NASA Technical Memorandum 86967 

Noise Constraints Effecting 

Optimal Propeller Designs 

Christopher J. Miller 

Lewis Research Center 

Cleveland, Ohio 

and 

John P. Sullivan 

Purdue University 

West Lafayette, Indiana 

Prepared for the 

SAE General Aviation Aircraft Meeting and Exposition 

sponsored by the Society of Automotive Engineers 

Wichita, Kansas, April 16-19, 1985 
ABSTRACT AERODYNAMIC ANALYSIS 

A preliminary design tool for advanced propellers has been developed combining a fast 

vortex lattice aerodynamic analysis, a fast subsonic point source noise analysis, and an optimization scheme using a conjugate directions 

method. Twist, chord and sweep distributions 

are optimized to simultaneously improve both the 

aerodynamic performance and the noise observed 

at a fixed relative position. 

for straight and advanced concept blades are 

presented. The techniques used include increasing the blade number, blade sweep, reducing the 

rotational speed, shifting the spanwise loading 

and diameter changes. Some methods yield improvement in both efficiency and noise. 

~ 

The optimal noise/performance trade-offs 

INTKODUCTIOh 

PREVIOUS WORK (1,2)* HAS IDENTIFIED THE performance gains and noise reductions available 

through the use of advanced concept propeller 

blades. These works optimized the performance 

only; noise was considered as a separate issue. 

To properly investigate advanced propeller concepts as well as conventional geometries, a 

design tool should be able to optimize both performance and noise at the same time. This paper 

describes the development of such a design tool 

and the noise/performance trade-offs available 

from optimal geometries. 

(1) a vortex lattice method for the aerodynamic 

performance prediction; (2) a subsonic compact 

source model in the time domain for the noise 

prediction; and (3) an optimization scheme using 

a conjugate directions method. 

Ihe design program consists of three parts: 

qiuunibers in parentheses designate references at 

end of paper. 

The vortex lattice method has been used by 

various researchers (2 to 7) in the study of 

propeller performance. The basic calculation 

geometry is shown in Fig. 1. There is a single 

curved lifting line divided spanwise into a set 

of horseshoe vortices. This lifting line is 

along the quarter-chord line of the blade. At 

control points on the three-quarter chord line, 

a boundary condition of no flow through the mean 

camber surface is imposed. The trailing vortices representing the wake are prescribed to 

lie on the surface swept out by the quarterchord line with pitch defined by the advance 

ratio. This method is a steady symmetric analysis, so the radial circulation distribution is 

the same for all blades. 

The equations to be solved are based on the 

boundary condition that there is no flow through 

the chord plane surface at the control points. 

The boundary condition for the jth vortex 

lattice element is therefore: 

V. * f;. = .vinflow, + r. xu + v. ;i. = o 

3 

< ’Jr J J 

(1) 

where 

-v V. total velocity at -;$ 

n. normal to the chord plane at the conJ J 

J trol point 

A 

--rc 

r. control point position J 

2 

,L shaft speed and axis [radians/sec) 

2 

V. total induced velocity at 2:. 

J J 

1 
The Biot-Savart Law is used to calculate the 

induced velocities due to the horseshoe vorticies. The induced velocity at the control 

point of lattice element j of blade 1 due to 

the vortex of element i of blade k is given 

byi 

where 

r. circulation strength of horseshoe 

vortex i per unit length 

Ktip propeller tip radius 

dhik 

2 

differential segment of the horseshoe 

vortex. 

Since the circulations are the unknowns, the 

boundary condition provides a set of simultaneous equations for the circulations. 

the forces on each element are calculated using 

the Kutta-Joukowski law. These forces are 

summed to provide the overall forces on the 

propeller. The blade forces for the jth 

element are given by: 

Once the circulations have been determined, 

cd o zero lift drag coefficient 

C chord length 

K drag polar constant. 

As a verification, Figs. 2 to 5 compare the 

predicted results with experimental test data. 

The data in bigs. 2 and 3 are for a representative general aviation propeller tested at the 

hASA Lewis kesearch Lenter (A). 

Figs. 4 and 5, taken from Ref. (?), are for a 

quarter scale general aviation propeller modeled 

after the blade on a Cessna 172: the PlcCauley 

lL160. In both cases, the predicted performance 

agrees well, but the predicted variation of 

power coefficient ( with advance ratio (J) 

results. The fixed wake model is partially to 

blame: the induced velocities would increase 

the wake helix angle, and hence would increase 

the induced drag at the higher loadings. This 

factor would raise the high loading prediction 

of Cp and straighten the curve. 

