high and low, in comparison to the frequency of maximum attenuation, the experimental
sound power attenuation is considerably h^her than that predicted by theory. This dif-
ference is probably greater than is indicated; it was pointed out in the previous section
that noise floors may have been reached except at the blade passage frequency and its
harmonics. Shorter liners might have yielded essentially the same experimental re-
sults, but then the associated theoretical curves would have been reduced.

Another common characteristic of the three figures is that the frequency for maxi-
mum sound power attenuation has been fairly well predicted. At approach speed
(60 percent), the magnitude is also fairly well predicted.

The differences between theory and experiment at high and low frequencies are not
easy to explain. For the low frequencies a different dependence of wall impedance on
frequency than that of equations (3) to (7) may provide better agreement between theory
and experiment. However, for the high frequencies the sound power attenuations are
near the theoretical maximums over a considerable frequency range. No real wall ma-
terial could have the impedance characteristics which are necessary for this behavior
(ref. 7). This would require a resistance which increases and a reactance which be-
comes more negative (stiffness controlled) with increasing frequency (see fig. 6).

A possible explanation, especially for the high frequencies, for the behavior of the
experimental data is as follows. The present theory assumes an axially propagating
sound wave (at the lined duct entrance) with no transverse wave motion. This is the
most conservative estimate available. Most of the acoustic power is directed axially,
while only transversely directed power can be absorbed at the lined wall. The turning
of the axial power into the walls must be accomplished by the proper Impedance match
between the duct and the wall. However, if some other mechanism exists which can
redirect the acoustic power into the walls, the attenuation might be greatly increased.
Such a mechanism exists in the form of gradients in the steady flow velocity. Propaga-
tion of sound in a duct with sheared flow has been investigated in references 10 to 13.
When the ratio of sound wavelength to boundary layer thickness is less than or nearly
equal to 1, the acoustic energy can be drastically redistributed. The presence of the
boundary layers alone in the exit duct may not be sufficient to account for the large high-
frequency attenuation. Some radial velocity gradients over the exit duct were observed
downstream from the stators. Application of the theory of reference 13 with these
velocity gradients may be sufficient to account for the larger attenuations.

For the inlet duct, refraction of sound in the boundary layers in the lined passages
will direct the acoustic energy toward the center of the passage and therefore, reduce
attenuation. However, the sound waves may be refracted in velocity gradients across
the duct, in the axial space between the rotor blades and the trailing edges of the inlet

splitter rings. The sound waves will then enter the lined splitter ring sections at an
angle rather than purely axially. This will result in increased higher-order transverse
mode content with resulting increased acoustic power attenuation.

The comparison of theory and experiment in figure 10(d) yields the same results as
the previous three parts only at high frequencies. At low frequencies there is virtually
no experimental sound power attenuation. At intermediate frequencies (centered around
1600 Hz) the theory greatly overpredicts the attenuation. Both these effects are prob-
ably caused by the emergence of the fan jet noise as the dominant source at low and in-
termediate frequencies. The rear-end power spectra for the treated ducts in f^re 8(d)
support this contention. The jet noise is seen to peak at about 125 hertz and then
steadily decrease to 2000 hertz, beyond which another noise source dominates. Once
the suppressor has reduced the internal intermediate frequency noise to the level of jet
noise which is produced externally, further reduction will not produce observable dif-
ferences. At low frequencies, where internal jet noise already dominates over inter-
nally generated noise, the suppressor can produce no observable effect in the far field.

There is a strong temptation to consider the sound power attenuations at low fre-
quencies as systematic errors in the data in figures 10(a) to (c). One might be justified
in shifting the attenuation curves downward by 3 to 5 decibels. This, however, raises
some very difficult questions. Why would the hard-cowl power spectra generally be
measured too high or else the soft-cowl data be measured too low? Why do these errors
generally disappear at high engine speeds in the rear hemisphere when jet noise domi-
nates? At present then, it must be assumed that the measurements are not in error and
that the noise suppressors are working over an extremely wide frequency range.

One further possibility for the difference in predicted and observed sound power
attenuations should be considered. Inlet splitters may have an effect on the rotor which
reduces noise production. This reduced noise should not be credited to the suppressor
as a simple noise absorption.

