JOINT INSTITUTE &IERO

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National Aeronautics and 

Space Administration 

Ames Research Center JIAA TR - 89 Stanford university 

DESIGN AND CALIBRATION OF THE 

MIXING LAYER AND WIND TUNNEL 

BY 

James H. Bell and Rabindra D. Mehta 

Stanford University 

Department of Aeronautics and Astronautics 

Stanford, CA 94305 

MAY 1989 

-- (IASA-CR-185472) DESXICI AND CALIB8ATI013 OF 1Y89 -268 66 TBB MIXIW LAYER AND UIblD TUHIEL 

Unir.) 36 p CSCL 01E 

[Staaford 

Unclas 

G3/09 0222706 
#- I 

ABSTRACT 

This report contains a detailed account of the design, assembly and calibration of a 

wind tunnel specifically designed for free-shear layer research. The construction of this 

new facility was motivated by a strong interest in the study of plane mixing layers with 

varying initial and operating conditions. The Mizing Luyer Wind Tunnel is located in the 

Fluid Mechanics Laboratory at NASA Ames Research Center. The tunnel consists of two 

separate legs which are driven independently by centrifugal blowers connected to variable 

speed motors. The blower/motor combinations are sized such that one is smaller than 

the other, giving maximum flow speeds of about 20 and 40 m/s, respectively. The blower 

speeds can either be set manually or via the Microvax I1 computer. The two streams are 

allowed to merge in the test section at the sharp trailing edge of a slowly tapering splitter 

plate. The test section is 36 cm in the cross-stream direction, 91 cm in the spanwise 

direction and 366 cm in length. One test section side-wall is slotted for probe access and 

adjustable so that the streamwise pressure gradient may be controlled. The wind tunnel 

is also equipped with a computer controlled, three-dimensional traversing system which 

is used to investigate the flow fields with pressure and hot-wire instrumentation. The 

wind tunnel calibration results show that the mean flow in the test section is uniform to 

within &0.25% and the flow angularity is less than 0.25'. The total streamwise free-stream 

turbulence intensity level is approximately 0.15%. Currently the wind tunnel is being used 

in experiments designed to study the three-dimensional structure of plane mixing layers 

and wakes. 

1 
NOMENCLATURE 

AR: 

CFM: 

C: 

cf: 

Cp: 

d: 

D: 

H: 

HP: 

He : 

Hi : 

K: 

L: 

I: 

P: 

Ap: 

P: 

4: 

RPM: 

ReL: 

u, v, w: 

ut, ut, wt: 

Ue : 

u, 21, w: 

x, Y, 2: 

P: 

699: 

e: 

U: 

P: 

-. 

( )ma=: 

Test section aspect ratio 

Contraction ratio 

Volume flow rate in cubic feet per minute 

Skin friction coefficient 

Pressure coefficient on contraction wall 

Screen wire diameter 

Honeycomb cell diameter 

Boundary layer shape factor 

Motor horse power 

Contraction height at exit 

Contraction height at inlet 

Screen pressure drop coefficient 

Contraction length or honeycomb cell length 

Screen mesh length 

Static pressure 

Static pressure drop or rise across element 

Total pressure 

Dynamic pressure 

Blower revolutions per minute 

Reynolds number, UL/v 

Mean velocity in the X,Y,Z directions, respectively 

Fkee-stream velocity in the wind tunnel test section 

Fluctuating velocity components in the X,Y,Z 

directions, respectively 

Instantaneous velocity in the X,Y,Z directions, 

respectively, e.g. u = U + ut 

Cartesian coordinates for streamwise, normal, 

and spanwise directions, respectively 

Screen open-area ratio 

Boundary layer thickness 

Boundary layer momentum thickness 

Kinematic viscosity 

Density 

(overbar) Time-averaged quantity 

Maximum value at station 

2 
1. INTRODUCTION 

Wind tunnels of various forms and designs have been widely used for model tests and 

basic fluid flow research since about the turn of the century. Since the 1930s, when the 

strong effect of free-stream turbulence on shear layers became apparent, emphasis has been 

laid on wind tunnels with low levels of turbulence and unsteadiness (Prandtl, 1933). A 

wind tunnel is still an essential tool in engineering, even with today's improved computing 

capabilities. Wind tunnel design always has been, and will be, a combination of art, science 

and commonsense. Traditionally, it has been difficult and unwise to make firm design 

rules for all the components of a wind tunnel, mainly due to the wide variety of tunnel 

designs and the lack of understanding of the flow, especially the turbulent boundary layers 

encountered in many of the tunnel components. Therefore, a more feasible and sensible 

approach was followed whereby approximate design boundaries were predicted based on 

exisiting "successful" designs. Such design guidelines for low-speed wind tunnels were first 

given by Bradshaw and Pankhurst (1964). 

In the late 1970s, a project was initiated by one of the present authors, designed to 

gain a better understanding of the performance of "blower" tunnel components (Mehta, 

1978). A blower tunnel is defined as an open-circuit wind tunnel driven by a centrifugal 

blower at the entry. The main aim of the study was to make the design of these components 

more scientific, logical and hence successful. 

Design rules for wide-angle diffusers were derived, based mainly on correlations of data 

collected from existing designs (Mehta, 1977). A wide-angle diffuser is defined as one which 

expands so rapidly that separation can only be avoided by some form of boundary layer 

control. Screens are by far the most popular means of preventing separation in wide-angle 

diffusers. This is in addition to the more accepted role of screens in improving the flow 

uniformity and reducing the turbulence intensity levels (Dryden and Schubauer, 1947). 

A detailed experimental investigation was carried out on the effects of different types of 

screens (metal and plastic) on turbulent flow, in particular turbulent boundary layers 

(Mehta, 1984 and 1985). The boundary layer turbulence was found to be reorganized 

and the thickness reduced, thus making the layer less susceptible to separation. The 

effects of screen inclination on pressure drop coefficient and deflection coefficient were also 

established. Tests on single-inlet centrifugal blowers revealed that the outlet flow was 

in the form of a large longitudinal swirl which may be used to advantage in preventing 

separation on the wide-angle diffuser walls (Mehta, 1977). Design guidelines based on the 

results of this research project were given in Mehta and Bradshaw (1979). 

In wind tunnels designed for shear layer research, the two-dimensionality of the flow (or 

the lack of it) is also of considerable concern (Morkovin, 1979). In a separate experimental 

study, the effect of screens, honeycombs and centrifugal blowers on the two-dimensionality 

of a boundary layer on the test section floor was investigated (Mehta and Hoffmann, 1986 

and 1987). Surveys of the spanwise variation in surface shear stress in three blower tunnels 

revealed that the main component responsible for altering the spanwise properties of the 

test section boundary layer was the last screen, thus confirming previous findings (e.g. 

Bradshaw, 1965). It was further confirmed that a screen with varying open-area ratio 

(due to dirt accretion, for example), produced an unstable flow. However, contrary to 

3 
popular belief, it was also found that for given incoming conditions and a screen free of 

imperfections, the screen open-area ratio done was not enough to describe its performance. 

More recently, an iterative design procedure was developed for two- or threedimensional contractions installed on small, low-speed wind tunnels (Bell and Mehta, 1988, 

1989a). The procedure consisted of first computing the potential flow field, and hence the 

pressure distributions, along the walls of a contraction of given size and shape using a 

three-dimensional numerical panel method. The pressure (or velocity) distributions were 

then fed into two-dimensional boundary layer codes to predict the behavior of the boundary 

layers along the walls. For small, low-speed contractions it was shown that the boundary layer behavior could be adequately predicted by assuming that a laminar boundary 

layer originated from stagnation conditions at the contraction entry and remained laminar 

throughout passage through the "successful" contraction designs. 