In these figures, the following symbols are used 

n blade number 

The data in 

has more curvature t Y an the experimental 

8314 blade pitch at the 3/4 radius 

J V/nD = advance ratio 

where 

po density 

s bound vortex length 

I. jth bound vortex segment circulation 

V total velocity at the bound segment midA 

J 

AIL1 ' point 

Dv viscous drag estimate. 

Here the viscous drag is calculated based on 

the section lift coefficient of the element, 

L, : 

= 1 sc (VM * VM, 

Dv 2 Po JJ 

(5) 

V axial inflow velocity (flight velocity) 

A V/iik = J/n = alternate advance ratio 

CT T/(p0n2D4) = thrust coeificient 

~p p/(p,n3~5) = power coefficient. 

hOISE ANALYSIS 

The noise analysis method is based on the 

work of Lowson (lo). 

that there is no mean flow (static operation) 

and that the continuous distribution of noise 

sources on the biade can be represented by point 

sources. These assumptions are met when considering low speed operation, as in take-off and 

climb, such that the helical tip speeds are less 

than about Plach 0.75. 

Lowson did not consider the noise produced 

by a radial force in his analysis. So, an additional term for the radial force noise has been 

derived for use in this study. The equation 

which must be solved for the normalized pressure 

perturbation produced by the propeller is: 

This analysis assumes 

where 

'd viscous drag coefficient 

2 
2 

9 angle from R axis. 

.[XT j -(X rMr - xs) (k? ++)I 

R - r sin($) sin(@) 

1 -M + 

s obs 

r 0 

(6) 

speed of sound 

ambient pressure 

iir/ao 

-Ms ( robs/ r) sin( $) s in(e ) 

iiRs/ao 

I Iril I = I Ixi - yil I 

(x,Y,o) = (robs cos($), robs sin($), 

0) = observer position 

(Xs, Rs cos(e), Rs sin(e)) = source 

po s it ion 

force coefficients on an element 

mass flux as force term = aoqm/ 

p .n2~4 

volume flux 

shaft speed [rps/s] 

propeller diameter 

specific heat ratio = 1.4 

azimuthal angle 

For the noise analysis, the x-axis is the forward axis of rotation (ii). This equation is 

evaluated at the normalized retarded time 6, 

for the observer time determined by: 

where 

ABS sweep introduced delay. 

The propeller is modeled as a distribution 

of point noise sources with strengths proportional to the local loading and thickness. Each 

of the vortex lattice elements is modeled as two 

noise elements in the spanwise direction (or 

along the hub to hub quarter-chord line for the 

bi-blade) to improve the assumption of compactness. The layout of noise sources is shown in 

Fig. 6. 

The force coefficients, (ACT,AC~,ACR), on 

the noise elements are half of the coefficient 

on the parent aerodynamic element, and the 

thickness source term is written in the form of 

a force coefficient (AK~) for calculation ease. 

The strength and positions of the source - sink 

pair used to represent thickness are chosen so 

that the Kankine body produced has the same 

chord as the local blade section, and the same 

frontal area as the noise element. 

sure perturbation (Eq. (6)) is calculated at 

equally spaced points in retarded time, e, along 

with the associated observer times, t. These 

pressures are then interpolated to an equally 

spaced set in observer time. The advantage of 

this method, due to llunro (G), is that the time 

consuming sdlution for the retarded time associated with a given observer time is eliminated. 

Also, the distribution of points in observer 

time is densest in the region of the pressure 

peaks whish results in an accurate interpolation 

of the peaks. The multiple blade waveform is 

created from this single blade waveform by 

superposition with the proper phase delays. 

has been produced, a Fast Fourier Transform is 

used to obtain the harmonic magnitudes. The 

overall sound pressure level (OASPL) noise measure is then calculated from these harmonics both 

in dB and using the A-weighted scale for dBA. 

ured near field noise, one tip radius away in 

the plane of rotation for the quarter scale 

propeller test data in Refs. (2) and (12) to 

(lit). 

