Aerodynamic performance and system effects of propeller fans (RP-1223).

INTRODUCTION

Little information exists for accurately predicting the aerodynamic and acoustical response of small propeller fans to common appurtenances at the fan inlet, which are referred to as system effects. Hence, ASHRAE RP-1223 (Darvennes et al. 2009a) was initiated to experimentally measure air and sound performance of propeller fans with systematic variation of conditions at the inlet plane of the fan. The intent was to simulate typical "in the field" installations of the fans, such as mitered elbows mounted at various angles at the fan inlet plane, inlet duct contractions of various area ratios, and walls perpendicular to the fan axis and located at various distances from the inlet. The tests were conducted in accordance with AMCA (1999). A complete description of the test program, including the experimental apparatus and test procedure, was provided in Young et al. (2009). Acoustical data resulting from this test program are presented in Darvennes et al. (2009a). In the present paper, aerodynamic system effects are presented in the form of dimensionless loss coefficients.

AERODYNAMIC DATA REDUCTION EQUATIONS

Measured aerodynamic performance of fans is generally presented in the form of a graph showing pressure, power, and efficiency as a function of volumetric flow rate. The energy transferred to the air by the impeller results in an increase in static and velocity pressure. The sum of the two pressures is referred to as the total pressure. Consider steady, adiabatic, one-dimensional flow of an incompressible flow through a fan where elevation changes are negligible. Let the subscript 1 denote an upstream location and the subscript 2 indicate a location downstream of the fan. In that instance, applying a first law energy balance to a control volume surrounding the fan yields the following expression for the mechanical energy added to the air by the fan per unit mass of fluid flow:

w = ([P.sub.2]/[rho] - [P.sub.1]/[rho]) + ([V.sub.2.sup.2]/2 - [V.sub.1.sup.2]/2) (1)

This can be interpreted in terms of the change of total pressure across the fan:

w = [[[P.sub.T2] - [P.sub.T1]]/[rho]] (2)

Multiplication of Equation 2 by the mass flow rate of air yields the following expression for the total power imparted to the air:

[H.sub.O] = [[m([P.sub.T2] - [P.sub.T1])]/[rho]] (3)

Applying the continuity equation, the fan total power can be written in terms of the volumetric flow rate

[H.sub.O] = Q([P.sub.T2] - [P.sub.T1]). (4)

In many fan applications the air cannot be considered incompressible, such that a compressibility coefficient [K.sub.P] must be applied to Equation 4. Therefore, in this study the fan power output was calculated by

[H.sub.O] = Q([P.sub.T2] - [P.sub.T1])[K.sub.P]. (5)

Equations necessary to calculate [K.sub.P] were taken from AMCA (1999).

When a calibrated electric motor is used to measure input, the fan shaft input power H can be evaluated in terms of the measured power input to the motor W and the motor efficiency [eta]. In that case,

H = W[eta]. (6)

The total fan efficiency is the ratio of the total fan power output to the shaft power input. Hence, by Equation 5 and referring to Equation 6, in this study the fan total efficiency was calculated per Equation 7:

[[eta].sub.T] = [[Q([P.sub.T2] - [P.sub.T1])[K.sub.P]]/H] (7)

Referring to Figure 1 (which was taken from AMCA [1999]) for flow through multiple flow nozzles located in an outlet chamber setup, the volumetric flow rate [Q.sub.5] can be expressed by Equation 8:

[FIGURE 1 OMITTED]

[Q.sub.5] = [[Y.sub.N]/[[rho].sub.5]][square root of [2([P.sub.5] - [P.sub.6])]][summation]([C.sub.N][A.sub.N]) (8)

where 5 denotes the section upstream of the nozzle, and 6 indicates the nozzle throat. Additional equations necessary to support the flow calculation per Equation 8 were obtained from AMCA (1999). The fan airflow rate at test condition [Q.sub.2] was obtained from the continuity equation:

[Q.sub.2] = [Q.sub.5]([[rho].sub.5]/[[rho].sub.2]) (9)

