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A comparison of the overall and broadband noise characteristics of full-scale and model helicopter rotors
Journal of Sound and Vibration (1973) 30(2), 135-152
A COMPARISON OF THE OVERALL AND BROADBAND NOISE CHARACTERISTICS OF FULL-SCALE AND MODEL HELICOPTER ROTOP, S J. W. LEVERTONt AND J. S. POLLARD
Applied Acoustics Group, Igestland ltelicopters Limited, Yeovil, Somerset, England (Received 17 February 1973) The broadband noise generated by full-scale and model rotors is compared in terms of spectral content and the dependence on tip speed and rotor thrust/pitch angle. Low frequency broadband noise and high frequency broadband noise are studied separately and blade "scaling" effects are outlined. The degree of agreement between measurements and theoretical and semi,empirical prediction methods is reviewed together with the directionality patterns. The parameters relating to the overall noise are also discussed. It is shown that in general good agreement is obtained between the full-scale and model rotors when considering spectral content and the dependency of the noise levels on tip speed and thrust. The scaling factors usually considered applicable to the low frequency broadband noise do not, however, appear to apply to either the model or full scale rotors.
I. INTRODUCTION A review of the literature suggests that while good agreement is often obtained between full-scale and model rotor test data, there are also occasions when the two sets of results exhibit large discrepancies. Since precise details of rotor noise spectrum shapes, etc., are rarely given, it is practically impossible to determine if these differences are genuine or are a function of the particular rotor or test environment. Full-scale testing is usually carried out on real helicopter rotors in an open-air environment while model rotors are in general run in either ordinary laboratories or anechoic rooms. Thus vastly different parameters exist from one rotor configuration to another and although blade aerofoil sections are often similar and the thrust coefficients, CT, are of the same order, important aerodynamic parameters such as blade tip speed can be very different. The Reynolds numbers associated with full-scale and model rotors are also significantly different and it is reasonable to assume that, while on a real rotor the flow is turbulent, it is often laminar on model rotors. Another difference between full-scale and model rotors is that, although the real rotors have blade twist, typically 8 ~, model rotors are usually made nontwisted for simplicity. For these and other reasons m a n y investigators are of the opinion that model results are unrepresentative of full-scale rotors. This may be true to some extent but on the other hand testing on small-scale rotors in the past has highlighted a number of important aspects of rotor noise and the understanding of rotor noise sources has been considerably advanced by model results. This has largely been due to the fact that new concepts can easily be tested on t Visiting Lecturer, Institute of Sound and Vibration Research, University of Southampton, Southampton SO9 5NH, England. 135
136
J.x.v. LEVERTONAND J. S. POLLARD
models and the test environment and conditions can be precisely controlled. Model testing also has other advantages in that it is relatively inexpensive and changes to the rotor design, which are not practicable on full-scale rotors, can be investigated. Even so the advantages of model testing would be nullified if the results were unrepresentative of that which occurs on full-scale rotors. In an attempt to clarify some of these points a detailed comparison of the noise characteristics of full-scale rotors and model rotors has commenced and the preliminary results of this study are presented in this paper. In view of the large amount of data obtained and the wide range of aspects studied it is obviously not possible to include all the results in a paper of this type. A representative selection has been chosen, therefore, to illustrate the areas ofagreement and disagreement between the full-scale and model results. Before any conclusions could be drawn from the comparisons, it was necessary, of course, to verify that the particular rotors chosen had results which were representative of full-scale and model rotors. This was achieved by comparing the results from the rotors under consideration with those of other similar investigations; except for a few minor details good agreement was obtained. Rotor noise can be considered from a practical viewpoint to consist of three noise sources: namely, rotational noise, low-frequency broadband noise and high-frequency broadband noise. To date the main emphasis of the comparative study has been placed on the lowfrequency broadband noise, since this dominates the spectrum and overall noise characteristics on three- and four-bladed full-scale rotors. A limited review of the high-frequency broadband noise has been conducted together with a study ofthe overall noise characteristics; in this context it should be noted that, although rotational noise is not discussed in detail, account is taken of some of the preliminary results presented in reference [l]. 2. DESCRIPTION OF ROTOR RIGS 2.1. FULL SCALE The full-scale data used in this paper has mainly been obtained from an investigation carried out on a rotor whirl tower at the Weston division of Westland Helicopters Limited in which a two-bladed, 56 ft diameter rotor was run in an "upside-down" mode so that the flow was directed upwards. This configuration gave as near as possible "clean" test conditions since the flow recirculation effects were removed. To allow the directivity characteristics of the noise to be assessed, and also to eliminate ground effects, a balloon arrangement was used to support a microphone at a distance of approximately five rotor diameters. Table I summarizes the rotor parameters and test conditions appropriate to this investigat!on while reference [2] details the results. It should be noted here that, although measurements were taken above the "inverted" rotor, the standard convention for indicating directions is used and thus the measuring positions are referred to as "below the rotor disc plane". The data obtained from this series of tests has been compared with the results measured on the rotor operating in the conventional mode and, except for a change in level of the broadband noise and the more repeatable nature of the "clean rotor" results, the trends are similar. 2.2. MODEL SCALE Much of the model-scale data described in this paper has been taken from the 9 ft diameter model rotor rig in a non-anechoic laboratory at the Institute of Sound and Vibration Research, Southampton University. As with the full-scale tests the rotor was operated with the flow upwards and measurements were made both in and "below" the rotor disc plane at a distance of three rotor diameters from the rig centre line, and the tests were conducted under normal laboratory conditions with three blades. Table 2 summarizes the rotor parameters,
FULL-SCALE AND MODEL HELICOPTERNOISE
137
TABLE 1
Details of S55 fidl-scale rotor (a) Blade parameters Number of blades (N):
2
Radius (R):
27.85 ft
Chord (c):
16"4 in
Area (NcR): Section: Twist :
76.2 ft 2 NACA 0012 8~
CLT =
Cr
T
S
89V~ pNcR
S = rotor solidity -
NcR ztRz
CT = thrust coefficient p = density of air
(b) Test conditions Roto{ Tip speed speed (Vr) (rev/min) fit/s) 140 160 180 205 230 260
408 466 525 598 670 758
0 0 0 0 0 0 0
700
875
Total thrust (T) (Ib) 1 1 2 5 1 4 5 0 1850 2375
0.046 0-075 0.096 0.123 0.036 0.045 0.057 0.074 0.094 0-121 0.028 0-047 0.077 0.095 0.022 0.027 0-035 0.045 0.057 0.074 0.017 0-028 0.046 0.014 0.017 0-022 0.028 0-036 0.046
3050
3900
0.155 0.126 0.094 0.075 0-059
0.20
5000
CLT 0.121 0.154 0.096 0-123 0-075 0.096
values
Reynolds number range: 3.57 to 6.62 • 106. (c) Measurement position Microphone at F7 (11-5 ~ below disc) at 250 ft radius
TABLE 2
Details of LS. V.R. model rotor (a) Blade parameters Number of blades (N): Radius (R): Chord (c): Area (NcR): Section: Twist:
3 4.5 ft 4 in 4"5 ft z NACA 0012 0
(b) Test conditions Rotor speed (rev/min) 500 600 700 800 900 CLT value
Tip speed (V-r) (ft/s) 236 283 330 377 424
0
2
0 0 0 0 0 0
4"76 6-84 9-4 12"!5 15-42 0'016
Reynolds number range: 0-503 to 0-904 x 106. (c) Measurement position Microphone at 9 ~ below the disc and 25 ft radius
Pitch angle (0) (degrees) 4 6 Estimated total thrust (Ib) 11.7 16.3 22-9 29-79 37"8 0"039
20.22 29.22 39"5 51-6 65"4 0-068
29 '43 42"45
0-098
10
39"15 56'43
0-131
138
J.w.