The design of the new wind tunnel was based on the results and analyses of all these 

investigations. The unique aspect of the tunriel% its ability to easily generate mixing 

layers with varying initial and operating conditions. Turbulent mixing layers play many 

important roles in practical aerodynamics. Mixing layers govern the rate of mixing in 

combustion chambers and flow reactors, for example, and are also responsible for most of 

the broadband noise generated in propulsion systems. Furthermore, the flow over modern 

aircraft wings with multiple airfoils is dominated by complex interactions involving freeshear flows, including mixing layers. The ability to control the structure and development, 

and hence mixing capabilities, of this shear flow would therefore have a vital impact on 

many engineering applications. 

Section 2 of this report contains a detailed description of the individual components of 

the wind tunnel. The wind tunnel performance characteristics are presented in Section 3 

while the flow quality measurements are given and discussed in Section 4. Some concluding 

remarks, including a brief description of ongoing research work in the new facility, are given 

in Section 5. 

2. DESCRIPTION OF THE WIND TUNNEL COMPONENTS 

The overall schematic of the Mixing Layer Wind Tunnel is shown in Fig. 1. Air is 

drawn-in by two separate blowers, one larger than the other, and passed through separate 

diffusers, flow conditioning elements and two separate but identical contraction sections 

into the test section. The selection or design of the individual components is now discussed 

starting from the upstream (intake) end. 

2.1 Drive Systems 

Over the past couple of decades, centrifugal blowers with backward-facing "airfoiltype" impellers have emerged as the optimum choice for driving open-circuit wind tunnels. 

Centrifugal blowers run with reasonable steadiness and efficiency over a wide range of flow 

conditions with varying tunnel power factor (defined as the ratio of shaft input power to 

tunnel output power) because the whole blade span operates at nominally the same lift 

4 
coefficient. The noise and pulsations generated by a centrifugal blower are generally found 

to be adequately low, even at off-design conditions. In blowers with single inlets, such 

as those used on the present tunnel, the impeller is asymmetrically positioned so as to 

accommodate the entry bellmouth within the blower casing. This asymmetry produces a 

swirling flow at the blower outlet which is found to be independent of the advance ratio; 

it is dependent on the ratio of rotor to casing width. The vortex-type outlet flow can 

be used to advantage to maintain wall flow attachment in the wide-angle diffuser. This 

compensates for the non-uniformity of the flow which is later improved by the screens in 

the wide-angle diffuser and the settling chamber. 

The "low-speed" leg of the present wind tunnel is driven by a size 223 Acoustafoil 

blower (New York Blower Company) - a 5,000 CFM centrifugal blower with a 56 cm 

impeller consisting of backward-facing, airfoil-type blades. The outlet dimensions of this 

blower are 43 x 64 cm and it is powered by a 5 HP electric motor (Eaton Corporation, 

Dynamatic Ajusto Spede, Model AC-186). The motor speed, and thus the flow rate, is set 

using an Eaton Dynamic Model 4000 Ajusto Spede motor contoller. A feedback circuit 

controlling the eddy-current clutch maintains the motor at a constant speed; mean flow 

velocity through the test section varies by less than 1% over the time of a typical run lasting 

about two hours. This blower/motor combination was retrived from the now discarded 

Shear Layer Wind Tunnel. Aside from considerations of cost, it was felt that the smaller 

blower would show better stability at low tunnel speeds, and so it would be advantageous 

to have a dedicated low-speed side. 

The high-speed side blower is a size 339 SWSI Class I1 Acoustafoil Centrifugal Fan 

(New York Blower Company). This blower delivers 15,000 CFM with a 86 cm impeller also 

consisting of backward-facing, airfoil-type blades. The blower has outlet dimensions of 61 x 

94 cm and is powered by a 20 HP electric motor (Eaton Corporation, Dynamatic Ajusto 

Spede Model ACM-280). As with the smaller motor, speed control is achieved by adjusting 

the eddy-current clutch through a feedback circuit. The flow rate of both blowers can be 

controlled either manually or through a D/A output signal from a Microvax I1 computer. 

The motors are physically isolated from the blowers so as to reduce vibration as much as 

possible, and the only physical connection between the blowers and the rest of the wind 

tunnel are duct tape couplings sealing the blower/diffuser junctions. The contributions 

to free-stream turbulence due to motor/blower vibration, blade passing frequency, and 

irregularities in the blower have been measured and found to be acceptably small, as 

discussed below in Section 4.3. 

A 122 x 122 x 122 cm filter box covers the inlet of the high-speed side blower, and a 

91 x 91 x 91 cm filter box covers the intake section of the low-speed side blower. The filter 

material consists of two layers of cheesecloth which has been found to effectively remove 

most of the dust in the incoming air while imposing only a negligible pressure drop; with 

the current set of flow-conditioning elements, the small and large blowers can generate 

flow speeds in the test section of up to 20 m/s and 40 m/s, respectively. The filters are 

considered to be effective since hot-wire drift, attributable to dirt accretion, has been found 

to be negligibly small. 

5 
2.2 Wide- Angle Diffusers 

Although diffusers form an itegral part of almost all wind tunnels, their flow characteristics have still to be fully understood. The flow through a diffuser inevitably depends 

on its geometry, defined by the area ratio, wall expansion angle, cross-sectional shape and 

wall contour. Other parameters such as geometric and flow conditions at entry and exit 

also dect the diffuser performance. With an arbitrary combination of these parameters, 

the flow through a diffuser becomes extremely difficult to predict in detail. The issue is 

further complicated by the occasional presence of boundary layer separation caused by the 

adverse pressure gradients necessarily present in a diffuser. It was in fact a separation 

in a conical diffuser that inspired Prandtl to come up with the boundary layer concept 

in 1904. A “wide-angle” diffuser is defined as a difiser in which the cross-sectional area 

increases so rapidly that separation can be avoided only by using boundary layer control. 

A wide-angle diffuser is used to reduce the length needed for a given area ratio rather than 

effecting a pressure recovery, and in blower tunnels such as the present one, it is installed 

between the blower and settling chamber. 

Screens are the most popular means of boundary layer control in diffusers. A screen, 

besides removing the direct effects of boundary layer growth and incipient separation, effectively gives the layer “a new lease of life” - the boundary layer thickness is reduced 

and the existing turbulence almost completely obliterated, thus forming a “new” boundary 

layer (Mehta, 1984 and 1985). In addition, screens also help to improve the flow nonuniformity and reduce the flow angularity - this is particularly useful for straightening the 

swirling type flow produced by a single-inlet centrifugal blower. For a diffuser using screens 

for boundary layer control, the important parameters are found to be the diffuser angle, 

area ratio, number of screens and the overall pressure drop coefficient of the screens. Mehta 

(1977 and 1978) collected data from over a hundred wide-angle diffuser designs (successful and unsuccessful) and plotted charts for relevant parameters from which design rules 

were derived. The number of screens required within the diffuser is determined from the 

diffuser angle and the area ratio. The area ratio also gives the overall required pressure 

drop coefficient from the screens in order to avoid separation. 

The present rectangular wide-angle diffuser designs were broadly based on these derived rules. The small diffuser expands from 43 x 64 cm to 91 x 137 cm over a length of 2.9 

m, giving a maximum included angle of 18.2” and an area ratio of 4.75. The larger diffuser 

expands from 61 x 94 cm to 91 x 137 cm over a length of 2.4 m, giving an included angle 

of 17.6” and an area ratio of 2.25. Each diffuser contains two relatively coarse screens with 

p = 64.6% (14 mesh, d = 0.36 mm), one at the upstream end and one at the downstream 

junction (Fig. 2). At the upstream end, the screen is “sandwiched” between the diffuser 

flange and a 15 cm long coupling flange, while the downstream screen is sandwiched between the diffuser and settling section flanges. In order to properly seal the junctions 

where a screen was sandwiched, the flange surface was lined with a rubber compound. 

2.3 Settling Chamber (Flow Conditioning Section) 

In open-circuit blower tunnels, settling chambers are created rather than designed 

since the frames containing the flow conditioning elements, clamped together, generally 

6 
form the settling chamber. The flow conditioning elements usually consist of a honeycomb 

and screens, the number and overall pressure drop depending on the turbulence level 

requirements in the test section. The settling chamber on the mixing layer wind tunnel 

consists of a 30.5 cm long “settling length” and a 44.5 cm stack of seven frames, each 

6.4 cm wide. Each side of the settling chamber (high-speed and low-speed) consists of a 

honeycomb and five screens, as illustrated in Fig. 2. 