Perceived Noise Levels

The perceived noise levels for simulated approach and takeoff conditions were cal-
culated according to reference 14 and are shown in figures 11(a) and (b). These figures
are for a single fan.

The single-fan approach condition can be seen in figure 11(a). The perceived noise
levels are given for 60 percent speed on a 375-foot (114. 3-m) sideline. Two data points
(120 and 130 ) appear to be in error for the standard nozzle conditiai with a short
acoustically treated inlet. The 10-percent-oversized-nozzle data are thus presented to
obtain the noise levels at these angles. The acoustic treatment reduces the maximum

104

100

S 68

£ 100

O
O O O

O O °

•«.

I I

■ ■ I

O standard nozzle

n 10-Percent-oversized nozzle

Open symbols denote hard cowl
Solid symbols denote acoustic
treatment

o- 92 —

(a) At 60 percent speed, on a 375-foot (114. 3-m) sideline.

O ° o

o o o

o o

56.

• •

40 60 80 100 120

Angular position from inlet, deg

140

160

(b) At 90 percent speed, on a 1000-foot (304. 8-m) sideline.

Figure 11. - Effect of acoustic treatment on perceived noise level. Short-
inlet design nozzle.

perceived noise level of the fan at approach by 13 PNdB (from 100 to 87 PNdB).

The single-fan takeoff condition is given in figure 11(b). The perceived noise levels
are for 90 percent speed on a 1000-foot (304. 8-m) sideline. The reduction in maximum
perceived noise level due to the noise suppressors is 12 PNdB (from 99 to 87 PNdB) at
takeoff.

An estimate of the perceived noise levels for an aircraft with four fans can be made
by adding 6 decibels to the maximum values in figure 11. Four fans with acoustically
treated inlet and exhaust ducts would thus produce 93 PNdB at both takeoff and approach.
The current DC-8 figures given by Pendley and Marsh (ref. 15) are 117 PNdB at takeoff
(1000 ft, 304. 8 m) and 120 PNdB at approach (370 ft, 112. 8 m). The reductions in per-
ceived noise level are thus 24 PNdB at takeoff and 27 PNdB at approach. These noise
reductions are obtained by considering the fan noise only, and will be realized only if
the core jet and turbine noise can be kept below the fan noise in an actual engine.

Estimates of the effective perceived noise levels have been calculated for an air-
craft with four turbofan engines. Again it is emphasized that these noise levels are
valid only if the core engine noise is less than the suppressed fan noise. Both tone and
time duration corrections are considered by use of the calculation procedure of refer-
ence 16. The effective perceived noise level calculations are made using the noise
spectra from the standard nozzle, short inlet configurations. The following comparisons
are made between the hard-cowl and acoustically treated versions. For simulated take-
off conditions, the effective perceived noise level was reduced by 17 EPNdB (102. 7 to
85. 7 EPNdB) and, for approach, by 14. 3 EPNdB (102. 5 to 88. 2 EPNdB). The effective
perceived noise level calculations were made using a relative humidity of 70 percent and
a temperature of 77° F (298. 2 K). For takeoff, the climbout angle was 5. 6°, the veloc-
ity along the flight path was 292 feet per second (89 m/sec), and the engine center line
was 9. 1° from horizontal. For approach, the glide ar^le was -3°, the velocity along the
flight path was 241 feet per second (73. 5 m/sec), and the engine center line was 0. 5°
from horizontal.

Aerodynamic Performance with Acoustic Treatment

The internal flow losses of the nacelle cowling were not measured directly, but the
installed fan performance was obtained with both hard and acoustic surfaces lining the
cowling. These data are presented in the following table for 60 and 90 percent fan
speeds, which represent the landing and takeoff operating conditions. Fan pressure
ratios shown include the inlet cowling losses because the pressure rise was measured
from ambient to the fan stator discharge. Thrust includes all the cowl losses because
it is the momentum measured at the nacelle exhaust. The data indicate that the acoustic

Percent
speed

Inlet

Wall
material

Airflow

Pressure
ratio

Thrust

Ibm/sec

kg/sec

Ibf

Short

Hard
Lined

505
493

229
224

1.14
1.14

7 200
7 400

32 000
32 900

Long

Hard
Lined

484
486

220
220

1.15
1.14

7 200
7 200

32 000
32 000

Short

Hard
Lined

741
749

336
340

1.36
1.36

16 700
16 800

74 300
74 700

Long

Hard
Lined

736
739

334
335

1.36
1.35

16 400
16 400

73 000
73 000

liner did not degrade the fan performance as much as did the long inlet cowling.