LEVERTON AND J. S. POLLARD
together with the typical operating conditions, and further information can be found in references [3] and [4]. Comparisons are also made with results obtained from a 51 inch diameter model rotor which was run in the free field conditions of the I.S.V.R. anechoic room to study the directivity patterns over a tip speed range of 0-400 ft/s and a pitch range of 0-12 ~ [5]. 3. C O M P A R I S O N
TECHNIQUES
It is apparent from both the model and full-scale tests that there are three sources--namely, rotational noise, low-frequency broadband noise and high-frequency broadband noise--and each of these have distinctive and well-defined characteristics which are a function of the blade tip speed, the rotor thrust and the measurement angle relative to the rotor disc plane. Because o f this latter aspect it was considered essential that direct comparisons between the full-scale and model results should be made at corresponding angles to the rotor disc plane and, if possible, at equivalent distances from the rotor. The former condition was satisfied by choosing an angle o f a b o u t - 1 0 ~ for both rotors but, since measurements were made at distances of about five rotor diameters and three rotor diameters, respectively, for the full-scale and model rotor, identical distances were not obtainable. This difference in distance is not considered to be significant, however, since both sets of measurements were made in the far field. In addition to the results obtained at this position, references have also been made, where appropriate, to results obtained at other measuring positions. Since investigations reported in the literature and reeent model results [6, 7] indicate that rotor noise generation and, in particular, the spectrum shape can be significantly affected by blade tip shape, it was decided that the comparisons should be conducted on rotors with as near as possible the same tip shapes. This was achieved in general by using results from a full-scale $55 bladed rotor and the I.S.V.R. "standard" rotor, since both of these rotors are fitted with the conventional "Sikorsky-type faired tip". In this comparison it should be noted that the full-scale results are for a two-bladed rotor while the model results were obtained from a three-bladed rotor. It can be argued theoretically that the number of blades is important, particularly in the case of the rotational noise generated by the rotor. In an experimental study of the first 20 rotational noise harmonics by Leverton [8], however, the effect of blade number on the harmonic content and level of the rotational noise was not detected. Since the investigation was limited to a study of broadband noise, and to a lesser extent overall noise, it seems reasonable to conclude that the blade number aspects can be ignored, except possibly when considering the influence of the rotational noise on the overall noise. It should also be borne in mind that although the full-scale and model rotors were both run at a set of constant velocity conditions (see Tables 1 and 2), the other test variable was thrust for the full-scale rotor and pitch angle for the model rotor. Thus, while the effects ofchanging the velocity at constant thrust can be studied on the full-scale rotor, a similar comparison cannot be made on the model rotor. Also the real helicopter rotor blades are twisted while the model rotor blades are non-twisted. Thus, although a true zero thrust case is possible on the model rotor, the zero thrust case on the full-scale rotor corresponds to the condition when the rotor has equal positive and negative lift components. In this paper the full-scale and model-scale results are compared by examining the noise sources separately in terms of the spectral content, the dircctivity patterns, and their dependence on tip speed and rotor thrust/pitch angle. The relative frequency shifts due to differences in rotor RPM/tip speed,l and blade dimensions are also discussed. In addition to this the overall noise characteristics of the full-scale and model rotors are reviewed. "I" R P M = rev/min.
FULL-SCALEAND ?slODELtlELICOPTERNOISE
139
4. ROTOR NOISE--GENERAL CHARACTERISTICS The noise produced by a rotor is a combination of the sounds generated by a number of individual source mechanisms. From a practical viewpoint rotor noise can be considered to consist of rotational or discrete frequency noise and "vortex" or broadband noise. Rotational noise consists of individual frequencies or tones which shoxv up on narrowband analysis as harmonically related discrete frequencies or "peaks". With the exception of the fundamental and the first few harmonics which are a function ofthe steady (mean) parameters ofthe rotor, discrete frequency noise is generated by the fluctuating lift forces. Broadband noise shows up on analysis as a band or " h u m p " of random noise spread over a relatively wide frequency range. It was originally thought that the broadband spectrum was dominated by a "peak" of energy in the 250-500 Hz range but improved measurement and analysis techniques have shown that in this region both rotational components and broadband components are present'and that at higher frequencies there is another narrowband " h u m p " of random noise. These two broadband regions exhibit different trends and thus can be treated as two separate sources. For convenience they are termed "low-frequency broadband noise" and "highfrequency broadband noise", respectively. Typical narrowband analysis traces are presented in Figures l and 2 for tile full-scale and model rotors, respectively, and also indicated on these traces are the three sources of interest. The traces shown are from a "quick look" analysis and although the filter bandwidths are changed at 200 Hz and 1.5 kHz this does not imply that the three noise sources are defined by these frequency ranges since, as can be seen in the figures, overlapping of the sources occurs. At low frequencies the spectrum for both rotors is dominated by rotational noise while the mid-frequency region consists of a combination of rotational noise and low-frequency broadband noise, with the relative levels depending on the blade operating conditions. It will also be observed from Figures I and 2 that, although the full-scale results exhibit, particularly at the low-speed conditions, a well defined high-frequency broadband noise " h u m p " at 3-5 kHz, the corresponding characteristic associated with the model results is less clear. Full-scale measurements show that this " h u m p " is a function of measurement angle and model results suggest that this source is more predominant at angles other than that applicable to Figure 2With the preceding comments taken into account, Figures 1 and 2 show that, except for the differences in frequency ranges of each noise source between the two rotors, the spectrum shapes are very similar in character. 5. LOW-FREQUENCY BROADBAND NOISE As mentioned earlier the low-frequency broadband noise region is often a combination of broadband noise and discrete frequency noise (see Figures 1 and 2) and it is necessary to select a filter bandwidth which gives a fairly accurate measure of the broadband energy, while at the same time allowing the spectrum characteristics to be determined. For both the full-scale and model results it was found that a 20 Hz bandwidth filter was most suitable. Typical 20 Hz bandwidth results are given in Figures 3 and 4 for the full-scale and model rotors, respectively. A study of these results and the 5 Hz narrowband results presented in Figures l and 2 show that although there is some indication of a " h u m p " in the broadband or "vortex" region, the peak is not well defined. It appears from reviewing corresponding 1/3 octave band data that in fact the well-defined peak referred to by other investigators is largely a result of the analysis methods used. Even so, it is clear from Figures 3 and 4, that both rotors exhibit the same type of low-frequency broadband spectrum shape. There is some indication of a fall-off in level at low frequencies but it is not possible to define the nature of
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this since the noise appears to fluctuate considerably with time and to be a combination of amplitude varying discrete frequency noise and broadband noise. To allow some meaningful results to be obtained it was necessary to establish an "analysis model" of the form illustrated in Figure 5. This is based on a log frequency scale and the SPL variation has a "fiat" portion and a constant dBloctave "fall off" portion with the latter extending over a range of two octaves. Since it has proved impossible to determine the centre frequency o f this broadband noise, the only frequency which could be studied was the breakpoint frequency shown on Figure 5. The analysis model has been superimposed on some o f the results on Figures 3 and 4 and as can be seen the model is a fair representation of the low-frequency broadband noise spectrum, particularly at the high speed and thrust/pitch conditions. Htgh frequency broaband noise ( 1 0 0 Hz b o n d w l d l h )
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In the context of this it should be noted that the actual method used to locate the "breakpoint" frequency is very sensitive to slight errors in superimposing the " m o d e l " on the analysis traces. Since there does not appear, however, to be any alternative method which can be used to study the frequency dependency, this approach has been adopted for the analysis.
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144
J . W . LEVERTON AND J. S. POLLARD
5.1. VARIATION WITtt BLADE VELOCITY
Figures 6 and 8 present the low-frequency broadband noise variations with blade tip speed (V) at constant thrust for the full-scale rotor (Figure 6) and at constant pitch for the model (Figure 8). The full-scale rotor results at a measurement angle of 11.5 ~ below the rotor disc (Figure 6) show typically a I/6 to Va variation at constant thrust while at angles nearer the rotor axis the dependency is less than V6, particularly at the high thrust conditions. The model results suggest that, except at very low speeds, the noise follows a V4 law at zero pitch/thrust and a V6 law at high pitch. Since thrust (at constant pitch) is proportional to V2, the relationship at high pitch is thus also a Ix~ law at constant thrust. Thus there are slight differences between the full-scale and model results. It should be noted that the values obtained above are contrary to the results of other investigators who have suggested that when thrust is held constant broadband noise varies as V' [9, 10] or V2'7 [11]. &-- 8O
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5.2. VARIATIONWITH THRUST/PITCII Figures 7 and 9 show the low-frequency broadband noise variations with thrust/pitch for full-scale and model respectively. It will be noted that both the thrust (T) and the pitch (0) scales are presented on a log base and that the zero thrust/pitch condition has also been added for completeness. The results show that on the full-scale rotor (Figure 7) the noise exhibits two trends, one which decreases with thrust and the other which increases at a rate of approximately T'. The "change over point" appears to be dependent on the blade (~L or CT value and the measuring
FULL-SCALE
145
AND MODEL IIELICOPTER, NOISE
angle relative to the rotor disc plane. The T z law occurs at parameters corresponding to the normal operating conditions for such a rotor and corresponds to that usually associated with broadband noise [9, 10]. It is interesting to note that, although the initial decrease of SPL with increasing thrust can not be explained, the variation with thrust shows very similar trends to the results compiled by Widnall [12], but in this case a curve was used instead ofthe "two law" approach of Figure 7. 70
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rs.
pitch angle9 Model tests. ~, 600 rev/min;
The model results (Figure 9) show a similar T 2 law at the high pitch/thrust settings but at low pitch there is a fundamental difference between the full-scale and model results. On the model (with zero blade twist) the noise initially decreases in level as the blade pitch is increased and a "minimum" level of noise occurs at a pitch angle of 2 to 3 ~ The measurement of the power absorbed by the rotor also shows similar characteristics. This is considered to be due to the fact that at zero pitch the blades are travelling in their own wake, while at 2 ~ pitch the wake is "pushed down", giving a "cleaner" inflow over the blade and hence lower noise levels. As the pitch is further increased effects similar to those associated with the full-scale rotor
occur.