aiction and the experimental data, but the waveform shape and magnitude are in excellent agreement. Additional data is available both ahead 

of and behind the plane of rotation. These data 

are compared with predictions in Ref. (E). The 

comparisons show the same phase shift along with 

a small magnitude difference. This magnitude 

To increase the execution speed, the presOnce the perturbation pressure time history 

Figure 7 compares the predicted and measThere is a phase shift between the pre3 
difference results in a few tenths of a dB difference in the OASPL. 

r(i) b 1 ad e e 1 eme 11 t c i rc u 1 a t ion s 

OPTIMIZATION MCTHOD 

An optimization method automates the process of choosing values of the independent variables to create a "best" result. In this work, 

the independent variables are the twist, chord, 

and sweep distributions, while the measure of 

the result is called the Index of Performance. 

The Index of Performance (IP) contains both the 

aerodynamic performance and noise of the propeller and is formulated so that an optimum 

geometry corresponds to a minimum of the funct ion. Different optimal geonLetries are obtained 

by changing the relative importance of noise 

versus efficiency in the IP. 

The conjugate directions method used for 

the optimization is that of Powell (16) as modified by Langwill (E). After searching along 

each of the current set of search directions, 

the technique searches along the direction resulting from the end points of the intermediate 

steps. This composite direction is the candidate conjugate direction. If the search proves 

rruitful, the conjugate direction replaces one 

of the previous search directions. 

ratic fit to three function evaluations. If the 

estimated minimum lies outside the fit interval, 

the interval is shifted towards the minimum 

until the minimum is bracketed. Then, if the 

prediction for the value of the objective function does not agree with the computed value, the 

fit interval width may be reduced. The minimum 

interval width allowable is based on the relative magnitude of the random portion of the 

objective function. This method is rapid and 

provides some immunity to numerical noise in the 

objective function. 

The Index of Performance (1P) incorporates 

the constraints in the manner of the SUMT fomillation developed by Fiacco and PlcCormick (2) : 

The line searches are done using a quadwhere 

LT propeller thrust coefficient 

(P propeller power coefficient 

($"'get parameter to obtain the desired Cp 

I\ OASPL noise measure in (dU) or (CiBA) 

m limit limit value of total sweep 

a the relative ivportance of efficiency 

versus noise 

b normalization for noise 

nels number of blade elements 

Wc Cp weighting 

wz weight on sweep 

12' weight on sweep "kink" constraint 

WS L weight to force hi-blade symmetric 

loading . 

lhe measure of sweep, S', is given by 

and the measure of inter-element angle is 

(9) 

where 

xh' 1,Yl ith node point coordinates 

tangent to the bound segment between 

nodes j and j + 1. 

?! 

Because of the numerical noise present in 

the calculation, the weightings in (8) are not 

decreased to zero as would be customary. Instead, the weightings are held constant and the 

offset target values are adjusted. This prevents the ainplification of the random noise in 

the IP. 

To further speed the optimization, the 

scheme is closely coupled with the aerodynamic 

analysis routine. This is done by recalculating 

only those portions of the model affected by 

geometry changes. For example, if an element is 

retwisted, only the induced velocities at the 

control point change. The induced velocities at 

the bound segment and all induced velocities on 

the other elements remain the same and are reused to calculate the performance. This is most 

beneficial initially when the search directions 

correspond to single geometric quantities. Compared with a complete solution for a design 

state, this technique is 2 to 24 times faster 

for evaluations of a 12 element blade model. 

method was to maximize straight blade ideal 

An initial test case for the optimization 

4 
efficiency as a function of the twist distribution. Goldstein's circulation distribution 

(g), which yields the maximum ideal efficiency 

under these assumptions, was obtained, verifying 

the method. 

For a more realistic prediction, a viscous 

drag estimate is included. This estimate 

assumes a laminar flow section operating in the 

laminar bucket with the same dependence of cdo 

on section thickness as for a Clark-Y airfoil 

(to be conservative). The Ci for stall was 

chosen to he 1.2, and stall was modeled by a 

sharp increase in the section drag coefficient. 

This drag model is justified because given the 

design lift coefficient, the corresponding airfoil shape can he found. By including viscous 

effects, the chord distribution can he a design 

variable in addition to the twist distribution. 

To investigate the trade-offs involved when 

noise is considered, various techniques for noise 

reduction are applied to a straight, a propleted 

and a hi-bladed propeller. These techniques 

include: reducing the diameter to reduce tip 

speed, shifting the loading inboard, and reducing the rotational rate to reduce tip speed. 