The differences shown are all within the experimental measurement error, which
was estimated to be about 2 percent of the full-speed values. These errors are
16 pounds per second (7. 3 kg/sec) on airflow, 0. 01 on pressure ratio, and 400 pounds
force (1780 N) on thrust.

SUMMARY OF RESULTS

The inlet and exhaust noise suppressors for a 6-foot- (183-m-) diameter fan have
been tested. Some of the more important results were as follows.

1. The suppressors provided more peak noise attenuation and much broader band
attenuation than was predicted.

2. The duct sound transmission theory predicted the frequency of peak attenuation
fairly well. The exception occurred when a fairly obvious jet noise floor dominated at
the frequency of predicted peak attenuation.

3. Some noise floors of unknown origin were apparently reached even with the short
treated inlet.

4. The noise suppressors provided a noise reduction of 13 PNdB at simulated ap-
proach conditions and 12 PNdB at takeoff, compared to the fan noise produced with hard
passages.

5. The suppressor and fan combination achieved significant noise reductions when
compared to the current DC- 8 aircraft. The simulated approach and takeoff noise re-
ductions were 27 and 24 PNdB, respectively. These results are obtained under the as-
sumption that the core engine noise is insignificant.

6, Degradations of the aerodynamic performance due to the noise suppressors were
within the experimental error of measurement.

Lewis Research Center,

National Aeronautics and Space Administration,
Cleveland, Ohio, September 14, 1970,
126-61.

APPENDIX -ACOUSTIC DATA

The acoustic data in the form of 1/3-octave-band sound-pressure-level frequency

"■ 4
spectra (in dB referenced to 2x10 microbar) are presented here. All data are for

nacelle configurations with acoustic treatment. The sound pressure levels have been
corrected to a 100-foot (30. 48-m) radius and to standard-day conditions (70 percent rela-
tive humidity; 59° F, 288. 2 K).

Four configurations are referred to in figures 12 to 75. These result from using
two inlet lei^ths and two nozzles. The short inlet is as shown in figure 2. The long inlet

is obtained by adding a 41-inch (1. 04-m) cylindrical section. The standard nozzle area

2
is 1895 square inches (1. 223 m ). The 10-percent-oversized nozzle has an area of 2150

2
square inches (1. 387 m ).

PERCENT

70
on

SPEED

6 8 10^ 2

FREQUENCY, HZ

FIGURE 12 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURfiTION 10. 10 -DEGREE ANGLE.
SHORT INLET: STflNDRRO NOZZLE; RCOUS 'CflLLT TREATED.

a
o

PERCENT

60
70,
80
90

SPEED

6 8 10

6 8 10'

FREQUENCT.HZ

FIGURE 13. -SOUND PRESSURE LEVEL ON lOG-FOQT RADIUS. CONFIGURflTION 10, 20 -DEGREE FlNaE.
SHORT INLET: STflNOflRD NOZHLE: RCOUSTICflLLT TREATED.

PERCENT

FREQUENCT.HZ

FIGURE 14. -SOUND PRESSURE LEVEL ON 100-FOOT RRDIUS. CONFIGURRTION 10.. 30 -DEGREE ANGLE.
SHORT INLET; STFINDHRO NOZZLE: flCOUSTICflLLT TREATED.

■=1 85.

PERCENT

60.
70,
— 81
90.

SPEED

6 8 10

2 1

FREQUENCr.HZ

FIGURE 15. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 10, 40 -DEGREE ANGLE.
SHORT INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

PER

60,

SPEED

FREQUENCY. HZ

FIGURE 16. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 10, 50-DEGREE ANGLE.
SHORT INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

60. ■■
10'

PERCENT

70
80
90

SPEED

e 10= 2 u

FREQUENCT.HZ

6 8 10'

FIGURE 17. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 10, 60 -DEGREE ANGLE.
SHORT INLET; STANDRRO NOZZLE: flCOUSTICRLLT TREATED.