On a real rotor with twist, however, the case corresponding to zero pitch on the roodel never occurs since at zero thrust part of the blade is generating positive thrust and part is generating negative thrust. Thus except for a small region on the blade there is always a flow away from the "blade path" and hence the next blade is passing into relatively "clean" air. As pitch/thrust is initially increased the negative component of the lift is decreased and the positive lift component increases. Thus it is possible to visualize that the level of the "selfgenerated wake" noise remains approximately constant, until external effects arising from the increase in the inflow velocity and the generation of the tip vortex have a major influence.
146
J . w . LEVERTON AND J. S. POLLARD
This most likely explains the character of the full scale results of Figure 7 and the "higher pitch" model results. 5.3. FREQUENCY CHARACTERISTICS
Although the low-frequency broadband noise "peak" is normally associated with a Strouhal number relationship, it is not really possible to detect a "peak" frequency on either the full-scale or model results. There is, however, from Figures 3 and 4, an indication of a " h u m p " shape. In an attempt to overcome this problem the "break-point frequency" (mentioned previously in this section) has been plotted as a function of rotor speed for both the full-scale and model data as shown in Figures l0 and 1 I, respectively. i
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P 700
//
"6
rr
I ll
6OO
50o
//"" I l/
// !
OBk
///c/ t
I
f~ 3 88Vr " 255 I
I
Ik. 1"2k 1"4k 1"6k "Break point* frequency ( H z )
I
18k
Figure 11. Low-frequency broadband noise: "'break-point" frequency Pitch: o, 0~; • 6 ~
2k
vs.
rotor rev/min. Model tests.
It is clear from the model results that the break-point frequency, and by implication the " h u m p " centre frequency, is a function of tip speed. On the model results of Figure 11 the continuous line represents the best straight line through the points and the dotted line corresponds to a Strouhal number relationship normally associated with helicopter rotors. Since the " h u m p " centre frequency is in the order of 200/500 Hz lower than the break-point frequency, it can be seen that, although the broadband noise is Strouhal number dependent, its centre frequency is over-predicted. On the full-scale rotor, however, the "break-point frequency" near the rotor disc plane is for all practical purposes independent of the rotor speed (Figure 10). A plot ofdata measured
147
F U L L - S C A L E A N D M O D E L H E L I C O P T E R NOISE
at 75 ~ to the rotor disc gives similar results, except at high thrusts where the frequency appears vaguely to be a function of tip velocity. This variation from the expected Strouhal number relationship could be associated with errors in determining the break-point frequency and/or a significant change in the broadband spectrum with tip speed. Even if these aspects were taken into account, however, it appears that the full-scale rotor results would still not follow a Strouhal number relationship. Further examination ofthis subject is obviously needed since Wilkes [13] also found that the broadband noise peak was insensitive to tip speed on a full-scale rotor. It is also of interest to note that if the accepted Strouhal number relationship is used then it overpredicts even the break-point frequencies. 6. HIGH-FREQUENCY BROADBAND NOISE For the high-frequency broadband noise study 100 Hz bandwidth analysis was used for the full-scale rotor (Figure 3) while 20 Hz analysis was used for the model rotor (Figure 4). In this context it should be noted that when analysing high-frequency noise these filters are essentially narrowband filters. From the narrowband analysis traces illustrated in Figure 3 it is clear that there is a " h u m p " o f high frequency noise at 4 to 5 kHz on the full-scale rotor. The prominence of this noise source is very dependent on tip speed and the angle relative to the rotor disc. This can be seen by comparing the results in Figures 1 and 3. Near the rotor disc (Figure 1) the " h u m p " is practically non-existent at high speeds, while at angles near the rotor axis (Figure 3) it can be detected at all test conditions. t
I
I
I
I
V 4 IGW~.._~ - ._._.... Z oJ 50 J 13_ 40
f 140
I 160
I r80 Rotor
f 205
f 230
260
rev/min
Figure 12. High-frequency broadband noise: full-scale tests "hump" SPL rs. rotor rev/min. Microphone 11"5~ below disc. Thrust: • 0 Ib; O, 700 lb; zx, 3050 lb. On the model results illustrated in Figures 2 and 4 the " h u m p " or Peak is not so well defined and only really exists at low speeds and small pitch angles. At other measurement angles to the rotor disc the " h u m p " is, however, more predominant but again becomes insignificant as the pitch and/or tip speed is increased. On the model it has only been possible, therefore, to study the noise source in detail at low tip speeds. It is not clear from the results available if the " h u m p " disappears or if the other broadband energy in the region rises to a level above the " h u m p " when the rotor conditions are changed. The high frequency broadband noise on the full-scale rotor follows a Y~ law at constant thrust and T -~/6 law at constant velocity as shown in Figures 12 and 13, respectively. It is not possible to determine the corresponding velocity law on the model rotor, but as the pitch (0) is increased a 0 u2 law is observed which corresponds approximately to a T 3/s law: this is illustrated in Figure 14. Thus on both rotor systems the level of the high-frequency broadband noise is for all practical purposes independent of thrust.
148
J . W . L E V E R T O N A N D J . S. P O L L A R D 70
I
il
I
I
I
l
I
I
I
I
I
E Z 0 oJ
-._~__ A _ "-ZX----A
CoO
A
"~
-0----0 X
v
A
_ A . . . .A
~'-
~ ---- -- "-O-X
X
---,-.--,--.--.__.__.__
~ 50
+ I II I 0 700
I
f
t
f125
I
I
1850
Thrust
I 3050
I
I 5000
(Ib)
Figure 13. lligh-frequency broadband noise: full-scale tests "hump" SPL rs. thrust. +, F7, 160 rev/min; x, B4, 160 rev/min; zx, B4, 260 rev/min; O, F7, 230 rev/min. 5O
,
II
.
.
.
.
.
.
.