The most efficient device for removing swirl and lateral mean velocity variations in 

the flow is a honeycomb, as long as the flow yaw angles are not greater than about 10”. 

Large yaw angles cause the honeycomb cells to “stall” which reduces their effectiveness 

besides increasing the pressure loss and causing flow nonuniformity. If severe yaw or swirl is 

expected in the incoming flow, it is advisable to install screens upstream of the honeycomb, 

so that the flow angles are reduced. A screen with K = 1.5 reduces yaw and swirl angles 

by a factor of 0.7 for swirl angles of about 40”, so multiple screens would be required 

upstream of the honeycomb in this case. The honeycomb should also be installed some 

distance downstream of the wide-angle diffuser exit, so that the flow static pressures and 

angles have had a chance to become more uniform. For this very reason, in the present wind 

tunnel, the diffuser sections are followed by 30.5 cm long, constant area “settling sections” 

(Fig. 2). Also, the honeycomb on each side is preceded by a screen with p = 62.4% (28 

mesh, d = 0.19 mm). 

An incidental effect of honeycombs is to reduce the turbulence level in the flow. Essentially, the lateral components of turbulence, like those of mean velocity, are inhibited by 

the honeycomb cells and almost complete annihilation is achieved in a length equivalent to 

about 5-10 cell diameters. Honeycombs themselves shed small-scale turbulence, the level 

of which is found to be higher when the cell flow is laminar than when it is turbulent: 

this is attributed to a basic instability of the laminar near wakes which leads to transition, and hence high turbulence intensities, in the free-shear layer (Mehta, 1978, Bell and 

Mehta, 1989b). Note that the cell flow in most small, low-speed wind tunnel honeycombs 

is laminar. Now since the shear layer instability in the near wake has a strength proportional to the shear layer thickness, the net reduction (ratio of turbulence suppressed to 

that generated) is greatest for the shortest honeycomb. 

For maximum overall benefit, the optimum cell length has been found to be about 6-8 

times its diameter. As for cell size, in general, approximately 150 cells per settling chamber 

“diameter” or 25,000 cells total are found to be adequate. Ideally, the cell size should be 

smaller than the smallest lateral wavelength of the velocity variation. This implies that 

the choice of cell size should also depend somewhat on the size of the blower generating 

the flow, since this determines the initial (input) flow scales and also on other elements 

such as screens which may be installed in the wide-angle diffuser. The selection of the two 

aluminum honeycombs for the present wind tunnel were based on all these criteria. The 

high-speed side honeycomb is 5.1 cm long with 6.4 mm cells (LID = 8) while the lowspeed side honeycomb is somewhat smaller with 3.2 mm cells over a 2.5 cm length (L/D 

= 8). The cross-sectional shape of the cells is not critical and it is usually determined by 

availabilty; the cell shape on both the present honeycombs is hexagonal. 

Screens have been used to improve flow quality in wind tunnels since the 1930s. 

Prandtl(l933) noted that the velocity distribution improved when screens were installed in 

7 
a wind tunnel and he gave a simple theory for this effect. The action of screens in reducing 

the turbulence intensity levels became apparent later. Dryden and Schubauer (1947) gave 

a physical explanation for this effect: turbulent eddies of larger scale than the screen mesh 

size are decreased, and although at the same time smaller scale eddies may be introduced, 

they decay much more rapidly than the original large-scale motions would in the absence 

of the screen. The expected overall effect is that at some distance downstream, both the 

turbulence intensity and scale have been reduced. 

Screens are normally made of metal wires interwoven to form square or rectangular 

meshes, and are usually installed in the settling chamber to improve the mean flow uniformity and to reduce the intensity of the oncoming turbulence. A screen also refracts the 

incident flow towards the local normal. The action of a screen is described in terms of 

two parameters, namely the pressure drop coefficient and the deflection coefficient. The 

factors of reduction of mean velocity variation and turbulence intensity by screens are 

tabulated in Mehta (1977). Screens make the flow uniform by imposing a static pressure 

drop proportional to velocity squared. A screen with a pressure drop coefficient of about 

two removes nearly all variation in the longitudinal mean velocity. Dryden and Schubauer 

(1947) showed that, for a given pressure drop, a greater reduction in turbulence intensity 

is obtained by using multiple screens, each with a comparatively low pressure drop coefficient. The most optimum formulation for the screen pressure drop coefficient (see Mehta, 

1978) is that due to Wieghardt (1953) and is defined as: 

In the present wind tunnel design, the four screens downstream of the honeycomb are 

identical on both sides of the settling chamber (Fig. 2). The first screen is 28 mesh, with 

0.19 mm wire diameter and /3 = 62.4%. The next two screens are 42 mesh, with 0.14 mm 

wire diameter and p = 59.1%. The last screen is 44 mesh, with 0.14 mm wire diameter 

and /.3 = 57.5%. The order of the screens is chosen such that the mesh size decreases 

with downstream position. This was done in an attempt to gradually decrease the scale of 

the oncoming turbulence since some recent results (Mehta, 1978) have further confirmed 

the belief that the scale of the turbulence is reduced by a screen. All the screens have 

an open area ratio of greater than 57% since low p screens have been known to produce 

instabilities resulting from a random coalition of jets and wakes amalgamating to form 

weak longitudinal vortices (Mehta and Hoffmann, 1986 and 1987). 

There are two important considerations when choosing the optimum spacing between 

the flow conditioning elements. For the pressure drops through the screens to be completely 

independent of each other, the spacing should be such that the static pressure has fully 

recovered from the perturbation before reaching the next screen. For full benefits from the 

turbulence-reduction point of view, the minimum spacing should perhaps be of the order of 

the large energy containing eddies. The basic design philosophy is that the screens must be 

placed sufficiently far apart so that they act as individual elements for which the behavior 

8 
is well understood. It has been found (empirically) that a spacing equivalent to 5% of the 

settling chamber diameter performs successfully, and was hence used in the present design. 

The last frames in the settling chamber do not contain any flow conditioning elements 

- they were left empty for two main reasons. Firstly, they allowed the flow to recover 

from the effects of the last screen before feeling the effects due to the contraction entry. 

Secondly, they enabled a smooth transition from the frame surface to the splitter plate; the 

splitter plate was designed to slide onto indentations machined on the last pair of frames. 

2.4 Contraction Section 

Contraction sections form an integral part of all wind tunnels, whether designed for 

basic fluid flow research or model testing. They are normally installed upstream of the test 

section and serve two main purposes. Firstly, since the total pressure remains constant 

through the contraction, both mean and fluctuating velocity variations are reduced to a 

smaller fraction of the average velocity at a given cross-section. Secondly, a contraction 

increases the flow mean velocity and this allows the flow conditioning elements to be placed 

in a lower speed region, thus reducing the pressure losses through these components. The 

most important single parameter in determining these effects is the contraction ratio, c. 

The factors of reduction of mean velocity variation and turbulence intensity for axisymmetric contractions were derived theoretically by Batchelor (1953) using the rapid distortion 

theory and are given as: 

(i) u-component mean velocity, U: I/c 

(ii) v or w-component mean velocity, V or W: ,/Z 

(iii) u-component r.m.s. intensity, u': &[3(zn4c3 - I)]'/~ 

(iv) o or w-component r.m.s. intensity, o' or w : 

' (341'2 

A contraction is less efficient in suppressing longitudinal turbulence than mean velocity 

variation. In absolute level, the lateral turbulence intensities (o' and 20') me enhanced 

while the streamwise intensity (u') is reduced. The reduction of fractional variation of 

the lateral component is therefore much less than that of the streamwise component, so 

it should be the desired reduction in the former (together with the desired mean flow 

uniformity) which decides the contraction ratio. Contraction ratios of between 6 and 10 

are found to be adequate for most small, low-speed wind tunnels - defined as tunnels 

with a test section cross-sectional area of less than about 0.5 m2 and free-stream velocities 

of less than about 40 m/s. 