The fan air density [[rho].sub.2] was calculated from measurements of the atmospheric air density [[rho].sub.0], the total pressure at the fan inlet [P.sub.T1], and the total (stagnation) temperature at the fan inlet [T.sub.T1], per AMCA (1999). However, for all outlet chamber setups, [P.sub.T1] was equal to 0 and [T.sub.T1] was equal to [T.sub.DO] per AMCA (1999), since the inlet air velocity was assumed to be negligible. The air velocity at the fan outlet [V.sub.2] was calculated based on the fan cross section, as shown in the following equation:

[V.sub.2] = [[Q.sub.2]/[A.sub.2]] (10)

Therein, the velocity pressure at the fan outlet was determined as follows:

[P.sub.V2] = [([V.sub.2]/[square root of 2]).sup.2][[rho].sub.2] (11)

The total pressure at the fan outlet plane [P.sub.T2] was calculated based on the measured static pressure inside the chamber (where the air velocity was insignificant). Likewise, the influence of the 3D short duct connecting the fan to the chamber was assumed to be negligible. Hence, referring to Figure 1:

[P.sub.T2] = [P.sub.S7] + [P.sub.V2] (12)

In the present study, the total pressure rise across the fan was calculated per Equation 13:

[P.sub.T] = [P.sub.T2] - [P.sub.T1] (13)

The aerodynamic performance of a fan can be reduced by inappropriate connections to the duct system. Duct connections to the fan should be designed such that the air enters and leaves the fan with a nearly uniform velocity profile, with no abrupt changes in flow direction. Fans are normally performance-tested with an open inlet, so that air flow entering the fan is approximately uniform. In this study, such a test configuration is referred to as the baseline case. In practice, space may be limited for fan installation (e.g., if there are cabinet walls in close proximity to the fan entrance). Otherwise, it may be necessary to mount elbows or contractions at the fan inlet. Such non-ideal installations incur a performance penalty (i.e., a reduction in total pressure rise at any given volumetric flow rate). This is termed a fan system effect factor, and its influence on fan performance is in addition to the typical pressure loss due to ductwork, fittings, equipment, etc.

The fan laws are approximate relationships that can be used either to estimate the aerodynamic performance of a particular fan under variable operating conditions, or to predict the performance of one fan using data from another geometrically similar fan. For example, during a fan performance test the air density or shaft speed may vary from one test condition to another. It is often desired to convert the test results to those that would prevail under conditions of constant air density and/or shaft speed. Hence, referring to AMCA (1999), actual volumetric flow rate data may be converted to standard test conditions using the following scaling law:

[Q.sub.C] = Q([N.sub.C]/N)([K.sub.P]/[K.sub.PC]) (14)

Likewise, total pressure rise data across the fan scales according to the following:

[P.sub.TC] = [P.sub.T][([N.sub.C]/N).sup.2]([[rho].sub.C]/[rho])([K.sub.P]/[K.sub.PC]) (15)

However, to a close approximation, the total efficiency of a scaled fan is assumed to remain constant, such that

[[eta].sub.TC] = [[eta].sub.T]. (16)

In this study, all volumetric flow rate and total pressure rise data were plotted by correcting back to conditions of standard air density (1.2 kg/[m.sup.3] or 0.075 [lb.sub.m]/[ft.sup.3]) and fan speed obtained at free delivery (for the ideal installation) by means of Equations 14 and 15. The ratio of the compressibility factor at standard and actual conditions was calculated per AMCA (1999). As a means of establishing whether any anomalies in the data existed, the dimensionless total pressure based upon impeller tip speed was likewise plotted as a function of the dimensionless fan airflow rate based upon tip speed, thereby yielding graphs of

[[P.sub.T]/[[1/2][rho][V.sub.[TIP].sup.2]]] versus [[Q.sub.2]/[[V.sub.TIP.sup.A]FAN].

For brevity these data are not shown.