Z
&
8 i,'~.~
oJ
X
40
/
X
~
J 13. i
0•}
I
I
I
I
l
I
I
I
2
3
5
6
7
8
.9
IO
PJl~
angre
(deg)
'Figure 14. ttigh-frequency broadband noise: "hump" SPL rs. pitch angle model tests, x, 600 rev#nin. The full-scale results show that the frequency of this " h u m p " is a function of tip speed (i.e., a Strouhal type dependency). Although this has not been studied in detail on the model, there are some indications of a similar trend. Even so there does not appear to be any direct link between the " h u m p " frequencies for the model rotors and those of the full-scale rotor. This is not surprising since, if the high-frequency broadband noise source is due to some characteristic "wake shedding" process, then it will be greatly influenced by the boundary layer/wake thickness, and whether the flow is laminar or turbulent. Since the Reynolds numbers for the full-scale and model rotors are very different, significant flow differences can be expected. From the model tests with different blade tip shapes [6] it is also clear that other high frequency broadband noise "humps" can be present in the spectrum when the thickness of the blade trailing edge is increased. Lowson [7] has shown similar results on a model with "clipped" tip shapes. 7. OVERALL NOISE As stated previously the overall noise generated by a rotor is a function of the three individual noise sources. On both full-scale and model rotors the high-frequency broadband noise is generally relatively unimportant in terms of the overall noise level. This does not imply that it is subjectively insignificant, but the overall noise is effectively controlled by the levels of rotational and low-frequency broadband noise. As can be seen from Figures I and 2 the rotational noise is higher in level relative to the broadband noise on the model rotor than on the full-scale rotor. This is generally true and at angles near the rotor disc plane the broadband nois~ is significantly reduced, such that even on the full-scale rotor the overall noise is very dependent on rotational noise. To obtain the desired tip speeds on the model rotor, the rotational frequencies are significantly higher (typically by a factor of 3-4) than on the full-scale rotor. Thus the model rotor has a greater
FULL-SCALE AND MODEL HELICOPTER NOISE
149
part o f the spectrum dependent on the rotational noise than has the full-scale rotor. The rotational noise also b e c o m e s m o r e predominant as the number of blades and the physical size o f the rotor are decreased and on s o m e small model rotors the noise is effectively all rotational in character. It follows, therefore, that on the full-scale rotor the O A S P L follows (for all practical purposes) the same trends as the low-frequency broadband noise, except at angles very near to the rotor disc plane where the rotational noise is significant. On the other hand the overall noise for the m o d e l rotor is linked essentially to the rotational noise characteristics or a combination o f rotational and low frequency broadband noise. 8. ROTOR NOISE DIRECTIVITY
Due to the lack of suitable model rotor data it has not been possible to compare directly the directional characteristics of the low-frequency broadband noise, the high-frequency broadband noise and the overall noise. Some data on the rotational noise directivity pattern for a model rotor is, however, available [5] and, by studying the relationship between the broadband noise and the rotational noise on the full-scale rotor, an indirect comparison has been possible. 60 7"0 80 60 70 80 O~
i "~ Z (M
60
. ~
O*
30 ~
30 9
9' ~0
90*
60*
60*
90*
(a)
(b)
Overall noise and low frequency
~
broadband noise directivity
El "o
40
50
) ) j'
40 50
::i 90*
40
60
50
60
O~~
k-i ~
40
30
5O
\
60*
60"
90*
(d)
Cc) Rotational noise directivity
Figure 15. Full-scale rotor noise directivities. (a) Zero lift; 260 rotor rev/min (758 ft/s); (b) 5000 Ib; 260 rotor rev/min; (c) zero lift; 260 rotor rev/min; (d) 5000 Ib; 260 rotor rev/min. I1 [3, "Flat" SPL (20 Hz analysis); z~---,',, dBA; x x, OASPL (dB LIN); o o, peak 1/3 octave band; 9 I , low harmonics (6-10); 9 9 mid harmonics (12-16); 9 O, high harmonics (16-20).
150
J. W. LEVERTON AND J. S. POLLARD o* Direction } of thrust 70
1
l.__l
5B
lOB" V
~,,.....~,,;..o~_~\,~
135" ~x,
,~^~ "y
'.d~
/ 115" "Gutin"
13,5"
Figure 16. R o t a t i o n a l noise directivity model tests. Blade passing h a r m o n i c s - , 1 0 ; - - . - - , 15; . . . . ,20. Pitch 12 deg, tip spced 400 ft/s.
, I; . . . .
,2;-----,5;
On the full-scale rotor the higher rotational noise harmonics follow, as expected, the same general pattern as the low-frequency broadband noise (Figure 15). Near the rotor disc plane there is a high level of low-frequency rotational noise [1], but apart from this region the low-frequency broadband noise and the overall noise exhibit the same characteristics, completing a "figure of eight" dipole shape around the rotor. The rotational noise directivity plots of Figure 16 for the model rotor show the same characteristics as on the full-scale rotor, except for the fundamental, and to some extent the second harmonic. Thus both model and full-scale rotors exhibit the same general directivity characteristics with a "dip" occurring in the rotor disc plane as expected. A "dip" of about 5 dB also occurs on the rotor axis of the full-scale rotor results at high thrust. This "dip" is independent of blade velocity and cannot be explained. 9. DISCUSSION OF RESULTS Tile scaling factors which are normally used for low-frequency broadband noise seem somewhat arbitrary and as shown are not really applicable. An interesting point to note is that while on the model rotor the frequency of the low-frequency broadband noise is, as expected, a function of the blade speed, the full-scale rotor results are for all practical purposes independent of blade speed. This aspect obviously needs further investigation. Although the thrust parameters given by the semi-empirical prediction methods are substantiated at representative rotor operating conditions, there is a significant discrepancy between the measured and theoretical velocity dependencies. Predictions of the level of broadband noise have been made by using the accepted semi-empirical methods [10, 1I] and these have been compared with both the full-scale and model results. In both cases the correct order of noise is predicted at the higher thrust and tip speed conditions but, since the suggested directivity factor disagrees with that measured, a direct comparison at a particular measuring position between theory and experiment is not possible. There is no theoretical explanation for the high-frequency broadband noise except for a suggestion by Lowson [7] that it is linked to the details of the blade tip shape and possible
F U L L - S C A L E A N D MODEL H E L I C O P T E R NOISE
151
blade/tip vortex interaction. The full-scale and model results both show, however, very similar characteristics although there is no well-defined scaling factor between the two. , The directivity and absolute levels predicted by the Davidson and tlargest formula [10] agree fairly well with the measured low-frequency broadband noise at high thrust/high tip speed conditions on the full-scale rotor, except in the rotor disc plane region where the predicted value is zero. A similar comparison on the model is not possible, but the directivity patterns derived from theoretical considerations by Morfey and Tanna [14] agree well with those illustrated in the figures. 10. CONCLUDING REMARKS The most significant points to emerge from this study have been the importance of blade tip speed on the noise generated and the insensitivity of the broadband noise sources to thrust.'lt is also clear that many of the accepted prediction methods are suspect and that large differences can result if the broadband noise sources are scaled on a simple "Strouhal relationship" from one rotor to another. It is hoped that the next stage of this study will clarify this situation and, at least, allow accurate empirical formulae to be derived for each source. This further work will include a review of aspects linked with "wake thickness", etc., to see ifallowances for the differences in, say, Reynolds number between full-scale and model results can be taken into account. If, however, the differences in the frequency ranges of the noise sources on the full-scale and model rotors are taken into consideration, then it is clear that the spectrum characteristics for both rotors are remarkably similar. The general dependency of the noise levels on parameters such as blade velocity and thrust also show good agreement. Thus it appears that although the precise spectrum details of a full-scale rotor cannot be "scaled", model rotors are well suited to studying rotor noise, particularly when the general trends are of interest. ACKNOWLEDGMENTS Tile full-scale "inverted rotor" investigation, carried out at Westland Helicopters Limited, was funded by the Ministry of Defence and the model test programme conducted by J. W. Leverton at the Institute of Sound and Vibration Research was supported by the National Aeronautics and Space Administration. The authors wish to thank all these organizations for permission to publish this paper and colleagues for their help in its preparation. The authors are particularly grateful to members of the Westland Research Experimental Department who, under the direction of Mr Tony lves, were responsible for the full-scale test work and tile majority of the analyses presented. Views expressed in this paper are, however, those of the authors. Tile authors also wish to thank N. J. Stainer for permission to publish Figure 16 which has been adapted from the results given in reference [5]. REFERENCES I. J. W. LEVERTONand J. S. POLLARD1972 American tlelicopter Society Mideast Region, Symposhtm on "Status ofTesthtg and Modelling Techniques", Philadelphia, 26-28 October 1972. A comparison of the noise characteristics of full-scale and model helicopter rotors. 2. J. W. LEVERTON 1972 AGARD Specialists gleethtg on "Aerodynamics of Rotary tVings", Marseille, France, 13-15 September 1972. The noise characteristics of a large "clean" rotor. 3. J. W. LEVERTON 1967 lnstitttte of Sound and Vibration Research glemorandum ISA V 194. Helicopter noise.
J. w. LEVERTONAND J. S. POLLARD
152
4. J. W. LEVER'rON 1969 htstitute of Sound and Vibration Research AIemorandtan 311. Helicopter rotor noise--Final Report : Part 1I. 5. N.J. STAINER1969 M.Sc. Dissertation, Institute of Sotozd and Vibration Research. An experimental investigation into rotational noise--for a low solidity rotor. 6. J. S. POLLARD and J. W. LEVr-R'rON 1972 IVestland ttelicopters Lhnited Research Paper 414. Effect of blade tip planform on the noise and aerodynamics of a helicopter rotor. 7. M.V. LowsoN, A. WHATMOREand C. E. WHITFIELD1972 Lottghborough Uni~'ersityof Technology Report TT7202. Source mechanisms for rotor noise radiation. 8. J. W. LEVERTON 1968 AGARD Meethlg on "tlelicopter eropldsion S)'stems", Ottawa, Canada, Jttne 1968. tlelicopter noise. 9. R. SC~ILEGEC,R. KIYG and I t. MULL 1966 USAA VLABS Technical Report 66-4. Helicopter rotor noise generation and propagation. 10. I. M. DAVIOSONand T. J. HARGEST1965 Jottrnal of the Royal Aeronatttieal Society 69, 325-336. Helicopter noise. II. J. O. GODDARD and T. J. STUCKEY 1964 IVestland Helicopters Lhnited Report AAD. 4/I. 9 Investigation and prediction of helicopter rotor noise. 12. S. E. WIDNALL 1969 American lnstitltte of Aeronautics and Astronautics Journal 6, 279-281. A correlation of vortex noise data from helicopter main rotors. 13. L. 1t. WIEKES ! 968 IVestland ttelicopters Lhnited Research Paper 349. Noise research on helicopter rotors.