Another contraction parameter that has to be selected, a priori, is the cross-sectional 

shape. In order to avoid crossflows and boundary layer separation in the corners, the ideal 

cross-sectional shape is circular. However, in the absence of separation, the secondary flows 

in the corners tend to remain localized, without any significant effect on the test section 

flow quality (Mehta, 1978). The cross-sectional shape for modern day contractions is, 

therefore, almost always chosen to match the other tunnel components which are normally 

square or rectangular. 

The wall shape design of a contraction of given area ratio and cross-section centers 

9 
on the production of a uniform and steady stream at its outlet. These conditions can be 

generally met by making the contraction section very long so that the length is equivalent to several times the inlet height. However, another desirable flow quality is minimum 

boundary layer thickness (in a laminar state) at the contraction exit. This suggests that 

the contraction length should be minimized so that the boundary layer growth is also 

minimized. Shorter contractions are also, of course, desirable for saving in cost and space. 

However, the risk of boundary layer separation increases as the contraction length is reduced. The danger of boundary layer separation results from the presence of regions of 

adverse pressure gradient on the walls at each end of a contraction of finite length. The 

adverse pressure gradients become stronger as the contraction length is decreased. In 

general, the boundary layer is less liable to separate at the contraction exit, due to its 

increased skin friction coefficient caused by passage through the strong favorable pressure 

gradient. Also, the concave curvature at the contraction inlet has a destabilizing effect on 

the boundary layer, in contrast to the convexcgrvature near the exit which has a stabilizing effect (Bradshaw, 1973). In addition to unnecessary thickening of the boundary layer, 

separation also generally leads to flow unsteadiness, which cannot be easily eliminated 

from the test-section flow. A design satisfying all criteria will be such that separation is 

just avoided (implying a minimum acceptable length) and the exit non-uniformity is equal 

to the maximum tolerable level for a given application (typically less than 1% variation in 

mean streamwise velocity outside the boundary layers). Since a satisfactory direct design 

method for the contraction wall shape was not readily available, a new design strategy was 

pursued. 

An iterative design procedure was developed for the contraction to be installed on the 

present mixing layer wind tunnel (Bell and Mehta, 1988 and 1989a). The procedure consisted of first computing the potential flow field and hence the pressure distributions along 

the walls of a contraction of given size and shape using a three-dimensional numerical panel 

method. The wall pressure or velocity distributions were then fed into two-dimensional 

boundary layer codes to predict the behavior of the boundary layers along the walls. For 

small, low-speed contractions it was shown that the assumption of a laminar boundary 

layer originating from stagnation conditions at the contraction entry and remaining laminar throughout passage through the “successful” designs was justified. This hypothesis 

was confirmed by comparing the predicted boundary layer data at the contraction exit 

with measured data in existing wind tunnels. The measured boundary layer momentum 

thicknesses at the exit of three existing contractions, two of which were 3-D, were found 

to lie within 10% of the predicted values, with the predicted values generally lower. 

Once the computational procedure had been validated, it was used to select the optimum wall shape for the new contraction. Other design considerations dictated a contraction ratio of about eight with a 2-D (or at most mildly 3-D) contraction. The main 

requirements were that a laminar boundary layer at low Reo (less than about 500) is obtained at the splitter plate edge at reasonable flow speeds (10 - 20 m/s). This of course 

implies that the contraction length must be minimized. This was in addition to the usual 

requirements of wanting to avoid boundary layer separation on the walls and obtaining a 

reasonable mean flow uniformity at the contraction outlet (less than 1% nonuniformity). 

After an extensive review of the existing literature on contraction design, four polynomial 

10 
shapes were selected for testing with the new computational scheme: Third Order Polynomial, Fifth Order Polynomial, Seventh Order Polynomial and Matched Cubics. The 

contraction area ratio and cross-section, together with the inlet and outlet dimensions had 

already been decided by other considerations. The contraction area ratio was fixed at a 

value of 7.7, contracting from overall inlet dimensions of 91 x 274 cm to 91 x 36 cm. The 

optimization process, therefore, consisted of choosing the ideal contraction wall shape and 

length. 

After running computations on a variety of designs, it was found that the minimum 

acceptable L/Hi for our requirements was 0.89, giving a length of 2.4 m. This particular 

length to height ratio was therefore used as a test case for all the contraction wall shapes. 

The contraction wall shape satisfying most of the requirements discussed above was the 

one given by the 5th order polynomial. It was free of separation (both on the centerline 

and in the corners) and gave a reasonable Reo (- 400) and flow uniformity (better than 

0.5% variation) at the operating conditions (Ue = 15 m/s). This wall shape was therefore 

used on the contraction for the mixing layer wind tunnel (Fig. 2). 

The contraction is split into two identical two-dimensional contraction sections, each 

with an area ratio of 7.7. The contraction sections are mirror-images, mounted side-byside, and separated by a splitter plate. The splitter plate separating the contractions tapers 

from 5 cm thick at the base to 0.5 mm thick at the trailing edge. The included angle thus 

formed is about 1.1". The splitter plate protrudes 15 cm into the test section in order to 

allow some distance for streamline curvature to relax after the flow has passed through the 

contraction section, and to dow some recovery distance, should it be desired to trip the 

splitter plate boundary layers. 

2.5 Test Section 

Test section design is totally dependent on the requirements of the particular experiiments to be performed in the wind tunnel. Blower tunnels are generally more flexible in 

accepting a wide variety of test sections, both with and without exit diffusers. The flow 

out of a contraction often takes a short distance (equivalent to about 0.5 diameters) before 

the flow nonuniformities and angularities are reduced to an acceptable level. Also, if a 

turbulence grid is installed at the test section inlet, it may take up to 10-15 mesh lengths 

before a homogeneous flow is obtained. These requirements often fix the minimum length 

of the test section. General purpose test sections normally have cross-sectional aspect 

ratios of approximately 1.5:l. However, tunnels designed for shear layer research normally 

have test sections with a width-to-height ratio of at least five. The main requirement, 

in this case, is that a reasonable thickness of irrotational flow remains between the shear 

layers by the end of the test section. 

The present test section is totally constructed out of Plexiglas and has the dimensions 

of 36 cm in the cross-stream direction, 91 cm in the spanwise direction, and 3.7 m in length. 

Since the contraction sections lie side-by-side, the mixing layer is oriented vertically. The 

overall test section aspect ratio is AR = 2.5; each half has an aspect ratio of just over 

five. It is more useful, however, to compare the size of the tunnel with the thickness of 

the shear layer being studied. In the case of a two-stream mixing layer of velocity ratio 

0.6, the maximum measured layer thickness (at 2.5 m downstream) was 11 cm (Bell and 

11 
Mehta, 1989b). So the spanwise extent of the mixing layer at this point was 8.3 times 

the mixing layer thickness, thus ensuring at least geometric two-dimensionality. By this 

station, the high- and low-speed sidewall boundary layers were about 2.5 cm thick, leaving 

a 10 cm potential flow core on each side of the mixing layer. The test section flow exhausts 

into the laboratory; there is no diffuser at the downstream end. Static pressure ports are 

drilled at 61 cm intervals in the test section bottom wd, roughly along the mixing layer 

centerline to monitor the streamwise pressure gradient. The sidewall on the low-speed side 

is movable and flexible, and can be adjusted to give a prescribed pressure gradient. 

The low-speed sidewall consists of a number of removable plexiglass sheets. The 

sheets can be bolted together in a variety of configurations, and axe held in place with 

toggle clamps which press against the sidewalls. The test section sidewalls extend outward 

on the low-speed side beyond the edge of the contraction section, forming a shelf on which 

the movable wall rests. The movable wall can be adjusted to any desired position and then 

locked in place with the toggle clamps; this enables the streamwise pressure gradient to be 

set to within < 1% of the dynamic pressure. One section of the movable wall has a series 

of spanwise slots cut into it. The slots are 1.6 cm wide and cover the center 76 cm of the 

wall. When not in use, the slots are closed with plexiglass plugs. By altering the movable 

wall configuration, probe access can be obtained at 7.6 cm intervals along the entire length 

of the test section. 