For these trade-offs, the design variables are 

the twist, chord and sweep distributions. 

Noise reductions are achieved through two 

mechanisms: phase cancellation and noise source 

strength reduction. Phase cancellation is 

accomplished either by adding blades or through 

sweep. The source strength dependence on propeller parameters depends on the type of noise 

source. The pressure perturbations due to 

thickness terms are proportional to relative 

Mach number squared, and the perturbations due 

to force terms are proportional to Mach number 

cubed. Loading noise is linearly proportional 

to the blade loadings though. Hence, reducing 

the relative Mach number through a reduction of 

the rotational speed or a reduction of the propeller diameter are expected to have a large 

effect on the noise. 

in Figs. 8 to 10 for the three blade types. The 

noise is calculated at a fixed position, 333 tip 

radii away, and 30" behind the plane of rotation. This angle is near the maximum of the 

directivity pattern, and at a distance appropriate for sideline measurements in take-oft. The 

trade-off curves were obtained by varying the 

parameters in the IP, Lq. (8). The weighting 

value, a, was shifted from consideration of 

only aerodynamic performance towards consideration of noise only. Eor each value of the 

weighting parameter, the propeller geometry was 

re-optimized. These optimal geometries produce 

the curves shown. 

the straight blade. A simple and very effective 

technique for performance improvement is to 

increase the blade number. The blade number as 

shown starts at three and increases in increments of one. When either ttle propeller solidity is held constant, or the blade planform is 

held constant, the noise is decreased while the 

efficiency goes up. Additional blades reduce 

The resulting optimized trade-offs are shown 

Figure 8 shows the trade-off results for 

the noise by introducing phase cancellation for 

certain harmonics while the strengths of the 

remaining harmonics are constant. The efficiency is increased because of reduced tip and 

swirl losses. The constant solidity cases have 

higher efficiencies than the constant planform 

cases because the viscous drq is lower. 

tive method shown. This trade-off starts with 

no sweep, and ends up with the tip swept hack 

about 45" from the initial geometry. The points 

along the curve represent equal increments in 

the weightings, but unequal changes in sweep. 

Sweep reduces the noise by introducing a phase 

shift and subsequent phase cancellation between 

noise from different sections of the blade. If 

the swept blade is concurrently retwisted to 

maintain the radial load distribution, the efficiency will remain nearly the same. A drawback 

to sweep is that for a blade with a low rotational Mach number, the amount of sweep needed 

for a useful reduction in noise is large. 

Reducing the rotational speed can reduce 

the noise with a small efficiency penalty. The 

changes shown correspond to 5 percent changes 

in rotational speed. The higher advance ratio 

results in higher blade loadings, hence greater 

tip and swirl losses, but there is a large 

decrease in the noise produced. 

the noise, but with a sizeable efficiency penalty. The efficiency drops because the effective diameter is reduced, increasing the disk 

loading. The inboard shift of the loading 

reduces the noise by reducing the strength of 

the high Mach number sources, and increasing 

the strength of the low Mach number sources. 

Although hard to quantify, the loadings shown 

are shifted so that the peak moves from the 

75 percent radius inboard to about the 

60 percent radius. 

A very ineffective method for reducing the 

noise is shown to be reducing the diameter. 

Changes here correspond to 5 percent reductions 

in diameter. The noise is reduced by the 

removal of the high Mach number tip sources. 

The remaining sources are increased in strength 

though, which offsets some of the reductions. 

The etficiency drops sharply because of the 

increased disk loading. This pertormance penalty is large and makes the method undesirable. 

Figure 9 shows the trade-off results for a 

propleted blade. The character of some of the 

trade-off curves has changed from the straight 

blade results. Most notable is the blade number 

trade-off which no longer shows an increase in 

efficiency with the decrease in noise. This 

curve is for a constant planform; for structural 

reasons, the tip thickness and chord are fixed. 

With increasing blade number, the constant blade 

planform propeller has increased viscous drag. 

Since the tip device is most beneficial at 

higher blade loadings, increasing the blade 

number will also reduce the efficiency gain. 