14. C. L. MORFEY and H . K. TANNA 1971 Journal of Sound and Vibration 15, 325-351. Sound radiation from a point force in circular motion.
A COMPARISON OF THE OVERALL AND BROADBAND NOISE CHARACTERISTICS OF FULL-SCALE AND MODEL HELICOPTER ROTOP, S J. W. LEVERTONt AND J. S. POLLARD
Applied Acoustics Group, Igestland ltelicopters Limited, Yeovil, Somerset, England (Received 17 February 1973) The broadband noise generated by full-scale and model rotors is compared in terms of spectral content and the dependence on tip speed and rotor thrust/pitch angle. Low frequency broadband noise and high frequency broadband noise are studied separately and blade "scaling" effects are outlined. The degree of agreement between measurements and theoretical and semi,empirical prediction methods is reviewed together with the directionality patterns. The parameters relating to the overall noise are also discussed. It is shown that in general good agreement is obtained between the full-scale and model rotors when considering spectral content and the dependency of the noise levels on tip speed and thrust. The scaling factors usually considered applicable to the low frequency broadband noise do not, however, appear to apply to either the model or full scale rotors.
I. INTRODUCTION A review of the literature suggests that while good agreement is often obtained between full-scale and model rotor test data, there are also occasions when the two sets of results exhibit large discrepancies. Since precise details of rotor noise spectrum shapes, etc., are rarely given, it is practically impossible to determine if these differences are genuine or are a function of the particular rotor or test environment. Full-scale testing is usually carried out on real helicopter rotors in an open-air environment while model rotors are in general run in either ordinary laboratories or anechoic rooms. Thus vastly different parameters exist from one rotor configuration to another and although blade aerofoil sections are often similar and the thrust coefficients, CT, are of the same order, important aerodynamic parameters such as blade tip speed can be very different. The Reynolds numbers associated with full-scale and model rotors are also significantly different and it is reasonable to assume that, while on a real rotor the flow is turbulent, it is often laminar on model rotors. Another difference between full-scale and model rotors is that, although the real rotors have blade twist, typically 8 ~, model rotors are usually made nontwisted for simplicity. For these and other reasons m a n y investigators are of the opinion that model results are unrepresentative of full-scale rotors. This may be true to some extent but on the other hand testing on small-scale rotors in the past has highlighted a number of important aspects of rotor noise and the understanding of rotor noise sources has been considerably advanced by model results. This has largely been due to the fact that new concepts can easily be tested on t Visiting Lecturer, Institute of Sound and Vibration Research, University of Southampton, Southampton SO9 5NH, England. 135
136
J.x.v. LEVERTONAND J. S. POLLARD
models and the test environment and conditions can be precisely controlled. Model testing also has other advantages in that it is relatively inexpensive and changes to the rotor design, which are not practicable on full-scale rotors, can be investigated. Even so the advantages of model testing would be nullified if the results were unrepresentative of that which occurs on full-scale rotors. In an attempt to clarify some of these points a detailed comparison of the noise characteristics of full-scale rotors and model rotors has commenced and the preliminary results of this study are presented in this paper. In view of the large amount of data obtained and the wide range of aspects studied it is obviously not possible to include all the results in a paper of this type. A representative selection has been chosen, therefore, to illustrate the areas ofagreement and disagreement between the full-scale and model results. Before any conclusions could be drawn from the comparisons, it was necessary, of course, to verify that the particular rotors chosen had results which were representative of full-scale and model rotors. This was achieved by comparing the results from the rotors under consideration with those of other similar investigations; except for a few minor details good agreement was obtained. Rotor noise can be considered from a practical viewpoint to consist of three noise sources: namely, rotational noise, low-frequency broadband noise and high-frequency broadband noise. To date the main emphasis of the comparative study has been placed on the lowfrequency broadband noise, since this dominates the spectrum and overall noise characteristics on three- and four-bladed full-scale rotors. A limited review of the high-frequency broadband noise has been conducted together with a study ofthe overall noise characteristics; in this context it should be noted that, although rotational noise is not discussed in detail, account is taken of some of the preliminary results presented in reference [l]. 2. DESCRIPTION OF ROTOR RIGS 2.1. FULL SCALE The full-scale data used in this paper has mainly been obtained from an investigation carried out on a rotor whirl tower at the Weston division of Westland Helicopters Limited in which a two-bladed, 56 ft diameter rotor was run in an "upside-down" mode so that the flow was directed upwards. This configuration gave as near as possible "clean" test conditions since the flow recirculation effects were removed. To allow the directivity characteristics of the noise to be assessed, and also to eliminate ground effects, a balloon arrangement was used to support a microphone at a distance of approximately five rotor diameters. Table I summarizes the rotor parameters and test conditions appropriate to this investigat!on while reference [2] details the results. It should be noted here that, although measurements were taken above the "inverted" rotor, the standard convention for indicating directions is used and thus the measuring positions are referred to as "below the rotor disc plane". The data obtained from this series of tests has been compared with the results measured on the rotor operating in the conventional mode and, except for a change in level of the broadband noise and the more repeatable nature of the "clean rotor" results, the trends are similar. 2.2. MODEL SCALE Much of the model-scale data described in this paper has been taken from the 9 ft diameter model rotor rig in a non-anechoic laboratory at the Institute of Sound and Vibration Research, Southampton University. As with the full-scale tests the rotor was operated with the flow upwards and measurements were made both in and "below" the rotor disc plane at a distance of three rotor diameters from the rig centre line, and the tests were conducted under normal laboratory conditions with three blades. Table 2 summarizes the rotor parameters,
FULL-SCALE AND MODEL HELICOPTERNOISE
137
TABLE 1
Details of S55 fidl-scale rotor (a) Blade parameters Number of blades (N):
2
Radius (R):
27.85 ft
Chord (c):
16"4 in
Area (NcR): Section: Twist :
76.2 ft 2 NACA 0012 8~
CLT =
Cr
T
S
89V~ pNcR
S = rotor solidity -
NcR ztRz
CT = thrust coefficient p = density of air
(b) Test conditions Roto{ Tip speed speed (Vr) (rev/min) fit/s) 140 160 180 205 230 260
408 466 525 598 670 758
0 0 0 0 0 0 0
700
875
Total thrust (T) (Ib) 1 1 2 5 1 4 5 0 1850 2375
0.046 0-075 0.096 0.123 0.036 0.045 0.057 0.074 0.094 0-121 0.028 0-047 0.077 0.095 0.022 0.027 0-035 0.045 0.057 0.074 0.017 0-028 0.046 0.014 0.017 0-022 0.028 0-036 0.046
3050
3900
0.155 0.126 0.094 0.075 0-059
0.20
5000
CLT 0.121 0.154 0.096 0-123 0-075 0.096
values
Reynolds number range: 3.57 to 6.62 • 106. (c) Measurement position Microphone at F7 (11-5 ~ below disc) at 250 ft radius
TABLE 2
Details of LS. V.R. model rotor (a) Blade parameters Number of blades (N): Radius (R): Chord (c): Area (NcR): Section: Twist:
3 4.5 ft 4 in 4"5 ft z NACA 0012 0
(b) Test conditions Rotor speed (rev/min) 500 600 700 800 900 CLT value
Tip speed (V-r) (ft/s) 236 283 330 377 424
0
2
0 0 0 0 0 0
4"76 6-84 9-4 12"!5 15-42 0'016
Reynolds number range: 0-503 to 0-904 x 106. (c) Measurement position Microphone at 9 ~ below the disc and 25 ft radius
Pitch angle (0) (degrees) 4 6 Estimated total thrust (Ib) 11.7 16.3 22-9 29-79 37"8 0"039
20.22 29.22 39"5 51-6 65"4 0-068
29 '43 42"45
0-098
10
39"15 56'43
0-131
138
J.w.