3. WIND TUNNEL PERFORMANCE 

A study was undertaken to analyze the performance of the mixing layer wind tunnel. 

The flow was assumed to be one-dimensional, steady and adiabatic. Furthermore, skin 

friction losses on the wind tunnel walls were neglected and empirical correlations were 

used to estimate the losses through the various wind tunnel components. The blower 

specifications were obtained from the manufacturer’s charts. The main pressure losses 

incurred by the flow are through the screens and the honeycomb. Wieghartdt’s (1953) 

formula (Eqn. 1) was used to estimate the pressure drop through the screens while Loerke 

and Nagib’s (1972) experimental data were interpolated to give the pressure losses through 

the honeycombs. There is an additional loss equivalent to the test section total head due 

to the absence of an exit diffuser. The measured performance of the two legs of the wind 

tunnel were then compared to the predicted performance over some selected ranges of 

operation. 

Figures 3 and 4 show the static pressure rise and brake horsepower plotted as a 

function of the flow rate for selected (constant) values of the blower RPM. These curves 

are plotted for both blowers from the performance charts supplied by the manufacturer. 

The RPM, pressure rise and power required by the blowers are plotted as a function 

of the flow rate (CFM) in Figs. 5, 6 and 7. Since the pressure drop coefficients of the 

screens and honeycombs are only weak functions of the Reynolds number, the functional 

relationships in Figs. 5, 6 and 7 should be first, second and third order, respectively, 

as is broadly the case. The measured RPM and power and are also compared with the 

predicted values in Figs. 5 and 7. The agreements are generally quite good, with the 

12 
actual performance somewhat lower than predicted. This is not too surprising since not 

all of the possible losses through the wind tunnel components have been accounted for in 

the present analyses. 

For the maximum test section flow speeds of 20 m/s and 40 m/s on the low- and 

high-speed sides of the wind tunnel, respectively, the CFMs are 7,000 and 15,000 giving 

RPMs of about 1300 and 1100, blower pressure rises of 1 inch and 4 inches and HPs of 2 

and 15, respectively. 

4. WIND TUNNEL FLOW CALIBRATION 

4.1 Mean Flow Uniformity 

The reasons for desiring a uniform core flow in a wind tunnel test section are obvious. 

Spatial non-uniformity of the mean velocity within the test section can be caused by 

asymmetric boundary layer growth or separation in one of the wind tunnel legs, or by a 

persisting asymmetric flow produced by the fan or blower. In a return-circuit wind tunnel, 

poor corner vane design can also lead to mean flow nonuniformity. The acceptable levels 

for mean flow uniformity are generally at least five times the r.m.s. turbulent fluctuation 

levels sought in a high quality wind tunnel. The flow conditioning installed in modern 

wind tunnels, such as a honeycomb, screens and a contraction with a reasonable area ratio, 

improves both the mean flow uniformity and the turbulence intensity levels. Therefore, 

fulfilling the low turbulence requirements will almost always ensure an acceptable level of 

mean flow uniformity as well (Bradshaw and Pankhurst, 1965). 

The mean flow uniformity in the test section was measured by scanning a pitot tube 

over portions of the potential core region. The flattened pitot tube with dimensions of 

about 0.5 x 1.0 mm and a local static port were connected to a Datametrics Inc type 

511-10 barocell pressure transducer, with its associated amplifier and signal conditioner. 

The pressure transducer range was 0-10 mm Hg, corresponding to a velocity range of 

0-47 m/s. For the mean flow uniformity measurements, the pressure transducer signal 

was first directed through the Precision Buck & Gain unit (typically set to a DC offset 

of 1 volt and a gain of 10) to expand the range and provide greater resolution with the 

A/D. The pressure reading at each point was averaged over 1750 samples obtained at a 

rate of 150 samples/second. The pressure transducer was calibrated against a Combist 

Precision Manometer, which had a range of 2.2 mm Hg and an accuracy of 0.00015 mm 

Hg. Pitot probe dynamic pressure data were reduced to velocity data by the FMDAS 

software (Hooper and Saunders, 1986), using the local temperature input by the operator. 

A typical example of the measured mean velocity variation is shown in Fig. 8 in the 

form of contour plots. The maximum variation of the mean streamwise velocity is less than 

f 0.25%. This level of uniformity was found to exist throughout the test section within 

the potential core regions, in the regions away from the shear layers. 

The mean flow angularity was obtained from the cross-wire measurements which were 

primarily conducted to evaluate the turbulence levels in the wind tunnel. Some typical 

measurements of the cross-flow velocities made in the free-stream are presented in Fig. 

13 
9. The normalized cross-flow velocities (V/Ue and W/Ve) were generally found to be less 

than f0.005; corresponding to a flow angularity of less than 0.25'. 

4.2 Boundary layer and Spanwise Skin Friction Measurements 

As in the present case, aspect ratios of five to six are typically used for test-sections in 

which a two-dimensional boundary layer is desired. However, it is now widely known that a 

uniform core-flow and a large aspect ratio does not necessarily guarantee a two-dimensional 

boundary layer. Large, quasi-periodic spanwise variations (20% or more) in boundary layer 

thickness, surface shear stress and turbulence intensity are often encountered in what is 

assumed to be a two-dimensional boundary layer (Bradshaw, 1965). Over the years, several 

researchers have observed and studied this spanwise variation (see Mehta and Hoffmann, 

1986 for a review). It has been attributed to a spatial instability which causes directional 

changes in the jets emerging from the pores of screens. The spatial instability leads to the 

formation of weak longitudinal vortices in the boundary layer. Longitudinal vortices, once 

formed, persist for long distances downstream, even in a turbulent boundary layer, and 

are effective at producing significant spanwise variations in surface shear stress. Since the 

local changes in flow direction are small and hence difficult to measure, the quantity most 

suited for this check is the spanwise surface shear stress, which is adequately measured by 

traversing a Preston tube across the test-section floor. 

In order to determine the two-dimensionality of the boundary layers on the splitter 

plate in the present wind tunnel, the spanwise surface shear stress distributions were 

measured for both untripped and tripped conditions. For the tripped cases, 2.5 cm wide 

roughness strips comprising of a random distribution of round glass beads were installed 

20 cm upstream of the splitter plate trailing edge. The beads used on the high-speed side 

had an average diameter of about 0.75 mm, whereas those for the low-speed side had an 

average diameter of about 1 mm. All the measurements of the boundary layer properties 

and surface shear stress were made with a 0.5 x 1 mm pitot tube. For the surface shear 

stress measurements, the front end of the pitot tube was placed on the splitter plate 

surface (acting as a Preston tube) before collecting the data. Contact between the pitot 

tube and the splitter plate was assured by completing an electrical circuit. Data acquisition 

and reduction procedures were the same as for the mean flow uniformity measurements 

described above in Section 4.1. 

The boundary layer properties measured on the splitter plate centerline and about 

1.2 cm upstream of the trailing edge are presented in Table 1 for both, the tripped and 

untripped cases. The repeatability in the measured boundary layer properties was generally 

better than 2%. 

The spanwise distributions of the surface shear stress are presented in Figs. 10 and 

11. The results are plotted in the form of the surface dynamic pressure reading normalized 

by the centerline value. The surface pressure reading is proportional to C,' for laminar 

boundary layers and proportional to Cf8I7 for turbulent boundary layers. Since the variation in spanwise surface shear stress is not significantly affected by Reynolds number 

(Mehta and Hoffmann, 1986), the fact that the flow speeds for the various cases are not 

the same is expected to have minimal effect on the conclusions. 