This reduced efficiency gain, along with the 

viscous drag increase, combine to reduce the 

efficiency of a propleted propeller. Further, 

Sweeping the blade is the next most effecShifting the loading inboard can reduce 

5 
while the total loading remains approximately 

constant, the additional tip thickness noise 

offsets the noise reductions from increased 

blade number. 

untavorable trends similar to those for the 

straight blade. Loading shift penalizes a propletrd propeller inure by unloading the tip and 

proplet, while the proplet's effect at higli 

loadings benefits the diameter reduction cases. 

able for the propleted propeller than the 

straight blade propeller. As the rotational 

speed is reduced to decrease the noise, the 

loading increases. Since proplets perform better at the higher loadings, there is less efficiency loss. 

blade propeller. Again, the loading shift and 

diameter reduction methods have unfavorable 

trends. Blade number increases are for a fixed 

planform, to support the tip device, and so 

increase the viscous drag while decreasing the 

effectiveness of the tip device. Rotational 

speed reduction is more favorable than for 

either the straight blade or propleted propeller. This is a result of the more effective 

tip device retaining more of the efficiency as 

the loading increases. Using a rotational speed 

reduction, the three way bi-blade propeller has 

about 1 dBA less noise at the same efficiency as 

the six way constant planform straight blade 

prope ller. 

The most interesting curve is for the 

effrct of sweep on a bi-blade propeller. After 

ail initial decrease in efficiency with decreasing noise, the trend flattens. If the curve is 

extrapolated to the efficiency of a three-way 

constant planform straight blade propeller, the 

Lliree-way bi-blade propeller would have a noise 

level that is 14 dBA lower than for the straight 

blade propeller. 

s UI'PlAKi 

Loading shift and diameter reduction show 

Rotational speed reduction is more favorFigure 10 presents the trade-offs for a biA unique preliminary design tool has been 

developed with the ability to simultaneously 

optirize the performance and noise of a propeller. liy using a vortex lattice aerodynamic 

analysis, a subsonic conipact source noise analysis, and an optimization scheme, an effective 

dnci fast preliminary design tool was created. 

This program was then used to investigate the 

noise/performance trade-offs for a straight, a 

propleted, and a bi-blade propeller. 

While the results demonstrate the utility 

of such a design tool, more importantly, the 

relative merits of the different noise reduction 

techniques are identified for optin,ized geometries. In short, for straight blades, increasing the blade number will increase efficiency 

and decrease the noise. Reducing the rotational 

speed will also reduce the noise hut will 

decrease the efficiency slightly. For a propleted lade, the most favorable method is to 

reduce the rotational speed, but again there is 

an erficiency penalty for noise reduction. 

Finally, for a hi-blade propeller, either a 

rotational speed reauction or sweeping the blade 

will reduce noise effectively. But, with its 

initially greater efficiency, there is a region 

where the hi-blade propeller has a higher efficiency arid a lower noise level than a similar 

straight blade p rope 1 ler. 

REFERENCES 

1. L.K. Chang, and J.P. Sullivan, "Optimization 

of Propeller Blade Shape by an Analytical 

I'lethod," AIAA Paper 82-1125, June 1982. 

2. J.P. Sullivan, L.K. Chang, and C.J. Miller, 

"'The Effect of Proplets and Ri-blades on 

the Pc,rformance and Noise of Propellers," 

SAE Paper 810600, April 1981. 

3. V.E. Baskin, L.S. Vildgruhe, Vozhdayev, 

Ye. S., and G.I. Maykapar, "Theory of the 

Lifting Airscrew," NASA TT-F-823, February 

1976. 

4. L.J. Bober, and L.K. Chang, "Factors 

Influencing the Predicted Performance of 

Advanced Propeller Designs," NASA TM-82676, 

1981. 

5. L.K. Chang, "The Theoretical Performance of 

High Efficiency Propellers," Ph.D. Thesis, 

Purdue University, 1980. 

6. T.A. Cgolf, O.L. Anderson, D.E. Edwards, and 

A.J. Landgrebe, "An Analysis for High Speed 

Propel ler-Nace 1 le Aerodynamic Performance 

Prediction." Volume 1, Theory and Initial 

Application, NASA CR-169450, and Volume 2, 

User's Xanual for the Computer Program, 

NASA CR-169451, (K79-912949-19, United 

'Technologies Research Center; KASA Contract 

NAS3-20961), June 1979. 

7. J.P. Sullivan, "The Effect of Blade Sweep 

on Propeller Performance," AZAA Paper 

77-716, June 1977. 