LEVERTON AND J. S. POLLARD
together with the typical operating conditions, and further information can be found in references [3] and [4]. Comparisons are also made with results obtained from a 51 inch diameter model rotor which was run in the free field conditions of the I.S.V.R. anechoic room to study the directivity patterns over a tip speed range of 0-400 ft/s and a pitch range of 0-12 ~ [5]. 3. C O M P A R I S O N
TECHNIQUES
It is apparent from both the model and full-scale tests that there are three sources--namely, rotational noise, low-frequency broadband noise and high-frequency broadband noise--and each of these have distinctive and well-defined characteristics which are a function of the blade tip speed, the rotor thrust and the measurement angle relative to the rotor disc plane. Because o f this latter aspect it was considered essential that direct comparisons between the full-scale and model results should be made at corresponding angles to the rotor disc plane and, if possible, at equivalent distances from the rotor. The former condition was satisfied by choosing an angle o f a b o u t - 1 0 ~ for both rotors but, since measurements were made at distances of about five rotor diameters and three rotor diameters, respectively, for the full-scale and model rotor, identical distances were not obtainable. This difference in distance is not considered to be significant, however, since both sets of measurements were made in the far field. In addition to the results obtained at this position, references have also been made, where appropriate, to results obtained at other measuring positions. Since investigations reported in the literature and reeent model results [6, 7] indicate that rotor noise generation and, in particular, the spectrum shape can be significantly affected by blade tip shape, it was decided that the comparisons should be conducted on rotors with as near as possible the same tip shapes. This was achieved in general by using results from a full-scale $55 bladed rotor and the I.S.V.R. "standard" rotor, since both of these rotors are fitted with the conventional "Sikorsky-type faired tip". In this comparison it should be noted that the full-scale results are for a two-bladed rotor while the model results were obtained from a three-bladed rotor. It can be argued theoretically that the number of blades is important, particularly in the case of the rotational noise generated by the rotor. In an experimental study of the first 20 rotational noise harmonics by Leverton [8], however, the effect of blade number on the harmonic content and level of the rotational noise was not detected. Since the investigation was limited to a study of broadband noise, and to a lesser extent overall noise, it seems reasonable to conclude that the blade number aspects can be ignored, except possibly when considering the influence of the rotational noise on the overall noise. It should also be borne in mind that although the full-scale and model rotors were both run at a set of constant velocity conditions (see Tables 1 and 2), the other test variable was thrust for the full-scale rotor and pitch angle for the model rotor. Thus, while the effects ofchanging the velocity at constant thrust can be studied on the full-scale rotor, a similar comparison cannot be made on the model rotor. Also the real helicopter rotor blades are twisted while the model rotor blades are non-twisted. Thus, although a true zero thrust case is possible on the model rotor, the zero thrust case on the full-scale rotor corresponds to the condition when the rotor has equal positive and negative lift components. In this paper the full-scale and model-scale results are compared by examining the noise sources separately in terms of the spectral content, the dircctivity patterns, and their dependence on tip speed and rotor thrust/pitch angle. The relative frequency shifts due to differences in rotor RPM/tip speed,l and blade dimensions are also discussed. In addition to this the overall noise characteristics of the full-scale and model rotors are reviewed. "I" R P M = rev/min.
FULL-SCALEAND ?slODELtlELICOPTERNOISE
139
4. ROTOR NOISE--GENERAL CHARACTERISTICS The noise produced by a rotor is a combination of the sounds generated by a number of individual source mechanisms. From a practical viewpoint rotor noise can be considered to consist of rotational or discrete frequency noise and "vortex" or broadband noise. Rotational noise consists of individual frequencies or tones which shoxv up on narrowband analysis as harmonically related discrete frequencies or "peaks". With the exception of the fundamental and the first few harmonics which are a function ofthe steady (mean) parameters ofthe rotor, discrete frequency noise is generated by the fluctuating lift forces. Broadband noise shows up on analysis as a band or " h u m p " of random noise spread over a relatively wide frequency range. It was originally thought that the broadband spectrum was dominated by a "peak" of energy in the 250-500 Hz range but improved measurement and analysis techniques have shown that in this region both rotational components and broadband components are present'and that at higher frequencies there is another narrowband " h u m p " of random noise. These two broadband regions exhibit different trends and thus can be treated as two separate sources. For convenience they are termed "low-frequency broadband noise" and "highfrequency broadband noise", respectively. Typical narrowband analysis traces are presented in Figures l and 2 for tile full-scale and model rotors, respectively, and also indicated on these traces are the three sources of interest. The traces shown are from a "quick look" analysis and although the filter bandwidths are changed at 200 Hz and 1.5 kHz this does not imply that the three noise sources are defined by these frequency ranges since, as can be seen in the figures, overlapping of the sources occurs. At low frequencies the spectrum for both rotors is dominated by rotational noise while the mid-frequency region consists of a combination of rotational noise and low-frequency broadband noise, with the relative levels depending on the blade operating conditions. It will also be observed from Figures I and 2 that, although the full-scale results exhibit, particularly at the low-speed conditions, a well defined high-frequency broadband noise " h u m p " at 3-5 kHz, the corresponding characteristic associated with the model results is less clear. Full-scale measurements show that this " h u m p " is a function of measurement angle and model results suggest that this source is more predominant at angles other than that applicable to Figure 2With the preceding comments taken into account, Figures 1 and 2 show that, except for the differences in frequency ranges of each noise source between the two rotors, the spectrum shapes are very similar in character. 5. LOW-FREQUENCY BROADBAND NOISE As mentioned earlier the low-frequency broadband noise region is often a combination of broadband noise and discrete frequency noise (see Figures 1 and 2) and it is necessary to select a filter bandwidth which gives a fairly accurate measure of the broadband energy, while at the same time allowing the spectrum characteristics to be determined. For both the full-scale and model results it was found that a 20 Hz bandwidth filter was most suitable. Typical 20 Hz bandwidth results are given in Figures 3 and 4 for the full-scale and model rotors, respectively. A study of these results and the 5 Hz narrowband results presented in Figures l and 2 show that although there is some indication of a " h u m p " in the broadband or "vortex" region, the peak is not well defined. It appears from reviewing corresponding 1/3 octave band data that in fact the well-defined peak referred to by other investigators is largely a result of the analysis methods used. Even so, it is clear from Figures 3 and 4, that both rotors exhibit the same type of low-frequency broadband spectrum shape. There is some indication of a fall-off in level at low frequencies but it is not possible to define the nature of
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this since the noise appears to fluctuate considerably with time and to be a combination of amplitude varying discrete frequency noise and broadband noise. To allow some meaningful results to be obtained it was necessary to establish an "analysis model" of the form illustrated in Figure 5. This is based on a log frequency scale and the SPL variation has a "fiat" portion and a constant dBloctave "fall off" portion with the latter extending over a range of two octaves. Since it has proved impossible to determine the centre frequency o f this broadband noise, the only frequency which could be studied was the breakpoint frequency shown on Figure 5. The analysis model has been superimposed on some o f the results on Figures 3 and 4 and as can be seen the model is a fair representation of the low-frequency broadband noise spectrum, particularly at the high speed and thrust/pitch conditions. Htgh frequency broaband noise ( 1 0 0 Hz b o n d w l d l h )
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Figure 3. Broadband noise spectra: full-scale tests. Spectrum analysis of rotor noise. Broadband noise study. Two $55 blades, thrust down, Test Series No. 1, Dircctivity study. Microphone B4 (75~ below disc).
In the context of this it should be noted that the actual method used to locate the "breakpoint" frequency is very sensitive to slight errors in superimposing the " m o d e l " on the analysis traces. Since there does not appear, however, to be any alternative method which can be used to study the frequency dependency, this approach has been adopted for the analysis.