14 
The laminar boundary layer results shown in Figs. 10a and b show a lasger and more 

systematic variation on the high-speed side of the splitter plate. The peak to peak variation 

of the surface pressure is about f25% on the high-speed side with a wavelength of about 

4 cm. The reasons for the difference between the two sides of the splitter plate are not 

obvious although it should be noted that the blowers supplying the air to the two sides 

are not identical. Also, the variation on the high-speed side is apparently not unusually 

large for luminur boundary layers (Wood, 1980). Boundary layer profiles measured on the 

centerline and at 2 = f 15 cm showed that the layer properties varied by less than 5% on 

the low-speed sied and by less than 10% on the high-speed side. Somewhat surprisingly, 

the results for the tripped case are comparable for the two sides (Figs. lla and b). The 

maximum peak to peak variation on both sides is about f5% which is within the acceptable 

limits given by Mehta and Hoffmann (1986) for turbulent boundary layers. Boundary layer 

profiles measured at three spanwise locations (2 = 0, f 13 cm) showed a variation of well 

within 5%. Although a distinct wavelength is not apparent, a nominal average of 2 cm is 

obtained on the low-speed side after careful examination of the data. Some recent spanwise 

surface shear stress measurements, with the boundary layers tripped by round wire trips, 

showed comparable peak to peak variations, although the details were somewhat different. 

The spanwise surface shear stress measurements therefore show that the initial boundary 

layers are nominally two-dimensional on both sides of the splitter plate for both tripped 

and untripped conditions. 

4.3 Turbulence Intensity Measurements 

The effects of free-stream fluctuations on the development and structure of a turbulent 

shear layer are now well appreciated. For example, boundary layer transition is found to 

be extremely sensitive to disturbances in the free-stream (Saric, 1986). More relevant to 

the present interest is the fact that mixing layers have also been shown to be affected by 

these fluctuations. It has been suggested that the breakdown to three-dimensionality of the 

mixing layer structure may well be determined by the level of the free-stream disturbances 

(Chandrsuda et al., 1978 and Wood and Bradshaw, 1982). 

Bradshaw and Pankhurst (1964) divided the free-stream fluctuations into two components, namely turbulence and unsteadiness. They defined unsteadiness as “a velocity 

fluctuation of low enough frequency to be noticeable on manometers and balances”. The 

distinction between the two was based on the sources of the irregularities - unsteadiness 

may be caused by an intermittent separation whereas “true” turbulence is generated by 

boundary layers and wakes of fan blades, for example. Wood and Westphal (1987) further 

split the unsteadiness into a one-dimensional (1-D) axial motion produced by, say the rotation of the fan or blower and the three-dimensional (3-D) unsteadiness which could be 

due to flow separation in some part of the wind tunnel. In most well designed wind tunnels 

where great care is taken to avoid boundary layer separations, the latter contribution may 

be ignored. It follows then that the unsteadiness will only contribute to p; the difference 

between u’2 and 3 or wt2 can then be attributed to unsteadiness. It is therefore important to measure all the components of turbulence intensity and, in particular, without high 

pass filtering so that the low frequency unsteadiness is also captured in the measurements. 

Also, the measurements should be made at several locations within the test section so that 

- 

15 
regions of local production, if any, and the importance of induced motions (potential flow 

fluctuations) from the shear layers are realized (Wood and Westphal, 1987). 

The present turbulence intensity measurements were made using a crossed hot-wire 

probe held on a 3-D traverse and linked to a fully automated data acquisition and reduction 

system controlled by a MicroVax I1 computer. The crossed hot-wire probe had 5 pm 

tungsten sensing elements about 1 mm long and positioned about 1 mm apart. The probe 

was calibrated statically in the potential core of the flow assuming a ‘cosine-law’ response to 

yaw, with the effective angle determined by calibration (Westphal and Mehta, 1983). The 

only filtering of the analog signals was to low pass filter at 30 Khz to provide immunity 

from high frequency electronic noise; the spectral measurements showed that all of the 

free-stream fluctuations resided well within this cut-off frequency. The analog signals were 

also DC offset, and amplified (x10) before being fed into a NASA-built computer interface 

containing a fast sample-and-hold A/D converter with 12 bit resolution. The offset/gain 

procedure allowed the input signal to be matched to the f10 volt range of the A/D, in 

order to maximize resolution of the signal. Th5 digital data were then multiplexed to the 

computer through a DMA board. 

In the present free-stream measurements, for reasonable repeatability in the Reynolds 

stresses (- 5%), it was found necessary to average the individual statistics over 50,000 

samples obtained at a rate of 400 samples per second; the data were actually averaged 

over 10 individual data points each evaluated from 5,000 samples. Data were obtained in 

two planes (uv and uw) by rotating the crossed hot-wire probe about its own axis. This 

method yielded all three components of mean velocity, five independent components of the 

Reynolds stress tensor and selected higher order products. 

Some preliminary turbulence measurements in the mixing layer wind tunnel with 

the sidewall positioned parallel to the fixed wall (i.e. normal mixing layer configuration) 

revealed that it was difficult to isolate the sources of the fluctuations because of “crosstalk 

problems” across the free-shear layer. The wind tunnel was therefore reconfigured such 

that each leg had its own test section; this was done by effectively extending the splitter 

plate all the way to the end of the test section. The two legs of the wind tunnel were then 

calibrated individually; some final checks with the test section in the normal (mixing layer) 

configuration were also made. The measurements were made in the Y and 2 directions 

along the nominal centerlines of the individual test sections. Most of the data presented 

below were obtained at free-stream velocities of 9 and 15 m/s on the low- and high-speed 

sides, respectively. 

The results for the low-speed side obtained at X = 11 cm are presented in Figs. 12a 

and b. The streamwise fluctuation (z) data from both planes (uw and uv) are presented 

for comparison. On the whole, reasonable agreement is observed between the results from 

the two planes with the uv-plane measurements slightly higher. There do not seem to be 

any systematic, significant variations in in either direction. The average (Flue2) level 

is about 3 x lo-‘, which correponds to a turbulence - intensity level, u’/Ue - 0.17%. In 

the spanwise (2) direction, the levels of 21’2 and wr2 are comparable and much lower, as 

expected. Again, there - is very little variation in the spanwise direction giving an average 

level for p)’2/Ue2 and wr2/Ue2 of about 0.5 x lo-‘, which corresponds to v’/Ue = w’/Ue 

- 0.07%. The results for these two components in the normal (Y) direction deserve more 

16 
comment. Both 2))2 and 20’2 are found to increase as the outer wall is approached, with 

increasing more rapidly. This is partly due to the induced fluctuations from the outer wall 

boundary layer and partly due to streamline curvature effects imposed by the contraction 

exit. 

The measurements for the high-speed side (Figs. 13a and b) exhibit similar qualitative 

trends. The average u’2/Ue2 level is somewhat lower at 2 x lo-‘, corresponding to d/Ue 

N 0.14%. The 0’2 levels are slightly higher than the w’2 levels in both - directions. The 

average 0’2/Ue2 level in figure lla is still about 0.5 x lo-‘, whereas the td2/Ue2 level is 

slightly lower at 0.25 x lo-‘. Once again, the turbulence levels are found to increase in 

the Y-direction as the outer wall is approached with 2)’2 increasing more rapidly. 

The measured turbulence levels on both sides of the mixing layer wind tunnel were 

well within the design specifications of u’/Ue 5 0.2%. As discussed above, the measured 

d/Ue levels of approximately 0.15% in both legs of the wind tunnel include contributions 

from the low frequency unsteadiness; the “true” turbulence levels are approximately half 

that (0.07%). 

4.4 Spectral Measurements 

Although the turbulence intensity levels were found to be adequately low in the mixing 

layer wind tunnel, some spectral measurements were made to investigate the nature or 

frequency content of the unsteadiness contributions in the hope of identifying its sources. 