^ 85.

60. I
10'

PERCENT

60.
70.
80.
90.

SPEED

6 8 10"

FREQUENCT.HZ

FIGURE 18. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 10, 70 -DEGREE ANGLE.
SHORT INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

□

^-^

<=> 85.

-J 80.

en 75.

9 7n

60.

PERCENT

60,
70.

on
uui.

90.

SPEED

2 t
FREQUENCY. HZ

FIGURE 19. -SOUND PRESSURE LEVEL ON 100-FOOT RHDIUS. CONFIGURATION 10, 80-DEGREE ANGLE.
SHORT INLET; STHNOHRO NOZZLE; flCOUSTICflLLY TREATED.

PERCENT

60
70

SPEED

8 10' 2 1
FREQUENCY, HZ

6 8 10"

FIGURE 20. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION ID, 90 -DEGREE ANGLE.
SHORT INLET: STANDARD NOZZLE; ACOUSTICALLY TREATED.

PERCENT

60
70
80,
90,

SPEED

8 10*

FREQUENCY, HZ

FIGURE 21. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURflllON 10. 100-OEGREE ANGLE.
SHORT INLET; STflNOHRD NOZZLE: flCOUSTICRLLY TREATED.

PERCENT

60,

70
80

SPEED

6 8 10

1 I 1
2 11 6 8 10 =

FREQUENCY. HZ

8 8 10"

FIGURE 22. -SOUND PRESSURE LEVEL ON 100-FOQT RADIUS. CONFIGURATION 10. 110-DEGREE ANGLE.
SHORT INLET; STANDARD NOZZLE; ACOUSTICALLY TREATED.

1=1 65.

PERCENT

60
70,
80
90,

SPEED

10'

8 10"

FREQUENCY, HZ

FIGURE 23. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURRTION 10. 120 -DEGREE ANGLE.
SHORT INLET; STflNDflRO NOZZLE; flCOUSTICfiLLT TREHTEO.

PERCENT

60
70

90.

SPEE3

FREQUENCY, HZ

FIGURE 24. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 10, 130-OEGREE ANGLE.
SHORT INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

5 100.1 ^

PERCENT

60,
70,
SO,
90,

SPEED

6 6 10' 2

FREQUENCY, HZ

6 8 10*

FIGURE 25. -SOUND PRESSURE LEVEL ON IGO-FOOT RfiOIUS. CONFIGURATION 10, mO -DEGREE ANGLE.
SHORT INLET: STflNOflRO NOZZLE; flCOUSTICflLLT TREHTEO.

■=> 85.

a
o

PERCENT

6 8 10'

FREQUENCT.HZ

FIGURE 26. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 10. 150-DEGREE ANGLE.
SHORT INLET: STANDARD NOZZLE; ACOUSTICALLY TREATED.

PERCENT

60,
70,
80,
90,

SPEED

1 l^-i

6 8 10= 2

FREQUENCY, HZ

6 8 10"

FIGURE 27. -SOUND PRESSURE LEVEL ON lOQ-FOOT RADIUS. CONFIGURATION 10, 160-DEGREE ANGLE.
SHORT INLET; STANDARD NOZZLE; ACOUSTICALLT TREATED.

PERCENT

o 85.

2 11

FREQUENCY, HZ

6 8 10"

FIGURE 28. -SOUND PRESSURE LEVEL ON 100-FOQT RADIUS. CONFIGURATION 11, 10-DEGREE ANGLE.
LONG INLET: STANDARD NOZZLE; ACOUSTICALLY TREATED.

a 65.

PERCENT

60,
7Q
SO
90

SPEED

2 l)

FREQUENCY. HZ

FIGURE 29. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURHTION 11, 20 -DEGREE RNGLE.
LONG INLET; STRNOflRO NOZZLE: HCOUST ICFILLT TREATED.

a 85.