AERODYNAMIC PERFORMANCE CURVES

A representative set of aerodynamic performance data obtained for the 610 mm (24 in.) and 914 mm (36 in.) diameter fans at 100% and 70% of the speed are presented in Figures 2 through 9. It is noted that at free delivery, the 100% nominal fan speed was 908 and 785 rpm, respectively. All other data are available in Darvennes et al. (2009b). Likewise, for brevity, the aerodynamic performance data for the 1219 mm (48 in.) fan at 70% and 47.5% of the speed are not presented in the present paper but are provided in Darvennes et al. (2009b). For the 1219 mm (48 in.) fan, the nominal speed at free delivery was 546 rpm. Due to insufficient air flow that could be admitted into the reverberant chamber, free delivery was only attainable at up to 47.5% of nominal fan speed, with the doors to the reverberant chamber open. Thus, the operational settings for the 1219 mm (48 in.) diameter fan were chosen to be 70% and 47.5% of the nominal fan speed (these values having the same ratio as for the 610 mm (24 in.) and 914 mm (36 in.) fans. Due to the limitations of the apparatus, the 1219 mm (48 in.) fan at 70% of nominal fan speed could not attain free delivery. In every instance, aerodynamic data reduction was performed per AMCA (1999), as outlined above. In these graphs, corrected fan total pressure rise is plotted at both fan speeds as a function of corrected volumetric flow rate. Likewise, Figures 2 through 9 present the measured efficiency as a function of corrected volumetric flow rate, at both fan speeds. Figures 2 through 9 are dimensionally correct for SI units only.

[FIGURE 2 OMITTED]

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[FIGURE 5 OMITTED]

[FIGURE 6 OMITTED]

[FIGURE 7 OMITTED]

[FIGURE 8 OMITTED]

[FIGURE 9 OMITTED]

All appurtenances were mounted at the entrance plane of the fans. In the tests performed with the 90[degrees] mitered elbow, the orientation was varied by 45[degrees] increments, such that 0[degrees] refers to the case where the elbow was mounted vertically. Separation distances for the wall were investigated until the wall had relatively little impact on the aerodynamic performance of the fans.

Figure 2 presents the baseline aerodynamic performance data obtained on the D = 610 mm (24 in.) fan, wherein there were no inlet fittings or other disturbances at the fan entrance. Figure 3 displays aerodynamic performance data that were measured with a contraction having an area ratio of 1.0. Figure 4 provides the measured aerodynamic performance data for a mitered elbow mounted vertically. Figure 5 depicts aerodynamic performance data obtained when a wall perpendicular to the fan axis was positioned at 0.25D. Figure 6 presents the baseline aerodynamic performance data obtained on the D = 762 mm (36 in.) fan, wherein there were no inlet fittings or other disturbances at the fan entrance. Figure 7 displays aerodynamic performance data that were measured with a contraction having an area ratio of 1.0. Figure 8 provides measured aerodynamic performance data for a vertical mitered elbow. Figure 9 depicts aerodynamic performance data obtained when a wall perpendicular to the fan axis was positioned at 0.25D.

SYSTEM EFFECT FACTORS

The concept of the system effect factor (SEF) has been described to account for the detrimental effect of such system connections on the performance of the fan. In the present study, air and sound performance of propeller fans was measured with systematic variation of inlet flow components so as to simulate typical "in the field" installations of the fans, and SEF was evaluated as herein described. The situation where fan performance was tested with an open inlet, so that airflow entering the fan was approximately uniform, was referred to as the baseline case. This yielded the ideal aerodynamic performance curve, plotted in terms of total pressure rise across the fan as a function of corrected volumetric flow rate. Therein, a second-order curve was fit to the data by means of the least-squares method, spanning the range from peak efficiency to free delivery, such that

[P.sub.t,ideal] = [A.sub.1] + [B.sub.1][Q.sub.c] + [C.sub.1][Q.sub.c.sup.2]. (17)

The coefficients [A.sub.1], [B.sub.1], and [C.sub.1] are the curve-fit parameters, and [Q.sub.c] is interpreted as the base-line/design flow rate. Similarly, performance data were obtained for the case where duct fittings/walls were mounted at the fan inlet, thereby causing a performance penalty (i.e., a reduction in total pressure rise at a given volumetric flow rate due to the resulting system effect). A second-order least-squares curve was fit to the data over the range from peak efficiency to free delivery, thereby yielding the curve-fit parameters [A.sub.2], [B.sub.2], and [C.sub.2]. Hence,

[P.sub.t,SEF] = [A.sub.2] + [B.sub.2]Q + [C.sub.2][Q.sup.2]. (18)