The signal from a single hot-wire was examined with a digital spectrum analyzer to 

characterize dominant frequencies of velocity fluctuations within the free stream. The output of the hot-wire anemometer was connected directly to the spectrum analyzer, so that 

a hot-wire voltage spectrum, rather than a velocity spectrum, was obtained. Nevertheless, 

the location and relative amplitude of frequency peaks should correspond. Frequency peak 

measurements were made by surveying the entire frequency spectrum, choosing the ranges 

which most accurately displayed any peaks found, and photographing the display with a 

Polaroid camera. The frequency range was not changed when measurements of relative 

amplitude were made, to avoid bandwidth effects on the amplitude. The measurements 

were made over a wide range of flow speeds, and hence blower RPMs, on both sides of 

the wind tunnel since the blower fan is the most obvious source of fluctuations. The 

speed range covered, and the corresponding blower RPMs and blade passing frequencies 

are tabulated in Table 2. 

Figure 14 shows the results for the low-speed side while those for the high-speed 

side are presented in Fig. 15. On both sides, and at all speeds, the most predominant 

(highest) peak occurs at the frequency corresponding to the revolutions per second. The 

blade passing frequency is not apparent in any of the results, although the results for the 

lowest flow speed (Ue = 5 m/s) show the presence of some extraneous peaks. The once 

per revolution unsteadiness is probably caused by a slight misalignment of the impeller 

which cannot be easily adjusted. Since the tubulence levels were within acceptable limits, 

no attempts were made to rectify this situation. 

17 
5. CONCLUDING REMARKS 

A new wind tunnel, specifically designed for mixing layer experiments, has been constructed and calibrated. The unique feature of the tunnel is that it is driven by two separate blowers which feed two legs separated up to the test section where the two streams 

are allowed to merge. The mean flow in the test section is uniform to within f0.25% 

and the flow angularity is less than 0.25’. The spanwise variation of the surface shear 

stress is less than 4~5% for the tripped boundary layers on the splitter plate indicating 

acceptable two-dimensionality. The total streamwise free-stream turbulence intensity level 

is approximately 0.15% with the cross-stream components equal to about 0.07%. The 

spectral measurements indicate that the main source of unsteadiness is the once per revolution contribution from the blower impeller. So overall, the wind tunnel flow quality 

is considered to be well within specifications and experiments designed to investigate the 

three-dimensional structure of plane mixing layers (Bell and Mehta, 1989b) and wakes are 

already being conducted in this facility. 

ACKNOWLEDGEMENTS 

We are grateful to Mat Russo of Oliver Johnson for his careful supervision of the tunnel 

component construction and to David Yaste for considerable help in assemblying the wind 

tunnel. We would also like to thank Susan Praskins for the test section fabrication and 

Bruce Wu for conducting the tunnel performance calculations. Helpful discussions with 

David Wood during the calibration phase are also gratefully acknowledged. 

This work was supported by the Fluid Dynamics Research Branch at NASA Ames 

Research Center under grant NCC-2-55. 

REFERENCES 

Batchelor, G.K. “The Theory of Homogeneous Turbulence,” Cambridge University Press, 

pp. 68-75, 1953. 

Bell, J.H. and Mehta, R.D. “Contraction Design for Small Low-Speed Wind Tunnels,” 

JIAA Report TR-64, Department of Aeronautics and Astronautics, Stanford University, 

April 1988. Also NASA CR-177488, August 1988. 

Bell, J.H. and Mehta, R.D. “Boundary Layer Predictions For Small Low-Speed Contractions,” AIAA Journal, Vol. 27, pp. 372-374, March 1989a. 

Bell, J.H. and Mehta, R.D. “Three-dimensional Structure of a Plane Mixing Layer,” Department of Aeronautics and Astronautics, JIAA TR-90, March 1989b. Also, AIAA Paper 

89-01 24, January 1 989 b . 

18 
Bradshaw, P. “Wind Tunnel Screens: Flow Instability and its Effect on Aerofoil Boundary 

Layers,” Journal of Royal Aeronautical Society, Vol. 68, pp. 198, 1964. 

Bradshaw P. and Pankhurst R.C. “The Design of Low-Speed Wind Tunnels,” Progress in 

Aerospace Sciences, Vol. 5, p. 1, 1964. 

Bradshaw, P. “The Effect of Wind Tunnel Screens on Nominally Two-Dimensional Boundary Layers,” Journal of Fluid Mechanics, Vol. 22, pp. 679, 1965. 

Bradshaw, P. “Two More Wind Tunnels Driven by Aerofoil-Type Centrifugal Blowers,” 

Imperial College Aero Report 72-10,1972. 

Bradshaw, P. “Effects of Streamline Curvature on Turbulent Flow,” AGARDograph 169, 

1973. 

Cebeci, T. and Bradshaw, P. “Momentum Transfer in Boundary Layers,” Hemisphere 

Publishing Co., 1977. 

Chandrsuda, C., Mehta, R.D., Weir, A.D. and Bradshaw, P. “Effect of Free-Stream Turbulence on Large Structure in Turbulent Mixing Layers,” Journal of Fluid Mechanics, Vol. 

85, pp. 693, 1978. 

Dryden, H.L. and Schubauer, G.B. “Use of Damping Screens for Reduction of Wind Tunnel 

Turbulence,” Journal of the Aeronautical Sciences, Vol. 14, p. 221, 1947. 

Hooper, C. and Saunders, D. “FMDAS: Fluid Mechanics Data Acquisition System User’s 

Guide,” Technical Report 7104-3-7, Sterling Software, September 1986. 

Laws, E.M. and Livesey, J.L. “Flow Through Screens,” Annual Review of Fluid Mechanics, 

pp. 247-266, Vol. 10, 1978. 

Loehrke, R.I. and Nagib, H.M. “Experimental Investigation on Management on FreeStream Turbulence,” AGARD Report 598, AD-749-891,1972. 

Mehta, R.D. “The Aerodynamic Design of Blower Tunnels with Wide-Angle Diffusers,” 

Progress in Aerospace Sciences, Vol. 18, No.1, p.59-120, 1977. 

Mehta, R.D. “Aspects of the Design and Performance of Blower Tunnel Components,” 

Ph.D. Thesis, Department of Aeronautics, Imperial College, University of London, November 1978. 

Mehta, R.D. and Bradshaw, P. “Design Rules for Small Low Speed Wind Tunnels,” The 

Aeronautical Journal, Vo1.83, No.827, November 1979. 

Mehta, R.D. “Turbulent Flow through Screens,” AIAA Paper 840538, January 1984. 

Mehta, R.D. “Turbulent Boundary Layer Perturbed by a Screen,” AIAA Journal, Vol. 23, 

No. 9, pp. 1335, September, 1985. 

19 
Mehta, R.D. and Hoffmann, P.H. “A Study of the Factors Affecting Boundary Layer TwoDimensionality in Wind Tunnels,” Department of Aeronautics and Astronautics, Stanford 

University, JIAA Report TR-66,1986. 

Mehta, R.D. and Hoffmann, P.H. “Boundary Layer Two-Dimensionality in Wind Tunnels,” 

Ezperiments in Fluids, Vol. 5, No. 5, pp. 358-360, 1987. 

Mokhtari, S. and Bradshaw, P. “Longitudinal Vortices in Wind Tunnel Wall Boundary 

Layers,” Aeronautical Journal, Vo1.87, 1983, pp. 233-236. 

Morkovin, M.V. Observations of Streamwise Vortices in Laminar and Turbulent Boundary 

Layers. NASA Contractor Report 159061, April, 1979. 

Patel, V.C. “Calibration of the Preston Tube and Limitations on its Use in Pressure 

Gradients,” Journal of Fluid Mechanics, Vol. 23, pp. 185,1965. 

Prandtl, L. “The Attainment of a Steady Air Stream in Wind Tunnels,” NACA-TM-726, 

1933. 

Saric, W.S. “Boundary Layer transition to Turbulence: The Last Five Years,” Proceedings 

of the Tenth Symposium on Turbulence, University of Missouri, Rolla, Missouri, pp. 19- 

Stratford, B.S. “The Prediction of Separation of the Turbulent Boundary Layer,” Journal 

of Fluid Mechanics, Vol. 5, pp. 1-16, 1959. 

1~19-18, 1986. 

I Westphal, R.V. and Mehta, R.D. “Crossed Hot-wire Data Acquisition and Reduction 

I System,” NASA TM-85871,1983. 