PERCENT SPEED

10' 2

FREQUENCT.HZ

FIGURE 30. -SOUND PRESSURE LEVEL ON 100-FOOT RflOIUS. CONFIGURRTION 11, 30 -DEGREE ANGLE.
LONG INLET; STHNDflRO NOZZLE: flCOUSTICflLLT TREATED.

6 8 10'

6 8 10"

FREQUENCY, HZ

FIGURE 31. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURRTION 11, 40 -DEGREE ANGLE.
LONG INLET; STANDARD NOZZLE; ACOUSTICALLY TREATED.

105.

D 100.

cc
o

31
X

rsi

UJ
UJ

9 7n

o
en

PERCENT

2 U
FREQUENCY, HZ

FIGURE 32. -SOUND PRESSURE LEVEL ON 100-FOQT RADIUS. CONFIGURATION 11, 50 -DEGREE ANGLE.
LONG INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

<=> 85.

CD
O

PERCENT

60
70
SG.
90.

SPEED

TOE PRSSflGE FREQL

1854.
2164.
24 76.
2791.

ENCT

FREQUENCY, HZ

6 8 10'

FIGURE 33. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 11. 60 -DEGREE ANGLE.
LONG INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

PERCENT

60
70
80
90

SPEED

FREQUENCY, HZ

6 8 10"

FIGURE 34. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 11, 70 -DEGREE ANGLE.
LONG INLET; STHNDHRD NOZZLE: ACOUSTICALLY TREATED.

a
o

PERCENT

2 1

FREQUENCY, HZ

FIGURE 35. -SOUND PRESSURE LEVEL ON lOG-FOOT RADIUS. CONFIGURflTiON 11, 80-DEGFlEE ANaE.
LONG INLET: STflNDflRD NOZZLE: flCOUSTICHLLT TREflTEO.

S loa.

PERCENT

60.
70.
BG.
90.

SPEE3

2 u

FREQUENCT.HZ

FIGURE 36. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURHTION 11. 90 -DEGREE ANGLE.
LONG INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

PERCENT

60
70
80
90

SPEE3

6 8 10

6 8 10«

FREQUENCT.HZ

FIGURE 37 . -SOUND PRESSURE LEVEL ON lOO-FOOT RADIUS. CONFIGURATION 11. 100 -DEGREE ANGLE.
LONG INLET; STHNDARD NOZZLE: flCOUSTICfiLLT TREATED.

□

PERCENT

SO
70
80
90

SPEED

6 8 10"

FREQUENCT.HZ

FIGURE 38 . -SOUND PRESSURE LEVEL ON lOO-FOOT RADIUS. CONFIGURATION 11, 110 -DEGREE ANGLE.
LONG INLET: STANDARD NOZZLE: ACOUSTICALLY TREATED.

PERCENT

105.

□

100.

•

95.

90.

70.

65. ■

60. ■
I

D' 2

i 1

SPEED

I I
6 e 10

FREQUENCY, HZ

FIGURE 39. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 11, 120-OEGREE ANGLE.
LONG INLET; STANDARD NOZZLE; ACOUSTICflLLT TREATED.

PERCENT

2 11
FREQUENCY, HZ

FIGURE 40. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 11, 130-DEGREE ANGLE.
LONG INLET; STANDARD NOZZLE; ACOUSTICALLY TREATED.

PERCENT

60
70,
80,
90

SPEED

6 8 10 =

FREQUENCY, HZ

FIGURE 41. -SOUND PRESSURE LEVEL ON IGO-FOOT RADIUS. CONFIGURATION 11, 140-DEGREE RNGLE.
LONG INLET: STRNDflRD NOZZLE: flCOUSTICflLLT TREATED.

PERCENT

60
70
60
90

SPEED

2 1
FREQUENCY. HZ

FIGURE 42. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION U. 150 -DEGREE RNGLE.
LONG INLET: STRNDflRD NOZZLE: RCOUSTICALLY TREATED.

PERCENT
60

SPEED

2 ij
FREQUENCY, HZ

FIGURE 43. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 11, 160-DEGREE ANGLE.
LONG INLET; STflNDflRD NOZZLE: fiCOUSTICALLT TREATED.

a 85.

Q
CD

PERCENT

60
70

SPEED

ic'

FREQUENCY, HZ

FIGURE ii . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12, 10-DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE: ACOUSTICALLY TREATED.