The difference in performance of the ideal baseline case and the case with duct fittings/walls mounted at the fan inlet represents the influence of the inlet appurtenance. The system effect is interpreted as that portion of the performance penalty due to the distortion of the fan inlet velocity profile caused by the presence of the appurtenance. In those instances where an elbow or contraction is mounted at the fan inlet plane, there is an additional performance penalty attributed to the total pressure loss through the fitting, herein designated as [[DELTA][P.sub.t]. Consistent with the data reduction approach employed by Clarke et al. (1978) and Zaleski (1988), the pressure loss of the elbow/contraction was subtracted from the total fitting performance penalty, thereby yielding the magnitude of the SEF. Hence, in this study, the fan SEF was defined in terms of the ideal flow rate, such that

SEF = [P.sub.t,ideal] - [P.sub.t,SEF] - [Delta][P.sub.t] or (19)

SEF = ([A.sub.1] - [A.sub.2]) + ([B.sub.1] - [B.sub.2])[Q.sub.c] + ([C.sub.1] - [C.sub.2])[Q.sub.c.sup.2] - [Delta][P.sub.t]. (20)

Loss coefficient data for ducted fittings taken from the ASHRAE Duct Fitting Database (2006) were used to calculate the total pressure loss across the elbows and contractions mounted at the inlet plane of the fan. The loss coefficient is defined as the ratio of the total pressure loss of a nonjunction fitting to that of velocity pressure, and is given by

C = [[[Delta][p.sub.t]]/[p.sub.v]]. (21)

The velocity pressure depends on the measured average velocity based on the downstream cross section of the fitting, which in this project was matched closely to that of the fan. Duct sections to provide for flow development upstream and downstream of the fitting were not present in the elbows and contractions tested in this project. For these particular entry fittings, reliable loss coefficient data were unavailable in the literature. Instead, the fittings were considered to be ducted, such that the total pressure is measured at prescribed locations in straight duct sections mounted upstream and downstream of the fitting, per ASHRAE (1999). Therein, the zero-length loss coefficient is determined by subtracting the duct tare pressure loss per unit length from the total pressure loss, thus giving a measure of the dynamic loss that occurs within a fitting.

The ASHRAE Duct Fitting Database (2006) indicated the loss coefficient for 90[degrees] round mitered elbows varying in diameter from 762 to 1372 mm (30 to 54 in.), ranging from 1.14 to 1.12, respectively. An average value of 1.13 was employed for all elbows tested in the project. For round contractions with a 90[degrees] included angle and an area ratio of 1.5, ASHRAE (2006) provided an average loss coefficient value of 0.24. Likewise, ASHRAE (2006) yielded an average loss coefficient of 0.12 for a round contraction with a 90[degrees] included angle and an area ratio of 1.25. By convention, the loss coefficient of a round contraction with an area ratio of unity was 0 (i.e., the dynamic pressure loss through such a fitting was neglected in this study).

It is customary to define the dimensionless loss coefficient as the ratio of SEF to the velocity pressure at the fan outlet, such that SEF is proportional to the square of the air velocity, as seen in Equation 22:

C = [[SEF]/[p.sub.v]] = [[SEF]/[[1/2][rho][([Q.sub.C]/A).sup.2]]] (22)

The slope of the SEF plotted against [p.sub.v] can be interpreted as the zero-length loss coefficient of the fitting, and a curve plotted through the data points will be a straight line if the loss coefficient is a constant. Likewise, Figure 10 (excerpted from AMCA [2002]) suggests that, in general, SEF is proportional to the square of the fan air velocity, that is,

SEF = a[([Q.sub.c]/A).sup.2], (23)

where a is an experimentally determined proportionality constant. If such a functional relationship is valid, SEF data plotted versus air velocity will be described by a single SEF class (e.g., curves F--X in Figure 10). Note that for positive SEF values, the constant a is readily related to the loss coefficient as follows:

a = [1/2][rho]C (24)

Table 1 lists the constants used to define the SEF class curves in Figure 10. The least squares method was employed to obtain an overall loss coefficient for each inlet appurtenance. For example, Equation 23 can be written as the following:

Table 1. System Effect Curves

SEF Curves in  Dynamic Pressure
Figure 10      Loss Coefficient C

F                      16.00
G                      14.20
H                      12.70
I                      11.40
J                       9.50
K                       7.90
L                       6.40
M                       4.50
N                       3.20
O                       2.50
P                       1.90
Q                       1.50
R                       1.20
S                       0.75
T                       0.50
U                       0.40
V                       0.25
w                       0.17
X                       0.10

SEF = C*([P.sub.v] (25)

Then, assuming a straight line can be fit through the data,

[y.sub.i] = M[x.sub.i] + B, (26)

where

[y.sub.i] = SEF, [x.sub.i] = [p.sub.v] (27)

M = C, (28)

The intercept B can be forced to 0 in the case of the fan appurtenances, since ideally the pressure loss associated with swirl at the fan inlet should be 0 when the air velocity at the fan entrance is 0. With the intercept forced to 0, it can readily be shown that

C = [[[n.summation over (i = 1)][x.sub.i][y.sub.i]]/[[n.summation over (i = 1)][x.sub.i.sup.2]]]. (29)

In this project the loss coefficient was obtained by fitting a straight line to the SEF versus velocity pressure, per Equation 25. Therein, the slope of the least squares curve yields the loss coefficient. The error in the slope of a least-squares curve fit is given by Beckwith et al. (1993) as the following:

[Delta][alpha] = [+ or -][t.sub.[a/2,n - 2]][[S.sub.yx]/[S.sub.xx]] (30)

where [t.sub.a/2,n - 2] is the t statistic with n-2 degrees of freedom (n is the number of points in the data set) and a = (1 - c) is the level of significance (in the present work c = 0.95). The value [S.sub.yx] is calculated using

[S.sub.yx] = [[(1/[n - 2][n.summation over (i = 1)][[[y.sub.i] - y([x.sub.i])].sup.2])].sup.[1/2]], (31)

where [y.sub.i] is the actual value at point i, and y ([x.sub.i]) is the value obtained by the least squares fit at point i. The quantity [S.sub.xx] in Equation 30 is found from the expression

[S.sub.xx.sup.2] = [n.summation over (i = 1)][[([x.sub.i] - [bar.x])].sup.2], (32)

where [bar.x] is the mean x value.

In this project, SEF values for each combination of test fan and inlet appurtenance were calculated by means of Equation 20. Therein, SEF values for each geometrically similar inlet condition were plotted as a function of the velocity pressure assessed at the inlet plane of the fan. The resulting data are shown in Figures 11 through 21. In these graphs no distinction is made regarding the fan impeller diameter (i.e., the figures exhibit SEF data obtained for the 610, 914, and 1219 mm [24, 36, and 48 in.] diameter fans, inclusively). Similarly, these graphs include data from the entire range of fan speeds tested. A least squares curve, as obtained from Equation 29, is displayed for each inlet appurtenance. Figures 11 through 21 are dimensionally correct for SI units only.

[FIGURE 11 OMITTED]

[FIGURE 12 OMITTED]

[FIGURE 13 OMITTED]

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[FIGURE 20 OMITTED]

[FIGURE 21 OMITTED]

Figures 11 through 14 depict the measured SEF values obtained when a wall perpendicular to the fan axis was positioned at 0.25, 0.5, 0.75, and 1D, respectively. The SEF data (as quantified in terms of loss coefficient) were significant for a separation distance of 0.25D. However, it is readily apparent that SEF values were much smaller for separation distances of 0.5D or greater, as the average loss coefficients for gaps [greater than or equal to] 0.5D were generally less than 10% of that for 0.25D. Figures 15 through 17 present SEF data that were measured when contractions having area ratios of 1.0, 1.25, and 1.5, respectively, were mounted at the entrance plane of the fan. There were positive SEF values for an area ratio of 1.0. By contrast, SEF values were negative for an area ratio of 1.5, implying that flow entering the fan was apparently aided by the presence of such a contraction. Therein, the measured SEF for an intermediate contraction ratio of 1.25 was correspondingly reduced. Figures 18 through 21 portray the measured SEF data that were acquired when a 90[degrees] mitered elbow was mounted at the fan entrance plane, and the orientation of the elbow was varied in 45[degrees] increments. Recall the indication 0[degrees] refers to the case where the elbow was mounted vertically. Moreover, the motors of these belt-driven fans were offset from the fan axis by approximately 254 mm (10 in.) and were mounted in the 180[degrees] position. For each fan tested, the motor blocked 2.1% to 8.5% of the fan inlet cross section and offered some impediment to the entering flow. The SEF values were observed to be a maximum at an elbow position of 0[degrees], and diminished as this angle was increased. Thereupon, the SEF values for the 135[degrees] case were negative, once more indicating that flow entering the fan was somewhat assisted by the presence of the elbow in that particular orientation.