Wieghardt, K.E.G. “On the Resistance of Screens,” Aeronautical Quarterly, Vol. 4, pp. 

186,1953. 

Wood, D.H. “A Reattaching, Turbulent, Thin Shear Layer,” Ph.D. Thesis, Department of I 

I Aeronautics, Imperial College, University of London, January 1980. 

Wood, D.H. and Bradshaw, P. “A Turbulent Mixing Layer Constrained by a Solid Surface,” 

Journal of Fluid Mechanics, Vol. 122, pp.265, 1984. 

Wood, D.H. and Westphal, R.V. “Measurements of the hee-Stream Fluctuations Above 

a Turbulent Boundary Layer,” NASA-TM 100036, November 1987. 

20 
Table 1 Boundary Layer Properties 

~ Condition ut 699 

~ (m/s) (cm) 

1 High-Speed Side, ITnt ripped 15.0 0.398 

, High-Speed Side. Tripped ~ 21.0 0.86 

I Low-Speed Side, Tripped 1 10.5 0.87 

I 

9.0 0.441 i 

1 Low-Speed Side. I1ntripped 

e Ree H Cf 

(cm) xi03 

0.0526 525 2.52 0.87 

0.0606 362 2.24 1-56 

0.096 1301 1.49 3.99 

0.098 686 1.53 4.8.5 

Table 2 Flow Speed versus Blower Setting 

Beed Side 1 

13.0 

I High-speed Side 

'U, (m/s) RPM'Rev/s 

5 167 ' 2.78 ' 10 I 317 I 5.28 

1 5 ' 466; 1 7.77 

20 j 611 j 10.18 
M 

E .- 
- SCREEN 

- !??FT!T ............ HONEYCOMB 

18.3 cm 

LENGTH E 2.5 cm L/D z 8 ....................... 

DlAM = 3.2 mm 

15 

1 7.0 cm ~- t 7.0 cm 

cm 

1 6.4 cm 

16.4 cm 

AREA = 0.28 m2 

AREA = 0.57 I 

/ 4 

I 

7 f 6.4 cm I I 

I* p,% MESH d,mm 

- 64.6 14 0.36 

- 64.6 14 0.36 

I LENGTH = 5.1 cm] :i L 8 

- 62.4 28 0.19 

- 59.1 42 0.14 

- 59.1 42 0.14 

- 57.5 44 0.14 

62.4 28 

CELL 

DlAM = 6.4 mm 

Figure 2. Block diagram of Mixing Layer Wind !Limnel, 

showing flow conditioning elements. 
9- 

FLOW RATE, CFM 

a) Low-Speed Side 

6- 

K 

w 

w 

v) 

E 

$4- 

W 

Y 

K 

a 

m 

FLOW RATE, CFM 

b) High-speed Side 

25 ,- 

a 

m 

I 1 1 I 10 1 I 1 

Figure 3. Static pressure rise versus flow rate (CFM). 

a) Low-Speed Side b) High-speed Side 

Figure 4. Brake horse power (BHP) versus CFM. 
2300 

I z1300 

a 

300 

--- PREDICTED 

-0- MEASURED 

1 3 5 7 9x103 

a) Low-Speed Side 

FLOW RATE, CFM 

21- 

1600 

1 loo 

--- PREDICTED 

4 MEASURED 

100 

1.5 6.5 11.5 

FLOW RATE, CFM 

b) High-speed Side 

Figure 5. Blower RPM versus CFM. 

16.5 x 1s 

W' 1 

pc 

7 W 

E 

0 

1 3 5 7 9x 103 

FLOW RATE, CFM 

2.5 7.5 12.5 17.5 x 103 

FLOW RATE, CFM 

a) Low-Speed Side b) High-speed Side 

Figure 6. Pressure rise versus CFM. 
5.0 

P 

s 

W 

d 

U 

3 

0 2.5 

W 

U 

0 

-0- COMPUTED 

*.***.* MANUFACTURER'S CURVE 

-Q- MEASURED 

FLOW RATE, CFM 

a) Low-Speed Side 

0-0 COMPUTED 

0. 

-.Q- MEASURED 

MAN U FACT U R E R 'S CURVE 

P 

c 

0.20 

W 

U 

0 

3 

W 

0 

1.5 6.5 11.5 16.5 x 

FLOW RATE, CFM 

b) High-speed Side 

Figure 7. Work required by blower versus CFM. 
CONTOUR LEVELS 

1 0.998500 

2 0.999000 

3 0.999500 

4 1.000000 

5 1.000500 

6 1.001000 

7 1.001500 

0 

0 I I 1 1 1 1 1 1 

-8.0 -6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 8.0 

Z, cm 

Figure 8. Contour plot of mean flow uniformity. The levels 

have been normalized by the average velocity. 
O= W/Ue 

O= V/Ue 

*c 0.010 

0 

d 

9 

30.005 

0 

I 

0 

L 

V 

a 

$0.005 

E: 

~0.000 

.r( 

w 

- 

- 

- 

k 

~0.010 I I I I I I 

0 - 6.0 -4.0 -2.0 0.0 2.0 4.0 6.0 2: 

Nominal Distance From Centerline (cm) 

Figure 9. Typical distribution of cross-flow velocities. The 

cross-flow velocities are measured along a line 

normal to the splitter plate, and are normalized 

by the average streamwise velocity. 
1.5 

-. : 1.0 L 

0) 

d 

Q) 

c 

\ G 

0.5 

0.0 

u 

-20.0 - 10.0 0.0 10.0 20.0 

z (cm) 

a) Low-Speed Side 

U' 0.5 

-20.0 -10.0 0.0 10.0 20.0 

(cm) 

b) High-speed Side 

Figure 10. Spanwise distributions of surface shear stress for 

laminar (untripped) boundary layers 
Q) 

c 

CI .c. 

L 

0) 

Y 

0) 

C 

\ G 

1.5 

1.0 

0.5 

0.0 

1.5 

1 .o 

0.5 

0.0 

-20.0 -10.0 0.0 

(cm> 

a) Low-Speed Side 

10.0 20.0 

-20.0 -10.0 0.0 

(cm) 

b) High-speed Side 

10.0 20.0 

Figure 11. Spanwise distributions of surface shear stress for 

turbulent (tripped) boundary layers 
Figure 12a. Turbulence intensity measurements on low-speed 

side. Data is shown along a line at X=11 cm, 

Y=5 cm. 

8X lo4 

0 

-20 -10 0 10 20 

2, cm 

- - 

0 22 ue2/U? UW-PLANE 0 v' /Up - " - " 

22 A U'~/U: UV-PLANE 0 w' /Ua 

8 

0 I 

v, an 

Figure 12b. Turbulence intensity measurements on low-speed 

side. Data is Shawn along a line at X=ll cm, Z=O 

cm. 
, 

Figure 13a. Turbulence intensity measurements on high-speed 

side. Data is shown along a line at X=ll cm, Y=5 

cm. 

8 X lod 

Z. cm 

- - 

22 0 u’*/U: UW-PLANE 0 v’ /Ue 

22 A uW2/Uf UV-PLANE 0 W’ /Ue 

- - 

0 

-18 -16 -14 -12 -10 -8 -6 -4 -2 0 

v, an 

Figure 13b. Turbulence intensity measurements on high-speed 

side. Data is shown along a line at X=ll cm, Z=O 

cm. 
U, = 15 m/wc I 

I I I 1 I I 1 I I I 

0 10 20 30 40 50 60 70 80 90 100 

FREQUENCY, Ht 

Figure 14. U-component Spectra Measurements, low-speed 

side. 

I I I 1 I I 1 I I I 

0 10 20 30 40 50 60 70 80 90 100 

FREQUENCY, Ht 

Figure 14. U-component Spectra Measurements, low-speed 

side. 
1 I I 1 I 

I 1 I I I I I I I I 

0 1020304050607080901 

FREQUENCY, Hz 

Figure 15. U-component Spectra Measurements, high-speed 

side.