□
o

PERCENT

60
70
30
90,

10'

SPEED

6 8 ID

2 « 6 B 10'

FREQUENCY. HZ

6 8 10"

FIGURE 45. -SQUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12, 20 -DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

<=> 85.

a
o

PERCENT

60
70

SPEED

+ h

5 8 10"

FREQUENCY, HZ

FIGURE 46 . -SOUND PRESSURE LEVEL ON lOO-FOOT RADIUS. CONFIGURATION 12, 30 -DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

PERCENT

60
70

SPEED

FREQUENCY, HZ

FIGURE 47 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12, UO -DEGREE ANGLE.
LONG INLET; lO-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

<=! 85.

PERCENT

60
70

- m.

SPEED

2 H
FREQUENCY, HZ

6 8 10"

FIGURE 48. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12. 50-OEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

O 85.

60. t
10'

ill

PERCENT

60
70
SO
90.

SPEED

FREQUENCY, HZ

FIGURE 49. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12, 60 -DEGREE ANGLE.
LONG INLET: 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

a 85.

□
O

-I-

PERCENT

60
70
80
90

SPEE3

2 U

FREQUENCY. HZ

FIGURE 50 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12, 70 -DEGREE ANGLE.
LONG INLET: 10-PERCENT OVERSIZE NOZZLE: ACOUSTICALLY TREATED.

PERCENT

60
70

SPEE3

10'

6 8 10^ 2 1

FREQUENCY, HZ

6 8 lO'

6 8 10"

FIGURE 51 . -SOUND PRESSURE LEVEL ON 100-FOQT RADIUS. CONFIGURHTIDN 12, 80 -DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

a 85. —

a
o

PERCENT

60.
70.

SPEED

FREQUENCY, HZ

FIGURE 52. -SOUND PRESSURE LEVEL ON lOO-FOOT RADIUS. CONFIGURATION 12, 90 -DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

PERCENT

60
70
SO
90,

SPEED

I I

6 B 10

6 8 10*

FREQUENCY, HZ

FIGURE 53. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONF IGURFIT ION 12. 100-DEGREE fiNOLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; HCOUSTICflLLT TREATED.

° 65.

PERCENT

60
70
80,
90

SPEED

6 e 10

6 8 10"

FREQUENCT.HZ

FIGURE 54. -SOUND PRESSURE LEVEL ON 100-FQOT RflOIUS. CONFIGURHTION 12. 110-DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACQUSTICALLT TREATED.

a 85.

PERCENT

60,
70-
80.
90

SPEED

2 <i

FREQUENCY, HZ

FIGURE 55 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12. 120-DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

PERCENT

6 8 10«

FREQUENCY. HZ

FIGURE 56. -SOUND PRESSURE LEVEL QN lOO-FOQT RADIUS. CONFIGURATION 12. 130-DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

O 65.

HI
O

PERCENT

60
70
80.

90,

SPEED BLWE PflSSdGE FREQLENCT.HZ

«—<►

v:)

a-*

FREQUENCY, HZ

FIOJRE 57 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 12, mO -DEGREE ANGLE.
LONG INLET: 10-PERCENT OVERSIZE NOZZLE: ACOUSTICALLY TREATED.

Q
O

PERCENT

60
70,
80
90

SPEED

FREQUENCY. HZ

FIGURE 58. -SOUND PRESSURE LEVEL ON lOO-FOOT RADIUS. CONFIGURATION 12. 150-DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE: ACOUSTICALLY TREATED.

PERCENT

60.
70.

go,

SPEED

2 1
FREQUENCY. HZ

8 io«

FIGURE 59. -SOUND PRESSURE LEVEL ON 100-FQOT RADIUS. CONFIGURflTION 12, 160 -DEGREE ANGLE.
LONG INLET; 10-PERCENT OVERSIZE NOZZLE: flCOUSTICRLLT TREATED.

° 85.

PERCENT

SO
70
80
90

SPEED

10'

2 u
FREQUENCY, HZ

6 8 10 =

6 8 10"

FIGURE 60. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13, 10-DEGREE ANGLE.
SHORT INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

PERCENT

60.
70.
80.
90.