The SEF data plotted in Figures 11 through 21 collectively exhibit considerable scatter about the least squares curve, as demonstrated by the uncertainty level that ranged from 0.038 to 0.512, located in Table 2. In an effort to determine whether a particular fan or fan speed contributed most significantly to the data scatter, selected SEF values were replotted in terms of SEF classes, per Figure 10. The resulting graphs are shown in Figures 22 through 24, which are dimensionally correct for SI units only. Figure 22 displays SEF data for the 0.25D wall case, and clearly distinguishes between the 610, 914, and 1219 mm (24, 36, and 48 in.) fans operated at the various speeds. Likewise, Figure 22 portrays the SEF values for the AR = 1.0 inlet contraction, whereas Figure 23 depicts SEF data for the inlet elbow oriented at

0[degrees] (i.e., opposite the location of the fan motor). It is noted that in Figure 23, several negative SEF values were not plotted. Similarly, the other inlet appurtenance cases considered in this project were not graphed in terms of SEF class, because of the prevalence of negative SEF data.

Table 2. Suggested SEF Classes

Inlet Appurtenance  Loss Coefficient  Uncertainty Level  SEF Class
                                                           Range

Wall 0.25D                     0.684      [+ or -]0.127    S--T


Wall 0.5D                      0.066      [+ or -]0.119  [greater
                                                          than or
                                                         equal to]
                                                             X

Wall 0.75D                     0.040      [+ or -]0.121  [greater
                                                          than or
                                                         equal to]
                                                             X

Wall 1D                        0.092      [+ or -]0.138  [greater
                                                          than or
                                                         equal to]
                                                             X

Contraction                    0.353      [+ or -]0.051    U--V
AR = 1

Contraction                    0.065      [+ or -]0.038  [greater
AR = 1.25                                                 than or
                                                         equal to]
                                                             X

Contraction                   -0.130      [+ or -]0.054  [greater
AR = 1.5                                                  than or
                                                         equal to]
                                                             X

Elbow 0[degrees]               0.563      [+ or -]0.286    S--T

Elbow 45[degrees]              0.301      [+ or -]0.305    U--V

Elbow 90[degrees]              0.171      [+ or -]0.214    V--W

Elbow 135[degrees]            -0.238      [+ or -]0.512  [greater
                                                          than or
                                                         equal to]
                                                             X

Table 2 presents the SEF and uncertainty level of the data portrayed in Figures 11 through 21 in terms of a range of suggested SEF classes. The SEF classes are based on the measured loss coefficients; refer to Table 1 for the relation between the SEF class and the loss coefficient. In those instances where the loss coefficient data were negative, the SEF class is therein reported as [greater than or equal to] X. This terminology implies either that the system effects associated with that appurtenance were negligible, or had a beneficial impact on the resulting pressure loss penalty.

CONCLUSION

In this study, air and sound performance was measured experimentally for three propeller fans having nominal impeller diameters of 610, 914, and 1219 mm (24, 36, and 48 in.) that were subjected to systematic variation of inlet flow components. The inlet conditions were intended to simulate installations of fans typically encountered in the field. The aerodynamic performance penalties associated with the various appurtenances were characterized in terms of SEF. The data reduction approach first required measurement of aerodynamic performance of a baseline case (with no inlet appurtenance). Thereafter, fan aerodynamic performance was measured when various appurtenances were situated at the inlet plane of the fans. These appurtenances consisted of 90[degrees] mitered elbows mounted at various angles, inlet duct contractions of various area ratios, or solid walls perpendicular to the fan axis located at various distances from the inlet. The ensuing data reduction was consistent with that of Clarke et al. (1978) and Zaleski (1988), whose data comprise the majority of the SEF information presented in AMCA (2002). The SEF data obtained in the present study displayed considerable scatter. Therein, a statistical study was performed to estimate the expected uncertainty. Under some circumstances, the uncertainty in SEF was approximately the same magnitude as the measured value, or even surpassed that value.