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144
J . W . LEVERTON AND J. S. POLLARD
5.1. VARIATION WITtt BLADE VELOCITY
Figures 6 and 8 present the low-frequency broadband noise variations with blade tip speed (V) at constant thrust for the full-scale rotor (Figure 6) and at constant pitch for the model (Figure 8). The full-scale rotor results at a measurement angle of 11.5 ~ below the rotor disc (Figure 6) show typically a I/6 to Va variation at constant thrust while at angles nearer the rotor axis the dependency is less than V6, particularly at the high thrust conditions. The model results suggest that, except at very low speeds, the noise follows a V4 law at zero pitch/thrust and a V6 law at high pitch. Since thrust (at constant pitch) is proportional to V2, the relationship at high pitch is thus also a Ix~ law at constant thrust. Thus there are slight differences between the full-scale and model results. It should be noted that the values obtained above are contrary to the results of other investigators who have suggested that when thrust is held constant broadband noise varies as V' [9, 10] or V2'7 [11]. &-- 8O
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5.2. VARIATIONWITH THRUST/PITCII Figures 7 and 9 show the low-frequency broadband noise variations with thrust/pitch for full-scale and model respectively. It will be noted that both the thrust (T) and the pitch (0) scales are presented on a log base and that the zero thrust/pitch condition has also been added for completeness. The results show that on the full-scale rotor (Figure 7) the noise exhibits two trends, one which decreases with thrust and the other which increases at a rate of approximately T'. The "change over point" appears to be dependent on the blade (~L or CT value and the measuring
FULL-SCALE
145
AND MODEL IIELICOPTER, NOISE
angle relative to the rotor disc plane. The T z law occurs at parameters corresponding to the normal operating conditions for such a rotor and corresponds to that usually associated with broadband noise [9, 10]. It is interesting to note that, although the initial decrease of SPL with increasing thrust can not be explained, the variation with thrust shows very similar trends to the results compiled by Widnall [12], but in this case a curve was used instead ofthe "two law" approach of Figure 7. 70
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The model results (Figure 9) show a similar T 2 law at the high pitch/thrust settings but at low pitch there is a fundamental difference between the full-scale and model results. On the model (with zero blade twist) the noise initially decreases in level as the blade pitch is increased and a "minimum" level of noise occurs at a pitch angle of 2 to 3 ~ The measurement of the power absorbed by the rotor also shows similar characteristics. This is considered to be due to the fact that at zero pitch the blades are travelling in their own wake, while at 2 ~ pitch the wake is "pushed down", giving a "cleaner" inflow over the blade and hence lower noise levels. As the pitch is further increased effects similar to those associated with the full-scale rotor
occur.
On a real rotor with twist, however, the case corresponding to zero pitch on the roodel never occurs since at zero thrust part of the blade is generating positive thrust and part is generating negative thrust. Thus except for a small region on the blade there is always a flow away from the "blade path" and hence the next blade is passing into relatively "clean" air. As pitch/thrust is initially increased the negative component of the lift is decreased and the positive lift component increases. Thus it is possible to visualize that the level of the "selfgenerated wake" noise remains approximately constant, until external effects arising from the increase in the inflow velocity and the generation of the tip vortex have a major influence.
146
J . w . LEVERTON AND J. S. POLLARD
This most likely explains the character of the full scale results of Figure 7 and the "higher pitch" model results. 5.3. FREQUENCY CHARACTERISTICS
Although the low-frequency broadband noise "peak" is normally associated with a Strouhal number relationship, it is not really possible to detect a "peak" frequency on either the full-scale or model results. There is, however, from Figures 3 and 4, an indication of a " h u m p " shape. In an attempt to overcome this problem the "break-point frequency" (mentioned previously in this section) has been plotted as a function of rotor speed for both the full-scale and model data as shown in Figures l0 and 1 I, respectively. i
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Figure 10. Low-frequency broadband noise: "break-point" frequency v s . rotor rev/min. Full-scale lests. Microphone F7 11"5~ below disc. Thrust: x, 0 Ib; o, 700 Ib; FI, 1850 Ib; z~, 3050 lb.
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rotor rev/min. Model tests.
It is clear from the model results that the break-point frequency, and by implication the " h u m p " centre frequency, is a function of tip speed. On the model results of Figure 11 the continuous line represents the best straight line through the points and the dotted line corresponds to a Strouhal number relationship normally associated with helicopter rotors. Since the " h u m p " centre frequency is in the order of 200/500 Hz lower than the break-point frequency, it can be seen that, although the broadband noise is Strouhal number dependent, its centre frequency is over-predicted. On the full-scale rotor, however, the "break-point frequency" near the rotor disc plane is for all practical purposes independent of the rotor speed (Figure 10). A plot ofdata measured
147
F U L L - S C A L E A N D M O D E L H E L I C O P T E R NOISE
at 75 ~ to the rotor disc gives similar results, except at high thrusts where the frequency appears vaguely to be a function of tip velocity. This variation from the expected Strouhal number relationship could be associated with errors in determining the break-point frequency and/or a significant change in the broadband spectrum with tip speed. Even if these aspects were taken into account, however, it appears that the full-scale rotor results would still not follow a Strouhal number relationship. Further examination ofthis subject is obviously needed since Wilkes [13] also found that the broadband noise peak was insensitive to tip speed on a full-scale rotor. It is also of interest to note that if the accepted Strouhal number relationship is used then it overpredicts even the break-point frequencies. 6. HIGH-FREQUENCY BROADBAND NOISE For the high-frequency broadband noise study 100 Hz bandwidth analysis was used for the full-scale rotor (Figure 3) while 20 Hz analysis was used for the model rotor (Figure 4). In this context it should be noted that when analysing high-frequency noise these filters are essentially narrowband filters. From the narrowband analysis traces illustrated in Figure 3 it is clear that there is a " h u m p " o f high frequency noise at 4 to 5 kHz on the full-scale rotor. The prominence of this noise source is very dependent on tip speed and the angle relative to the rotor disc. This can be seen by comparing the results in Figures 1 and 3. Near the rotor disc (Figure 1) the " h u m p " is practically non-existent at high speeds, while at angles near the rotor axis (Figure 3) it can be detected at all test conditions. t
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Figure 12. High-frequency broadband noise: full-scale tests "hump" SPL rs. rotor rev/min. Microphone 11"5~ below disc. Thrust: • 0 Ib; O, 700 lb; zx, 3050 lb. On the model results illustrated in Figures 2 and 4 the " h u m p " or Peak is not so well defined and only really exists at low speeds and small pitch angles. At other measurement angles to the rotor disc the " h u m p " is, however, more predominant but again becomes insignificant as the pitch and/or tip speed is increased. On the model it has only been possible, therefore, to study the noise source in detail at low tip speeds. It is not clear from the results available if the " h u m p " disappears or if the other broadband energy in the region rises to a level above the " h u m p " when the rotor conditions are changed. The high frequency broadband noise on the full-scale rotor follows a Y~ law at constant thrust and T -~/6 law at constant velocity as shown in Figures 12 and 13, respectively. It is not possible to determine the corresponding velocity law on the model rotor, but as the pitch (0) is increased a 0 u2 law is observed which corresponds approximately to a T 3/s law: this is illustrated in Figure 14. Thus on both rotor systems the level of the high-frequency broadband noise is for all practical purposes independent of thrust.
148
J . W . L E V E R T O N A N D J . S. P O L L A R D 70
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Figure 13. lligh-frequency broadband noise: full-scale tests "hump" SPL rs. thrust. +, F7, 160 rev/min; x, B4, 160 rev/min; zx, B4, 260 rev/min; O, F7, 230 rev/min. 5O
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'Figure 14. ttigh-frequency broadband noise: "hump" SPL rs. pitch angle model tests, x, 600 rev#nin. The full-scale results show that the frequency of this " h u m p " is a function of tip speed (i.e., a Strouhal type dependency). Although this has not been studied in detail on the model, there are some indications of a similar trend. Even so there does not appear to be any direct link between the " h u m p " frequencies for the model rotors and those of the full-scale rotor. This is not surprising since, if the high-frequency broadband noise source is due to some characteristic "wake shedding" process, then it will be greatly influenced by the boundary layer/wake thickness, and whether the flow is laminar or turbulent. Since the Reynolds numbers for the full-scale and model rotors are very different, significant flow differences can be expected. From the model tests with different blade tip shapes [6] it is also clear that other high frequency broadband noise "humps" can be present in the spectrum when the thickness of the blade trailing edge is increased. Lowson [7] has shown similar results on a model with "clipped" tip shapes. 7. OVERALL NOISE As stated previously the overall noise generated by a rotor is a function of the three individual noise sources. On both full-scale and model rotors the high-frequency broadband noise is generally relatively unimportant in terms of the overall noise level. This does not imply that it is subjectively insignificant, but the overall noise is effectively controlled by the levels of rotational and low-frequency broadband noise. As can be seen from Figures I and 2 the rotational noise is higher in level relative to the broadband noise on the model rotor than on the full-scale rotor. This is generally true and at angles near the rotor disc plane the broadband nois~ is significantly reduced, such that even on the full-scale rotor the overall noise is very dependent on rotational noise. To obtain the desired tip speeds on the model rotor, the rotational frequencies are significantly higher (typically by a factor of 3-4) than on the full-scale rotor. Thus the model rotor has a greater
FULL-SCALE AND MODEL HELICOPTER NOISE
149
part o f the spectrum dependent on the rotational noise than has the full-scale rotor. The rotational noise also b e c o m e s m o r e predominant as the number of blades and the physical size o f the rotor are decreased and on s o m e small model rotors the noise is effectively all rotational in character. It follows, therefore, that on the full-scale rotor the O A S P L follows (for all practical purposes) the same trends as the low-frequency broadband noise, except at angles very near to the rotor disc plane where the rotational noise is significant. On the other hand the overall noise for the m o d e l rotor is linked essentially to the rotational noise characteristics or a combination o f rotational and low frequency broadband noise. 8. ROTOR NOISE DIRECTIVITY
Due to the lack of suitable model rotor data it has not been possible to compare directly the directional characteristics of the low-frequency broadband noise, the high-frequency broadband noise and the overall noise. Some data on the rotational noise directivity pattern for a model rotor is, however, available [5] and, by studying the relationship between the broadband noise and the rotational noise on the full-scale rotor, an indirect comparison has been possible. 60 7"0 80 60 70 80 O~
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Figure 15. Full-scale rotor noise directivities. (a) Zero lift; 260 rotor rev/min (758 ft/s); (b) 5000 Ib; 260 rotor rev/min; (c) zero lift; 260 rotor rev/min; (d) 5000 Ib; 260 rotor rev/min. I1 [3, "Flat" SPL (20 Hz analysis); z~---,',, dBA; x x, OASPL (dB LIN); o o, peak 1/3 octave band; 9 I , low harmonics (6-10); 9 9 mid harmonics (12-16); 9 O, high harmonics (16-20).