8. R.J. Jeracki, and G.A. fititchell, "Low and 

High Speed Propellers for General Aviation 

- Performance Potential and Recent Wind 

Tunnel Test Results," SAE Paper 811090, 

April 1981. Note: This reference describes 

the tests in which the 0.566 scale data were 

obtained. The data contained herein have 

not yet been published and were obtained in 

a private communication with Robert J. 

Jeracki, Propeller Research Section, 

Propulsion Aerodynamics Division, NASA 

Lewis Research Center, Cleveland, Ohio, 

January 1983. 

9. G.P. Succi, D.H. Munro, J.A. Zimmer, P.D. 

Dunbeck, E.E. Larabee, K.U. Ingard, and 

.J.L. Kerrebrock, "Noise and Performance of 

Propellers for Light Aircraft ," KASA 

CK-165732, 1981. 

6 
10. M.V. Lowson, "The Sound Field for 

Singularities in Motion," Proceedings of 

the Royal Society (London), Vol. A286, 

1965, pp. 559-572. 

11. D.H. Munro, "The Production of Sound by 

Moving Objects," Ph.D. Thesis, MIT, June 

1980. 

12. G.P. Succi, "Computed and Experimental 

Pressure Signatures from Two 1/4 Scale 

General Aviation Propellers," NASA 

CR-165 728, 1980. 

13. G.P. Succi, "Noise and Performance of 

General Aviation Aircraft: A Review of the 

MIT Study," SAE Paper 810586, April 1981. 

14. G.P. Succi, D.H. Munro, and J.A. Zimmer, 

"Experimental Verification of Propeller 

Noise Prediction," AIAA Journal, Vol. 20, 

No. 11, November 1982, pp. 1483-1491. 

15. C.J. Miller, "Optimally Designed Propellers 

Constrained by Noise," Ph.D. Thesis, Purdue 

University, 1984. 

16. M.J.D. Powell, "An Efficient Method for 

Finding the Minimum of a Function of 

Several Variables without Calculating 

Derivatives," Computer Journal, Vol. 7, No. 

2, July 1964, pp. 155-162. 

17. W.I. Zangwill, W.I., "Minimizing a Function 

without Calculating Derivatives," Computer 

Journal, Vol. 10, 1967, pp. 293-296. 

18. A.V. Fiacco, and G.P. McCormick, "Nonlinear 

Programming," New York, John Wiley & Sons, 

1968. 

19. S. Goldstein, "On the Vortex Theory of 

Screw Propellers," Proceedings of the Royal 

Society of London, Series A, Vol. 123, 

1929, pp. 440-465. 

7 

R & 

0 NODE POINTS 

0 CONTROL POINT 

HELICAL HORSESHOE 

STRENGTH r---- 

Figure 1. - Lifting line geometry. 

.25 

.20 

+ 

0 . 15 

+- 

m 

3 

CC 

I + 

s .10 

+ 

z 

w 

I=! 

LL 

.05 

0 

-. 05 i 

- 

A NASA TEST DATA, MACH = 0.11, B = 3 

0 PREDICTION, p3/4 = 28O, INVISCID 

0 PREDICTION, p3/4 = 28O, VISCOUS ESTIMATE 

n 

A 

32 

.6 .8 1.0 1.2 1.4 1.6 

ADVANCE RATIO, J = VlnD 

Figure 2. - Comparison of analytical and experimental 

thrust coefficients for the LT10282+4 propeller. 

Experiniental data fron, reference [jl. 
.25 

.20 

a 

0 

cf 

0 

a 

LL 

+ 

z 

- .15 

s 

0 .10 

w 

0 

8 .05 U 

0 

0 

NASA TEST DATA, M4CH = 0.11, B = 3 

0 PREDICTION, B314 = 28O, INVISCID 

0 PREDICTION, p3/4 = 28O, VISCOUS ESTIMATE - 

A 

A 

A 

a 

a 

A 

32 

-. 05 

ADVANCE RATIO, J = VinD 

Figure 3. - Com2arison of analytical and experimental 

power coefficients for the LT10282+4 propeller. 

Experimental data frorr reference @I. 