The aerodynamic effects of the wall were quite dependent on distance from the entrance plane of the fan. For the fans tested in this program, when the distance from the wall to the fan entrance exceeded 0.25D, the SEF values were considerably reduced in magnitude. Expressed in terms of SEF, the aerodynamic influence of the contractions for the fans decreased as the area ratio increased from 1.0 to 1.5. At the latter area ratio, the measured SEF values implied the presence of the contraction aided the flow slightly. The orientation of the elbow relative to the position of the fan motor influenced the aerodynamic performance. Maximum SEF values for all fan sizes were obtained when the inlet of the elbow was oriented at 0[degrees]. For the belt-driven fans considered in this study, the fan motor was offset from the fan axis and was oriented in the 180[degrees] position. Presumably, the presence of the fan motor impeded the airflow entering the fan. The SEF values progressively decreased as the position of the elbow inlet approached 180[degrees]. The latter test condition could not be achieved in this test program because of limitations on the size of the test facility.

In general, when the Reynolds number is sufficiently large the flow is dominated by inertial effects, and the influence of viscosity is relatively less important. In those instances, the SEF values should scale directly with the velocity pressure. This would further imply the corresponding loss coefficients would be a function of geometry only. Although Figures 11 through 21 were plotted without regard to the particular fan size or impeller speed, in some cases it was apparent that SEF data tended to plot in fairly distinct bands, where each data group corresponded to a certain combination of fan size and speed. This was further evident in Figures 22 through 24 where the particular test conditions were clearly distinguished. In particular for the SEF values sketched in Figures 11 through 14 pertaining to the influence of a wall adjacent to the inlet of a fan, much of the scatter evident in the graphs for separation distances [greater than or equal to] 0.5D was attributable to data obtained on the 1219 mm (48 in.) fan at fan speeds corresponding to 70% and 47.5% of the nominal value. Similarly, for Figures 18 through 21, which examined the influence of elbow orientation on SEF, there were also reasonably distinct bands associated with each specific fan size and speed. This implied the SEF data related to solid walls and mitered elbows situated at the fan inlet plane were not entirely independent of impeller diameter or fan speed (i.e., to some extent there was some Reynolds number dependency). In contrast, Figures 15 through 17, which examined the effects of contractions mounted at the fan inlet, exhibited no such banding of the SEF data (i.e., the loss coefficient values obtained therein were functions solely of the contraction geometry). In those instances each condition of fan size and speed yielded SEF values that were apparently randomly scattered about a central tendency.

It is important to note that SEF values reported in this study for elbows or contractions mounted at the entrance plane of the fan had the total pressure loss of the fitting [[DELTA][P.sub.t] subtracted from the performance penalty, per Equation 20. That pressure loss was difficult to ascertain exactly, and was therefore estimated as described previously. It was not possible to accurately measure fitting loss coefficients in situ in the experimental apparatus employed in this study. Typically, [[DELTA][P.sub.t] comprised anywhere from 50% to more than 175% of the calculated SEF pressure loss/gain (i.e., it was the dominant term in Equation 20). Hence, uncertainty in the elbow or contraction fitting pressure loss coefficient contributed directly to uncertainty in the SEF loss coefficients reported herein. In some instances, the pressure loss of an inlet fitting can be significant, and should be carefully taken into account by system designers when attempting to predict installed fan performance.