150
J. W. LEVERTON AND J. S. POLLARD o* Direction } of thrust 70
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Figure 16. R o t a t i o n a l noise directivity model tests. Blade passing h a r m o n i c s - , 1 0 ; - - . - - , 15; . . . . ,20. Pitch 12 deg, tip spced 400 ft/s.
, I; . . . .
,2;-----,5;
On the full-scale rotor the higher rotational noise harmonics follow, as expected, the same general pattern as the low-frequency broadband noise (Figure 15). Near the rotor disc plane there is a high level of low-frequency rotational noise [1], but apart from this region the low-frequency broadband noise and the overall noise exhibit the same characteristics, completing a "figure of eight" dipole shape around the rotor. The rotational noise directivity plots of Figure 16 for the model rotor show the same characteristics as on the full-scale rotor, except for the fundamental, and to some extent the second harmonic. Thus both model and full-scale rotors exhibit the same general directivity characteristics with a "dip" occurring in the rotor disc plane as expected. A "dip" of about 5 dB also occurs on the rotor axis of the full-scale rotor results at high thrust. This "dip" is independent of blade velocity and cannot be explained. 9. DISCUSSION OF RESULTS Tile scaling factors which are normally used for low-frequency broadband noise seem somewhat arbitrary and as shown are not really applicable. An interesting point to note is that while on the model rotor the frequency of the low-frequency broadband noise is, as expected, a function of the blade speed, the full-scale rotor results are for all practical purposes independent of blade speed. This aspect obviously needs further investigation. Although the thrust parameters given by the semi-empirical prediction methods are substantiated at representative rotor operating conditions, there is a significant discrepancy between the measured and theoretical velocity dependencies. Predictions of the level of broadband noise have been made by using the accepted semi-empirical methods [10, 1I] and these have been compared with both the full-scale and model results. In both cases the correct order of noise is predicted at the higher thrust and tip speed conditions but, since the suggested directivity factor disagrees with that measured, a direct comparison at a particular measuring position between theory and experiment is not possible. There is no theoretical explanation for the high-frequency broadband noise except for a suggestion by Lowson [7] that it is linked to the details of the blade tip shape and possible
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blade/tip vortex interaction. The full-scale and model results both show, however, very similar characteristics although there is no well-defined scaling factor between the two. , The directivity and absolute levels predicted by the Davidson and tlargest formula [10] agree fairly well with the measured low-frequency broadband noise at high thrust/high tip speed conditions on the full-scale rotor, except in the rotor disc plane region where the predicted value is zero. A similar comparison on the model is not possible, but the directivity patterns derived from theoretical considerations by Morfey and Tanna [14] agree well with those illustrated in the figures. 10. CONCLUDING REMARKS The most significant points to emerge from this study have been the importance of blade tip speed on the noise generated and the insensitivity of the broadband noise sources to thrust.'lt is also clear that many of the accepted prediction methods are suspect and that large differences can result if the broadband noise sources are scaled on a simple "Strouhal relationship" from one rotor to another. It is hoped that the next stage of this study will clarify this situation and, at least, allow accurate empirical formulae to be derived for each source. This further work will include a review of aspects linked with "wake thickness", etc., to see ifallowances for the differences in, say, Reynolds number between full-scale and model results can be taken into account. If, however, the differences in the frequency ranges of the noise sources on the full-scale and model rotors are taken into consideration, then it is clear that the spectrum characteristics for both rotors are remarkably similar. The general dependency of the noise levels on parameters such as blade velocity and thrust also show good agreement. Thus it appears that although the precise spectrum details of a full-scale rotor cannot be "scaled", model rotors are well suited to studying rotor noise, particularly when the general trends are of interest. ACKNOWLEDGMENTS Tile full-scale "inverted rotor" investigation, carried out at Westland Helicopters Limited, was funded by the Ministry of Defence and the model test programme conducted by J. W. Leverton at the Institute of Sound and Vibration Research was supported by the National Aeronautics and Space Administration. The authors wish to thank all these organizations for permission to publish this paper and colleagues for their help in its preparation. The authors are particularly grateful to members of the Westland Research Experimental Department who, under the direction of Mr Tony lves, were responsible for the full-scale test work and tile majority of the analyses presented. Views expressed in this paper are, however, those of the authors. Tile authors also wish to thank N. J. Stainer for permission to publish Figure 16 which has been adapted from the results given in reference [5]. REFERENCES I. J. W. LEVERTONand J. S. POLLARD1972 American tlelicopter Society Mideast Region, Symposhtm on "Status ofTesthtg and Modelling Techniques", Philadelphia, 26-28 October 1972. A comparison of the noise characteristics of full-scale and model helicopter rotors. 2. J. W. LEVERTON 1972 AGARD Specialists gleethtg on "Aerodynamics of Rotary tVings", Marseille, France, 13-15 September 1972. The noise characteristics of a large "clean" rotor. 3. J. W. LEVERTON 1967 lnstitttte of Sound and Vibration Research glemorandum ISA V 194. Helicopter noise.
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4. J. W. LEVER'rON 1969 htstitute of Sound and Vibration Research AIemorandtan 311. Helicopter rotor noise--Final Report : Part 1I. 5. N.J. STAINER1969 M.Sc. Dissertation, Institute of Sotozd and Vibration Research. An experimental investigation into rotational noise--for a low solidity rotor. 6. J. S. POLLARD and J. W. LEVr-R'rON 1972 IVestland ttelicopters Lhnited Research Paper 414. Effect of blade tip planform on the noise and aerodynamics of a helicopter rotor. 7. M.V. LowsoN, A. WHATMOREand C. E. WHITFIELD1972 Lottghborough Uni~'ersityof Technology Report TT7202. Source mechanisms for rotor noise radiation. 8. J. W. LEVERTON 1968 AGARD Meethlg on "tlelicopter eropldsion S)'stems", Ottawa, Canada, Jttne 1968. tlelicopter noise. 9. R. SC~ILEGEC,R. KIYG and I t. MULL 1966 USAA VLABS Technical Report 66-4. Helicopter rotor noise generation and propagation. 10. I. M. DAVIOSONand T. J. HARGEST1965 Jottrnal of the Royal Aeronatttieal Society 69, 325-336. Helicopter noise. II. J. O. GODDARD and T. J. STUCKEY 1964 IVestland Helicopters Lhnited Report AAD. 4/I. 9 Investigation and prediction of helicopter rotor noise. 12. S. E. WIDNALL 1969 American lnstitltte of Aeronautics and Astronautics Journal 6, 279-281. A correlation of vortex noise data from helicopter main rotors. 13. L. 1t. WIEKES ! 968 IVestland ttelicopters Lhnited Research Paper 349. Noise research on helicopter rotors.
14. C. L. MORFEY and H . K. TANNA 1971 Journal of Sound and Vibration 15, 325-351. Sound radiation from a point force in circular motion.