SPEED

6 8 10

6 8 10 =

FREQUENCY. HZ

FIGURE 61. -SOUND PRESSURE LEVEL UN 100-FOOT RADIUS. CQNFIGLIRflTION 13, 20-DEGfiEE ANGLE.
SHORT INLET: 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

Q
O

PERCENT

60. k

10'

FREQUENCY, HZ

FIGURE 62. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13, 30-DEGflEE ANGLE.
SHORT INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

PERCENT

2 1
FREQUENCY, HZ

FIGURE 63. -SOUND PRESSURE LEVEL QN 100-FQOT RADIUS. CQNFIGURRTIQN 13. 14D-0EGREE ANGLE.
SHORT INLET; 10-PERCENT OVERSIZE NOZZLE; flCOUSTICRLLT TREATED.

FREQUENCY. HZ

FIGURE 64. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13, 50-DEGREE ANGLE.
SHORT INLET; lO-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

60. t
ID'

PERCENT

60,
70,
80,
90.

SPEE3

FREQUENCT.HZ

FIGURE 65 . -SOUND PRESSURE LEVEL ON 100-FOOT RRDIUS. CONFIGURATION 13, 60 -DEGREE ANGLE.
SHORT INLET; 10-PERCENT OVERSIZE NOZZLE; flCOUSTICHLLT TREfiTEO.

PERCENT

60
70
80
90

SPEED

6 e 10

6 8 10"

FREQUENCT.HZ

FIGURE 66 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13. 70 -DEGREE ANGLE.
SHORT INLET: 10-PERCENT OVERSIZE NOZZLE: RCOUSTICALLT TREATED.

a 85.

a
o

-A-

PERCENT

70
_ on

90,

SPEED

1 I -,4+4

2 U 6 S 10'

FREQUENCY. HZ

6 8 10"

FIGURE 67 . -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFluURRTION 13, 80-DEGFiEE ANGLE.
SHOR" INLET; 10-PERCENT OVERSIZE NOZZLE; flCOUSTICflLLT TREHTED.

2 1
FREQUENCY. HZ

FIGURE 58. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIC-URATION 13. 90-DEGfiEE ANGLE.
SHORT INLET; lO-PERCENT OVERSIZE NOZZLE: ACOUSTICALLY TREATED.

t=l 85.

60. '■
10'

PERCENT

60,
70.
80.
90.

SPEE3

6 8 10"

FREQUENCY, HZ

FIGURE 69 . -SOUND PRESSURE LEVEL ON lOG-FOOT RHOIUS. CONF ICURRT ION 13. 100-DEGREE ANGLE.
SHORT INLEl; lO-PERCENT OVERSIZE NOZZLE; flCOUSTICflLIT TREATED.

<=> 85.

PERCENT

60
7C
8G
90

SPEED

FREQUENCY. HZ

FIGURE 70. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13, 110 -DEGREE ANaE.
SHORT INLET; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

1=1 85.

PEf

a
o

60
70
80
90

SPEEO

2 1

FREQUENCY. HZ

6 8 10"

FIGURE 71. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13. 120 -DEGREE ANGLE.
SHORT INLEl. 10-PERCENT OVERSIZE NOZZLE; flCOUSTICflLI T TREflTED.

1=1 85.

PERCENT

60.
70.
80.
90.

SPEED

10'

FREQUENCT.HZ

FIGURE 72. SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURATION 13, 130 -DEGREE ANGLE.
SHORT INLEl; 10-PERCENT OVERSIZE NOZZLE; ACOUSTICfll I T TREATED.

'^ 85.

PFRCENl

60
70
SC
90

SPEED

■^

1 ) -

B— ■'

— 1 —

— (-

6 8 10

6 8 10"

FREQUENCY. HZ

FIGURE 73 . -SOUND PRESSURE 1 EVEL ON lOO-FOOT RADIUS. CQNFIGURHT ION 13, mO -DEGREE ANGLE.
SHORT INLET: 10-PERCENT OVERSIZE NOZZLE. ACOUSTICflLIT TREATED.

a
o

1=1 85.