.10 

.08 

k0 

+- .06 Ln 

3 

CY 

I 

tLL 

k0 .04 

0 

- 5 

0" .02 U LL 

0 

0 

-. 02 

0 .1 .2 .3 .4 

ADVANCE RATIO, A = V/RR 

Figure 4. - Comparison of analytical and experimental 

thrust coefficients for the 1C160 propeller. Experimental data from reference [Q]. 
.7 

.6 

.5 

a 

0 

m- .4 

s 

0 

a 

U 

+ 

z 

w 

0 .3 

0 

u .2 

5 

LL 

0 

.1 

0 

-. 1 0 

p 

- 

'RED ICTION 

INVISCID 

- \\ \ FnnMrln 

J 

.1 .2 .3 .4 

ADVANCE RATIO, h = V/RR 

Figure 5. - Comparison of analytical and experimental 

power coefficients for the 1C160 propeller. Experimental data frm reference [u] reduced in value by 

ha If. 

I MASS SINKS: 

Z 

Figure 6. - Noise source locations. 
PRED ICTIGN, 

VISCOUS -.. 

‘L THEORY 

- EXPERIMENT 

-5 0 

-1.5 

I 

-1. 0 -. 5 D .5 1.0 1.5 

TIME, ms 

Figure 7. -Analytical and experir,ental comparison of 

near-field noise for propeller 1C160 in the plane of 

rotation. Experimental data a n d t h e o r y frolrl references [&I, la51. 
m81 r 

.79 *801 BLADE NUMBER 

C 0 N S TAN T 

SOLIDITY \ 

BLADE NUM3ER 

C ON STA N T 

SHIFT / 

s3 

63 65 67 69 71 

OASPL, dBA 

.74 

Figure 8. - Optimized noiselperformance trade-offs 

starting frorr the three way straight blade 63) 

baseline configuration. Design point: J = 0.8, 

Cp = 0.154. 
.79 

.78 

t 

0 

z 

- W 

0 .77 

W U U 

.76 

.75 

s B3 .79 r 

- 

R REDU - 

- 

- 

.78 

& .77 

- 5 

0 

LL 

.76 

.75 

.74 

SPL, dBA 

Figure 10. - Optimized noiselperformance trade-offs 

starting frorr, the three way syrnm?trically loaded 

bi-blade (SB3) baseline configuration. Design point; 

J = 0. 8, Cp = 0.154. 

0 REDU 

- 

- 

- 

- 

LOADING SHIFT 

1. Report No. 

NASA TM-86967 

Noise Constraints Effecting Optimal Propeller Designs 

2. Government Accession No 

7. Author(s) 

Christopher J. Miller, Lewis Research Center, and 

John P. Sullivan, Purdue University, West Lafayette, 

Indiana 

9. Performing Organization Name and Address 

National Aeronautics and Space Administration 

Lewis Research Center 

Cleveland, Ohio 44135 

2. Sponsoring Agency Name and Address 

National Aeronautics and Space Administration 

Washington, D.C. 20546 

Security Classif (of this report) 20 Security Classif (of this page) 21 No of pages 

Unclassified Unclassified 

3. Recipient's Catalog NO I 

22 Price' 

5. Report Date 

~ 

6 Performing Organization Code 

505-45-58 _____I 

a P+rforrning Organizntror~ Report No 

E-2449 

Work Unit No. 

_____- .____-- 

11 Contract or Grant No 

13. Type of Report and Period Covered 

Technical Memoranduiri _.. ___- 

14. Sponsoring Agency Code 

FSupplementary Notes 

Prepared for the SAE General Aviation Aircraft Meeting and Exposition sponsored 

by the Society of Automotive Engineers, Wichita, Kansas, April 16-19, 1985. 

j. Abstract 

A preliminary design tool for advanced propellers has been developed combining a 

fast vortex lattice aerodynamic analysis, a fast subsonic point source noise 

analysis, and an optimization scheme using a conjugate directions method. 

chord and sweep distributions are optimized to simultaneously improve both the 

aerodynamic performance and the noise observed at a fixed relative position. 

optimal noise/performance trade-offs for straight and advanced concept blades 

are presented. The techniques used include increasing the blade number, blade 

sweep, reducing the rotational speed, shifting the spanwise loading and diameter 

changes. Some methods yield improvement in both efficiency and noise. 

Twist, 

The 

Key Words (Suggested by Author@)) 18. Distribution Statement 

Performance; Aircraft; Aerodynamics; 

Blade; Optimization 

Unclassified - unlimited 

STAR Category 02