In this study, the tip velocities for the three fans ranged from 24.4 m/s (4803 ft/min) for the 1219 mm (48 in.) diameter fan operating at 70% of nominal speed to 37.6 m/s (7402 ft/min) for the 914 mm (36 in.) diameter fan operating at 100% speed. It is suggested that further investigations of the influence of inlet appurtenances on aerodynamic and acoustic performance be conducted at fixed values of the dimensionless tip speed (Q/A)/[V.sub.tip]. This could readily be achieved by selecting several sizes among geometrically similar fans, and adjusting their speeds accordingly. Therein, it would be necessary to forego the strict adherence to the 70% and 100% fan speed protocol employed in the present study for the 610 and 914 mm (24 and 36 in.) fans.

ACKNOWLEDGMENTS

The work reported in this paper is the result of cooperative research (RP-1223) between ASHRAE and Tennessee Tech University. The project was sponsored by TC 5.1, Fans. The technical assistance provided by David Carroll, John Cermak, and John Murphy is gratefully acknowledged.

NOMENCLATURE

[A.sub.FAN] = cross-sectional area of fan

[A.sub.N] = cross-sectional area of nozzle(s)

[C.sub.N] = nozzle discharge coefficient(s)

D = fan impeller diameter

H = fan shaft power input

[H.sub.O] = fan power output

[K.sub.P] = compressibility coefficient

m = mass flow rate

N = fan rotational speed

P =pressure

Q = volumetric flow rate

T = temperature of air

V = velocity of air

[V.sub.TIP] = impeller tip velocity

w = mechanical energy added to the air

W = measured power input of fan motor

[Y.sub.N] = nozzle expansion coefficient(s)

Greek Symbols

[eta] = calibrated fan motor efficiency

[[eta].sub.T] = total fan efficiency

[rho] =density of air

Subscripts

1 = static pressure upstream of fan

2 = static pressure downstream of fan

5 = section upstream of nozzle

6 = nozzle throat

C = converted to standard test conditions

DO =dry bulb

S7 = measured static pressure in nozzle chamber

T = total pressure rise across fan

T1 = total pressure upstream of fan

T2 = total pressure downstream of fan

V2 = velocity pressure downstream of fan

REFERENCES

AMCA. 1999. ANSI/AMCA Standard 210-99, Laboratory Methods of Testing Fans for Aerodynamic Performance Rating. Arlington Heights, IL: Air Movement and Control Association.

AMCA. 2002. AMCA Publication 201-2002, Fans and Systems. Arlington Heights, IL: Air Movement and Control Association.

ASHRAE. 2001. ASHRAE Standard 120-99, Method of Testing to Determine Flow Resistance of HVAC Ducts and Fittings. Atlanta: American Society of Heating, Refrigerating and Air-Conditioning Engineers, Inc.

ASHRAE. 2006. ASHRAE Duct Fitting Database, Version 4.0.3. Atlanta: American Society of Heating, Refrigerating and Air-Conditioning Engineers, Inc.

Beckwith, T.J., R.D. Marangoni, and J.H. Lienhard V. 1993. Mechcanical Measurements, 5th ed. New York: Addison-Wesley Publishing Company, Inc.

Clarke, M.S., J.T. Barnhart, F.J. Bussey, and E. Neitzel. 1978. The effects of system connections on fan performance. ASHRAE Transactions 82(2):227-63.

Darvennes, C., M.N. Young, and S. Idem. 2009a. Acoustic system effects of propeller fans due to inlet installations. ASHRAE Transactions 115(2).

Darvennes, C., S. Idem, and M.N. Young. 2009b. Inlet installation effects on propeller fans, air and sound. ASHRAE RP-1223, American Society of Heating, Refrigerating and Air-Conditioning Engineers, Inc., Atlanta.

Young, M.N., C. Darvennes, and S. Idem. 2009. Test apparatus and procedure to measure inlet installation effects of propeller fans. ASHRAE Transactions 115(2).

Zaleski, R.H. 1988. System effect factors for axial flow fans. Air Movement and Control Association Engineering Conference, Phoenix, AZ.

Received March 5, 2008; accepted September 3, 2008

This paper is based on findings resulting from ASHRAE Research Project RP-1223.

M.N. Young, PhD

C. Darvennes, PhD

S. Idem, PhD

Member ASHRAE

M.N. Young is an engineer with the Tennessee Valley Authority, Knoxville, TN. C. Darvennes and S. Idem are professors in the Department of Mechanical Engineering at Tennessee Tech University, Cookeville, TN.