PFRCENl

6 a 10"

FREQUENCT.HZ

FIGURE 74 . -SOUND PRESSURE I EVEL ON IDQ-FOOT RADIUS. CONFIGURATION 13. iSO-DEGREE ANGLE.
SHORT INLET: 10-PERCENT OVERSIZE NOZZLE; ACOUSTICALLY TREATED.

PFRCENl

60
70

CD
O

SPFED

2 I
FREQUENCY. HZ

FIGURE 75. -SOUND PRESSURE LEVEL ON 100-FOOT RADIUS. CONFIGURRT ION 13. 160 -DEGREE ANGLE.
SHORT INLET: 10-PERCENT OVERSIZE NOZZLE; flCOUSTICflLLY TREATED.

REFERENCES

1. Leonard, Bruce R. ; Schmiedlin, Ralph F. ; Stakolich, Edward G. ; and Neumann,

Harvey E. : Acoustic and Aerodynamic Performance of a 6- Foot-Diameter Fan
for Turbofan Engines. I - Design of Facility and QF-1 Fan. NASA TN D-5877,
1970.

2. Goldstein, Arthur W. ; Lucas, James G. ; and Balombin, Joseph R. : Acoustic and

Aerodynamic Performance of a 6-Foot-Diameter Fan for Turbofan Engines.
n - Performance of QF-1 Fan in Nacelle Without Acoustic Suppression. NASA
TN D-6080, 1970.

3. Progress of NASA Research Relating to Noise Alleviation of Large Subsonic Jet

Aircraft. NASA SP-189, 1968.

4. NASA Acoustically Treated Nacelle Program. NASA SP-220, 1969.

5. Rice, Edward J. : Attenuation of Sound in Soft-Walled Circular Ducts. Aerody-

namic Noise. H. S. Ribner, ed. , Univ. Toronto Press, 1969, pp. 229-249.

6. Rice, Edward J. : Propagation of Waves in an Acoustically Lined Duct with a Mean

Flow. Basic Aerodynamic Noise Research. NASA SP-207, 1969, pp. 345-355.

7. Groenweg, John F. : Current Understanding of Helmholtz Resonator Arrays as Duct

Boundary Conditions. Basic Aerodynamic Noise Research. NASA SP-207, 1969,
pp. 357-368.

8. Anon. : Standard Values of Atmospheric Absorption as a Function of Temperature

and Humidity for Use in Evaluating Aircraft Fly-Over Noise. Aerospace
Recommended Practice 866, SAE, 1964.

9. Morse, Philip M. : Vibration and Sound. Second ed. , McGraw-Hill Book Co. , Inc. ,

1948.

10. Pridmore-Brown, D. C. : Sound Propagation in a Fluid Flowing Through an Attenu-

ating Duct. J. Fluid Mech., vol. 4, pt. 4, Aug. 1958, pp. 393-406.

11. Tack, D. H. ; and Lambert, R. F. : Influence of Shear Flow on Sound Attenuation in

a Lined Duct. J. Acoust. Soc. Am,, vol. 38, no. 4, Oct. 1965, pp. 655-666.

12. Mungur, P.; and Gladwell, G. M. L. : Acoustic Wave Propagation in a Sheared

Fluid Contained in a Duct. J. Sound Vib. , vol. 9, 1969, pp. 28-48.

13. Mungur, P. ; and Plumblee, H. E. : Propagation and Attenuation of Sound in a Soft-

Walled Annular Duct Containing a Sheared Flow. Basic Aerodynamic Noise Re-
search. NASA SP-207, 1969, pp. 305-327.

14. Anon. : Definitions and Procedures for Computing the Perceived Noise Level of

Aircraft Noise. Aerospace Recommended Procedure 86 5A, SAE, Aug. 1969.

15. Pendley, Robert E. ; and Marsh, Alan H. : Noise Prodictions and Economic Effects

of Nacelle Modifications to McDonnell Douglas DC-8 Airplanes. Progress of
NASA Research Relating to Noise Alleviation of Large Subsonic Jet Aircraft.
NASA SP-189, 1968, pp. 173-195.

16. Anon. : Noise Standards: Aircraft Type Certification. Vol. m, Part 36 of Federal

Aviation Regulations.

60 NASA-Langley, 1971 28 E-5878

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